ML081280096

WCAP-15885, Rev. 0, R.E. Ginna Heatup and Cooldown Limit Curves for Normal Operation.

ML081280096
Person / Time
Site:	Ginna
Issue date:	07/31/2002
From:	Laubham T Westinghouse
To:	Office of Nuclear Reactor Regulation
References
WCAP-15885, Rev 0
Download: ML081280096 (59)

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Westinghouse Non-Proprietary Class 3 WCAP-15885 July 2002 Revision 0 R. E. Ginna Heatup and Cooldown Limit Curves for Normal Operation Swo inghouse

WESTINGHOUSE NON-PROPRIETARY CLASS 3 WCAP-15885, Revision 0 R.E. Ginna Heatup and Cooldown Limit Curves for Normal Operation T. J. Laubham JULY 2002 Prepared by the Westinghouse Electric Company LLC for the Rochester Gas and Electric Corporation

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Approved:A rehn d[ G h Manager Engineering and Materials Technology Westinghouse Electric Company LLC Energy Systems P.O. Box 355 Pittsburgh, PA 15230-0355

©2002 Westinghouse Electric Company LLC All Rights Reserved

ii PREFACE This report has been technically reviewed and verified by:

J.H Ledgef A~~4 Sections 1, 2, 4, 5, 6 and 7 S.L. Anderson Section 3 Only i)

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iii TABLE OF CONTENTS LIST O F TA B L ES .................................................................................................................................. iv LIST OF FIGURES ........... ......................................................................................... ...... vi EX ECUTIV E SU MMA RY .................................................................................................................... vii 1 INT RO D UC T ION ......................................................................................... ............................ I 2 FRACTURE TOUGHNESS PROPERTIES ............................................................................ 2 3 RADIATION ANALYSIS AND NEUTRON DOSIMETRY .................................................. 9 4 CRITERIA FOR ALLOWABLE PRESSURE-TEMPERATURE RELATIONSHIPS ............. 34 5 CALCULATION OF ADJUSTED REFERENCE TEMPERATURE ..................................... 38 6 HEATUP AND COOLDOWN PRESSURE-TEMPERATURE LIMIT CURVES .................. 43 7 RE FER EN CE S ......................................................................................................................... 50

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iv LIST OF TABLES Table 1 Summary of the Best Estimate Cu and Ni Weight Percent and Initial RTNDT Values for the R.E. Ginna Reactor Vessel Materials ............................................................... 3 Table 2 Inlet (Tcold) Operating Temperatures ......................................................................... 4 Table 3 Calculated Integrated Neutron Exposure of the Surveillance Capsules @ R.E. Ginna, Turkey Point Units 3 &4, Davis-Besse ....................................................................... 5 Table 4 Calculation of Chemistry Factors using R.E. Ginna, Turkey Point Units 3 &4 and Davis-Besse Surveillance Capsule Data ..................................................................... 6 Table 5 Summary of the R.E. Ginna Reactor Vessel Beltline Material Chemistry Factors ...... 8 Table 6 Summary of Calculated Maximum Pressure Vessel Exposure Clad/Base Metal Interface .............................................................................................................. 14 Table 7 Summary of Calculated Maximum Exposure ofthe Intermediate to Nozzle Shell Circumferential Weld and the Nozzle Shell Course Clad/Base Metal Interface ........... 15 Table 8 Calculated Surveillance Capsule Exposure ................................................................ 16 Table 9 Calculated Surveillance Capsule Lead Factors .......................................................... 16 Table 10 Measured Sensor Specific Activities and Reaction Rates (Capsule V) ........................ 20 Table 11 Measured Sensor Specific Activities and Reaction Rates (Capsule R) ........................ 21 Table 12 Measured Sensor Specific Activities and Reaction Rates (Capsule T) ........................ 22 Table 13 Measured Sensor Specific Activities and Reaction Rates (Capsule S) ........................ 23 Table 14 Summary of Sensor Reaction Rates from Capsules V, R, T, and S ............................. 24 Table 15 Comparison of Measured, Calculated, and Best Estimate Reaction Rates at the Surveillance Capsule Center ..................................................................................... 29 Table 16 Comparison of Calculated and Best Estimate Exposure Parameters at the Surveillance C apsule C enter ........................................................................................................ 31 Table 17 Comparison of Measured and Calculated Neutron Sensor Reaction Rates For In-Vessel Surveillance Capsules V, R, T, and S ....................................................................... 33 Table 18 Comparison of Best Estimate and Calculated Fast Neutron Exposure Rates For In-Vessel

!.) Surveillance Capsules V, R, T, and S ........................ ................................................ 33

v LIST OF TABLES Table 19 Summary of the Vessel Surface, 1/4T and 3/4T Fluence Values used for the Generation of the 52 EFPY Heatup/Cooldown Curves .............................................. 39 Table 20 Summary of the Calculated Fluence Factors used for the Generation of the 52 EFPY Heatup and Cooldown Curves .................................................................................. 39 Table 21 Calculation of the ART Values for the 1/4T Location @ 52 EFPY ............................. 40 Table 22 Calculation of the ART Values for the 3/4T Location @ 52 EFPY ............................. 41 Table 23 Summary of the Limiting ART Values Used in the Generation of the R.E. Ginna H eatup/Cooldown Curves ........................................................................................ 42 Table 24 52 EFPY Heatup Curve Data Points Using 1996 App. G & ASME Code Case N-641 (without Uncertainties for Instrumentation Errors) ........................................................ 47 Table 25 52 EFPY Cooldown Curve Data Points Using 1996 App. G & ASME Code Case N-641 (without Uncertainties for Instrumentation Errors) ........................................................ 49

vi LIST OF FIGURES Figure 1 R.E. Ginna r, 0 Reactor Geometry at the core Midplane ............................................ 12 Figure 2 R.E. Ginna r,z Geometry .......................................................................................... 13 Figure 3 R.E. Ginna Reactor Coolant System Heatup Limitations (Heatup Rates of 60 &

