CNS-13-005, License Amendment Request for Methodology Report DPC-NE-3001-P, Revision 1, Multidimensional Reactor Transients and Safety Analysis Physics Parameters Methodology

License Amendment Request for Methodology Report DPC-NE-3001-P, Revision 1, Multidimensional Reactor Transients and Safety Analysis Physics Parameters Methodology

ML13325B142
Person / Time
Site:	Mcguire, Catawba, McGuire
Issue date:	11/14/2013
From:	Henderson K Duke Energy Carolinas
To:	Document Control Desk, Office of Nuclear Reactor Regulation
References
CNS-13-005 DPC-NE-3001-P, Rev 1
Download: ML13325B142 (105)

Text

Kelvin Henderson DUKE Vice President ENERGY. Catawba Nuclear Station 803-701-4251 Duke Energy CNS-13-005 CNO1VP I 4800 Concord Rd.

York, SC 29745 November 14, 2013 10 CFR 50.90 U.S. Nuclear Regulatory Commission Attention: Document Control Desk Washington, DC 20555-0001

Subject:

Duke Energy Carolinas, LLC (Duke Energy)

McGuire Nuclear Station, Units 1 and 2; Docket Nos. 50-369, 50-370 Catawba Nuclear Station, Units 1 and 2; Docket Nos. 50-413 and 50-414 License Amendment Request for Methodology Report DPC-NE-3001-P, Revision 1, MultidimensionalReactor Transients and Safety Analysis Physics Parameters Methodology (Proprietary)

In accordance with the provisions of 10 CFR 50.90, Duke Energy is submitting a License Amendment Request (LAR) for the Renewed Facility Operating Licenses (FOLs) for McGuire Nuclear Station, Units 1 and 2 and Catawba Nuclear Station, Units 1 and 2. Specifically, Duke Energy requests NRC review and approval of proposed changes to the FOLs based on DPC-NE-3001 -P, MultidimensionalReactor Transientsand Safety Analysis Physics Parameters Methodology.

The proposed changes constitute Revision 1 to DPC-NE-3001-P. Major changes contained in Revision 1 include a new methodology for the main steam line break transient from the full power initial condition, revisions to the rod ejection methodology and the main steam line break methodology from the hot zero power initial condition, and a revised single failure assumption for the dropped rod transient. The analysis impacts associated with NRC Information Notice 2011-21, "Realistic Emergency Core Cooling System Evaluation Model Effects Resulting From Nuclear Fuel Thermal Conductivity Degradation", are also addressed.

It is Duke Energy's intent to publish an approved version of DPC-NE-3001-P following NRC approval of this LAR.

Enclosure 1 provides Duke Energy's evaluation of the LAR which contains a description of the proposed changes, the technical analysis, the determination that this LAR contains No Significant Hazards Consideration, and the basis for the categorical exclusion from performing an Environmental Assessment/Impact Statement.

This LAR does not request any changes to the McGuire or Catawba Technical Specifications (TS). Consequently, no proposed revised TS pages are included within this submittal.

Attachments 1a and lb in Enclosure 1 provide proprietary and non-proprietary versions of the revised methodology report sections. All changes are relative to Revision Oa of DPC-NE-3001-P.

Attachment la of Enclosure 1 to this letter contains proprietary information.

Withhold from public disclosure under 10 CFR 2.390.

Upon removal of Attachment la of Enclosure 1, this letter is uncontrolled. wwwduke-energy com

U.S. Nuclear Regulatory Commission Page 2 November 14, 2013 DPC-NE-3001-P contains information that is proprietary to Duke Energy. In accordance with 10 CFR 2.390, Duke Energy requests that this information be withheld from public disclosure. An affidavit is included (Enclosure 2) from Duke Energy attesting to the proprietary nature of the information in the report. The specific information that is proprietary to Duke Energy is identified in the report.

Duke Energy is requesting that the NRC review and approve this LAR within one year from the date of submittal.

Future revisions to the McGuire and Catawba Updated Final Safety Analysis Reports (UFSARs) necessary to reflect approval of this LAR will be made in accordance with 10 CFR 50.71(e).

In accordance with Duke Energy administrative procedures and the Quality Assurance Program Topical Report, the proposed amendments have been previously reviewed and approved by the McGuire and Catawba Plant Operations Review Committees.

Pursuant to 10 CFR 50.91, a copy of this LAR has been forwarded to the appropriate State of North Carolina and State of South Carolina officials.

There are no regulatory commitments contained in this letter or its enclosures/attachments.

If you have any questions or need additional information on this matter, please contact L.J.

Rudy at (803) 701-3084.

I declare under penalty of perjury that the foregoing is true and correct.

Executed on November 14, 2013.

Very truly yours, Kelvin Henderson LJR/s Enclosures/attachments Attachment la of Enclosure 1 to this letter contains proprietary information.

Withhold from public disclosure under 10 CFR 2.390.

Upon removal of Attachment la of Enclosure 1, this letter is uncontrolled.

U.S. Nuclear Regulatory Commission Page 3 November 14, 2013 xc (with enclosures/attachments):

V.M. McCree Regional Administrator U.S. Nuclear Regulatory Commission - Region II Marquis One Tower 245 Peachtree Center Ave., NE Suite 1200 Atlanta, GA 30303-1257 J. Zeiler Senior Resident Inspector (McGuire)

U.S. Nuclear Regulatory Commission McGuire Nuclear Station G.A. Hutto, III Senior Resident Inspector (Catawba)

U.S. Nuclear Regulatory Commission Catawba Nuclear Station J.C. Paige (addressee only)

NRC Project Manager (McGuire and Catawba)

U.S. Nuclear Regulatory Commission One White Flint North, Mail Stop 8 G9A 11555 Rockville Pike Rockville, MD 20852-2738 W.L. Cox, III Section Chief Division of Environmental Health, Radiation Protection Section North Carolina Department of Environment and Natural Resources 1645 Mail Service Center Raleigh, NC 27699 S.E. Jenkins Manager Radioactive & Infectious Waste Management Division of Waste Management South Carolina Department of Health and Environmental Control 2600 Bull St.

Columbia, SC 29201 Attachment la of Enclosure 1 to this letter contains proprietary information.

Withhold from public disclosure under 10 CFR 2.390.

Upon removal of Attachment l a of Enclosure 1, this letter is uncontrolled.

Enclosure 1 LICENSEE EVALUATION

Subject:

License Amendment Request for Methodology Report DPC-NE-3001-P, MultidimensionalReactor Transientsand Safety Analysis Physics Parameters Methodology (Proprietary) 1.0

SUMMARY

DESCRIPTION 2.0 DETAILED DESCRIPTION

4.0 REGULATORY EVALUATION

4.1 Applicable Regulatory Requirements/Criteria 4.2 Precedent 4.3 Significant Hazards Consideration 4.4 Conclusions

5.0 ENVIRONMENTAL CONSIDERATION

Page 1 of 5

Enclosure 1 LICENSEE EVALUATION 1.0

SUMMARY

DESCRIPTION This evaluation supports a request to amend Renewed Facility Operating Licenses (FOLs) NPF-9 and NPF-1 7 for McGuire Nuclear Station Units 1 and 2, and NPF-35 and NPF-52 for Catawba Nuclear Station Units 1 and 2.

The proposed changes would revise the FOLs to reflect the approval of DPC-NE-3001-P, MultidimensionalReactor Transientsand Safety Analysis Physics ParametersMethodology.

2.0 DETAILED DESCRIPTION DPC-NE-3001 -P, MultidimensionalReactor Transients and Safety Analysis Physics Parameters Methodology, describes Duke Energy's methodologies for: 1) simulating the UFSAR Chapter 15 events characterized by multidimensional reactor transients, and 2) systematically confirming that reload physics parameters important to UFSAR Chapter 15 transients and accidents are bounded by values assumed in the licensing analyses. The multidimensional reactor transients described are the rod ejection accident, the main steam line break, and the dropped rod transient. The analytical approaches combine neutronics calculations with system and core thermal-hydraulics simulations. Important physics parameters are identified for each event and selected to produce a bounding transient response. For each reload core, the reference analysis is confirmed to be bounding by showing that event-specific key parameters for the reload core are within the conservative envelope of values assumed in the reference analysis, and that event-specific acceptance criteria are met. This methodology report is applicable to the McGuire and Catawba Nuclear Stations.

On November 15, 1991, the NRC approved Revision 0 of DPC-NE-3001-P. Revision Oa of DPC-NE-3001-P was subsequently generated under the 10 CFR 50.59 process. Revision 0a included the replacement of the CASMO-3/SIMULATE-3 methodology with the NRC-approved CASMO-4/SIMULATE-3 methodology from DPC-NE-1 005-PA, Duke Power Nuclear Design Methodology Using CASMO-4/SIMULATE-3 MOX, for the generation of nuclear data. The rod ejection methodology was updated to replace the ARROTTA calculation methodology with the NRC-approved SIMULATE-3K methodology from DPC-NE-2009-PA, Westinghouse Fuel Transition Report. References to the EPRI-ARMP methodology were removed, and the content related to radiological dose analysis results was deleted. Other revisions addressed issues that were identified based on experience gained in applying the original methodologies. In addition, content from the NRC-approved methodology report DPC-NE-2009-PA was merged into Revision Oa so that the methodologies were all included in this report.

Revision 1 of DPC-NE-3001 -P includes a new methodology for modeling the main steam line break transient from the full power initial condition, revisions to the rod ejection methodology, main steam line break methodology at the hot zero power initial condition, and a revised single failure assumption for the dropped rod transient. Input to the Rod Control System for the dropped rod transient is based on the newly installed Distributed Control System which replaces the 7300 Process Control System. The CASMO-4 computer code replaces the CASMO-3 code for generating nuclear data for SIMULATE-3K, and the SIMULATE-3K computer code is revised. The regulatory issue associated with fuel thermal conductivity degradation is also addressed. Other revisions include enhancements and clarifications that have been identified based on experience gained in applying the original methodologies, or are error corrections.

Page 2 of 5

Enclosure 1 The new main steam line break methodology from the full power initial condition is being added because this scenario has the potential to be more limiting than the main steam line break accident from the hot zero power initial condition. The methodology used by Duke Energy to analyze the hot zero power main steam line break accident was previously approved by the NRC as part of the original submittal of DPC-NE-3001-P.

An update to the main steam line break accident from the hot zero power initial condition is being performed to revise overly conservative assumptions in the original analysis. Duke Energy is also requesting NRC approval for the use of the WLOP critical heat flux correlation for use in evaluating the departure from nucleate boiling ratio.

Revisions to the rod ejection accident analysis methodology are being made to update the versions of the CASMO and SIMULATE-3K computer codes used in the evaluation of the transient, and to incorporate methodology enhancements identified as a result of experience gained in applying the methodology. Updates to the CASMO and SIMULATE-3K computer codes used in this analysis are being pursued to include model refinements and because the code versions referenced in the original submittal are no longer supported by the code vendor.

The dropped rod methodology is being revised to be consistent with the current configuration of the plant where the 7300 Process Control System was replaced by the Distributed Control System. The methodology is also revised to include a single failure assumption determined consistent with the selection of single failures being assumed in protection and safeguards systems for other UFSAR Chapter 15 transients. Additional methodology changes are also made based on experience gained in applying the methodology.

Demonstration analyses for the main steam line break from hot zero power, rod ejection, and dropped rod accidents are included in Enclosure 1.

Fuel Thermal Conductivity Degradation (TCD) is a physical phenomenon in which the material properties of the fuel (pellets) are affected over the course of in-core operation (burnup),

resulting in a reduced ability to transfer energy from the pellet to the coolant. Consequently, stored energy in the pellets will be higher at burnups when fuel TCD is considered than when fuel TCD is not considered (other effects, such as power falloff with burnup, must also be considered to account for the overall impact of fuel TCD). Appendix B is being added to DPC-NE-3001-P to demonstrate how current Duke Energy non-LOCA safety analysis models and analyses account for this effect.

3.0 TECHNICAL EVALUATION

The technical justification supporting this amendment request is included in the attached revised methodology report sections.

4.0 REGULATORY EVALUATION

4.1 Applicable Regulatory Requirements/Criteria The applicable regulatory requirements for Reactor Design are defined in 10 CFR 50, Appendix A, Criterion 10. This LAR is being submitted in accordance with 10 CFR 50.90.

Page 3 of 5

Enclosure 1 4.2 Precedent By letter dated November 15, 1991, the NRC found methodology report DPC-NE-3001-P to be acceptable for referencing in licensing analyses for the McGuire and Catawba Nuclear Stations (TAC Nos. 75954/75955/75956/75957).

In addition, Duke Energy has submitted and the NRC has approved other methodology reports and revisions to these reports to support reload core design licensing analyses for the McGuire and Catawba Nuclear Stations.

4.3 Significant Hazards Consideration Duke Energy has evaluated whether or not a significant hazards consideration is involved with the proposed amendments by focusing on the three standards set forth in 10 CFR 50.92, "Issuance of amendment," as discussed below:

1. Does the proposed amendment involve a significant increase in the probability or consequences of an accident previously evaluated?

Response: No.

The proposed amendments involving methodology report DPC-NE-3001-P, Multidimensional Reactor Transients and Safety Analysis Physics Parameters Methodology, support the use of revised methodologies for simulating the Updated Final Safety Analysis Report (UFSAR) Chapter 15 events characterized by multidimensional reactor transients, and for systematically confirming that reload physics parameters important to UFSAR Chapter 15 transients and accidents are bounded by values assumed in the licensing analyses. The methodology report revision will be approved by the NRC prior to implementation. The proposed amendments will have no impact upon the probability of occurrence of any design basis accident. The proposed amendments will not affect the performance of any plant equipment used to mitigate the consequences of an analyzed accident. There will be no significant impact on the source term or pathways assumed in accidents previously evaluated. No analysis assumptions will be violated and there will be no adverse effects on offsite or onsite dose as the result of an accident.

Therefore, the proposed amendments do not involve a significant increase in the probability or consequences of an accident previously evaluated.

2. Does the proposed amendment create the possibility of a new or different kind of accident from any accident previously evaluated?

Response: No.

The proposed amendments do not change the methods governing normal plant operation; nor are the methods utilized to respond to plant transients altered. In addition. the proposed methodology changes will not create the potential for any new initiating events or transients to occur in the actual physical plant.

Therefore, the proposed amendments do not create the possibility of a new or different kind of accident from any accident previously evaluated.

Page 4 of 5

Enclosure 1

3. Does the proposed amendment involve a significant reduction in a margin of safety?

Response: No.

Margin of safety is related to the confidence in the ability of the fission product barriers to perform their design functions during and following an accident.

These barriers include the fuel cladding, the reactor coolant system, and the containment system. The proposed methodology revision will assure the acceptability of analytical limits under normal, transient, and accident conditions.

The use of the proposed methodology revision once it has been approved by the NRC will ensure that all applicable design and safety limits are satisfied such that the fission product barriers will continue to perform their design functions.

Therefore, the proposed amendments do not involve a significant reduction in a margin of safety.

Based on the preceding discussion, Duke Energy concludes that the proposed amendments do not involve a significant hazards consideration under the standards set forth in 10 CFR 50.92(c), and, accordingly, a finding of "no significant hazards consideration" is justified.

4.4 Conclusions In conclusion, based on the considerations discussed above, (1) there is reasonable assurance that the health and safety of the public will not be endangered by operation in the proposed manner, (2) such activities will be conducted in compliance with the Commission's regulations, and (3) the issuance of the amendments will not be inimical to the common defense and security or to the health and safety of the public.

5.0 ENVIRONMENTAL CONSIDERATION

Pursuant to 10 CFR 51.22(b), an evaluation of this license amendment request has been performed to determine whether or not it meets the criteria for categorical exclusion set forth in 10 CFR 51.22(c)(9) of the regulations. Implementation of this amendment will have no adverse impact upon the McGuire or Catawba units; neither will it contribute to any additional quantity or type of effluent being available for adverse environmental impact or personnel exposure.

It has been determined there is:

1. No significant hazards consideration;

2. No significant change in the types or significant increase in the amounts of any effluents that may be released offsite; and

3. No significant increase in individual or cumulative occupational radiation exposure.

Therefore, these proposed amendments to the McGuire and Catawba Nuclear Station Renewed FOLs meet the criteria of 10 CFR 51.22(c)(9) for categorical exclusion from an environmental impact statement.

Page 5 of 5

Enclosure 1 ATTACHMENT lb Revised Methodology Report Sections (Non-Proprietary)

Attachment lb Contents:

Appendix D: Proposed Methodology Revisions and Technical Justifications (includes report markups except for Chapter 5)

Appendix E: Chapter 5 Markups Appendix F: Fuel Thermal Conductivity Degradation Impact to Non-LOCA Safety Analysis (new Appendix B to DPC-NE-3001-P)

Appendix D Non-Proprietary Version of Proposed Methodology Revisions to DPC-NE-3001-PA and Technical Justifications D-I

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Front of Report Changes

1. Section: Cover Page

Description:

The cover page is updated to Revision 1 dated October 2013.

McGuire Nuclear Station Catawba Nuclear Station MULTIDIMENSIONAL REACTOR TRANSIENTS AND SAFETY ANALYSIS PHYSICS PARAMETERS METHODOLOGY DPC-NE-3001-A Revision 1 October 2013 Non-Proprietary Nuclear Engineering Division Nuclear Generation Department Duke Energy Carolinas, LLC D-2

Appendix D DPC-NE-3001-P, Revision I Changes (Non-Prop. Version)

2. Section: Abstract

Description:

The abstract is revised by adding the following new paragraph to sunmmarize the contents of Revision 1.

Revision 1 Revision 1 includes a new methodology for modeling the main steam line break transient from the full power initial condition, revisions to the rod ejection methodology, main steam line break methodology at the HZP initial condition, and a revised single failure assumption for the dropped rod transient. Input to the Rod Control System for the dropped rod transient is based on the newly installed Distributed Control System which replaces the 7300 Process Control System. The CASMO-4 code replaces the CASMO-3 code for generating nuclear data for SIMULATE-3K, and the SIMULATE-3K code is revised. The regulatory issue associated with fuel conductivity degradation is also addressed. Other revisions include enhancements and clarifications that have been identified based on experience gained in applying the original methodologies, or are error corrections.

3. Section: Table of Contents

Description:

The Table of Contents is revised to reflect the following changes associated with Revision 1.