100 0F/hr) Applicable for the First 52 EFPY (Without Margins for Instrumentation Errors) Using 1996 App. G Methodology ..... 45 Figure 4 R.E. Ginna Reactor Coolant System Cooldown Limitations (Cooldown Rates up to 100°F/hr) Applicable for the First 52 EFPY (Without Margins for Instrumentation Errors) Using 1996 App. G Methodology ..... 46

vii EXECUTIVE

SUMMARY

This report provides the methodology and results of the generation of heatup and cooldown pressure temperature (PT) limit curves for normal operation of the R.E. Ginna reactor vessel. The PT curves were generated based on the latest available reactor vessel information and updated fluences. The new R.E.

Ginna heatup and cooldown pressure-temperature limit curves were generated using the "axial flaw" methodology of 1995ASME Code,Section XI through the 1996 Addenda. 6kithe PT curves'were*

deveope odeCas N-41, hic al~wsthe ~s&f t&Kkmetodlgy (ASM -Code Case, usng.SME NIý-640) and he rlaxed "circ. flaw" method1ol"ý (ASMECode CsqNý 8)

The material with the highest adjusted reference temperature (ART) was the intermediate to lower shell girth weld seam. However, as it turns out, the intermediate shell forging is limiting for the lower temperature portion of the PT curves despite having a lower set of ART values. This is due to the fact that the higher ART values come from a circumferential weld and ASME Code Case N-641 (or N-588) allows for less restrictive methodology when a circumferential weld has the higher ART values. Thus, composite heatup and cooldown PT curves were developed from two different sets of PT curves: 1) PT curves using the highest ART values from the Intermediate to lower shell girth weld seam with the "circ flaw" methodology (Code Case N-641 or N-588), and 2) PT curves using the highest "axial flaw" ART values (which are lower than the "Circ flaw ART values) with the "axial flaw" methodology (1996 ASME Code).

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1 INTRODUCTION Heatup and cooldown limit curves are calculated using the adjusted RTNDT (reference nil-ductility temperature) corresponding to the limiting beltline region material of the reactor vessel. The adjusted RTNDT of the limiting material in the core region of the reactor vessel is determined by using the unirradiated, reactor vessel material fracture toughness properties, estimating the radiation-induced ARTNDT, and adding a margin. The unirradiated RTNDT is designated as the higher of either the drop weight nil-ductility transition temperature (NDTI) or the temperature at which the material exhibits at least 50 ft-lb of impact energy and 35-mil lateral expansion (normal to the major working direction) minus 600 F.

RTNDT increases as the material is exposed to fast-neutron radiation. Therefore, to find the most limiting RTNDT at any time period in the reactor's life, ARTNDT due to the radiation exposure associated with that time period must be added to the unirradiated RTNDT (IRTNDT). The extent of the shift in RTNDT is enhanced by certain chemical elements (such as copper and nickel) present in reactor vessel steels. The Nuclear Regulatory Commission (NRC) has published a method for predicting radiation embrittlement in Regulatory Guide 1.99, Revision 2, "Radiation Embrittlement of Reactor Vessel Materials."B 11 Regulatory Guide 1.99, Revision 2, is used for the calculation of Adjusted Reference Temperature (ART) values (RTNDT + ARTNDT + margins for uncertainties) at the 1/4T and 3/4T locations, where T is the thickness of the vessel at the beltline region measured from the clad/base metal interface.

The heatup and cooldown curves documented in this report were generated using the most limiting ART values and the NRC approved methodology documented in WCAP- 14040-NP-A, Revision 2E21, "Methodology Used to Develop Cold Overpressure Mitigating System Setpoints and RCS Heatup and Cooldown Limit Curves" with exception of the following: 1) The fluence values used in this report are calculated fluence values (i.e. comply with Reg. Guide 1.190), not the best estimate fluence values. 2)

The Ki, critical stress intensities are used in place of the Kia critical stress intensities. This methodology is taken from approved ASME Code Case N-64P13 (which covers Code Cases N-640 and N-588). 3) The 1996 Version of Appendix G to Section XIE41 will be used rather than the 1989 version.

The purpose of this report is to present the calculations and the development of the R.E. Ginna heatup and cooldown curves for 52 EFPY This report documents the neutron fluence evaluation, the calculated ART values and the development of the PT limit curves for normal operation. The PT curves herein were generated without instrumentation errors. The PT curves include a hydrostatic leak test limit curve from 2485 psig to 1500 psig, along with the pressure-temperature limits for the vessel flange region per the requirements of 10 CFR Part 50, Appendix Gd].

2 2 FRACTURE TOUGHNESS PROPERTIES The fracture-toughness properties of the ferritic materials in the reactor coolant pressure boundary are determined in accordance with the NRC Standard Review Plan 161. The beltline material properties of the R.E. Ginna reactor vessel are presented in Table 1.

Best estimate copper (Cu) and nickel (Ni) weight percent values used to calculate chemistry factors (CF) in accordance with Regulatory Guide 1.99, Revision 2, are provided in Table 1. Additionally, surveillance capsule data is available for four capsules (Capsules V, R, T and S) already removed from the R.E. Ginna reactor vessel. This surveillance capsule data was also used to calculate CF values per Position 2.1 of Regulatory Guide 1.99, Revision 2 in Table 4. These CF values are summarized in Table 5. It should be noted that in addition to R.E. Ginna, surveillance weld data from Turkey Point Unit 3 and 4 and Davis-Besse was used in the determination of CF for the nozzle to intermediate shell girth weld of heat # 71249.

Per WCAP-15092, Revision 3[71, the weld heat # 71249 was determined to be not credible. It should be noted here that the intermediate shell forging was determined not to be credible, while the lower shell forging and the intermediate to lower shell girth weld seam were determined to be credible.

The Regulatory Guide 1.99, Revision 2 methodology used to develop the heatup and cooldown curves documented in this report is the same as that documented in WCAP-14040, Revision 2.

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3 TABLE 1

) Summary of the Best Estimate Cu and Ni Weight Percent and Initial RTNDT Values for the R.E. Ginna Reactor Vessel Materials Material Description Cu (%)(8) Ni(%)(a) Initial RTNDTb)

Closure Head Flange n/a n/a -75 0 F Vessel Flange n/a n/a -520 F Nozzle Shell Forging 123P118(a) 0.07 0.68 30OF Intermediate Shell Forging 125S255(a) 0.07 0.69 20°F Lower Shell Forging 125P666(a) 0.05 0.69 40OF Nozzle to Intermediate Shell Girth Weld (Heat # 0.23 0.59 10°F 71249)w Intermediate Shell to Lower Shell Girth Weld 0.25 0.56 -4.80F (Heat # 6 1 7 8 2 )(d)

Ginna Surveillance Weld 0.23 0.53 -- -

(Heat # 61782)(e)

Notes:

(a) The Cu & Ni for the forgings were taken from material, WCAP-725418 1,WCAP-14684191or RVID2. For the inter. & lower shell forgings, RVID2 has 0.68 Ni, however, the material cert. has 0.69 Ni. Thus, the higher Ni value will be used in the calculations. The nozzle forging copper value was not reported on the material cert, thus, per RG. 1.99 one should assume 0.35 unless justification is provided. Since the nozzle forging is made from the same material as the inter. & lower shell forgings at the same time period, it is a safe assumption to say the copper value would equal the highest Cu value from the known forgings on the vessel (i.e. 0.07).

(b) The Initial RTNDT values are measured values unless otherwise noted. The values were obtained from RVID2 or WCAP-14684191.

(c) The nozzle shell to inter, shell girth weld was fabricated from weld wire heat # 71249 Linde 80, flux Lot 8445.

This is the exact heat and flux as the inter, to lower shell girth weld on the Turkey Point Units 3 & 4 Reactor Vessel. It is also identical to the Turkey Point 3 & 4 Surv. weld material. The best estimate Cu & Ni was taken from WCAP-15092 R.31'3. This differs from RVID2 for Ginna but not for Turkey Pt. (Ref. BAW 2325).

(d) The intermediate shell to lower shell girth weld was fabricated from weld wire heat number 61782 Linde 80, flux Lot 8350. This is the same heat surveillance weld material (flux lot is 8346), but differ flux lot. The best estimate Cu & Ni was taken from WCAP-146841 91 and RVID2.

8 (e) The Ginna surveillance weld best estimate average Cu & Ni was determined from one unirradiated samplet Or 10,]one irradiated sample from Cap. T (Specimen W26) (0o, two irradiated samples from reconstituted specimens Cap. T, and 4) Eight irradiated samples from Cap. S (W22, W23, W27, W28, W32, W35, W36 and W37) [9].

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4 The chemistry factors were calculated using Regulatory Guide 1.99 Revision 2, Positions 1. 1 and 2.1.

) Position 1.1 uses the Tables from the Reg. Guide along with the best estimate copper and nickel weight percents. Position 2.1 uses the surveillance capsule data from all capsules withdrawn to date. The fluence values used to determine the CFs in Table 4 are the calculated fluence values at the surveillance capsule locations. Hence, the calculated fluence values were used for all cases.

In order to account for operating temperature differences, the measured ARTNDT values from the Turkey Point Units 3 and 4 and Davis-Besse surveillance weld data (heat #71249) were adjusted so it can be applied to the R.E. Ginna nozzle to intermediate shell girth weld, which is of the same heat. No adjustment for chemistry was necessary since the overall best estimate Cu and Ni for heat 71249 is higher than the surveillance specimen average Cu and Ni (i.e. a ratio less than 1.0)11. The measured ARTNDT values from the intermediate to lower shell girth weld, which is contained in the R.E. Ginna surveillance program, were adjusted for chemistry using the ratio procedure given in Position 2.1 of Regulatory Guide 1.99, Revision 2. See Table 2 below for the Tcold operating temperatures at R.E. Ginna, Turkey Point and Davis-Besse.