Chapter 5

• Section 5.3.2.2 Response Times (HZP Case) is renamed

• Section 5.3.2.5 Boron Injection Modeling (HZP Case) is renamed

• Section 5.3.2.6 Core Kinetics Modeling (HZP Case) is renamed

• Section 5.3.2.7 Core Kinetics Modeling (HFP Case) is new

" Section 5.4 Results and Conclusions (HZP Case) is renamed

• Section 5.4.2.2 Minimum DNBR and CFM Results is renamed

" Section 5.5 Results and Conclusions (HFP Case) is new

" Section 5.6 Cycle Specific Evaluation is renumberedfirom 5.5 Appendix B - Fuel TCD Impact to Non-LOCA Safety Analysis List of Figures 5-28 Main Steam Line Break I-FP Steam Line Pressure 5-29 Main Steam Line Break HFP Cold Leg Temperature 5-30 Main Steam Line Break HFP Power and Indicated Power 5-31 Main Steam Line Break HFP Minimum DNBR 5-32 Main Steam Line Break HFP Maximum Allowable Radial Peaks List of Tables 5-3 Sequence of Events for 4.9 ft2 Split Break Full Power Initial Condition D-3

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Section 1.0 Changes - Introduction 1-1 Chapter 1.0 Introduction (p. 1-1)

Description:

Reference to the thermal power level of McGuire and Catawba is removed since this power level is subject to change.

This report is applicable to the McGuire and Catawba Nuclear Stations, which are 344-MWt 4 loop Westinghouse units.

Technical Justification: The reference power level of 3411 MWth is removed because it is subject to change and is not needed to define report applicability.

1-2 Chapter 1.0 Introduction (p. 1-1) and Figure 1

Description:

The paragraph describing the mitigating options when one or more plant parameters assumed in the reference analysis is found to be non-conservative in the reload analysis is revised to clarify that the non-conservatism can be dispositioned via reanalysis, evaluation or by revising a loading pattern.

If, however, one or more of the plant parameters assumed in the reference analysis are found to be non-conservative for the reload cycle, those accident analyses which are affected by the non-conservative parameters must be reevaluated, reanalyzed, or the loading pattern must be revised to demonstrate applicable acceptance criteria are met for the reload core.

Technical Justification: This revision clarifies that if a key parameter, or parameters, is not bounded by the reference analysis, acceptability of the applicable acceptance criteria can be demonstrated through evaluation or re-analysis. Evaluation is the process where the non-conservative parameters can be accommodated by an evaluation or limited scope analysis, whereas, re-analysis consists of rerunning the accident or transient analysis in question and updating the licensing basis analysis. If applicable acceptance criteria cannot be met through evaluation or re-analysis, then the core loading pattern must be revised.

1-3 Chapter 1.0 Introduction (p. 1-3)

Description:

The paragraph describing the steam line break analysis methodology is revised to include a description of the new methodology for the hot full power initial condition case, and the centerline fuel melt (CFM) check is added for the HZP case.

The steam line break accident analysis methodology is presented in Chapter 5. The system thermal-hydraulic analysis is performed with RETRAN-02. Separate methodologies for the zero power initial condition and the full power initial condition are presented. For the zero power initial condition the limiting TFhe w.rst case scenar.io, whi.h e..urs at zero power at end-of-cycle case is presented.

Cases both with and without offsite power are analyzed. The core power peaking at the return-to-power statepoint condition and including the worst stuck rod is determined.

D-4

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

The approach to DNBR is then predicted with VIPRE-0 1. The results sho':' that the DNBR lim.it is not ex.eeded for the limiting eases. The return-to-power statepoint is also evaluated to demonstrate centerline fuel melt (CFM) margin. The results show that DNBR and CFM limits are not exceeded for the limiting cases. For the full power initial condition reactor power increases due to the decrease in moderator temperature until a trip setpoint terminates the event. The limiting statepoint during the power excursion is evaluated against the DNBR and centerline fuel melt acceptance criteria. The results show that these limits are not exceeded for the limiting break size, location, and time in cycle. It is noted that the steam line break from a full power initial condition is not an asymmetric event, unlike the other events described in this report.

Technical Justification: Both the DNBR and centerline fuel melt (CFM) acceptance criteria must be satisfied for the steam line break accident. Additional details concerning these changes are summnarized in the Section 5 (Steam Line Break changes).

1-4 Chapter 1.0 Introduction (p. 1-4 References)

Description:

The list of references is revised.

1-1 Nuclear Physics Methodology for Reload Design, DPC-NF-20 10-A, Revision 2a, Duke Energy Carolinas, February December 2009 1-2 Thermal-Hydraulic Transient Analysis Methodology, DPC-NE-3000-PA, Revision 35a, Duke Energy Carolinas, Septeniber-2004 October 2012 1-3 SIMULATE-3K Models & Methodology, SSP-98/13, Revision 06, Studsvik Scandpower, JM:!y- 998, January 2009 Technical Justification: References 1-1 (DPC-NF-2010-A) and 1-2 (DPC-NE-3000-PA) are updated to the current versions of these reports. Reference 1-3 is updated to reflect the methodology associated with version 2 of the SIMULATE-3K code.

D-5

Appendix D DPC-NE-3001-P, Revision I Changes (Non-Prop. Version)

Figure 1 Reload Core Safety Analysis Verification Process All values conservative D-6

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Section 2.0 Changes - Determination of Safety Analysis Physics Parameters 2-1 Section 2.3.5 Steam System Piping Failure (15.1.5) (p. 2-4)

Description:

The following paragraph is expanded to describe the key parameters for the new methodology for the steam line break from a hot full power initial condition.

Steam Line Break Initiated From a Zero Power Initial Condition (15.1.5)

This transient is initiated by a rupture of a main steam line at zero power initial conditions. The resulting primary system overcooling causes a loss of core shutdown and a return-to-power condition occurs. This transient is analyzed by assuming a conservatively large reactivity insertion as the core cools down. The power increase is exacerbated by assuming a least negative Doppler coefficient. The boron concentration in the safety injection flowpath and the boron worth are both minimized. Due to the assumption of a stuck rod, the core power distribution at the limiting statepoint will be highly peaked. Consequently, the core power distribution must be evaluated to quantify the number of fuel pins exceeding the DNBR limit.

Steam Line Break Initiated From a Full Power Initial Condition (15.1.5)

The transient is initiated by a rupture of a main steam line at full power initial conditions. The decrease in moderator temperature causes an increase in reactivity and core power. Reactor power increases until a trip setpoint is reached, the control rods insert, and the event is terminated. The limiting case is determined by analyzing a range of break sizes and times-in-cycle. Kinetics parameters and moderator temperature coefficients for the full range of times-in-cycle are evaluated. The core power distribution at the limiting statepoint, consistent with the time-in-cycle, must be evaluated to determine the margin to the DNBR and centerline fuel melt limits. Minimum Doppler feedback consistent with the time in life of the MTC is necessary for a conservative analysis, and is therefore a key parameter.

Table 2-1 Revision:

2.3.5 Steam line break (HZP case) 2.3.5 Steam line break (HFP case) 15.1.5

• MTC

• BOC to EOC values

• DTC

• Least negative Technical Justification: The new methodology for the steam line break initiated from a full power initial condition has determined that a range of break sizes and times-in-cycle must be analyzed to identify the limiting case. This occurs because for different cases analyzed different Reactor Protection System trip functions come into play. As a result, a range of kinetics parameter values and moderator temperature coefficients that span BOC to EOC must be considered. The least negative Doppler temperature coefficient is assumed, consistent with the time in life of the MTC. The resulting core power distribution for the limiting cases and times-in-cycle are then evaluated for the DNBR and centerline fuel melt limits to ensure that positive margin exists to each of the limits.

D-7

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) 2-2 Section 2.3.23, Moderator Dilution Accident (15.4.6) (p. 2-8)

Description:

The last sentence is revised to include an alternative equivalent approach to including a conservative allowance.

The boron worth used to convert the shutdown margin to ppmb should be conservatively large, or if a best estimate boron worth is used to convert the shutdown margin to ppmb, an allowance is applied to the initial boron concentration to maintain conservatism.

Technical Justification: The intent of the original statement is to ensure that the value of the differential boron worth is selected in a conservative manner to ensure overall conservatism of the methodology. Application of the methodology has shown that it is more convenient to include an allowance in the initial boron concentration rather than to adjust the differential boron worth to a conservative value. A conservative allowance in the initial boron concentration is used to justify using a best estimate differential boron worth. This approach is essentially equivalent to the original approach.

2-3 Section 2.4 Reload Cycle Evaluation (p. 2-10)

Description:

The second sentence in the first paragraph and the last sentence is the second paragraph are revised to clarify that accidents for which physics parameters are not bounded or thermal limits exceeded are evaluated, reanalyzed or the core is redesigned to ensure acceptable accident consequences.

The important physics parameters in Table 2-1 are evaluated each reload cycle to ensure that values assumed in the current licensing analyses bound the reload core.

Accidents for which the physics parameters are not bounded would be reevaluated or re-analyzed to ensure acceptable accident consequences or, the core would be redesigned so the physics parameters fall within the limits assumed in the reference analysis.

Accidents for which either thermal limit acceptance criteria or technical specification limits are exceeded would be reevaluated or reanalyzed to ensure acceptable accident consequences, or the core would be redesigned to obtain acceptable peaking factors and accident consequences.

Technical Justification: Refer to the technical justification for change 1-2.

2-4 Section 2 References (p. 2-10)

Description:

The list of references is revised.

2-4 SIMULATE-3K Models & Methodology, SSP-98/13, Revision 06, Studsvik Scandpower, July 1999, January 2009 Technical Justification: Reference 2-4 is updated to reflect version 2 of the SIMULATE-3K code.

D-8

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Section 3.0 Changes - Determination of Safety Analysis Physics Parameters 3-1 Section 2 References (p. 3-6)

Description:

Reference 3-2 is revised.

3-2 Nuclear Design Methodology for Core Operating Limits of Westinghouse Reactors, DPC-NE-2011-PA, Revision la, Duke Power, December 20 June 2009 Technical Justification: Reference 2-4 is updated to correct the report date.

D-9

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Section 4.0 Changes - Rod Ejection Accident Analysis 4-1 Section 4.2.1.2, SIMULATE-3K (p 4 -3 )

Description:

The SlIMULATE-3K version that is used for the rod ejection analysis includes new models. CASMO-3 is replaced with CASMO-4 for the generation of cross sections and nuclear data for use in SIMULATE-3K.

The SIMULATE-3K code (Reference 4-4) is a three-dimensional transient neutronic version of the SIMULATE-3 code (Reference 4-8). SIMULATE-3K uses the QPANDA full two-group nodal spatial model developed in SIMULATE-3, with the addition of six delayed neutron groups. The program employs a fully-implicit time integration of the neutron flux, delayed neutron precursor, and heat conduction models. Subcritical neutron sources may be modeled. Decay heat is based on the ANSI/ANS 5.1-1994 standard. Delayed neutron fractions are fully functionalized similar to other cross sections to provide an accurate value of beta for the time-varying neutron flux. The control of time step size may be determined either as an automated feature of the program or by user input. Use of the automated feature allows the program to utilize larger time steps (which may be restricted to a maximum size based on user input) at times when the solution is changing slowly and smaller time steps when the solution is changing rapidly.

Additional capability is provided in the form of modeling a reactor trip. The trip may be initiated at a specific time in the transient or following a specified excore detector response. Use of the excore detector response model to initiate the trip allows the user to specify the response of individual detectors as required to initiate the trip, as well as the time delay prior to release of the control rods. The model also allows for the impact of coolant density changes to be accounted for at the option of the user. The velocity of the control rod movement is also controlled by user input.

The SIMULATE-3K thermal-hydraulic model includes a spatial heat conduction model and a 5-equation hydraulic channel model. The heat conduction model solves the conduction equation on a multi-region mesh in cylindrical coordinates.

Temperature-dependent values may be employed for the heat capacity, thermal conductivity, and gap conductance. Burnup dependent models may be employed for thermal conductivity gap conductance, and the pellet radial power profile.

A single characteristic pin conduction calculation is performed consistent with the radial neutronic node geometry, with an optional calculation of the peak pin behavior available to monitor local maxima. A single characteristic hydraulic channel calculation is performed based on the radial neutronic node geometry. The model allows for direct moderator heating at the option of the user. This thermal-hydraulic model is used to determine fuel temperature and moderator density for updating the cross-sections, and may additionally be used to provide edits of fuel temperature and enthalpy throughout the transient.

The SIMULATE-3K program utilizes a CASMO-34 (Reference 4-24-7) cross-section library and reads a CASMO--34/SIMULATE-3 (Reference 4-2-09) restart file (exposure and burnup-related information). Executed in the static or transient mode, SIMULATE-3K performs the same solution techniques, pin power reconstruction, and cross -section development as SIMULATE-3.

D-10

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Technical Justification: SIMULATE-3K version 1 was reviewed and approved for performing rod ejection analyses by the NRC in Reference 4-17 (SIMULATE-3K analyses are documented in Section 6.6 of this reference). Version 2 (Reference 4-4) is used in this revision. The following new models are available in Version 2:

• 5-equation hydraulics model

• Bumup-dependent gap conductance

• Burnup-dependent fuel pellet conductivity

• Burnup-dependent pellet radial power profile

• ANSFANS-5.1-1994 decay heat model

• Subcritical neutron sources

• Option to account for coolant density changes in the excore detector model

• Explicit pin-by-pin conduction model

• Waterside corrosion model

• Revised CHF correlation Each of these new models is considered to be an enhancement to the version 1 code with the intent of obtaining better simulation of the physical processes. These are all standard models in version 2. Version 2 is applied in Revision 1 applications with the intent of improving accuracy by replacing the more simplistic models that existed in Version 1. In addition, CASMO-4 replaces CASMO-3 as the cross section and nuclear data source for SIMULATE-3K. This change provides consistency in the cross section and nuclear data sources used for the steady state SIMULATE-3 model and the transient SJMULATE-3K model. NRC approval of the steady state CASMO-4 based SIMULATE-3 model was obtained in Reference 4-9.

Demonstration SIMULATE-3K transient analysis results using version 2 of the code are presented in Figures 1 and 2 for rod ejection cases analyzed at BOC HFP and HZP, and EOC HFP and HZP conditions. Initial conditions assumed for the ejected rod worth, Doppler temperature coefficient and moderator temperature coefficient are shown in Table 4-1. Values for each parameter were selected to conservatively bound expected reload values. The ejected rod worth includes conservatism in both beta-effective and rod worth.

Table 4-1 Rod Ejection Analysis Initial Condition Physics Parameters Parameter BOC HFP BOC HZP EOC HFP EOC HZP Ejected Rod Worth($) 0.188 1.321 0.255 1.447 DTC (pcm/°F) -1.15 -1.45 -1.30 -1.60 MTC (pcm/°F) -0.29 7.0 -30.0 -15.0 The HFP calculations were initiated with the Control Bank D at its HFP rod insertion limit less 12 steps to account for the bank being mis-positioned. Initial reactor power was set at rated thermal power plus uncertainty.

The HFP transient is initiated with the ejection of a Control Bank D rod at core location D-12. The rate of rod ejection is scaled to be consistent with an ejection time of 0.1 seconds from fully inserted to fully withdrawn. Core power increases rapidly as the rod is ejected and continues to increase until the Doppler feedback offsets the positive D-1 1

Appendix D DPC-NE-3001-P, Revision I Changes (Non-Prop. Version) reactivity insertion from the ejected rod. Fuel temperature continues to rise until it reaches an equilibrium value. Consequently, power decreases due to increased Doppler feedback and then decreases more rapidly after Control and Shutdown Banks begin to insert. The initial rapid power excursion produces a reactor trip on positive flux rate.

Control and Shutdown Banks begin to insert after the peak power level is achieved due to trip delay. The HFP BOC and EOC core average power response is presented in Figure 4-1 based on the physics inputs from Table 4-1. The BOC power response is less peaked due to the lower ejected rod worth.

HZP rod ejection calculations are performed with the reactor critical at a very low power level with Control Banks positioned at their HZP rod insertion limits. Control Bank D is the only bank fully inserted. The HZP transient is initiated with the ejection of the Control Bank D rod at core location D- 12 from its fully inserted to fully withdrawn position in 0.1 seconds. Because the worth of the rod is greater than one dollar, the reactor becomes prompt critical resulting in a rapid power excursion before Doppler feedback turns the power excursion around. Power decreases rapidly as the fuel continues to heat-up until near-equilibrium conditions are reached. The reactor is tripped by the positive rate trip and is shutdown after a trip delay as rods are inserted. The power response for the EOC HZP case is more severe than the BOC case because of the larger ejected rod worth. Figure 4-2 shows the HZP BOC and EOC core average power response based on the physics inputs from Table 4-1.

The results presented in Figures 4-1 and 4-2 are typical of bounding analysis results that would be presented in the UFSAR for HFP and HZP rod ejections.

Figure 4-1 HFP BOC and EOC Core Average Power Versus Time 140.0 120.0 100.0 EOC 80.0 60.0 40.0 20.0 0.0 0.0 0.5 1.0 1.5 2.0 2.5 3,0 315 4.0 nime (seac D-12

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Figure 4-2 HZP BOC and EOC Core Average Power Versus Time 500.0 I=*uf" 450,0 400.0 350Z0 300.0 BOc 250.0 S200.0 150.0 100.0 Rod Inser~tIon Strt 50.0 0.0 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 Time (sec)

Results from the SIMULATE-3K analysis are used subsequent calculations to confirm the acceptability DNBR, fuel enthalpy and reactor coolant system pressure limits are shown in Figure 4-3.

D-13

Appendix D DPC-NE-3001-P, Revision I Changes (Non-Prop. Version)

Figure 4-3 Rod Ejection Accident Analysis Flow Diagram SIMULATE-3 CASMIO-4 Key Physics Parameters Initial Conditions r - Cross Sections and (Ejected rod worth, beta Few Group Constants (Initial power level, flow, pressure, inlet temperature, effective, Doppler and Moderator rod position, etc..)

temperature coefficients)

SIMULATE-3K Transient Neutmn Power Transient Neutron Power Power Distribution

............. ... . .. . .. . . . . . . . . . . . . - . .-. .

VIPREVIPRE Volumetric Coolant DNB J Expansion Flow Rates

.................... ....

L -..................... . ....... .. .............. .......................... .... ........