TABLE 2 Inlet (Tcold) Operating Temperatures IGinna I Turkey Point I Davis Besse Average of the Tcold values for each Capsule removed to date for R.E. Ginna documented in the E900 Database. Current Tcold is 531°F, which started in 1996. This value will be used in future evaluations with subsequent capsule withdrawals

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5 Contained in Table 3 are the calculated fluence values for the four capsules removed from the Ginna reactor vessel to date. Also included are the Calculated fluence values for Turkey Point Units 3 and 4 and Davis Besse. The Ginna fluence values are documented in Section 3 of this report. They were determined using ENDF/B-VI cross-sections and followed the guidance in Regulatory Guide 1.1900"'. The best available fluence information for Turkey Point Unit 3 and 4 comes from WCAP-14044t 21 . The calculated fluences irrWCAP-14044 were determined using ENDF/B-IV cross-sections. Thus, for conservatism the calculated fluences were increased 15% to account for going to ENDF/B-VI. [Note that WCAP-15092 Rev. 3 used the higher measured fluences from WCAP- 14044] The Davis Besse material was partially irradiated at Turkey Point and Davis Besse. Per Turkey Point document PTN-ENG-SESJ 0118 Revision 01 (which is a reference in WCAP-15092 Rev. 3), the cumulative measured fluence was 2.57 E19 n/cm 2. Since this too was determined using ENDF/B-IV cross-sections, then it will be increased 15% as well. This should be conservative since 15% times the calculated fluences from WCAP-14044 will not increase above the measured fluences.

TABLE 3 Calculated Integrated Neutron Exposure of the Surveillance Capsules

@ R.E. Ginna, Turkey Point Units 3 & 4, Davis-Besse Capsule I Fluence

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NOTES:

(a) Per WCAP-14044 and increased 15% to account for ENDF/B-VI Cross-sections.

(b) Per Turkey Point document PTN-ENG-SESJ 0118 Revision 011, the fluence is combined between Turkey Point and Davis Besse equivalent to 2.57 EM9. This value was also increased 15%

) to account for ENDF/B-VI cross-sections.

6 TABLE 4 Calculation of Chemistry Factors using R.E. Ginna, Turkey Point & Davis Besse Surveillance Capsule Data m m Material Capsule Capsule P) FF b) ARTNDT'C) FF*ARTNDT FFa Lower Shell V 0.587 0.851 25 21.275 0.724 Forging 125P666 R 1.02 1.006 25 25.150 1.012 T 1.69 1.144 30 34.320 1.309 S 3.64 1.335 42 56.070 1.782 SUM: 136.815 4.827 CFLsF ISP6m= 7(FF

• RTlNDT) + Z( FF 2) (136.815) (4.827) = 28.3 0F Intermediate Shell V 0.587 0.851 0 0 0.724 Forging 125S255 R 1.02 1.006 0 0 1.012 T 1.69 1.144 0 0 1.309 S 3.64 1.335 60 80.1 1.782 SUM: 80.1 4.827 CFispx2s5s5 = Z(FF

• RT~NT) + Z( FF2) = (80.1) + (4.827) = 16.6 0F Ginna Surveillance V 0.587 0.851 149.8 (140) 127.480 0.724 Weld Metal R 1.02 1.006 176.6 (165) 177.660 1.012 (Heat # 61782) T 1.69 1.144 160.5 (150) 183.612 1.309 S 3.64 1.335 219.4 (205) 292.899 1.782 SUM: 781.651 4.827 2

CFHt L61792 = (FF

• RT*T) + ( FF ) = (781.65 1) + (4.827) = 161.9"F Turkey Point Davis 2.956 1.287 221 (215) 284.427 1.656 Surveillance Weld T (TP3) 0.699 0.900 163 (166) 146.700 0.810 Material(d) V (TP3) 1.484 1.109 176 (179) 195.184 1.230 (Heat # 71249) T(TP4) 0.673 0.889 208 (211) 184.912 0.790 SUM: 811.223 4.486 CF Ht71249 = Y.(FF

• RTT) -- Z( FF2) = (811.223-F) - (4.486) = 180.8 0 F See Next Page for Notes

7 Notes:

(a) f= fluence. See Table 3, (x 10'9 n/cm 2, E > 1.0 MeV).

(b) FF = fluence factor = f 0.1 log ).

-_21" (c) ARTNT values are the measured 30 ft-lb shift values taken from the following documents:

- Ginna Plate and Weld... WCAP-1468410 1 .

- Turkey Point & Davis Besse... WCAP-15092 R-3171.

(d) Ginna operates with an average inlet temperature of approximately 549 0F, Turkey Point 3&4 operate with an average inlet temperature of approximately 546 0F, and Davis Besse operates with an average inlet temperature of approximately 5550F. The measured ART*T values from the Turkey Point 3&4 surveillance program were adjusted by subtracting 3°F to each measured ARTrDT and the Davis Besse surveillance program data was adjusted by adding 6°F to the measured ARTrr value before applying the ratio procedure.

The surveillance weld metal ARTNDT values have been adjusted by a ratio factor of:

Ratio Ginna = 1.07, Ratio Turkey Point = 1.0 (conservative), Ratio Davis Besse = 1.0 (conservative)

The pre-adjusted values are in parenthesis.

8 TABLE 5 Summary of the R.E. Ginna Reactor Vessel Beltline Material Chemistry Factors Material Chemistry Factor Position 1.1 Position 2.1 Nozzle Shell Forging 123P118 44.00 F ---

Intermediate Shell Forging 125S255 44.0 0 F 16.6 0 F Lower Shell Forging 125P666 31.0 0 F 28.3 0 F Nozzle to Intermediate Shell Girth 167.6 0F 180.8°F(&)

Weld (Heat # 71249)

Intermediate Shell to Lower Shell 170.4 0F 161.9 0F Girth Weld (Heat # 61782)

Ginna Surveillance Weld Seams 158.9 0 F - - -

(Heat # 61782)

(a) Using Surveillance Data from Turkey Point.

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.9 3 RADIATION ANALYSIS AND NEUTRON DOSIMETRY

3.1 INTRODUCTION

This section describes a discrete ordinates S. transport analysis performed for the R. E. Ginna reactor to determine te neutron radiation environment within the reactor pressure vessel and surveillance capsules. In this evaluation, fast neutron exposure parameters in terms of fast neutron fluence (E > 1.0 MeV) and iron atom displacements (dpa) were established on a plant and fuel cycle specific basis for the first twenty nine reactor operating cycles. In addition, neutron dosimetry sensor sets from the first four surveillance capsules withdrawn from the R. E. Ginna reactor were re-analyzed using current dosimetry evaluation methodology.

The results of these dosimetry re-evaluations provided a validation of the plant specific neutron transport calculations. The validated calculations were then used to project future fluence accumulation through operating periods extending to 54 effective full power years (efpy).

The use of fast neutron fluence (E > 1.0 MeV) to correlate measured material property changes to the neutron exposure of the material has traditionally been accepted for development of damage trend curves as well as for the implementation of trend curve data to assess vessel condition. In recent years, however, it has been suggested that an exposure model that accounts for differences in neutron energy spectra between surveillance capsule locations and positions within the vessel wall could lead to an improvement in the uncertainties associated with damage trend curves as well as to a more accurate evaluation of damage gradients through the reactor vessel wall.

Because of this potential shift away from a threshold fluence toward an energy dependent damage function

) for data correlation, ASTM Standard Practice E853, "Analysis and Interpretation of Light-Water Reactor Surveillance Results," recommends reporting displacements per iron atom (dpa) along with fluence (E > 1.0 MeV) to provide a data base for future reference. The energy dependent dpa function to be used for this evaluation is specified in ASTM Standard Practice E693, "Characterizing Neutron Exposures in Iron and Low Alloy Steels in Terms of Displacements per Atom." The application of the dpa parameter to the assessment of embrittlement gradients through the thickness of the reactor vessel wall has already been promulgated in Revision 2 to Regulatory Guide 1.99, "Radiation Embrittlement of Reactor Vessel Materials." Therefore, in keeping with the philosophy espoused in the current standards governing pressure vessel exposure evaluations, dpa data is also included in this section.

All of the calculations and dosimetry evaluations described in this report were based on the latest available nuclear cross-section data derived from ENDF/B-VI and made use of the latest available calculational tools. Furthermore, the neutron transport and dosimetry evaluation methodologies follow the guidance and meet the requirements of Regulatory Guide 1.190, "Calculational and Dosimetry Methods for Determining Pressure Vessel Neutron Fluence."'"I Additionally, the methods used to determine the pressure vessel neutron exposure are consistent with the NRC approved methodology described in WCAP-14040-NP-A, "Methodology Used to Develop Cold Overpressure Mitigating System Setpoints and RCS Heatup and Cooldown Limit Curves," January 1996.121

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10 3.2 NEUTRON TRANSPORT CALCULATIONS In performing the fast neutron exposure evaluations for the R. E. Ginna surveillance capsules and reactor vessel, plant specific forward transport calculations were carried out using the following three-dimensional flux synthesis technique:

0(r,9,z) = 0(r, 0)

• r-(rz) q0(r) where 4(rO,z) is the synthesized three-dimensional neutron flux distribution, ý(rO) is the transport solution in r,0 geometry, ý(rz) is the two-dimensional solution for a cylindrical reactor model using the actual axial core power distribution, and ý(r) is the one-dimensional solution for a cylindrical reactor model using the same source per unit height as that used in the r,0 two-dimensional calculation.

For the R. E. Ginna analysis, all of the transport calculations were carried out using the DORT two-dimensional discrete ordinates code Version 3. 1131 and the BUGLE-96 cross-section library1 14]. The BUGLE-96 library provides a 67 group coupled neutron-gamma ray cross-section data set produced specifically for light water reactor application. In these analyses, anisotropic scattering was treated with a P5 legendre expansion and the angular discretization was modeled with an S16 order of angular quadrature.

A plan view of the r,0 model of the R. E. Ginna reactor geometry at the core midplane is shown in Figure 1.

Since the reactor exhibits octant symmetry only a 0° to 450 sector is depicted. In addition to the core, reactor internals, pressure vessel and primary biological shield, the model also included explicit

. =>' representations of the surveillance capsules, the pressure vessel cladding, and the insulation located external to the pressure vessel.

From a neutronic standpoint the inclusion of the surveillance capsules and associated support structure in the analytical model is significant. Since the presence of the capsules and structure has a marked impact on the magnitude of the neutron flux as well as on the relative neutron and gamma ray spectra at dosimetry locations within the capsules, a meaningful evaluation of the radiation environment internal to the capsules can be made only when these perturbation effects are properly accounted for in the analysis.

In developing the rO analytical model of the reactor geometry shown in Figure 1, nominal design dimensions were employed for the various structural components. Likewise, water temperatures and, hence, coolant density in the reactor core and downcomer regions of the reactor were taken to be representative of full power operating conditions. The reactor core itself was treated as a homogeneous mixture of fuel, cladding, water, and miscellaneous core structures such as fuel assembly grids, guide tubes, etc. The rO geometric mesh description of the reactor model shown in Figure 3.2-1 consisted of 170 radial by 67 azimuthal intervals. Mesh sizes were chosen to assure that proper convergence of the inner iterations was achieved on a pointwise basis. The pointwise inner iteration flux convergence criterion utilized in the rO calculations was set at a value of 0.001.