Allowable RETRAN Peaking Pressure Response

.. .... -.............. -......... -.... -....-....-...... .........

SIMULATE-3 Cycle-Specific Checks (Key Physics Parameters and Post Ejected Power Distribution)

Failure of Cycle 1 Specific Check I

-- -- - - -- -----------------------------

D- 14

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) 4-2 Section 4.2.1.3 SIMULATE-3K Code Verification (p. 4-4)

Description:

The code validation description is updated to be consistent with version 2 of the code.

The SIMULATE-3K code has been benchmarked against many numerical steady state and transient benchmark problems by the code vendor. The results of these benchmarks are for light water reactor reactivity initiated transients are described in Reference 4-413, and show excellent agreement between SIMULATE-3K and the reference solutions. Some of the SIMULATE-3K benchmarks which have been performed are as follows. : The fuel conduction and thermal hydraulics model has becn benchmfar-ked against the TRXC code (Refer-ence 4 10). The transient neutronics model has been benchmarked, using standard LWR problems, to reference solutions generated by QUANDRY (Reference 4-11), SPANDEX (Referencc 1 12), NEM (Refercncc 13), and CUBBOX (Reference 4-14). Finaly, a benchmark of the couipled per-formance of thc transient neutronics anid thermfal hydr-aulic models was provided by cmaioof results from a standard NEACRP rod eje p*rbem to*, and the PANTHER codes (Reference 4-15). The fuel conduction and thermal-hydraulics models have been benchmarked against the TRAC code (Reference 4-10) as described in Reference 4-12. Steady-state components of the SIMULATE-3K model are implemented consistent with the CASMO-4/SIMULATE-3 methodology and performance benchmarks which were approved for use in Reference 4-9.

Technical Justification: The code validation description is updated to be consistent with version 2 of the code. Model descriptions and the benchmark of individual and coupled models are described in References 4-12 and 4-13. The results of these benchmarks show excellent agreement between SIMULATE-3K and the reference solutions. A demonstration analysis using version 2 of the code is shown as part of the technical justification for change 4-1.

4-3 Section 4.2.2.2, Fuel Temperature and Enthalpy Calculation (pp. 4-5 and 4-6)

Description:

The calculation of fuel temperatures and enthalpy will use a single channel VIPRE-01 model as an alternative to the [ ] VIPRE-01 model.

Page 4-6: The [ I model with fuel conduction (Figure 4-2) is used to simulate the peak fuel pin in the hot assembly during the transient.

Technical Justification: The current methodology uses the [ I channel VIPRE-01 model to calculate the fuel temperatures and enthalpy. Based on additional experience it is also acceptable to use a [ ] for this purpose. There are no important modeling aspects associated with any interactions with adjacent channels so the additional details involved with applying the [ ] channel model are not necessary. Therefore, the use of either model is appropriate.

D-15

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) 4-4 Section 4.2.2.2, Fuel Temperature and Enthalpy Calculation (Fuel Conduction Model)

(p. 4-7)

Description:

The specificity of the fuel performance code used to match gas gap conductivity is removed.

Technical Justification: The fuel performance code used to determine gas gap conductivity is dependent upon fuel type being irradiated. The PAD fuel performance code is applicable for the Westinghouse RFA fuel product (which is currently being irradiated) but not for other vendors fuel products.

4-5 Section 4.2.2.2, Fuel Temperature and Enthalpy Calculation (Fuel Enthalpy Calculation)

(p. 4-9)

Description:

The VIPRE-01 code has been modified to calculate UO2 enthalpy as a function of temperature.

VIPRE-01 is used to does-net perform a fuel enthalpy calculation. Thus The fuel enthalpy for a given fuel temperature during the transient is calculated separately-40m NlP-RE-based on the equation obtained from MATPRO (Reference 4-23).

Technical Justification: In the original methodology the U0 2 enthalpy value was calculated manually based on VIPRE-01 temperature results using the equation from MATPRO. VIPRE-01 has been modified to do this calculation internal to the code using the same equation from MATPRO. This revision is a user convenience.

4-6 Section 4.2.2.4, DNBR Evaluation (Fuel Conduction Model) (p. 4-16)

Description:

The gap closure time is [

]

[

]]

D-16

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) 4-7 Section 4.2.2.4, DNBR Evaluation (Fuel Conduction Model) (p. 4-16)

Description:

The value for the [

I Technical Justification: I I

4-8 Section 4.3.1, Initial Conditions (p. 4-18)

Description:

[

I based on additional experience with applying the methodology.

I I

Technical Justification: The original methodology included a modeling approach that was intended to introduce conservatism that would ensure that future core designs would remain bounded for the rod ejection accident. [

D-17

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) 4-9 Section 4.3.1, Initial Conditions (p. 4-19)

Description:

As an additional method, the moderator temperature coefficient at beginning-of-cycle is adjusted to a conservative value by changing the core boron concentration.

As an additional method for adjusting the MTC at BOC, the MTC can be adjusted to a conservative value by changing the boron concentration.

Technical Justification: The current methodology for adjusting the moderator temperature coefficient uses a multiplier on the cross sections. Changing the boron concentration is an additional method available for adjusting the MTC to a conservative value for beginning-of-cycle cases.

4-10 Section 4.3.1, Initial Conditions (p. 4-20)

Description:

The logic for only modeling the D-12 core location is updated.

The ejected control rod is located at core location D-12 or H-08. Figure 4-4 shows these this locations in the reactor. The control rod in location D-12 is part of Control Bank D (hereafter referred to as Bank D). At the HZP rod insertion limit, Bank D is the only bank fully inserted. At the HFP insertion limit, it is the only bank in the core. For the HFP rod ejection cases, the ejected rods at core locations H-08 and D-12 are evaluated. For the HZP rod ejection cases, the central control rod (location H-08, also a member of Bank D) is not chosen as the ejected rod because sensitivity studies have shown that for a given ejected rod worth at core location D-12, the corresponding ejected rod worth at core location H-0 issignificantly lower due to the asymmetry produced when D-12 is ejected. This reduction in ejected rod worth results in a decrease in the transient response. For this reason the control rod at core location H-08 is not evaluated at HZP conditions.

Technical Justification: In the original methodology, the HFP rod ejection case was only evaluated with an ejected rod at core location D-12. Core location H-08 was not evaluated because the D- 12 case resulted in a higher post ejected transient Fq relative to the H-08 case. In addition, the ejected rod worth in core location D-12 in most instances was higher than the ejected rod worth at core location H-08 which would lead to a more severe transient response. However, additional experience with the methodology has indicated that Fq alone is not always a good indicator of whether or not a case is conservative because the calculation of both the DNB ratio and fuel enthalpy are dependent upon the integrated energy deposited in the fuel over the course of the transient. As a result, the value of post-ejected peaking factor Fq is not an absolute indicator of the level of conservatism in the transient. The ejected rod worth in dollars and Doppler temperature coefficient are the two most important parameters. In addition, D-18

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) the evolution of core designs has resulted in the HFP ejected rod worth at H-08 approaching and in some instances exceeding the ejected rod worth at D-12. As a result, HFP rod ejection calculations will be performed at both the H-08 and D-12 core locations in future applications of the methodology. The HZP rod ejection cases will continue to be performed for the D-12 core location only because the ejected rod worth at core location D-12 remains higher than the ejected rod worth at core location H-08. To confirm ejected rod worth input assumptions in the reference analysis remain valid, reload checks are performed to evaluate the worth of ejected control rods from core locations H-08, D-12, and from control rods in control banks that are not fully inserted in the reactor core at the HZP initial condition.

4-11 4.3.1, Initial Conditions (p.4-21)

Description:

The modeling of the motion of the ejection rod with constant velocity or with constant acceleration is deleted. The ejection time of 0.058 seconds is deleted.

The control rod is ejected in 0.1 seconds at constant velocity for the HZP cases.

The control rod at D-12 is ejected in 0.058 scoe at constant acceleration proportional to the fully inserted to fully withdrawn. This time is consistent with the-ejection time of 0.1 seconds used in the HZP cases.

Technical Justification: The original methodology specified that the motion of the ejected rod should be modeled with constant velocity for the HZP cases, and with constant acceleration for the HFP cases. The important aspect of the modeling of the ejected rod is the ejection time of 0.1 seconds for a fully inserted rod to move from an all rods in condition to an all rods out condition. The additional detail regarding constant velocity or constant acceleration is excessive detail and is deleted. The ejection time is deleted and replaced with the basis for the calculation of this ejection time. The deletion of the HFP ejection time avoids the condition where the distance between the rod insertion limit and top of the active fuel region changes, and necessitates a change in the full power ejection time. Changes in the distance between the rod insertion limit and top of the active fuel region can result from a vendor design change or implementation of a new fuel product.

4-12 Section 4.3.1, Initial Conditions (p. 4-21)

Description:

The statement that sensitivity studies show that the peak power level attained during the transient is slightly higher for faster ejection time is removed.

The control rod is ejected in 0. 1 seconds at constant velocity for the HZP cases. This is significantly faster than physically possible, even when friction is ignored in the ejection time calculation. Sensitivity studies show that the peak power level attained during the transient is slightly higher-(more consenvative) for-a faster- ejectien time.

Thus, the control rod ejection results are conservative with respect to control rod ejection time.

Technical Justification: Additional experience with the methodology has shown that for rod ejection times around the 0.1 second threshold, changes in the ejected rod speed primarily affect the time of the peak power, but has little or no impact on the magnitude of the peak power level attained. As a result, this statement is removed as excessive detail.

D-19

Appendix D DPC-NE-3001-P, Revision I Changes (Non-Prop. Version)

The important aspect of the methodology is the assumption that the rod is ejected in 0. 1 seconds.

4-13 Section 4.3.1, Initial Conditions (p. 4-21)

Description:

The initial condition power level is changed to reflect a change in power level uncertainty.

For the HFP statepoints, the reactor is initially at 102% ef rated power plus power level uncertainty with Bank D at 149 swd.

Technical Justification: The power level uncertainty has been reduced from 2% to approximately 0.3% as part of the measurement uncertainty recapture (MUR) uprate at McGuire and Catawba. Application of the appropriate power level uncertainty will be applied to the rated thermal power condition.

4-14 Section 4.3.2, Boundary Conditions (p. 4-21)

Description:

The methodology is revised to credit the power range high flux positive rate trip function. Specific values for the high flux reactor trip signals are deleted.

The reactor trip signal is generated when the third highest excore channel reaches either [12%Jf-eo- the HZP eases or Ill ;1 f* r the W-IFP eases its high flux or high flux positive rate trip setpoint. This modeling is based on a single failure of the highest channel and a two-out-of-the-remaining-three trip coincidence logic.

Technical Justification: The methodology is revised to credit both the high flux and high flux positive rate trip functions. Specific values for the high flux trip function are deleted to avoid a future situation where the numerical value changes and then the methodology report requires revision. The methodology remains the same without the numerical values.

4-15 Section 4.3.2, Boundary Conditions (pp. 4-21 and 4-22)

Description:

The modeling of the negative reactivity insertion associated with the insertion of the control rods following reactor trip is simplified as an alternative.

During the reactor trip the ejected rod and a second rod with the highest worth are assumed not to fall into the reactor. To conservatively model the reactor trip, not all of the control rod banks are allowed to drop, and some of the banks that are dropped have their worth reduced by a cross section adjustment. The rod worth adjustment is made in SIMULATE-3K by [

]. Also, negative reactivity inserted due to the reactor trip is not allowed to exceed the conservative trip reactivity curve. The integral worth of the falling control rods is computed for several different axial positions of the rods at the initial conditions.

An alternative approach to this detailed modeling is to explicitly model the reactor trip assuming both top and bottom peaked xenon distributions. The reactor trip assumes that the ejected rod and an adjacent control rod do not fall into the reactor core. The worth of the D-20

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) remaining control banks that drop are reduced by a cross section adjustment. The trip insertion time is assumed to be at the technical specification limit.

Technical Justification: The original methodology specifies in detail how the negative reactivity associated with the insertion of the control rod following reactor trip (i.e. the trip reactivity curve) is modeled. The important concept is that the start of the negative reactivity insertion is conservatively delayed, and that the rate of negative reactivity insertion is conservatively small relative to the rod insertion time and the differential worth of the control rods less the worth of a high worth stuck rod and ejected rod. In the alternative approach, the negative reactivity inserted following a reactor trip is reduced by assuming both the ejected rod, and an adjacent rod do not fall into the core, and by reducing the worth of the control rods that do fall into the reactor core by a cross section adjustment.

The rate of negative reactivity insertion is determined by the initial condition xenon distribution (either top or bottom peaked), and by the technical specification rod insertion time. This more realistic approach results in a negative reactivity insertion that is slower for the top peaked xenon case relative to the bottom peaked case. The explicit modeling of the axial shapes removes the need to model the trip reactivity curve. Both concepts maintain a conservative approach to modeling control rod insertion following reactor trip.

4-16 Section 4.4.4, DNBR and Fuel Pin Census (p. 4-26)

Description:

Reference to the offsite failed fuel fraction limit is removed and replaced with a statement that the limit for the number of fuel experiencing DNB is set to a value that maintains offsite dose consequences within applicable regulatory limits.

The above results show that the HFP, BOC case has the largest number of pins experiencing DNB. The limit for the number of fuel pins experiencing DNB is set to a value that maintains offsite dose consequences within applicable regulatory limits. off-ite dose enseguenees are analyzed based on 50-% of the fuel pin expeiening DNB to eonservatively bound the above results.

Technical Justification: Failed fuel fractions that must be satisfied to meet applicable regulatory dose limits are established based on the methodology defined in Chapter 15 of the UFSAR. The specific failed fuel fraction value is subject to change based on changes in source term, plant modification or other methodology changes. As a result, the specific value is eliminated. The important aspect of the analysis remains unchanged in that the rod ejection accident results must satisfy the dose related acceptance criterion.

4-17 Section 4.5.2, Boundary Conditions (Pressurizer Safety Valve Modeling) (p. 4-29)

Description:

The pressurizer safety valve blowdown value is decreased from 5% to 1%.

The pressurizer code safety valves function as overpressure mitigation equipment. The nominal lift setpoint is increased by 3% to account for calibration allowance. The valves are assumed to open linearly until they are fully open at a pressure 3% above the adjusted lift setpoint. The valve modeling then includes a hysteresis effect that keeps the valves fully open until the pressure decreases to 1% 51% below the adjusted lift setpoint.

D-21

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Technical Justification: The pressurizer safety valve blowdown value determines the pressure at which the valve re-closes. The blowdown is decreased from 5% to 1% in this revision. The 5% value is more of a typical value, whereas the 1% value is a conservative minimum value, which is consistent with the intent of a conservative analysis.

4-18 Section 4.5.2, Boundary Conditions (Pressurizer Safety Valve Modeling) (p. 4-29)

Description:

An alternative method is added to allow for crediting a pop-open model approach rather than a linear ramping open approach.

Alternatively, the valve can be assumed to pop-open to the full open position in 0.5 seconds after the drifted lift setpoint is reached.

Technical Justification: The pop-open model approach rather than a linear ramping open approach is added as an alternative modeling approach to be consistent with modeling of other UFSAR chapter 15 accidents. This approach was approved by NRC in Revision 2 to DPC-NE-3002-A, dated April 26, 1996.

4-19 Section 4.6, Cycle Specific Evaluation (p. 4-29)

Description:

Combine the ejected rod worth and Beta-effective into a single parameter, ejected rod worth in dollars ($), in Tables 2-1 and Table 4-4, and add a footnote to Table 4-4.

Table 2-1 Summary of Safety Analysis Physics Parameters Report Transient Or FSAR Conservative Section Accident Section Key Parameters Direction 2.3.24 Rod ejection 15.4.8

• MTC 0 Most positive

• DTC 0 Least negative

• Ejected rod worth ($)

• Maximum SBet effee Mirim'

• Core power distribution 9 Maximum total peak with ejected rod 0 Maximize number of pins in DNB D-22

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Table 4-4 Rod Ejection Reload Checklist*

Parameter BOC HFP BOC HZP EOC HFP EOC HZP Ejected Rod less than 2-00 0.364 7-20 1.307 2200 0.50 900 2.25 Worth (pem $)

g r-eate-tha 0005 0.00554 0.004 0.004 DTC (pcnrd°F) less than -0.9 -0.9 -1.2 -1.2 MTC (pcim/'F) less than 0.0 +7.0 -10.0 -10.0 DNB Census less than 50% 50% 50% 50%

FQ less than .4.12 16.62 4.88 23.55

• The values in this table are for illustrative purposes only. Actual analysis-of-record values are contained in the Updated Final Safety Analysis Report.

Technical Justification: The rod ejection methodology utilizes conservative initial conditions and boundary conditions as a means to introduce conservatism into rod ejection transient analysis results at the HFP and HZP initial conditions. The results from the transient analysis are used to confirm that the peak coolant system pressure, radial average fuel peak enthalpy and the percentage of the fuel rods that exceed the DNBR limit satisfy the acceptance criteria for each parameter. Current reload checks compare cycle-specific values of ejected rod worth and Beta-effective against the ejected rod worth and Beta-effective parameters used in the rod ejection reference (licensing) analysis. However, the ejected rod worth in dollars (calculated by dividing worth by Beta-effective) is the key parameter that determines the severity of the transient. Neither the ejected rod worth nor Beta-effective, by themselves, completely determine the transient response. Therefore, the reload check is updated to perform a cycle-specific check of the ejected rod worth in dollars. A footnote is added to Table 4-4 for clarification stating that the values in this table are for illustrative purposes only, and that analysis of record values can be found in the updated final safety evaluation report.

4-20 References (pp. 4-30a and 4-30b)

Description:

The list of references is revised.

4-4 SIMULATE 3 Kineti* s Ther..y and Moedel Des.ription, 9SA 96/26 3K Models & Methodology, SSP-98/13, Revision 06, Studsvik of-Ameriea, April 1996 Scandpower, January 2009 4-8 SIMULATE-3 Methodology, Advanced Three Dimensional Two-Group Reactor Analysis Code, SOA-95/4418, Studsvik of America, Oetobei: 1995 4-12 Development of a Variable Time Step Transient bým Code: SPANEX, Trans. Am. Nuel. Soe. 68 125, B. N. Aviles, 1993 SIMULATE-3 Kinetics and Theory And Model Description, SOA-96/26, Studsvik of America D-23

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) 4-13 NEM: A Thfee Di n .ensionai T ransient e'".tr-.m.s Koutine tor the ,.%

W

• PF1 Reacter Thermal Hydr-aulic Compuiter-Code, B. R. Bandini, Pennsylvania State U .versity, 1990 LWR Core Reactivity Transients.

SIMULATE-3K Models and Assessment, SSP-04/443, Revision 2, Studsvik Scandpower 4-17 Duke Power Company Westinghouse Fuel Transition Report, DPC-NE-2009-PA, Revision 23a, December- 200 September 2011 4-19 Thermal-Hydraulic Transient Analysis Methodology, DPC-NE-3000-PA, Revision -35a, Duke Power Company, September2004 October 2012 4-20 Nuelear Design Methodolegy Using CASMO4 3/SIhLULATE 312, DPCG NE 1001 A, Revision 1, Duke Energy Carolinas, Januar.y 2009 Intentionally Blank 4-21 CASMO 3: A Fuel Assembly Bumup Program User's Manual, STeUPSVntion A 88/48, Studsvik of Aleriea, September 1988.

Intentionally Blank Technical Justification: Reference 4-4 is updated to reflect the current version of the SIMULATE-3K methodology report. Reference 4-8 is updated to reflect the current version of SIMULATE-3 methodology report. References 4-12 and 4-13 are replaced per change 4-2. References 4-17 and 4-19 are updated to the current versions of each methodology report. References 4-20 and 4-21 are deleted per change 4-1.

D-24

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Section 5.0 Changes - Steam Line Break Analysis Appendix B contains a markup-up copy of the Chapter 5 changes. The page numbers referred to in each change are relative to revision Oa of the DPC-NE-3001-P.

5-1 Chapter 5, Steam Line Break Analysis (Entire Section)

Description:

New methodology for a main steam line break from a full power initial condition is included.

Technical Justification: The original methodology for the main steam line break is for the zero power initial condition and focuses on the core power distribution response associated with a return-to-power with the highest worth control rod stuck in its fully withdrawn position. The modeling and assumptions provide a conservative result for the zero power initial condition steam line break scenario. This revision is applicable to the full power initial condition and focuses on the pre-trip power excursion resulting from the reactivity insertion caused by the overcooling of the moderator. The transient response is of interest relative to the DNB and centerline fuel melt acceptance criteria.

The modeling and assumptions result in a conservative prediction of the maximum core power level prior to the event being terminated by reaching one of the Reactor Protection System setpoints. Application of the methodology has shown that a range of break sizes and locations and times in core life must be analyzed to determine the limiting case. A key aspect of the methodology is the effect of the decrease in reactor vessel downcomer water temperature on the excore power range flux indication. This effect causes the indicated flux to be lower than the core average, which effectively raises the high flux trip setpoint and delays tripping the reactor on high flux. This places more reliance on the other trip functions to protect the fuel. The modeling of the reactor vessel downcomer temperature effect was previously reviewed and approved by the NRC for Duke's Oconee Nuclear Station in methodology report DPC-NE-3005-PA ("UFSAR Chapter 15 Transient Analysis Methodology" - Refer to Section 16.1.3 for modeling details). Other elements of the methodology are consistent with Duke's previously approved methods for analyzing the main steam line break and other UFSAR Chapter 15 non-LOCA transients. A flow diagram of the HFP steam line break analysis is shown in Figure 5-1. Elements of the methodology that are applicable to the Westinghouse fuel design, such as the CHF correlation, were previously reviewed and approved by the NRC in DPC-NE-2009-PA (Reference 5-15). Appendix B details the revisions to Chapter 5 that are necessary to distinguish between the original zero power initial condition analysis methodology, and the new methodology for the full power initial condition.

D-25

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Figure 5-1 I1iFP Steam Line Break Accident Analysis Flow Diagram

.. . ................. -.... -.... -..............-....-..............-..... ............. "':"' '11.................... ..... '.---.-..---.

SIMULATE-3 Boundary Conditions Key Physics Parameters r (Initial power level, flow, (Doppler and Moderator pressure and moderator temperature coefficients) temperature, etc.)

..................... a

............. .. .. ......................................

RETRAN Limit ing DNBR Statepoint (defined by: power level, moderator temper"ature, pressure, and flow)

................ ................ ........................... . ... . . . ... . . ...

, ..... ... . ...... . ..... . .... ... ..

,.

I _____

... ....... .... ___

... VIPRE DNB Fuel Melt Limits vable Peaking Limits j'

.. M u .u..... . ... .. ... . .. . .

Failure of Cycle Specific Check SIMULATE-3

---

-- -- - -- - - - - - C vqcle-Specific Checks (Kuey physics parameters an.d power distribution)

D-26

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) 5-2 Chapter 5, Steam Line Break Analysis

Description:

Chapter 5 is revised to separate discussions of the HZP case and the HFP case.

5-3 Section 5.1.2, Acceptance Criteria (p. 5-1); Section 5.1.3, Analytical Approach (p. 5-2);

Section 5.2.2.2 Power Distributions (p. 5-5); Section 5.4.2.2, Minimum DNBR Results (p. 5-18); Section 5.5, Cycle Specific Evaluation (p. 5-18)

Description:

Text in sections 5.1.2 and 5.1.3 is revised to include the centerline fuel melt (CFM) limit as an acceptance criterion and to refer to the current NRC-approved code used for the centerline fuel melt temperature limit calculations. Text in sections 5.2.2.2, 5.4.2.2 and 5.5 is revised to specify that predicted peaking factors for the I-FP and HZP SLB transients are compared against applicable CFM linear heat rate limits to confirm fuel melt limits are not exceeded.

Technical Justification: Both the DNBR and centerline fuel melt (CFM) acceptance criteria must be satisfied for the steam line break accident. This revision includes a reference to the PAD code (Reference 5-9) that is used to calculate centerline fuel melt linear heat rate limits for Westinghouse RFA fuel. Duke's use of the PAD code is described in the NRC-approved methodology report DPC-NE-2009-PA (Section 4 of Reference 5-15).

5-4 Section 5.2.1.3, Break Modeling (p.5-4)

Description:

Clarification to remove implication that the 1.4 ft2 break size is the limiting break size.

Technical Justification:

The text as originally written could imply that the limiting break size is 1.4 ft2. The change removes this potential implication.

5-5 Section 5.2.2.1, Core Physics Parameters (HZP Case) (p. 5-4)

Description:

This section is revised to delete the specific value of the Doppler temperature coefficient and to delete a statement that is revised in Section 5.3.2.6. A clarification of how the reactivity versus temperature curve is calculated is also added.

Technical Justification: The specific value of the Doppler temperature coefficient (DTC) in the original methodology was stated as -3.5 pcm/°F. The DTC will be replaced by an integral check that compares the SIMULATE-3 predicted reactivity at the RETRAN-02 state point for each reload core to ensure that the limiting state point remains conservative (i.e. demonstrated by showing the SIMULATE-3 return to power is lower than the RETRAN state point power level or by showing SIMULATE-3 is subcritical at the RETRAN state point). In this methodology revision, the DTC is implicit in the D-27

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) calculated reactivity along with the moderator temperature coefficient. The revision methodology is included in Section 5.3.2.6.

5-6 Section 5.2.3.2, Analysis Methodology (p. 5-6) Section 5.4.2.2 Minimum DNBR Results (p.5-18)

Description:

The WLOP critical heat flux (CHF) correlation is added as an alternate correlation for use in evaluating DNBR for Westinghouse fuel.

Technical Justification: The WLOP CHF correlation is used to evaluate DNBR for Westinghouse fuel with system pressures between 185 and 1800 psia and was developed in Reference 5-16. The acceptability of referencing the WLOP CHF correlation in licensing application was approved by the NRC in Reference 5-17. The correlation limit of 1.18 was determnined in Reference 5-16. This CHF correlation limit was confirmed applicable with Duke's version of VIPRE-01 and is applicable for use in future steam line break analyses with low pressure state points.

5-7 Section 5.3.2.6, Core Kinetics Modeling (HZP Case) (Temperature Feedback) (p. 5-14)

Description:

The second and third sentences are revised to clarify how reactivity feedback is calculated.

Technical Justification: [

5-8 Section 5.3.2.6, Core Kinetics Modeling (HZP Case) (Temperature Feedback) (p. 5-14);

Section 5.5 (New section 5.6), Cycle-Specific Evaluation (p. 5-18)

Description:

Adjustments are made to the reactivity inputs to the RETRAN-02 point kinetics model to account for spatial reactivity effects.

Technical Justification: [ -

D-28

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

I The core power distribution at the limiting heat flux state point is evaluated to confirm that DNBR and centerline fuel limits are not exceeded.

A demonstration calculation showing the application of the new method is performed for the offsite power maintained case. This case was selected because it is limiting relative to the offsite power lost case. Table 5-1 shows the sequence of events following the steam line break. The k-effective versus temperature curve assumed in the analysis is presented in Figure 5-2 and corresponds to the end-of-cycle all rods inserted condition with the highest worth control rod stuck in its fully withdrawn position. The variation of the moderator temperature coefficient with temperature and pressure is included in the generation of this curve.

The response of the primary and secondary systems is shown in Figures 5-3 through 5-11.

Primary and secondary responses progress in a similar manner as described in Section 5.4.1. A SIMULATE-3 reactivity check is performed at the core conditions corresponding to the peak heat flux state point conditions to confirm reactivity in the RETRAN systems analysis is greater than the reactivity calculated by SIMULATE-3. This reactivity check is performed for each reload core, and by showing the return to power at the core conditions corresponding to the peak heat flux state point condition is lower than the RETRAN power level, provides cycle-specific verification the systems analysis is conservative.

A flow diagram of the HZP steam line break analysis is shown in Figure 5-12.

Table 5-1 Sequence of Events for Offsite Power Maintained Case Zero Power Initial Condition Break occurs / Operator manually trips reactor 0.01 Pressurizer level goes offscale low 12 SI actuation on low pressurizer pressure 19 Steam line isolation on low steam line pressure 24 Criticality occurs 25 SI pumps begin to deliver unborated water to RCS 38 Peak heat flux occurs 118 High-head SI injection lines purged of unborated 119 water / One train of SI fails D-29

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Figure 5-2 K-effective Versus Moderator Temperature 1.07 1.06 1.05 1.04 1.03

• 1.02 1.01 1.00 0.99 0.98 300 350 400 450 500 550 600 MIoderator Temperature (F)

Figure 5-3 MNS/CNS UFSAR Section 15.1.5 - SLB With OSPM 1200 1000

" 800

- 600

,.2

-INTACTLOOPS]

-FAULTED LOOP c1 400 200 0

0 25 50 75 100 125 150 175 200 Time (Sec)

D-30

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Figure 5-4 MNS/CNS UFSAR Section 15.1.5 - SLB With OSPM 600 550 500 450 400 350 300 25 50 75 100 125 150 175 200 Time (Sec)

Figure 5-5 MNS/CNS UFSAR Section 15.1.5 - SLB With OSPM 600 550 500 E 450 CL Z 400 350 300 25 50 75 100 125 150 175 200 Time (Sec)

D-31

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Figure 5-6 MNS/CNS UFSAR Section 15.1.5 - SLB With OSPM 20 18 j16 14 12

,-10

• 8

<6 4

0 0 25 50 75 100 125 150 175 200 Time (Sec)

Figure 5-7 MNS/CNS UFSAR Section 15.1.5 - SLB With OSPM 1000 500 0

E

-500

-1000

-1500 25 50 75 100 125 150 175 200 Time (Sec)

D-32

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Figure 5-8 MNS/CNS UFSAR Section 15.1.5 - SLB With OSPM 40 35 30 25 20 15 10 25 50 75 100 125 150 175 200 Time (Sec)

Figure 5-9 MNS/CNS UFSAR Section 15.1.5 - SLB With OSPM 18 16 14 12 10 6

4 25 50 75 100 125 150 175 200 Time (Sec)

D-33

Appendix D DPC-NE-3001-P, Revision I Changes (Non-Prop. Version)

Figure 5-10 MNS/CNS UFSAR Section 15.1.5 - SLB With OSPM 2500 2000 1500 1000 500 0

0 25 50 75 100 125 150 175 200 Time (See)

Figure 15-11 MNS/CNS UFSAR Section 15.1.5 - SLB With OSPM 5000 4500

"*4000 3500 3000 2500 c 2000 1500 1000 500 0 25 50 75 100 125 150 175 200 Time (See)

D-34

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Figure 5-12 HZP Steam Line Break Accident Analysis Flow Diagram SIMULATE-3 Initial and Key Physics Parameters Boundary Conditions r - - - (Shutdown margin, reactivity versus temperature curve, (Initial power level, flow, pressure differential boron worth, Doppler and moderator temperature, etc.)

temperature coefficient) dk RET RAN Limiting DNI BR Statepoint (defined by: power level, moderator t emperature, pressure, and flow)

Iterate to Nlormalize RETRAN and SIMULATlE-3 Power Response SIMUL/ATE-3 Execui ted at Limiting DNB RStatepoint Power Distribution at Limiting DNBR Statepoint S............. .... :..... 1.... . .. ........ -.... ,........ .

VIPRE DNB

..........................

Me .. . . . . . .

... .... ... .. , .*,....

, .** ,. * .. . ., .

T...

Fuel Melt Limits SIMUL ATE-3 Cycle-Spec ific Checks (Key physics para meters, and core subcriticality and p(ower distribution at Failure of Cycle limiting DNB R statepoint)

I Specific Check L-------------------------------------I D-35

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) 5-9 Cycle-Specific Evaluation (p. 5-18)

Description:

The option to perform an evaluation to address non-conservative reload checks is added to the last paragraph.

Technical Justification: Refer to the technical justification for change 1-2.

5-10 References (p. 5-19)

Description:

References are revised and new references are added.

5-3 SIMULATE-3 Methodology, Advanced Three Dimensional Two-Group Reactor Analysis Code, SOA-95/158, Studsvik of America, hIc., October 1995.

5-7 Thermal-Hydraulic Transient Analysis Methodology, DPC-NE-3000-PA, Revision 35a, Duke Poeer-C-empa*y Energy Carolinas, September-2004 October 2012 5-9 Westinghouse Improved Performance Analysis and Design Model (PAD 4.0), WCAP-15063-P-A, Revision 1, Westinghouse, July 2000 5-10 Thermal-Hydraulic Statistical Core Design Methodology, DPC-NE-2005-PA, Revision 4a, December 2008 5-11 WCAP-15025-P, Modified WRB-2 Correlation, WRB-2M, for Predicting Critical Heat Flux in 17x17 Rod Bundles with Modified LPD Mixing Vane Grids, Westinghouse Energy Systems, February 1998.

5-12 0. W. Hermann and C. V. Parks, "SAS2H: A Coupled One-Dimensional Depletion and Shielding Analysis Module," NUREG/CR-0200, Volume 1, Section S2, March 2000 5-13 0. W. Hermann and R. M. Westfall, "ORIGEN-S: SCALE System Module to Calculate Fuel Depletion, Actinide Transmutation, Fission Product Buildup and Decay, and Associated Radiation Source Terms," NUREG/CR-0200, Volume 2, Section F7, March 2000 5-14 Judith F. Briesmeister, Ed., "MCNP - A General Monte Carlo N-Particle Transport Code," Los Alamos National Laboratory Report, LA-13709-M, March 2000 5-15 Westinghouse Fuel Transition Report", DPC-NE-2009-PA, Revision 3a, September 2011 5-16 Addendum 2 to WCAP-14565-P-A Extended Application of ABB-NV Correlation and Modified ABB-NV Correlation WLOP for PWR Low Pressure Applications, Revision 0, Westinghouse Electric Company LLC, April 2008 D-36

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) 5-17 February 14, 2008 letter from H. K. Nieh (NRC) to J. A. Gresham (Westinghouse), "Final Safety Evaluation for Westinghouse Electric Company Topical Report WCAP-14565-P, Addendum 2, Revision 0, "Addendum 2 to WCAP-14565-P-A, Extended Application of ABB-NV Correlation and Modified ABB-NV Correlation WLOP for PWR Low Pressure Applications" Technical Justification: References 5-3 and 5-7 were updated to the current version of these reports. Reference 5-9 (PAD) was added as the source for centerline fuel melt limits for Westinghouse RFA fuel. Reference 5-10 was added as part of the new HFP steam line break methodology. References 5-12 through 5-14 pertain to the modeling of the effect that changes in downcomer moderator temperature have on the neutron flux incident on the excore detectors which were approved in DPC-NE-3005-PA.

Reference 5-15 is the NRC-approved methodology report describing Duke's use of the PAD code in reload applications. References 5-16 and 5-17 were added as part of the WLOP CHF correlation.

D-37

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Section 6.0 Changes - Dropped Rod Analysis 6-1 Section 6.1.3 Analytical Approach (p. 6-2)

Description:

The specificity of the fuel performance code used to calculate centerline fuel melt limits is removed.

The SIMULATE-3 post-drop power distributions are also compared to the centerline fuel melt limits from the appropriate fuel performance code (e.g. PAD (Reference 6-5) for Westinghouse fuel) to demonstrate margin.

Technical Justification: The fuel performance code used to determine centerline fuel melt limits is dependent upon the fuel type being irradiated. The PAD fuel performance code is applicable for Westinghouse fuel product only (Westinghouse RFA fuel is currently being irradiated), but not for other vendors fuel products.

6-2 Section 6.3.1 Initial Conditions (Average Fuel Temperature) (p. 6-5)

Description:

The initial average fuel temperature is revised from maximum to minimum and the numerical values are deleted.

Ma-m--m Minimum average fuel temperatures at beginning, middle, and end-of-cycle are used. Minimum values have been shown to be conservative.

Technical Justification: The initial average fuel temperature is revised from maximum to minimum based on additional experience in applying the methodology. Minimum values have been shown to be conservative.

6-3 Section 6.3.2.1 Physics Parameters (Control Bank D Worth) (p. 6-6)

Description:

The minimum Control Bank D worth available for withdrawal is reduced to 250 pcm, and includes consideration of the dropped rod(s) being in Control Bank D.

The Control Bank D worth available for withdrawal ranges from [ I pcm at the rod insertion limit as a function of burnup, including consideration for the dropped rod(s) being in Control Bank D.

Technical Justification: This revision adds clarification that the Control Bank D worth available for withdrawal may be as low as j I pcm for the situation where the dropped control rod(s) is in Control Bank D.