A section view of the rz model of the R. E. Ginna reactor is shown in Figure 2. The model extended radially from the centerline of the reactor core out to a location interior to the primary biological shield and over an axial span from an elevation 1 foot below the active fuel to approximately 1 foot above the active fuel. As in the case of the rO model, nominal design dimensions and full power coolant densities were

11 employed in the calculations. In this case, the homogenous core region was treated as an equivalent

) cylinder with a volume equal to that of the active core zone. The stainless steel former plates located between the core baffle and core barrel regions were also explicitly included in the model. The rz geometric mesh description of the reactor model shown in Figure 2 consisted of 153 radial by 90 axial intervals.

Mesh sizes were chosen to assure that proper convergence of the inner iterations was achieved on a pointwise basis. The pointwise inner iteration flux convergence criterion utilized in the r,z calculations was also set at dvalue of 0.001.

The one-dimensional radial model used in the synthesis procedure consisted of the same 153 radial mesh intervals included in the rz model. Thus, radial synthesis factors could easily be determined on a meshwise basis throughout the entire geometry.

The core power distributions used in the plant specific transport analysis for the R. E. Ginna reactor were taken from the appropriate fuel cycle design reports for Cycles I through 29. The data extracted from the design reports represented cycle average relative assembly powers, burnups, and axial distributions.

Therefore, the calculated results provided data in terms of fuel cycle averaged neutron flux which, when multiplied by the appropriate fuel cycle length, in turn, yielded the incremental fast neutron exposure for each fuel cycle. In constructing, the core source distributions, the energy distribution of the source was based on an appropriate fission split for uranium and plutonium isotopes; and from that fission split, composite values of energy release per fission, neutron yield per fission, and fission spectrum were determined. Fluence projections beyond the end of Cycle 29 were based on the assumption that the core power distribution averaged over Cycles 26 through 29 would be representative of future plant operation.

Cycles 26 through 29 were designed as 18 month fuel cycles using the low leakage fuel management

) concept.

The maximum calculated fast neutron fluence (E > 1.0 MeV) and dpa exposure values for the R. E. Ginna pressure vessel are provided in Table 6. As presented, these data represent the maximum exposure of the pressure vessel clad/base metal interface at azimuthal angles of 0, 15, 30, and 45 degrees relative to the core cardinal axes. The data tabulation includes the plant specific calculated fluence at the end of cycle twenty nine (the last cycle completed at the R. E. Ginna plant) and projections for future operation to 28, 32, 36, 40, 44, 48, 52, and 54 EFPY. Similar data applicable to the intermediate shell to nozzle shell circumferential weld as well as to the nozzle shell course located above the top of the active fuel stack are given in Table 7.

The results of the updated fluence calculations for the four surveillance capsules withdrawn to date from the R. E. Ginna reactor are provided in Table 8. These calculated values of neutron fluence should be used to specify the neutron exposure of the irradiated test specimens for use in materials damage correlations.

Updated lead factors for the R. E. Ginna surveillance capsules are provided in Table 9. The capsule lead factor is defined as the ratio of the calculated fluence at the geometric center of the surveillance capsule to the corresponding maximum calculated fluence at the pressure vessel clad/base metal interface.

In Table 9, the lead factors for capsules that have been withdrawn from the reactor (V, R, T, and S) were based on the calculated fluence values for the irradiation period corresponding to the time of withdrawal for the individual capsules. For the capsules remaining in the reactor (P and N), the lead factors correspond to the calculated fluence values at the projected end of cycle twenty nine, the last fuel cycle completed at the time of analysis. The lead factors provided in Table 9 should be used as the basis for the development of future capsule withdrawal schedules for the R. E. Ginna reactor.

b

.t 0) 40 30 120 ISO 200 240 R Axts (cm)

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13 Figure 2

) R. E. Ginna r,z Geometry Int

)

0 40 80 120 ISO 200 240 R Axis (CM)

14 Table 6 Summary of Calculated Maximum Pressure Vessel Exposure Clad/Base Metal Interface Neutron Fluence [E > 1.0 MeV]

Cumulative Neutron Fluence [n/cm2]

Operating Time

[EFPYJ 0.0 Degrees 15.0 Degrees 30.0 Degrees 45.0 Degrees 24.8 (EOC 29) 2.68e+19 1.69e+19 1.22e+19 1.09e+19 28 2.94e+19 1.85e+19 1.34e+19 1.20e+19 32 3.26e+19 2.05e+19 1.48e+19 1.33e+19 36 3.57e+19 2.25e+19 1.63e+19 1.46e+19 40 3.89e+19 2.45e+19 1.77e+19 1.60e+19 44 4.21e+19 2.65e+19 1.92e+19 1.73e+19 48 4.53e+19 2.85e+19 2.07e+19 1.86e+19 52 4.85e+19 3.05e+19 2.21e+19 2.00e+19 54 5.01e+19 3.15e+19 2.28e+19 2.06e+19

) Iron Atom Displacements Cumulative Iron Atom Displacements [dpa]

Operating Time

[EFPYI 0.0 Degrees 15.0 Degrees 30.0 Degrees 45.0 Degrees 24.8 (EOC 29) 4.37e-02 2.85e-02 2.Ole-02 1.77e-02 28 4.79e-02 3.12e-02 2.20e-02 1.94e-02 32 5.3 le-02 3.46e-02 2.44e-02 2.16e-02 36 5.82e-02 3.79e-02 2.68e-02 2.37e-02 40 6.34e-02 4.12e-02 2.9le-02 2.59e-02 44 6.86e-02 4.46e-02 3.15e-02 2.80e-02 48 7.38e-02 4.79e-02 3.39e-02 3.02e-02 52 7.89e-02 5.12e-02 3.63e-02 3.23e-02 54 8.15e-02 5.29e-02 3.75e-02 3.34e-02