6-4 Section 6.3.2.1 Physics Parameters (FAH Versus Worth) (p. 6-6)

Description:

The effect of a dropped rod on FAH includes the location of the dropped rod(s).

The effect of a dropped rod on FAH is a function of bumup, dropped rod worth, and the number and location of dropped rods. FAn responses can be derived as a function of any of these variables. Enveloping FAH responses derived from D-38

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) assembly average power as a function of dropped rod worth and burnup are presented in Figures 6-2 to 6-4.

Technical Justification: This revision adds clarification that the location of the dropped rod(s) has an impact on the resulting FAH, and that responses can be developed as a function of the number of dropped rods, worth, location or burnup.

6-5 Section 6.3.2.1 Physics Parameters (FAH Versus Worth) (p. 6-6)

Description:

An alternative approach to using an enveloping FAH response as a function of dropped rod worth, and the number and location of dropped rod(s) in the DNBR calculation is added. This approach uses the post-drop FAH and axial shape in the DNBR analysis for each case evaluated.

An alternative to using enveloping FAHl responses in the DNBR calculation is to use the post-drop power distribution (FAHl and axial shape) for each dropped rod combination in the DNBR analysis.

Technical Justification: The characterization of the post dropped rod FAH response as a function of burnup and dropped rod worth, and the use of a single bounding axial shape at each burnup in the dropped rod analysis methodology was developed to produce a conservative analysis result, and to eliminate the necessity of having to evaluate the entire dropped rod case matrix on a cycle-specific basis. This approach was largely adopted because of the large computational overhead necessary to evaluate the post-drop power distribution. However, while this approach is conservative, it produces unrealistic combinations of radial and axial peaking factors. With the computational efficiency gained with the evolution of computer hardware and analysis software, it is now practical to evaluate all post-drop rod power distributions on a cycle-specific basis. The proposed alternative methodology uses the post-drop FAH and axial shape to calculate DNBR margin relative to the design DNBR limit. The coupling of the post-drop FAH and axial shape results in a DNBR margin gain relative to the original methodology by removing unrealistic combinations of radial and axial peaking factors used in the DNBR calculation.

6-6 Section 6.3.2.1 Physics Parameters (Axial Shape) (p. 6-6)

Description:

An alternative approach to using a bounding axial shape in the DNBR calculation is added. The alternative approach is to use the post-drop axial shape for each dropped rod combination in the DNBR analysis.

An alternative approach to using a bounding axial shape in the DNBR calculation is to use the post-drop axial shape (and FAR) for each dropped rod combination in the DNBR analysis.

Technical Justification: Refer to the technical justification for change 6-5.

D-39

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) 6-7 Section 6.3.2.1 Physics Parameters (Core Tilt Following Rod Drop) (p. 6-6)

Description:

Excore detector tilts can be characterized as a function of dropped rod worth and the number and location of dropped rods.

Insert the following after the second sentence:

Excore detector tilts can be characterized as a function of dropped rod worth and the number and location of the dropped rods.

Technical Justification: The proposed methodology will associate the excore detector tilt with the number and location of the rods that are dropped, rather than just the dropped rod worth. This revision will enable a more realistic value of excore tilt to be used in the methodology.

6-8 Section 6.3.2.1 Physics Parameters (Core Tilt Following Rod Drop) (p. 6-7)

Description:

The last sentence is modified to acknowledge that the total flux incident on the excore detectors used to calculate core tilt is not only from the fuel assemblies closest to the excore detector, but from other fuel assemblies in the core, and to clarify that the flux incident on the excore detectors is calculated using appropriate weighting factors.

The excore tilt for the four quadrants is modeled by using weighting factors to account for the relative importance of fuel assemblies closest to the exeore deteete in generateing a detector response.

Technical Justification: The original methodology calculated tilt based on the change in power following a dropped rod using the fuel assemblies closest to the excore detector to generate a detector response. The revision acknowledges that the flux incident on the excore detectors is not only from the fuel assemblies closest to the excore detector, but from other fuel assemblies in the core. This allows for a more realistic value of excore tilt to be used in the methodology by calculating flux incident on the excore detectors using appropriate weighting factors.

6-9 Section 6.3.2.2 Reactor Protection System (p. 6-8)

Description:

The methodology is revised to credit the high flux, the over-temperature AT, and the over-power AT trip functions, and to identify the limiting single failure.

The RETRAN-02 analysis takes credit for a reactor trip on low pressurizer pressure, high flux, over-temperature AT, and over-power AT. I I The low pressurizer pressure trip setpoint is reached only for higher dropped rod worth cases at beginning-of-cycle. The over-temperature and over-power AT trip setpoints may be reached for some dropped rod events. No credit is taken for the lo',' low steam generator level trip or-main steamn isolation on low steam line pressur D-40

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Technical Justification: The original methodology only credited the low pressurizer pressure trip function to trip the reactor following a dropped rod transient. The other reactor trips, specifically the high flux, the over-temperature AT, and the over-power AT trip functions were not credited. This revision allows crediting these additional trip functions should the setpoints be reached in the analysis. The low-low steam generator level trip and the main steam isolation on low steam line pressure setpoints are not approached in the analysis and are therefore deleted.

6-10 Section 6.3.2.3 Power Range Nuclear Instrumentation (p. 6-8)

Description:

The methodology is revised to include a more detailed model for determining the effect of a decrease in reactor vessel downcomer temperature on the flux signal indicated by the excore power range instrumentation.

The negative reactivity insertion following the dropped rod causes a mismatch between the primary heat source and the secondary heat sink. The resulting decrease in cold leg temperature upon entering the reactor vessel downcomer attenuates the neutron flux leakage exiting the reactor. This attenuation effect reduces the flux incident on the excore neutron detectors, thereby creating an error in the indicated flux value (indicated excore detector power less than true reactor power). A conservative attenuation factor is assumed as a function of the change in reactor vessel downcomer density resulting from the change in temperature.

The effect of a change in reactor vessel downcomer water temperature (density) on the excore flux detector signal is modeled by consideration of the relevant source, material composition, and detector geometry details. The fuel region is considered as a homogeneous mixture inside a cylinder of equivalent volume as the core region inside the baffle plates. The balance of the geometry is modeled as a series of concentric cylinders, representing the baffle plates, flow channel, core barrel, neutron pads, downcomer region, and reactor vessel. Particle tallies at the detector account for the details of detector geometry and design.

Fuel characterization is performed with the SAS2H/ORIGEN-S modules of the SCALE Code System, as necessary (References 6-10 and 6-11). Transport and tallying of particles is performed with the MCNP computer code (Reference 6-12). Variance reduction is performed in MCNP as necessary to achieve reliable D-41

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version) statistics for the tallies. The particle tally results are used to characterize the relative effects of Reactor Coolant System temperature (density) on detector response.

Technical Justification: A decrease in reactor vessel downcomer water temperature will attenuate the flux leakage incident on the excore power range flux indication. This effect causes the indicated flux to be lower than the core average. I I This modeling of the reactor vessel downcomer temperature effect was previously reviewed and approved by the NRC for Duke's Oconee Nuclear Station in methodology report DPC-NE-3005P-A ("UFSAR Chapter 15 Transient Analysis Methodology" - Refer to chapter 16.1.3 of DPC-NE-3005P-A for methodology details).

6-11 Section 6.3.2.4 Rod Control System (p. 6-8)

Description:

The Rod Control System is explicitly modeled in the RETRAN-02 analysis. The controller uses a power mismatch signal and a temperature error signal to determine Control Bank D insertion or withdrawal and rod speed. The power mismatch signal is a difference between turbine power (impulse or inlet pressure) and the I

] The temperature error signal is a difference between a reference temperature based on turbine power and the aue...nee..ed.high medial select primary loop T-ave indication. Due to the impertance of Control Bank D withdrawal on the dropped red analysis, the worst ease single failure has been detennined to result -

the M flux indieation auctioneer-ing low. Tkis failure causes the maximum post drop power-levels by aeceler-ating the onset and increasing the rate of Control Ban withd.awa..

Technical Justification: Changes to the turbine power, NI flux and loop Tave signals input to the Rod Control System are the result of plant modifications. Turbine power is now based on inlet pressure versus impulse pressure. The auctioneered high NI flux indication signal and primary loop Tave signal are replaced with median select signals.

(Median select is essentially equivalent to the second highest indication of each parameter.) The change in NI and loop Tave signals are the result of installation of the Distributed Control System (DCS). The turbine power signal change is the result of installation of a new high pressure turbine.

I D-42

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

The NI response input to the Rod Control System [

I is shown in Figure 6-1. The results presented are typical, and consider all combinations of dropped rods in each group of Control and Shutdown Banks. A conservative input NI signal to the Rod Control System is selected to maximize the mismatch between indicated reactor power and turbine power, which in-turn maximizes the positive reactivity addition to the reactor core from the withdrawal of the Control Bank.

The transient response of the primary system for typical BOC and EOC 100 and 400 pcm dropped rod cases based on the revised methodology that models the actual design of the Rod Control System plus the limiting single failure of one NI channel are presented.

Results for the BOC cases are shown Figures 6-2 through 6-6, and results for the EOC cases are shown in Figures 6-5 through 6-9.

BOC TransientResponse The transient is initiated when a single or multiple rods drop into the reactor core.

Reactor power initially rapidly decreases as the result of negative reactivity from the dropped rod. In response to the power mismatch between indicated power (from the NIs) and turbine power, and the temperature error between the reference temperature based on turbine power (T-ref) and the average loop temperature (Tave), the Rod Control System withdraws Control Bank D. The positive reactivity inserted from the withdrawal of Control Bank D causes reactor power to recover and exceed its initial power level.

Reactor power stabilizes and plateaus as the negative reactivity from Doppler and moderator temperature feedback offset the positive reactivity inserted from the withdrawal of Control Bank D. Reactor power decreases as the core continues to heat-up. Pressurizer pressure initially decreases due to the impact of the dropped rod on reactor power, and then increases with power. Pressurizer level follows Tave. Tave initially decreases as the result of the reactor power decrease from the dropped rod, and then starts to increase once there is a balance between reactor power and steam load.

Most cases do not result in a reactor trip. However, a reactor trip on low-low pressurizer pressure is produced for high dropped rod worth cases.

Figure 6-2 shows reactor power and the NI signal input to the Rod Control System. As the dropped rod worth increases, the initial reduction in reactor power increases, and consequently the initial reduction in Tave and pressure also increase (Figures 6-3 and 6-4). Tave and pressure recover, and increase with increasing reactor power. Figure 6-5 shows pressurizer level follows the change in Tave.

D-43

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

EOC TransientResponse Reactor power initially decreases rapidly following a dropped rod or rods followed by a reactor power increase due to the positive reactivity addition from the withdrawal of Control Bank D, and from the initial temperature decrease in the presence of a negative moderator temperature coefficient (MTC). The more negative Doppler temperature coefficient and MTC at EOC limits the magnitude of the power overshoot relative to BOC conditions.

Figure 6-6 shows the reactor power response and the NI input to the Rod Control System.

The positive reactivity from the Control Bank D withdrawal is eventually offset from negative reactivity from Doppler and moderator temperature feedback as the reactor coolant system heats-up. Figures 6-7 and 6-8 show the Tave and pressurizer pressure responses.

Tave decreases until core power exceeds steam load. Pressurizer pressure initially decreases with the decrease in reactor power, and then increases following power. Figure 6-9 shows the pressurizer level response which follows Tave.

Thermal-HydraiulicAnalysis The RETRAN core statepoint conditions that correspond to the time of minimum DNBR establish the thermal-hydraulic boundary conditions that are used in combination with post-drop peaking factors as a function of dropped rod worth to calculate DNBR. The [

I VIPRE-01 model from Reference 6-2 is used for this analysis. The DNBR calculated is compared against the statistical core design (SCD) DNBR limit to ensure the SCD DNBR limit is not exceeded for each combination of thermal condition and peaking factor.

A flow diagram of the dropped rod analysis is shown in Figure 6-10.

D-44

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Figure 6-2 BOC 100 and 400 pcm Dropped Rod Reactor Power and NI Input to the Rod Control System 140.0 120.0 100.0 80.0 60.0 40.0 20.0 0.0 0.0 10.0 20.0 30.0 40.0 50.0 60.0 70.0 80.0 90.0 100.0 110.0 Figure 6-3 BOC 100 and 400 pcm Dropped Rod Reactor Coolant System T-ave 595 590 585 580 575

-Tave (100 pcro)

- -Tave (400 pcm) 570 565 560 555 0 10 20 30 40 s0 60 70 s0 90 100 110 Tlme(sec)

D-45

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Figure 6-4 BOC 100 and 400 PCM Dropped Rod Pressurizer Pressure 2,400 2,300 2,200 U

2,100 Pressurizer Pressure 1100pcm)

- -Pressurizer Pressure (400 pcm) 2,000  %.5 1,900 1,800 0 '0 10.0 20.0 30,0 40.0 50.0 60.0 70.0 80.0 90.0 100.0 110.0 Time (sec)

Figure 6-5 BOC 100 and 400 PCM Dropped Rod Pressurizer Level 70 60 50 40 30 5- - Pressuriter Level (100 20  %

%

10 0

0.0 10.0 20.0 30.0 40.0 50.0 60.0 70,0 80.0 90.0 100.0 110.0 Time (sec)

D-46

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Figure 6-6 EOC 100 and 400 pcm Dropped Rod Reactor Power and NI Input to the Rod Control System 120 110 100 t

u 90 80 70 s60o 40 40 10 20 30 40 50 60 70 80 90 100 110 Time (sec)

Figure 6-7 EOC 100 and 400 pcm Dropped Rod Reactor Coolant System Tave 592 590 588 586 584 I-582 580 578 576 10 20 30 40 50 60 70 80 90 100 110 Time (sec)

D-47

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Figure 6-8 EOC 100 and 400 pcm Dropped Rod Pressurizer Pressure 2,400 2,300 2,200

• 2,100 2,000 - - Pressurizer Pressure (100 pcm) 1 - Pressurizer Pressure (400 pcm) 1,900 1,800 0 10 20 30 40 50 60 70 80 90 100 11(

Time (sec)

Figure 6-9 EOC 100 and 400 pcm Dropped Rod Pressurizer Level 60 55 so 45 40 35

- Pressurizer Level (100 pcm)

-- Pressurizer Level (400 pcm) 30 25 20 0 10 20 30 40 50 60 70 80 90 100 110 T"me(sec)

D-48

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

Figure 6-10 Dropped Rod Accident Analysis Flow Diagram

..............

. . . .. ................................................

. .... . . . . . . . , . , .. . . . . ,......................................

................ .... . . .

SIMULATE-3 Initial and Key Physics Parameters Boundary Conditions (Doppler and Moderator (Initial power level, flow, temperature coefficients, pressure and moderator dropped rod worth, available worth temperature, etc.)

for withdrawal, and excore tilt)

RETRAN Limiting DNBR Statepoint (defined by: power level, moderator temperature, pressure, and flow)

...........................

.......

.............

..........

........................

I a

SIMULATE-3 VIEPRE DNB Post Drop a Power Distribution

.. . ... . .. b.. ...

Failure of Cycle q Specific Check SIMULATE-3

- ---------------- Cycle-Specific Checks a (Key physics parameters and power distribution)

D-49

Appendix D DPC-NE-3001-P, Revision I Changes (Non-Prop. Version) 6-12 Section 6.3.2.5 Pressurizer Pressure and Level Control (p. 6-9)

Description:

Pressurizer heaters are not credited.

Since the dropped rod transient is a DNB transient, pressurizer sprays and the pressurizer PORV are assumed to function in order to minimize primary pressure.

Pressurizer heaters are assumed not credited to-Fuetion. Pressurizer level control is assumed to be in manual, and is not important for this transient.

Technical Justification: The original methodology states that the pressurizer heaters are assumed not to function, which seems to be conservative given that the dropped rod transient is evaluated against the DNBR acceptance criterion, and low pressure is the conservative direction. However, one operating mode has the pressurizer heaters energized to offset the depressurization effect of using the sprays to keep the boron concentration in the pressurizer near the boron concentration in the Reactor Coolant System. Additional experience in applying the methodology has determined that for some dropped rod cases this operating mode is slightly more conservative. For other cases the original methodology is conservative. The revision covers both of these modeling approaches by stating that the pressurizer heaters are not credited.

6-13 Section 6.3.2.6 Main Feedwater and Turbine Control (p. 6-9)

Description:

[

II I

Technical Justification: I I

6-14 Section 6.4 Results and Conclusions (p. 6-9)

Description:

A clarification was made that the dropped rod demonstration analysis results are typical. A reference to the UFSAR was added to identify current analysis of record results.

D-50

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

The demonstration analysis results presented are based on analysis of the AREVA Mk-BW fuel product, and are representative of typical dropped rod transient results. Results for the analysis of record can be found in Section 15.4.3 of the UJFSAR.

Technical Justification: The note was added to denote the history of the demonstration analysis results, that the results are typical and the analysis of record can be found in the UFSAR. The technical content of the report is unaffected by this change.

6-15 Section 6.6 Cycle Specific Evaluation (p. 6-14)

Description:

The cycle-specific evaluation is revised to include insights gained with experience in applying the methodology.

InTitial F*

• Aalflux shape

• Post-drop FAh

• Post-drop axial power distribution

• Post-drop excore tilt

• Moderator and Doppler temperature coefficients

• Maximum dropped rod worth

• Available Control Bank D worth for withdrawal Sever-al r-eload cycles were analyzed in order-to detenaine bounding inputs fcrF th-e r-efer-ence dropped road analysis. The results of these analyses established bounding cunves, ver-sus the number and worth of dropped rods, defuntng both the incesm radial pealing and limiting excore detector r-esponses. These inpt ar considered independent of the reload cor-e design and will not be checked on a cycle specific basis.

While the above physic paamtes are net expeceted to change fcr-a reload coere, they ar~e checked to ensure. that the r-efer-ence dropped rod analysi rmin valid. For each reload core, the m cre F for-the pre dropped condition is ver"ified t be less than 1.50 for allowed rod inser*ions. post-drop FAh, axial power distribution and excore tilt are verified to be less than the values assumed in the reference analysis for all dropped rod combinations. Moderator and Doppler temperature coefficients are also verified to be conservative by comparison against the coefficients used assumed in the reference analysis. The axial shapes assumed in the r-efer-ence analysis will be checked for all dropped rod combinations. The maximum allowable dropped rod worth is verified to be less than the maximum worth analyzed ( ,8001 pem).

Technical Justification: The original methodology includes the initial FAh and the axial flux shape as cycle-specific key parameter checks to verify the acceptability of the post-drop core power peaking coupled with the use of a bounding curve defining the increase in radial peaking (FAh) as a function of the number and worth of dropped rods. The initial FAh check is replaced with a post-drop FAh check to confirm the FAh analysis assumptions. This check captures cycle-specific peaking variations that may occur in the post-drop power distribution that may not be present in the pre-drop power distribution.

D-51

Appendix D DPC-NE-3001-P, Revision 1 Changes (Non-Prop. Version)

The second change is made because the axial shape used in the DNBR analysis is a normalized quantity (Fq/FAh) and it is therefore possible for the reference axial shape to be exceeded in a cycle-specific check without exceeding the reference axial power distribution. The proposed methodology revision allows comparison against the reference analysis axial power distribution to verify the acceptability of the DNB analysis. This additional check ensures that the limiting core power distributions have been checked for the limiting dropped rod cases, rather than performing additional analyses for core power distributions and cases that are known to be non-limiting. A post-drop excore tilt check is added to validate the values assumed in the system transient analysis. The maximum worth value verified is removed to preclude updating the report if this value changes in the future.

The important aspect of the method is retained which requires confirmation that the maximum dropped rod worth assumed in the analysis-of-record bounds the reload core.

Additional editorial revisions are included to delete content that is no longer needed to describe the methodology.

6-16 References (p. 6-15a)

Description:

The list of references is revised.

6-2 Thermal-Hydraulic Transient Analysis Methodology, DPC-NE-3000-PA, Revision 35a, Duke Power Company, September*-2004 October 2012 6-3 SIMULATE-3 Methodology Advanced Three Dimensional Two-Group Reactor Analysis Code, SOA-95/158, Studsvik of America, October 1995 6-6 Thermal-Hydraulic Statistical Core Design Methodology, DPC-NE-2005-PA, Revision 34a, September 204 December 2008 6-9 The BWU Critical Heat Flux Correlations, BAW- 101 99P-A, AREVA NP, Nevembef 199 August 1996 6-10 0. W. Hermann and C. V. Parks, "SAS2H: A Coupled One-Dimensional Depletion and Shielding Analysis Module," NUREG/CR-0200, Volume 1, Section S2, March 2000 6-11 0. W. Hermann and R. M. Westfall, "ORIGEN-S: SCALE System Module to Calculate Fuel Depletion, Actinide Transmutation, Fission Product Buildup and Decay, and Associated Radiation Source Terms,"

NUREG/CR-0200, Volume 2, Section F7, March 2000 6-12 Judith F. Briesmeister, Ed., "MCNP - A General Monte Carlo N-Particle Transport Code," Los Alamos National Laboratory Report, LA-13709-M, March 2000 Technical Justification: References 6-2, 6-3, 6-6 and 6-9 are updated to the current versions of the reports. New References 6-10 through 6-12 are associated with change 6-10.

D-52

Appendix E Non-Proprietary Marked-up Version of Chapter 5 of DPC-NE-3001-P, Revision 1 Note: Appendix E contains all text markups. It does not contain Figures that did not change.

5.0 STEAM LINE BREAK ANALYSIS 5.1 Overview 5.1.1 Description of Steam Line Break Accident The steam line break transient is described in UJFSAR Section 15.1.5. The steam release arising from a break in a main steam line would result in an initial increase in steam flow, with a subsequent decrease during the accident as the steam pressure falls. The energy removal from the Reactor Coolant System (RCS) causes a reduction of coolant temperature and pressure. In the presence of a negative moderator temperature coefficient, the cooldown results in an insertion of positive reactivity-4-f and a subsequent increase in reactor power that can challenge fuel design limits. Two steam line break scenarios are analyzed. The first scenario focuses on the core power distribution response during the return to power with a stuck control rod. This scenario is limiting from a hot zero power (HZP) initial condition, and is referred to as the "HZP case". At the HZP initial condition, if the most reactive control rod is assumed stuck in its fully withdrawn position after reactor trip, the core might become critical and return to power. A return to power following a steam line rupture at HZP is a potential problem mainly because of the high power peaking factors which exist assuming the most reactive control rod to be stuck in its fully withdrawn position. The core is ultimately shut down by the boric acid injection delivered by the Safety Injection System. The second scenario focuses on the pre-trip core power excursion and is limiting from the full power initial condition. This scenario is referred to as the "I-FP case". The pre-trip power excursion for the full power case is a potential problem because the excore flux signal is attenuated producing a deviation between indicated and actual core power, which can result in a delay in the reactor trip signal, and a higher power level. Where the methodology is different for the two cases it is designated as such.

5.1.2 Acceptance Criteria A major steam line break is classified as an ANS Condition IV event, a limiting fault. Minor secondary system pipe breaks are classified as ANS Condition III or infrequent events. The analysis is performed assuming a stuck control rod, a single failure in the engineered safety features, and with consideration of both offsite power maintained and offsite power lost. The 5-1 (LAR Page No.: E-l)

following two criteria must be satisfied. First, the core must remain in place and intact. The analysis submitted herein meets this criterion by showing that the 95/95 DNB limit of Reference 5-1, Section 4.4-is and the centerline fuel melt limit of Reference 5-1, Section 4.2 are satisfied.

Future analyses using these same methods might meet the criterion by demonstrating continued core cooling capability based on an acceptable fuel damage model and result. Second, radiation doses must not exceed applicable regulatory limits. The Condition III and rV criteria regarding overpressurization are not challenged by a steam line break transient.

5.1.3 Analytical Approach The steam line break transient requires a limiting set of physics parameters to be determined for use as initial and boundary conditions. These parameters are input to a McGuire/Catawba RETRAN-02 (Reference 5-2) model for the system thermal-hydraulic analysis. The RETRAN-02 analysis generates the core statepoint conditions which correspond to the transient time of minimum DNBR or maximum centerline fuel temperature. The SIMULATE-3 code (Reference 5-3) based on the methodology described in Reference 5-5 is used to generate core power distributions corresponding to the statepoint conditions. The core power distribution along with the core thermal-hydraulic boundary conditions from the RETRAN-02 analysis are then input to a McGuire/Catawba VIPRE-01 (Reference 5-4) model to calculate the minimum DNBR. If this value is below the DNBR limit, then a fuel rod census would be performed to determine the number of fuel rods in DNB (assumed to experience cladding failure) and therefore the fraction of gap activity released. The dose consequences of this release would then be evaluated. The core power distribution from SIMULATE-3 is also compared to the centerline fuel melt linear heat rate limit from the applicable fuel performance code (Reference 5-9) to demonstrate margin.

5.2 Simulation Codes and Models 5.2.1 System Thermal-Hydraulic Analysis 5.2.1.1 Selection of a Bounding Unit 5-2 (LAR Page No.: E-2)

Differences between the McGuire and Catawba units are discussed in Section 3.1.6 of Reference 5-7 for the steam line break transient. The most important differences with respect to steam line break are the steam generator type and the differences in the Auxiliary Feedwater System flowrates. McGuire and Catawba Unit 1 steam generators have been replaced with Babcock &

Wilcox Canada feedring steam generators. Catawba Unit 2 has Westinghouse Model D5 preheater steam generators. The steam generators influence the transient response due to design differences such as heat transfer areas, tube alloys, tube bundle height, and initial liquid inventory. The Auxiliary Feedwater System flowrates are different due to pump discharge piping resistance and throttle valve positions, and pump capacity. Both steam generator designs are analyzed separately.

The Auxiliary Feedwater System flowrates used are conservative for the unit for which the analysis is applicable.

5.2.1.2 Modifications to Base Plant Model Renodalization of Reactor Vessel 5-3 (LAR Page No.: E-3)

I Renodalization of Steam Generator Secondary (HZP Case)

Since the main feedwater piping contains only subcooled water connected to the steam generator, this inventory would remain inactive during a steam line break and would not affect the transient analysis. In order to save computational costs, the main feedwater piping nodes are eliminated and the feedwater is added directly to the steam generator via a positive fill junction, similar to the junction already used for auxiliary feedwater.

L I

1_

I 5-4 (LAR Page No.: E-4)

5.2.1.3 Break Modeling The full cross-sectional area of the 34" main steam line is 5.4 ft 2 . The area of the flow restrictor at the steam generator outlet is 1.4 ft 2. [

This analysis uses the Moody critical flow model. For the timeframe of interest, the break flow is always limited by critical flow.

5.2.2 Nuclear Analysis The transient system response during a steam line break accident is sensitive to core reactivity versus temperature and the Doppler Temperature Coefficient. The core thermal-hydraulic response is sensitive to the three dimensional core power distribution. Therefore, the nuclear analysis for this event must specify pre-break core physics characteristics and post-break power distributions based on the calculated system response.

5.2.2.1 Core Physics Parameters HZP Case The k-effective versus temperature curve (Figure 5-2) is calculated at EOC including the effects of temperature and pressure with all rods inserted and the highest worth rod stuck in its fully withdrawn position. The k-effective versus temperature curve and the Doppler temperature coefficient are selected such that a limiting return to power occurs in the RETRAN-02 analysis.

This cure r-epr-,e.nts the. effet on rx-eaeivt, of a asmmtr coldown from the technical specification shutdow.n mar-gin limit. The Peppier-coefficient was chosen to be 3.5 peý' 0 F for th-s analysis. The conservatism of the k-effective versus temperature curve and Doppler coefficient will be confirmed each cycle as described in Section 5.6.

5-5 (LAR Page No.: E-5)

HFP Case The limiting core physics parameters are dependent on the break size. The moderator temperature coefficient is determined by sensitivity analyses within the range of the least negative beginning-of-cycle value and the most-negative end-of-cycle value. The Doppler coefficient is selected to be a least negative value consistent with (or conservative for) the time-in-cycle of the moderator temperature coefficient.

5.2.2.2 Power Distributions HZP Case

] SIMULATE-3 is used to explicitly calculate the peak pin value at limiting RETRAN-02 statepoints. The three-dimensional power distribution and the system analysis results are then combined for the thermal-hydraulic evaluation. The three-dimensional power distribution is also used in the centerline fuel melt evaluation.

HFP Case The RETRAN-02 limiting statepoint for the HFP case is input to SIMULATE-3 to predict the power distribution including the peak pin. The three-dimensional power distributions and system analysis results are used in the DNBR and centerline fuel melt evaluations.

5.2.3 Core Thermal-Hydraulic Analysis 5.2.3.1 VIPRE-O1 Code Description The VIPRE-01 code (Reference 5-4) is used for the steam line break core thermal-hydraulic analyses. VIPRE-O1 is a subchannel thermal-hydraulic computer code. With this subchannel analysis approach, the nuclear fuel element is divided into a number of quasi one-dimensional 5-6 (LAR Page No.: E-6)

channels that communicate laterally by diversion crossflow and turbulent mixing. Given the geometry of the reactor core and coolant channels and the boundary conditions or forcing functions, VIPRE-01 calculates core flow distributions, coolant conditions, fuel rod temperatures and the minimum departure from nucleate boiling ratio (MDNBR) for steady-state conditions and for transients. VIPRE-01 accepts all necessary boundary conditions that originate either from the RETRAN-02 system transient simulation or the core neutronics simulation. Included is the capability to impose different boundary conditions on different regions of the core model. For example, different core region inlet temperatures, flow rates, heat flux, and even different assembly and pin radial powers or axial flux shapes can be modeled in steady-state or transient modes.

5.2.3.2 Analysis Methodology HFP Case The VIPRE-01 model from Reference 5-7 is used along with the statistical core design (SCD) methodology (Reference 5-10). With the SCD methodology the uncertainties in many of the initial conditions (e.g., power, pressure, temperature, flow) are included in the statistical DNBR limit rather than in the initial conditions of the analysis. The DNBR correlation used for Westinghouse fuel is the WRB-2M correlation (Reference 5-11) with an SCD limit of I The CHF correlation used below the first mixing vane grid is the BWU-N correlation (Reference 5-8) with a correlation limit of 1.21. A transient VIPRE-01 analysis is performed to determine the limiting DNBR case and the limiting statepoint. Using this statepoint the maximum allowable pin radial peaks (MARPs) versus axial peaks for different peak locations are determined. The peaking margin is then computed for each assembly from the MARPs and the SIMULATE-3 predicted core power distribution. Positive peaking margin is then confirmed for each pin.

HZP Case VIPRE-O1 model is used. Given the RETRAN-02 5-7 (LAR Page No.: E-7)

statepoint core quadrant inlet temperatures, core quadrant inlet flow rates, core exit pressure, core average surface heat flux, and the assembly axial and radial power distributions from the neutronics code, this [ ] model calculates the statepoint local coolant properties and the DNBR.

The critical heat flux (CHF) correlation used to evaluate the DNBR for Westinghouse fuel is either the Westinghouse W-3S correlation (Reference 5-4, Appendix D) or the Westinghouse WLOP correlation (Reference 5-16). The W-3S CHF correlation has been approved by the NRC for analysis with system pressures as low as 500 psia (Reference 5-6). The W-3S correlation limit is 1.45 in the pressure range of 500-1000 psia.

The WLOP CHF correlation has been approved by the NRC for analysis over the following ranges with a correlation limit of 1.18 (Reference 5-17).

Pressure (psia) 185 to 1800 Local Coolant Quality < 0.75 Local Mass velocity (Mlbm/hr-ft 2) 0.23 to 3.07 Heated Hydraulic Diameter Ratio 0.679 to 1.00 Heated Length (inches) 48* to 168 Grid Spacing Term 27 to 115

• Set as the minimum heated length value, applied at all elevations below 48 inches Two steady-state cases are analyzed: the first case with offsite power available, and the second case with offsite power unavailable. A statepoint DNBR calculation is performed instead of a transient DNBR calculation since the steam line break accident is a slow transient and a statepoint consisting of the limiting surface heat flux and inlet boundary conditions provides conservative DNBR results.

Model Description (HZP Case) 5-8 (LAR Page No.: E-8)

I Axial Power Distributions I

I Radial Power Distributions I

I 5.3 Transient Analysis 5.3.1 Initial Conditions Pressurizer Pressure (HZP Case)

Since this transient is being evaluated for minimum DNBR, a low initial pressurizer pressure is used. The low initial pressure causes an earlier safety injection actuation since the transient starts 5-9 (LAR Page No.: E-9)

closer to the setpoint. This is compensated for in the safety injection setpoint as described below.

Nominal pressurizer pressure less an allowance for uncertainty is assumed.

Pressurizer Pressure (HFP Case)

Since this transient is analyzed using the statistical core design methodology, nominal full power pressurizer pressure is assumed.

Pressurizer Level A low initial pressurizer level minimizes RCS inventory during the transient. This minimizes core outlet pressure and is, therefore, conservative for evaluation of minimum DNBR. This effect more than compensates for the slightly quicker boration when the safety injection fluid mixes with the smaller RCS mass. Nominal pressurizer level less an allowance for uncertainty is assumed.

RCS Temperature (HZP Case)

Since this transient is being evaluated for minimum DNBR, a high initial RCS temperature is used.

A slightly greater reactivity insertion results from starting from a high initial temperature since the slope of the k-effective vs. temperature curve is greater at higher temperatures. The hot zero power programmed RCS temperature of 557°F plus an allowance for uncertainty is assumed.

RCS Temperature (HFP Case)

Since this transient is analyzed using the statistical core design methodology, nominal full power RCS temperature is assumed.

RCS Flow (HZP Case)

Since this transient is being evaluated for minimum DNBR, a low initial RCS flow is used. The effect of lower flow on DNBR more than offsets the decrease in primary-to-secondary heat transfer. The minimum measured flow less an allowance for uncertainty is assumed. High core bypass flow is assumed.

5-10 (LAR Page No.: E-10)

RCS Flow (HFP Case)

Since this transient is analyzed using the statistical core design methodology, RCS thermal design flow is assumed.

Steam Generator Water Inventory Since the primary-to-secondary heat transfer is the driving force behind the excessive RCS cooldown and depressurization, steam generator inventory is maximized to provide the largest cooldown capacity and to prolong the time prior to U-tube uncovery and heat transfer degradation.

The nominal value plus an allowance for uncertainty is assumed.

Core Power (HZP Case)

Initial core heat output would result in a lower temperature decrease since this energy would have to be removed in addition to that stored in the RCS fluid and metal. This would result in a milder transient and would be non-conservative. The core is, therefore, initially at hot zero power, here defined as I x 10-9 times full power.

Core Power (HFP Case)

Since this transient is analyzed using the statistical core design methodology, nominal full power is assumed.

Steam Generator Tube Plugging Assuming no steam generator tube plugging maximizes the steam generator heat transfer area and minimizes the RCS loop flow resistance. Both of these effects enhance primary-to-secondary heat transfer and are, therefore, conservative. These effects more than offset the slight decrease in RCS inventory which would result from plugged tubes. Therefore, no tube plugging is assumed for this analysis.

5-11 (LAR Page No.: E- 11)

5.3.2 Boundary Conditions 5.3.2.1 Availability of Systems and Components Reactor Protection System (IHFP Case)

The Reactor Protection System reactor trips that can actuate are high flux, high positive flux rate, and overpower AT. An appropriate trip delay time is assumed.

Excore Flux Detector Error Due to Overcooling (HFP Case)

As the steam generators depressurize, the saturation temperature decreases and causes excessive primary-to-secondary heat transfer. The resulting decrease in cold leg temperature upon entering the reactor vessel downcomer attenuates the neutron flux leakage. This attenuation effect reduces the flux incident on the excore neutron detectors, thereby creating an error in the indicated flux value (indicated excore detector power less than true reactor power). A conservative attenuation factor is assumed as a function of the change in reactor vessel downcomer density resulting from the change in temperature.

The effect of a change in reactor vessel downcomer water temperature (density) on the excore flux detector signal is modeled by consideration of the relevant source, material composition, and detector geometry details. The fuel region is considered as a homogeneous mixture inside a cylinder of equivalent volume as the core region inside the baffle plates. The balance of the geometry is modeled as a series of concentric cylinders, representing the baffle plates, flow channel, core barrel, neutron pads, downcomer region, and reactor vessel. Particle tallies at the detector account for the details of detector geometry and design.