15 Table 7

)

Summary of Calculated Maximum Exposure of the Intermediate to Nozzle Shell Circumferential Weld and the Nozzle Shell Course Clad/Base Metal Interface Neutron Fluence [E > 1.0 MeV]

Cumulative Neutron Fluence fn/cm 2 l Operating Time JEFPYJ 0.0 Degrees 15.0 Degrees 30.0 Degrees 45.0 Degrees 24.8 (EOC 29) 1.05e+18 6.64e+17 4.78e+17 4.26e+17 28 1.16e+18 7.28e+17 5.25e+17 4.68e+17 32 1.28e+18 8.07e+17 5.83e+17 5.21e+17 36 1.41e+18 8.86e+17 6.41e+17 5.74e+17 40 1.54e+18 9.65e+17 6.98e+17 6.27e+17 44 1.66e+18 1.04e+18 7.56e+17 6.80e+17 48 1.79e+18 1.12e+18 8.14e+17 7.32e+17 52 1.92e+18 1.20e+18 8.72e+17 7.85e+17 54 1.98e+18 1.24e+18 9.00e+17 8.12e+17

)

Iron Atom Displacements Cumulative Iron Atom Displacements [dpal Operating Time

[EFPY] 0.0 Degrees 15.0 Degrees 30.0 Degrees 45.0 Degrees 24.8 (EOC 29) 1.83e-03 1.19e-03 8.38e-04 7.35e-04 28 2.Ole-03 1.3 1e-03 9.20e-04 8.09e-04 32 2.23e-03 1.45e-03 1.02e-03 9.Ole-04 36 2.45e-03 1.59e-03 1.12e-03 9.92e-04 40 2.67e-03 1.73e-03 1.22e-03 1.08e-03 44 2.89e-03 1.87e-03 1.33e-03 1.17e-03 48 3.1 le-03 2.02e-03 1.43e-03 1.27e-03 52 3.33e-03 2.16e-03 1.53e-03 1.36e-03 54 3.44e-03 2.23e-03 1.58e-03 1.40e-03

,)

16 Table 8

) Calculated Surveillance Capsule Exposure Irradiation Time Fluence (E > 1.0 MeV) Iron Displacements Capsule [EFPY] In/cm2 l [dpal V 1.4 5.87e+18 1.07e-02 R 2.6 1.02e+19 1.85e-02 T 6.9 1.69e+19 2.94e-02 S 17.0 3.64e+19 6.38e-02 Table 9 Calculated Surveillance Capsule Lead Factors Capsule ID And Location Status Lead Factor V (130) Withdrawn EOC 1 2.96 R (130) Withdrawn EOC 3 2.97 T (230) Withdrawn EOC 9 1.82 S (330) Withdrawn EOC 22 1.79 P (23o) In Reactor 1.91 N (33-) In Reactor 1.81 Note: Lead factors for capsules remaining in the reactor are based on cycle specific exposure calculations through fuel cycle twenty nine.

)

17 3.3 NEUTRON DOSIMETRY EVALUATIONS i) 3.3.1 Sensor Reaction Rate Determinations In this section, the results of the evaluations of the four neutron sensor sets withdrawn as a part of the R. E.

Ginna Reactor Vessel Materials Surveillance Program are presented. The capsule designation, location within the ijactor, and time of withdrawal of each of these dosimetry sets were as follows:

Azimuthal Withdrawal Irradiation Capsule ID Location Time Time [efps]

V 130 End of Cycle 1 4.46e+07 R 130 End of Cycle 3 8.05e+07 T 230 End of Cycle 9 2.17e+08 S 330 End of Cycle 22 5.36e+08 The type and radial locations of the neutron sensors within the capsules are summarized as follows:

Radius Sensor Type [cm]

Copper 158.11

44) 158.11 Iron - Core Side Charpy 159.11 Iron - Vessel Side Charpy 158.91 Nickel 158.35 Uranium 238 158.35 Neptunium 237 159.11 Bare Cobalt-Aluminum 159.11 Cd Cov. Cobalt-Aluminum The copper, nickel, and cobalt-aluminum monitors, in wire form, were placed in holes drilled in spacers at several axial levels within the capsules. The cadmium shielded uranium and neptunium fission monitors were accommodated within the dosimeter block located near the center of the capsule. The iron sensors were obtained by cutting small samples from individual charpy specimens taken from several locations within the surveillance capsules.

18 The use of passive monitors such as those listed above does not yield a direct measure of the energy

- : dependent neutron flux at the point of interest. Rather, the activation or fission process is a measure of the integrated effect that the time and energy dependent neutron flux has on the target material over the course of the irradiation period. An accurate assessment of the average neutron flux level incident on the various monitors may be derived from the activation measurements only if the irradiation parameters are well known. In _articular, the following variables are of interest:

" The measured specific activity of each monitor,

• The physical characteristics of each monitor,

" The operating history of the reactor,

" The energy response of each monitor, and

• The neutron energy spectrum at the monitor location.

The radiometric counting of each of the R. E. Ginna dosimetry data sets was accomplished by Westinghouse using established ASTM procedures. Following sample preparation and weighing, the activity of each monitor was determined by means of a high resolution gamma spectrometer. For the copper, iron, nickel, and cobalt-aluminum sensors, these analyses were performed by direct counting of each of the individual samples. In the case of the uranium and neptunium fission sensors, the analyses were

-) carried out by direct counting preceded by disolution and chemical separation of cesium from the sensor material.