Fuel characterization is performed with the SAS21VORIGEN-S modules of the SCALE Code System, as necessary (References 5-12 and 5-13). Transport and tallying of particles is performed with the MCNP computer code (Reference 5-14). Variance reduction is performed in MCNP as necessary to achieve reliable statistics for the tallies. The particle tally results are used to characterize the relative effects of Reactor Coolant System temperature (density) on detector response.

5-12 (LAR Page No.: E-12)

Rod Control System (HFP Case)

The Rod Control System is assumed to be in manual to maximize the power increase.

Reactor Coolant Pumps The reactor coolant pumps are assumed to trip when offsite power is lost. For portions of the analysis during which offsite power is maintained, all reactor coolant pumps are assumed to be operating.

Pressurizer Pressure Control No credit is taken for pressurizer heater operation. This assumption enhances the RCS depressurization and is therefore conservative for the evaluation of minimum DNBR.

Pressurizer Level Control No credit is taken for the automatic operation of the Chemical and Volume Control System (CVCS) to attempt to increase RCS mass and thereby maintain pressurizer level and pressure. The charging and letdown flows are assumed to isolate simultaneously and to be balanced prior to isolation. Not taking credit for CVCS action to maintain pressure is conservative for the evaluation of minimum DNBR.

Turbine Control (HFP Case) 5-13 (LAR Page No.: E-13)

Condenser Steam Dump (HZP Case)

The condenser steam dump valves are initially assumed to be open slightly to release the steam generated by the relatively small heat input to the RCS from the reactor coolant pumps. These valves are assumed to be closed after reactor trip. However, since the flow through these valves is very small compared to break flow, the opening or closing these valves has an insignificant effect on the analysis.

Main Feedwater The main feedwater pumps take suction from the hotwell pumps via the condensate booster pumps.

Both of the latter sets of pumps are run from offsite power. When offsite power is lost, both of these types of pumps trip, causing the main feedwater pumps to trip on low suction pressure, condensate booster pump trip, or safety injection. It is assumed that this process takes no more than 5 seconds. For events in which offsite power is maintained, no main feedwater pump trip is assumed. For all cases, no credit is taken for feedwater isolation on low-low RCS average temperature coincident with reactor trip.

Auxiliary Feedwater (HZP Case)

All three auxiliary feedwater pumps are assumed to start on loss of offsite power and deliver flow to all four steam generators. This is conservative since it maximizes the secondary heat sink.

Offsite Power (HZP Case)

As instructed by Section 15.1.5 of Reference 5-1, the assumptions regarding the loss of offsite power and the timing of such a loss were studied to determine their effects on the consequences of the accident. Analyses were performed with offsite power both maintained throughout the transient and lost during the transient. The core is ultimately shut down by borated water from the high and intermediate-head safety injection pumps. In the absence of offsite power, the pumps are powered from emergency buses energized by diesel generators. The diesels start on either a safety injection signal or an undervoltage condition on the emergency buses (indicative of the loss of offsite power). Since delaying diesel generator start delays borated water delivery, and is therefore conservative, the loss of offsite power is timed to coincide with the safety injection actuation.

5-14 (LAR Page No.: E-14)

Offsite Power (HFP Case)

The results of a sensitivity analysis indicate that an assumption of loss of offsite power coincident with turbine trip results in conservative DNBR results.

Safety Injection Pumps (HZP Case)

The injection of borated water introduces negative reactivity and is therefore a benefit. The injection of cold, unborated water is a penalty, however, since it makes the cooldown more severe.

Because of this, the single failure, the loss of one train of safety injection, is timed to coincide with the point at which the high-head safety injection piping is purged of unborated water.

5.3.2.2 Response Times (HZP Case)

Pumped Safety Injection Flow A delay is assumed from the SI setpoint being reached until the SI signal is generated. An additional delay is assumed from the diesel generator start signal until the first load group, which includes the high-head safety injection pump discharge valves, is sequenced onto the emergency bus. A third delay is assumed from the sequencing of the first load group onto the emergency bus until delivery of unborated water to the RCS. The total of these three delays is 33 seconds. For the case in which offsite power is maintained, the corresponding delay is 19 seconds.

Feedwater Isolation Valves Following the receipt of a safety injection signal, an additional delay is assumed to generate a feedwater isolation signal and complete closure of the isolation valves. The total response time for the feedwater isolation function is 12 seconds.

Main Steam Isolation Valves Following the receipt of a steam line isolation signal, an additional delay is assumed to close the main steam isolation and main steam isolation bypass valves. The total response time for the steam line isolation function is 10 seconds.

5-15 (LAR Page No.: E-15)

Auxiliary Feedwater Pumps Since cold auxiliary feedwater flow into the steam generator makes the cooldown more severe, no time delay is assumed between the loss of offsite power and the delivery of flow to the steam generators.

5.3.2.3 Flow From Interfacing Systems Safety Iniection (HZP Case)

Safety injection flow is varied as a function of RCS pressure. The limiting head-flow curves among the high and intermediate head pumps are adjusted to conservatively account for pump head degradation.

Main Feedwater (HZP Case)

At hot zero power, the main feedwater control valve is closed, and the feedwater is delivered to the steam generator upper nozzle through the main feedwater control bypass valve. In assessing the amount of main feedwater flow during a steam line break, the following aspects must be considered: automatic control of pump speed, automatic control of bypass valve position, and line resistance of the piping to the upper nozzle. The speed controller will initially attempt to reduce pump speed. No credit is taken in the analysis for a flow reduction due to this effect. Rather than model the bypass valve controller in detail, the analysis conservatively assumes that the valve instantaneously travels to its full open position. A lower limit is placed on the upper nozzle piping resistance in this configuration. The flow boundary condition is then conservatively increased as steam generator pressure decreases, by assuming that feedwater pump discharge pressure remains constant at the initial value corresponding to the low resistance limit.

Main Feedwater (HFP Case)

Main feedwater pump speed and control valves are assumed to be in manual to maximize the flowrate after the initiation of the break. Flow increases from the full power initial condition value as steam generator pressure decreases. This modeling continues through to the time of turbine trip and the limiting statepoint.

5 -16 (LAR Page No.: E-16)

Auxiliary Feedwater (HZP Case)

Auxiliary feedwater flow is varied as a function of steam generator pressure. The limiting head-flow curves among the motor and turbine-driven pumps are adjusted to conservatively account for installed pump performance being better than the curves and for pump motor speed being higher than predicted.

5.3.2.4 Engineered Safety Features Actuation Setpoints Safety Iniection (HZP Case)

Safety injection is assumed to be actuated at the low pressurizer pressure technical specification setpoint less an allowance for uncertainty.

Steam Line Isolation (HZP Case)

Steam line isolation is assumed to occur at the low secondary pressure technical specification setpoint less an allowance for uncertainty. No credit is taken for steam line isolation on high containment pressure for breaks inside containment.

5.3.2.5 Boron Injection Modeling (HZP Case)

Transport The boron transport model used is described in Section VII.2.5 of Reference 5-2. The boron is assumed to be soluble in the transport medium and to have no direct effect on the fluid equations.

The basic equation computes the time rate of change of boron mass in a control volume from the net inflow from connected volumes plus the net generation within that volume.

Purge Volumes Purge volumes from the outlet of the refueling water storage tank to the inlet of the RCS are separately calculated for both the high and intermediate-head safety injection pumps. These piping volumes are assumed to be initially at a concentration of 0 ppm. Borated water is assumed to reach 5-17 (LAR Page No.: E-17)

the RCS only after an amount of unborated water equal to the purge volume has been injected.

This purging is done separately for the high and intermediate head pumps.

Concentration The boron concentration in the injection water is a conservatively low refueling water storage tank value including an allowance for measurement error.

5.3.2.6 Core Kinetics Modeling (HZP Case)

Point Kinetics The RETRAN-02 point kinetics model is used for the system thermal-hydraulic analysis. The particular option employed uses one prompt neutron group, six delayed neutron groups, eleven delayed gamma emitters, plus U-239 and Np-239. The point kinetics model is adequate for this application since the system analysis does not require detailed modeling of power distribution effects. The power distributions used in the system analysis are determined to be conservative as discussed below. The effective delayed neutron fraction is chosen to minimize Doppler feedback (maximum beta-effective) in units of dollars of reactivity, while the prompt neutron lifetime value is chosen to minimize the ratio of beta-effective to the prompt neutron lifetime. This ratio is a RETRAN-02 input. Minimizing it increases the neutron power spike when prompt criticality is achieved.

Temperature Feedback The basis for the temperature feedback is a relationship between the reactivity vs. temperature curve and [

1, which is input to the point kinetics model. [

5-18 (LAR Page No.: E-18)

I Axial Power Distribution The axial power distribution for the RETRAN-02 analysis is simply the energy deposition fraction for each of the three axial core conductors. These fractions approximate the axial power distribution calculated by the three dimensional core model described in Section 5.2.2.2. The RETRAN-02 axial power distribution at the peak heat flux statepoint is more top-peaked than the distribution calculated by the three dimensional model. This approach is conservative since it results in a more severe return to power.

Radial Power Distribution This approach is conservative since it results in a more severe return to power.

Control Rod Reactivity Since the steam line break transient is a concern chiefly because of power peaking in the vicinity of a stuck rod, the control rods are assumed to begin the transient outside of the core; i.e., the reactor is initially not tripped. Manual action by the operator is assumed to immediately trip the reactor.

This assumption is conservative since any cooldown prior to rod insertion would introduce positive reactivity which would increase core power. This would increase RCS stored energy and cause decay heat generation, both of which cause a less severe cooldown. The amount of negative reactivity introduced by rod insertion is sufficient to make the core subcritical by the technical specification shutdown margin.

5-19 (LAR Page No.: E- 19)

Boron Reactivity The negative reactivity inserted by boration is modeled by [

I core boron concentration. This concentration is multiplied by a boron worth to give a reactivity.

5.3.2.7 Core Kinetics Modeling (IFP Case)

The RETRAN-02 point kinetics model is used for the system thermal-hydraulic analysis. The particular option employed uses one prompt neutron group, six delayed neutron groups, eleven delayed gamma emitters, plus U-239 and Np-239. The point kinetics model is adequate for this application since the system analysis does not require detailed modeling of power distribution effects. The key parameters in this analysis are the break size and the moderator temperature coefficient. The limiting case is determined by analyzing a range of break sizes and varying the moderator temperature coefficient, which is essentially varying the time-in-cycle. Values of the other kinetics parameters are then selected consistent with the time-in-cycle. A minimum Doppler temperature coefficient is assumed to conservatively maximize the core power excursion.

5.4 Results and Conclusions (HZP Case) 5.4.1 Primary and Secondary System Response Sensitivity studies are performed to determine the limiting break size. The steam line break transient is analyzed both with offsite power maintained and with offsite power lost coincident with safety injection actuation. Typical event sequences for the two cases are presented in Tables 5-1 and 5-2 for the 1.4 ft2 break size. Figures 5-4 through 5-14 correspond to the case with offsite power maintained and 5-15 through 5-25 to the case with offsite power lost.

Offsite Power Maintained Steam line pressure in the faulted steam line (Figure 5-4) decreases after the break occurs. The depressurization rate initially increases after steam line isolation occurs, since beyond this point 5-20 (LAR Page No.: E-20)

only the faulted steam generator is supplying steam to the break. The depressurization rate then decreases as the steam line continues to blow down towards atmospheric pressure. Steam line pressure in the intact steam line (shown on the same figure) also decreases until steam line isolation occurs. Beyond this point the intact steam generators, and therefore their associated steam lines, experience a slight pressurization.

The cold leg temperatures (Figure 5-5) closely follow the pressures in the respective steam lines.

The hot leg temperatures (Figure 5-6) follow the cold leg temperatures until the return to power occurs. A larger difference between the hot and cold leg temperatures develops beyond this point due to the core heat output.

Core boron concentration (Figure 5-7) is zero until after the unborated water is purged from the safety injection piping. Thereafter, it slowly increases as the borated safety injection water mixes with the unborated RCS inventory.

The temperatures drive the core reactivity transient shown in Figure 5-8. Reactivity initially drops to the tecntical specification shutdown margin on reactor trip as the rods fall into the core. The positive reactivity inserted due to the decreasing temperatures causes total reactivity to increase until prompt criticality is momentarily achieved. The fuel temperature feedback caused by the sudden power increase causes reactivity to decrease rapidly to near zero. Reactivity decreases slowly as power increases due to increasing fuel temperature feedback. Reactivity decreases further with the addition of borated water from the Safety Injection System.

The neutron power transient (Figure 5-9) caused by this reactivity transient, is zero until prompt criticality occurs. At this point power spikes up and then immediately decreases sharply due to the negative Doppler feedback. Power then increases in equilibrium with reactivity until just after boron reaches the core. This is followed by a slow decrease toward shutdown. The core heat flux (Figure 5-10) is similar to the core power with two exceptions. First, there is some heat flux generated prior to prompt criticality by removal of stored energy from the fuel. Second, the power spike at prompt criticality is too brief to be reflected in the heat flux.

Pressurizer level (Figure 5-11) decreases rapidly until the pressurizer empties. It stays at zero until enough water inventory is added by the Safety Injection System to offset the contraction of the original inventory due to the cooldown. Pressurizer pressure (Figure 5-12) decreases relatively slowly until the pressurizer empties. The decrease is more rapid until the saturation pressure is 5-21 (LAR Page No.: E-21)

reached in the hottest parts of the RCS. Thereafter, pressure increases slowly as inventory addition from the Safety Injection System offsets inventory contraction from the cooldown.

Break flow (Figure 5-13) initially decreases as the steam line pressure decreases. After steam line isolation, flow from the intact loops stops. Beyond this point flow decreases with decreasing pressure.

The core mass fluxes (Figure 5-14) increase with time since the reactor coolant pumps provide essentially constant volumetric flow which, with the decreasing RCS temperatures, is equivalent to an increasing mass flow rate.

Offsite Power Lost Although Figures 5-4 through 5-14 depict the case in which offsite power is maintained, the discussion is generally applicable to Figures 5-15 through 5-25, the case in which offsite power is lost at safety injection. Important exceptions are noted below.

Neutron power (Figure 5-22) does not begin a sustained decrease until after boron from both the high-head and intermediate-head safety injection pumps has reached the core.

The core mass fluxes (Figure 5-25) decrease beyond the point at which offsite power is lost due to the coastdown of the reactor coolant pumps.

The system transient response for each case is reviewed to select the statepoint(s) for the power peaking and DNBR analysis. Values provided for each statepoint include neutron power, core heat flux, core outlet pressure, l

5.4.2 Core Response 5.4.2.1 Axial and Radial Power Distributions Using the limiting statepoints from the RETRAN-02 analyses discussed in Section 5.4.1, axial and radial power distributions are calculated as described in Section 5.2.2.2. The axial power distribution for the offsite power maintained case has a top peaked shape, whereas the offsite 5-22 (LAR Page No.: E-22)

power lost case has a bottom peaked shape. This effect is caused by the difference in moderator temperature feedback resulting from the large difference in RCS flow. Typical values of the maximum axial peaking factors for the peak radial location are [ I for the offsite power maintained and offsite power lost cases, respectively. Figures 5-26 and 5-27 show the asyrmnetric core assembly radial power distributions. The cold quadrant, which contains the stuck rod, has a more highly peaked assembly radial power distribution than the rest of the core. Typical maximum hot assembly pin radial power peaking factors are ] for the offsite power maintained and offsite power lost cases, respectively.

5.4.2.2 Minimum DNBR and CFM Results Using the limiting statepoints from the RETRAN-02 analyses discussed in Section 5.4.1, together with the power distributions discussed in Section 5.4.2.1, the VIPRE-01 [ I model is used to calculate the core local fluid properties and MDNBR. The MDNBRs predicted by either the W-3S ClIHF correlation, or the WLOP CHF correlation, are greater than 1-4-5 their respective CHF limits of 1.45 and 1.18, respectively, for both the offsite power maintained and offsite power lost cases. Margin to the centerline fuel melt limit is also confirmed. Therefore, the criterion that the core remain in place and intact, as discussed in Section 5.1.2, is met. Because this criterion is met, the current UFSAR dose analysis, which assumes no DNBR-related fuel failures, remains valid.

5.5 Results and Conclusions (HFP Case)

The sequence of events for a typical limiting HFP case is shown in Table 5-3. The main steam line depressurization (Figure 5-28) that follows the opening of the break causes an increase in primary-to-secondary heat transfer and a resulting decrease in cold leg temperature. The cold water enters the reactor vessel downcomer (Figure 5-29) where the excore flux signal is attenuated and a deviation between excore detector indicated core power and actual average core power develops (Figure 5-30). The reactor trips on high flux at 13.15 seconds following a 0.50 second delay. The peak core power is 130.83% at 13.31 seconds. These results are evaluated to confirm margin to the centerline fuel melt limit. Figure 5-31 shows a typical trend of minimum transient DNBR. A minimum DNBR value of 1.56 is indicated, which is greater than the WRB-2M statistical DNBR limit of I I (Reference 5-15). Typical MARP limits based on the minimum DNBR statepoint are shown in Figure 5-32. Positive DNBR 5-23 (LAR Page No.: E-23)

margin is confirmed by comparing SIMULATE-3 power distributions calculated at the RETRAN-02 statepoint against the MARP limits.

5.56 Cycle-Specific Evaluation HZP Case The cycle-specific reload evaluation for the steam line break accident focuses on the conservative

• corephysics parameters input to the system transient modeling. Each reload cycle is evaluated to determine whether confirm that the reaete-In addition, the core power distribution at the limiting peak heat flux statepoint must be evaluated each reload to confirm that DNBR and CFM limits are not exceeded. There is a high degree of confidence that each reload core will be bounded since the system model was developed with:

• The minimum shutdown margin allowed by the technical specifications

• A conservative reactivity versus temperature response

" A conservative Doppler coefficient.

if the cycle specific r.a.ivity .heck shows the reaet. r to be subcr.itical with respect to the core assumed in the existing licensing basis analysis, inluding a stuck rod, then the response pr.edicted by the system analysis bounds the reload cor-e. If the reload core is not subcftical at these eendi*tien cycle-specific reactivity check fails, two the following approaches are available to obtain acceptable steam line break analysis results: redesign the reload core, evaluate or reanalyze the transient.

HLFP Case For the HFP case the range of moderator temperature coefficient (MTC) values is confirmed to bound the reload core. The values of the Doppler temperature coefficient are confirmed to be minimum values and consistent with the time-in-cycle of the MTC. Core power 5-24 (LAR Page No.: E-24)

distributions at the limiting statepoint are compared against VIPRE MARP limits to confirm margin to the DNBR limit. Core power distributions are also compared against centerline fuel melt limits to confirm positive margin exists. If any of these checks fail, then further evaluation or reanalysis of the accident is performed, or the reload core is redesigned to demonstrate applicable safety criteria are satisfied.

References 5-1 Standard Review Plan, Volume III, NUREG-0800, NRC, Revision 2, July 1981.

5-2 RETRAN-02: A Program for Transient Thermal-Hydraulic Analysis of Complex Fluid Flow Systems, EPRI NP-1850-CCM, Revision 6, EPRI, December 1995 5-3 SIMULATE-3 Methodology, Advanced Three Dimensional Two-Group Reactor Analysis Code, SOA-95/--518, Studsvik of America, Inc., October 1995.

5-4 VIPRE-0 1: A Thermal-Hydraulic Code for Reactor Cores, EPRI NP-25 11 -CCM-A, Revision 4, EPRI, February 2001.

5-5 Nuclear Design Methodology Using CASMO-4/SIMULATE-3 MOX, Revision 1, DPC-NE-1005-PA, Duke Energy Carolinas, November 2008 5-6 January 31, 1989 letter from A. S. Thadani (NRC) to W. J. Johnson (Westinghouse),

"Acceptance for Referencing of licensing Topical Report, WCAP-9226-P/9227-NP,

'Reactor Core Response to Excessive Secondary Steam Releases."

5-7 Thermal-Hydraulic Transient Analysis Methodology, DPC-NE-3000-PA, Revision 35a, Duke Power Company Energy Carolinas, September-2004 October 2012 5-8 The BWU Critical Heat Flux Correlations, BAW-10199P-A, AREVA NP, August 1996 5-9 Westinghouse Improved Performance Analysis and Design Model (PAD 4.0), WCAP-15063-P-A, Revision 1, Westinghouse, July 2000 5-10 Thermal-Hydraulic Statistical Core Design Methodology, DPC-NE-2005-PA, Revision 4a, December 2008 5-11 WCAP-15025-P, Modified WRB-2 Correlation, WRB-2M, for Predicting Critical Heat Flux in 17x17 Rod Bundles with Modified LPD Mixing Vane Grids, Westinghouse Energy Systems, February 1998 5-25 (LAR Page No.: E-25)

5-12 0. W. Hermann and C. V. Parks, "SAS2H: A Coupled One-Dimensional Depletion and Shielding Analysis Module," NUREG/CR-0200, Volume 1, Section S2, March 2000 5-13 0. W. Hermann and R. M. Westfall, "ORIGEN-S: SCALE System Module to Calculate Fuel Depletion, Actinide Transmutation, Fission Product Buildup and Decay, and Associated Radiation Source Terms," NUREG/CR-0200, Volume 2, Section F7, March 2000 5-14 Judith F. Briesmeister, Ed., "MCNP - A General Monte Carlo N-Particle Transport Code," Los Alamos National Laboratory Report, LA-13709-M, March 2000 5-15 "Westinghouse Fuel Transition Report", DPC-NE-2009-PA, Revision 3a, September 2011 5-16 Addendum 2 to WCAP-14565-P-A Extended Application of ABB-NV Correlation and Modified ABB-NV Correlation WLOP for PWR Low Pressure Applications, Revision 0, April 2008 5-17 February 14, 2008 letter from H. K. Nieh (NRC) to J. A. Gresham (Westinghouse),

"Final Safety Evaluation from Westinghouse Electric Company Topical Report WCAP-14565-P, Addendum 2, Revision 0, "Addendum 2 to WCAP-14565-P-A, Extended Application of ABB-NV Correlation and Modified ABB-NV Correlation WLOP for PWR Low Pressure Applications" 5-26 (LAR Page No.: E-26)

Table 5-1 Sequence of Events for 1.4 ft2 Split Break With Offsite Power Maintained Zero Power Initial Condition Event Time (seconds)

Break occurs / Operator manually trips reactor 0.01 Pressurizer level goes offscale low 22 SI actuation on low pressurizer pressure 35 Steam line isolation on low steam line pressure 36 Criticality occurs 46 SI pumps begin to deliver unborated water to RCS 52 High-head SI lines purged of unborated water / 119 One train of SI fails Peak heat flux occurs 120 Intermediate-head SI lines purged of unborated water 191 5-27 (LAR Page No.: E-27)

Table 5-2 Sequence of Events for 1.4 ft2 Split Break With Offsite Power Lost at SI Actuation Zero Power Initial Condition Event Time (seconds)

Break occurs / Operator manually trips reactor 0.01 Pressurizer level goes offscale low 22 SI actuation on low pressurizer pressure / 35 Offsite power lost Reactor coolant pumps begin to coast down Steam line isolation on low steam line pressure 36 Criticality occurs 52 SI pumps begin to deliver unborated water to RCS 66 High-head SI lines purged of unborated water/ 134 One train of SI fails Intermediate-head SI lines purged of unborated water 223 Peak heat flux occurs 228 5-28 (LAR Page No.: E-28)

Table 5-3 Sequence of Events for 4.9 ft 2 Split Break Full Power Initial Condition Event Time (seconds)

Break occurs 0 High flux trip setpoint reached 12.65 Reactor trips and control rod insertion begins 13.15 Time of peak power 13.31 Turbine trip on reactor trip, Loss of offsite power, reactor 13.45 coolant pumps trip and begin to coast down End of simulation 20 5-29 (LAR Page No.: E-29)

Note: Figures 5-1 through 5-27 are not revised Figure 5-28 Main Steam Line Break HFP Steam Line Pressure 1100 1000 900 R-ý uJ 800 U)

ILI a.

zLU 700

-J 600 Faulted Loop U)

Intact Loops 500 400 0 5 10 15 20 TIME (SECONDS) 5-52 (LAR Page No.: E-30)

Figure 5-29 Main Steam Line Break HFP Cold Leg Temperature 560

.555

• 550 O* 545 -Faulted Loop LU

-- Intact Loops CL U.I 535 U.I o 530 525 520 0 5 10 15 20 TIME (SECONDS) 5-53 (LAR Page No.: E-3 1)

Figure 5-30 Main Steam Line Break HFP Power and Indicated Power 140 120 100 208 oz 80 z

U-0 40 60

---

Attenuated Neutron Power 20 20 0 5 10 15 20 TIME (SECONDS) 5-54 (LAR Page No.: E-32)

Figure 5-31 Main Steam Line Break HFP Minimum DNBR 2.5 uJ 0

0 2 LI.

Q

,h 1.5 0 2 4 6 8 10 12 14 16 TIME (SECONDS) 5-55 (LAR Page No.: E-33)

Figure 5-32 Main Steam Line Break HFP Maximum Allowable Radial Peaks 5-56 (LAR Page No.: E-34)

Appendix F Fuel Thermal Conductivity Degradation Impact to Non-LOCA Safety Analysis (Constitutes Appendix B to DPC-NE-3001)

Non-Proprietary Version B-I (LAR Page No.: F-i)

Appendix B (DPC-NE-3001-P)

Fuel Thermal Conductivity Degradation Impact to Non-LOCA Safety Analysis B.1 Background Fuel thermal conductivity degradation (TCD) is a physical phenomenon in which the material properties of the fuel (pellets) are affected over the course of in-core operation (burnup) resulting in a reduced ability to transfer energy from the pellet to the coolant. Consequently, stored energy in the pellets will be higher at burnups when fuel TCD is considered than when fuel TCD is not considered (other effects, such as power fall-off with burnup must also be considered to account for the overall impact of fuel TCD).

I B.2 Evaluation of Impact of U0 2 Fuel Thermal Conductivity Degradation on Non-LOCA Analyses Duke Energy employs a series of computer codes to analyze the Updated Final Safety Analysis Report (UFSAR) Chapter 15 postulated transients and accidents to demonstrate compliance with NRC's regulations. The analyses simulate postulated events and compare the predicted safety parameter results to key analysis input assumptions and to applicable regulatory criteria.

Simulation of the reactor fuel and its response during a transient is an integral part of the safety analyses.

Within the analysis, the fuel pellet thermal conductivity model deternines the rate at which heat is transferred from the fuel pellet to the fuel cladding through the gas gap, and subsequently to the coolant.

A degraded (lower) fuel pellet conductivity results in higher fuel temperatures at a given linear heat-generation rate. Consequently, the analytical prediction of the fuel thermal conductivity affects the results of certain safety parameters. The affected parameters are:

An evaluation of the impact of fuel TCD on non-LOCA safety analyses for McGuire and Catawba is discussed in subsequent sections of this appendix. The evaluations conclude that existing safety analysis methodologies appropriately account for the effects of fuel TCD, the effects are negligible or margin to the applicable regulatory limit is sufficient to account for the effects of fuel TCD.

The impact of fuel TCD on fuel design and safety analyses are addressed through discussions in the following areas:

Core Design and Analysis o Effect of fuel TCD on Doppler Reactivity Feedback o Potential Effect on the Core Design Reload Design Safety Analysis Review Checklist Evaluations Fuel Rod Design for McGuire and Catawba o [ I Chapter 15 Analyses of Non-LOCA Transients B-2 (LAR Page No.: F-2)

Appendix B (DPC-NE-3001-P)

Fuel Thermal Conductivity Degradation Impact to Non-LOCA Safety Analysis o Thermal-Hydraulic Analysis - DNB o Rod Ejection Accident Fuel Enthalpy Calculation B2.1 Core Design and Analysis The SIMULATE-3 core design model is used to confirn the acceptability of safety related physics inputs assumed in the safety analyses. Fuel temperatures used by SIMULATE-3 are functionalized against power and burnup, and [

The reactivity results presented confirm the predicted reactivity feedback between full power and no load conditions is consistent with plant measurements, [

B2.1.1 Effect of Fuel TCD on Doppler Reactivity Feedback Doppler reactivity feedback in the RETRAN model is calculated using bounding values of the Doppler temperature coefficient and core average fuel temperature. These parameters are selected to minimize or maximize Doppler reactivity feedback in transients sensitive to changes in these input parameters to produce a conservative result. Cycle-specific reload calculations are performed to confirm the acceptability of Doppler temperature coefficient and core average fuel temperature Chapter 15 accident analysis input assumptions.

The effect of fuel TCD is to increase the amount of Doppler feedback for a given transient power level change. [

B2.1.2 Potential Effect on the Core Design Reload Design Safety Analysis Review Checklist Evaluations Key accident analysis input assumptions are compiled and summarized into a single document titled reload design safety analysis review (REDSAR) checklist.

B-3 (LAR Page No.: F-3)

Appendix B (DPC-NE-3001-P)

Fuel Thermal Conductivity Degradation Impact to Non-LOCA Safety Analysis B2.2 Fuel Rod Design for McGuire and Catawba Duke performs the fuel mechanical analysis for McGuire and Catawba using vendor fuel rod design methodology. The methodology and fuel performance code is dependent upon the fuel product being irradiated. Since both the McGuire and Catawba reactors currently contain Westinghouse fuel, current Duke mechanical analyses are performed using the NRC-approved fuel rod design methodology described in Reference 5, and the PAD fuel performance code, Reference 2. [

I 9

I B2.2.1 [

B2.3 Chapter 15 Analyses of Non-LOCA Transients UFSAR Chapter 15 analyses are based on systems and core thermal hydraulic analyses performed using the RETRAN and VIPRE codes. The thermal conductivity model used in the systems analysis is initialized to either a maximum or minimum fuel temperature which bounds the fuel temperatures calculated by the core neutronics model or fuel performance code. The initialization performed minimizes or maximizes reactivity feedback depending upon the event being analyzed. [

B-4 (LAR Page No.: F-4)

Appendix B (DPC-NE-3001-P)

Fuel Thermal Conductivity Degradation Impact to Non-LOCA Safety Analysis

] The effect of fuel conductivity degradation would be to increase stored energy and Doppler feedback which would limit the overall transient response (i.e. heat flux) for over-power events, resulting in a DNB benefit.

B2.3.1 Thermal-Hydraulics Analysis - DNB Duke core thermal-hydraulic analyses are performed using the VIPRE code. VIPRE is executed in either steady state or transient mode as applicable.

I B2.3.2 Rod Ejection Accident Fuel Enthalpy Calculation Rod ejection accident fuel enthalpy calculations are performed using the transient VIPRE fuel rod conduction model. Fuel enthalpy is calculated using transient neutron core power and power distribution forcing functions from a 3-D kinetics analysis. As noted in the previous section, the default fuel rod conduction model in VIPRE does not account for fuel thermal conductivity degradation. The impact of this limitation is the reduction in heat transfer from inside of the fuel pellet to the outer surface during the transient is not modeled for fuel with high bumup where fuel thermal conductivity degradation becomes important resulting in an underestimation of fuel enthalpy. As a result, not accounting for fuel thermal conductivity degradation in high burnup fuel is non-conservative. The effect in fresh fuel is negligible.

Sensitivity studies were performed to determine the increase in fuel enthalpy resulting from fuel thermal conductivity degradation.

B-5 (LAR Page No.: F-5)

Appendix B (DPC-NE-3001-P)

Fuel Thermal Conductivity Degradation Impact to Non-LOCA Safety Analysis Current analysis of record (AOR) enthalpy results for McGuire and Catawba are shown in Table B-1.

Table B-1 McGuire and Catawba AOR Peak Enthalpy Results (No Fuel TCD)

Parameter HFP BOC HZP BOC HFP EOC HZP EOC Initial Power level, % 102 0 102 0 Maximum Enthalpy, cal/gm 94 59 64 97 The peak fuel enthalpy results for McGuire and Catawba are sufficiently low to accommodate a fuel thermal conductivity degradation enthalpy [ ] without challenging the peak fuel enthalpy acceptance criteria.

B3.0 Summary This appendix addressed the impact of fuel TCD for Duke non-LOCA analyses. The affects of fuel TCD are included in the core neutronics model used to set and verify accident analysis fuel temperature and Doppler fuel temperature coefficient assumptions. The impact of fuel TCD to fuel rod design criteria margins can be accommodated through analysis margins and by more rigorous analyses that take into account actual fuel rod power histories as a function of burnup. The impact to the centerline fuel melt (CFM) design criterion is influenced the most, With the exception of the rod ejection accident fuel enthalpy calculation, the Duke systems analysis and thermal analysis methodology is conservative relative to the fuel conductivity phenomena. The effects of fuel TCD in the rod ejection accident fuel enthalpy calculation is accommodated by margin to the acceptance criteria.

B5.0 References

1. DPC-NE-1005-PA, Rev. 1, "Nuclear Design Methodology Using CASMO-4/SIMULATE-3 MOX",

Duke Energy Carolinas, November 2008

2. WCAP-15063-P-A, "Westinghouse Improved Performance Analysis and Design Model (PAD 4.0),

Revision 1, July 2000

3. Letter from J. A. Gresham (Westinghouse) to U.S.NRC, "Westinghouse Response to December 16, 2011 NRC Letter Regarding Nuclear Fuel Thermal Conductivity Degradation", February 17, 2012

4. NUREG/CR-7024, "Material Property Correlations: Comparisons between FRAPCON-3.4, FRAPTRAN1.4, and MATPRO", March 2011 B-6 (LAR Page No.: F-6)

Appendix B (DPC-NE-3001-P)

Fuel Thermal Conductivity Degradation Impact to Non-LOCA Safety Analysis

5. DPC-NE-2009-PA, Rev. 3a, "Duke Power Company Westinghouse Fuel Transition Report",

September 2011 B-7 (LAR Page No.: F-7)

Enclosure 2 AFFIDAVIT of Garry D. Miller

1. I am Senior Vice President of Nuclear Engineering, Duke Energy Corporation, and as such have the responsibility of reviewing the proprietary information sought to be withheld from public disclosure in connection with nuclear plant licensing and am authorized to apply for its withholding on behalf of Duke Energy.

2. I am making this affidavit in conformance with the provisions of 10 CFR 2.390 of the regulations of the Nuclear Regulatory Commission (NRC) and in conjunction with Duke Energy's application for withholding which accompanies this affidavit.

3. I have knowledge of the criteria used by Duke Energy in designating information as proprietary or confidential.

4. Pursuant to the provisions of paragraph (b) (4) of 10 CFR 2.390, the following is furnished for consideration by the NRC in determining whether the information sought to be withheld from public disclosure should be withheld.

(i) The information sought to be withheld from public disclosure is owned by Duke Energy and has been held in confidence by Duke Energy and its consultants.

(ii) The information is of a type that would customarily be held in confidence by Duke Energy. The information consists of analysis methodology details, analysis results, supporting data, and aspects of development programs, relative to a method of analysis that provides a competitive advantage to Duke Energy.

(iii) The information was transmitted to the NRC in confidence and under the provisions of 10 CFR 2.390, it is to be received in confidence by the NRC.

(iv) The information sought to be protected is not available in public to the best of our knowledge and belief.

(v) The proprietary information sought to be withheld in the submittal is that which is marked in the proprietary version of DPC-NE-3001 -P, MultidimensionalReactor Transients and Safety Analysis Physics ParametersMethodology. This information enables Duke Energy to:

(a) Support license amendment requests for its McGuire and Catawba reactors.

(b) Perform nuclear design calculations on McGuire and Catawba reactor cores containing low enriched uranium fuel.

(vi) The proprietary information sought to be withheld from public disclosure has substantial commercial value to Duke Energy.

(a) Duke Energy uses this information to reduce vendor and consultant expenses associated with supporting the operation and licensing of nuclear power plants.

(b) Duke Energy can sell the information to nuclear utilities, vendors, and consultants for the purpose of supporting the operation and licensing of nuclear power plants.

Page 1 of 2

Enclosure 2 (c) The subject information could only be duplicated by competitors at similar expense to that incurred by Duke Energy.

5. Public disclosure of this information is likely to cause harm to Duke Energy because it would allow competitors in the nuclear industry to benefit from the results of a significant development program without requiring a commensurate expense or allowing Duke Energy to recoup a portion of its expenditures or benefit from the sale of the information.

Garry D. Miller affirms that he is the person who subscribed his name to the foregoing statement, and that all the matters and facts set forth herein are true and correct to the best of his knowledge.

a^rry D. Miller Subscribed and sworn to me: 2_ 2.

Date Notary Public My commission expires: ,oaOi.

SEAL Page 2 of 2