The irradiation history of the reactor over the irradiation period experienced by Capsules V, R, T, and S was obtained on a monthly basis from reactor startup to the end of the dosimetry evaluation period. For the sensor sets utilized in the surveillance capsules, the half-fives of the product isotopes are long enough that a monthly histogram describing reactor operation has proven to be an adequate representation for use in radioactive decay corrections for the reactions of interest in the exposure evaluations.

Having the measured specific activities, the operating history of the reactor, and the physical characteristics of the sensors, reaction rates referenced to full power operation were determined from the following equation:

R=. nP A

• NoFYZ"*-!-!-C p. (1-e-t I)e'd i=1 ref where:

A = measured specific activity (dps/g)

R = reaction rate averaged over the irradiation period and referenced to operation at a core power level of Pf (rps/nucleus).

• ) No = number of target element atoms per gram of sensor.

19

. F = weight fraction of the target isotope in the sensor material.

Y = number of product atoms produced per reaction.

Pj = average core power level during irradiation period j (MW).

Pr* = maximum or reference core power level of the reactor (MW).

CJ = calculated ratio of ý(E > 1.0 MeV) during irradiation period j to the time weighted average ý(E > 1.0 MeV) over the entire irradiation period.

X- = decay constant of the product isotope (s-).

t = length of irradiation period j (s).

td = decay time following irradiation period j (s).

and the summation is carried out over the total number of monthly intervals comprising the irradiation period.

In the above equation, the ratio Pj/P,, accounts for month by month variation of power level within a given fuel cycle. The ratio Cj is calculated for each fuel cycle using the methodology described in Section 3.2 of this report and accounts for the change in sensor reaction rates caused by variations in flux level due to changes in core power spatial distributions from fuel cycle to fuel cycle. For a single cycle irradiation Cj =

1.0. However, for multiple cycle irradiations, particularly those employing low leakage fuel management, the additional C1 correction must be utilized. This additional correction can be quite significant for sensor sets that have been irradiated for many fuel cycles in a reactor that has transitioned from non-low leakage to low leakage fuel management.

Prior to using the measured reaction rates in the least squares adjustment procedure discussed later in this

) section, additional corrections were made to U23' measurements to account for the presence of U3 5 impurities in the sensors as well as to adjust for the build-in of plutonium isotopes over the course of the irradiation. These corrections were location and fluence dependent and were derived from the plant specific discrete ordinates analysis described in Section 3.2. Corrections were also made to the U23' and Np137 sensor reaction rates to account for gamma ray induced fission reactions that occurred over the course of the irradiation. These photo-fission corrections were, likewise, location dependent and were based on the transport calculations described in Section 3.2.

Results of the sensor reaction rate determinations for Capsules V, R, T, and S are given in Tables 10 through 14. In Tables 10 through 13, the measured specific activities, gradient corrected specific activities, and decay corrected reaction rates are listed for Capsules V, R, T, and S, respectively. A summary of the reaction rates for each capsule is provided in Table 14. The data listed in Table 14 are indexed to the geometric center of the respective capsules and included all corrections for U23 impurities, Pu build-in, and photo-fission effects.

20 Table 10 Measured Sensor Specific Activities and Reaction Rates Capsule V Radially Radially Adjusted Adjusted Average Measured Saturated Saturated Reaction Reaction Radius Activity Activity Activity Rate Rate Sample ID Foil ID [cm.] [dps/g] [dps/g] [dps/g] [rps/atom] [rpslatom]

CU Top 158.11 7.38E+04 4.63E+05 4.43E+05 6.77E-17 CU Top-Mid 158.11 6.77E+04 4.25E+05 4.07E+05 6.21E-17 CU Bot-Mid 158.11 7.48E+04 4.70E+05 4.49E+05 6.86E-17 CU Bottom 158.11 8.13E+04 5.1OE+05 4.89E+05 7.45E-17 6.82E-17 FE W-1 158.11 2.47E+06 5.OOE+06 4.82E+06 7.64E-15 FE R-1 158.11 2.57E+06 5.20E+06 5.02E+06 7.95E-15 FE S-6 158.11 2.18E+06 4.41E+06 4.25E+06 6.74E-15 FE P-7 158.11 2.57E+06 5.20E+06 5.02E+06 7.95E-15 FE W-2 159.11 2.04E+06 4.13E+06 4.78E+06 7.58E-15 FE R-3 159.11 1.95E+06 3.95E+06 4.57E+06 7.25E-15 FE S-8 159.11 2.02E+06 4.09E+06 4.74E+06 7.51E-15 FE P-9 159.11 2.1OE+06 4.25E+06 4.92E+06 7.80E-15 7.55E-15 NI Middle 158.11 2.38E+07" 6.51E+07 6.21E+07 -, 8.90E-1.5 8.90E-15 U Middle 158.35 2.30E+05 7.26E+06 7.26E+06 - 4.77E-14 3.91E-14 NP Middle 158.35 1.23E+06 3.88E+07 3.88E+07 2.48E-13 2.44E-13

) Notes:

" The average U-238(n,f) reaction rate of 2.91E-14 includes the correction of a factor of 0.861 to account for plutonium build-in and an additional factor of 0.950 to account for photo-fission effects in the sensor.

" The average Np-237(n,f) reaction rate of 2.44E-13 includes the correction of a factor of 0.983 to account for the photo-fission effects in the sensor.

)

21 Table 11 Measured Sensor Specific Activities and Reaction Rates Capsule R Radially Radially Adjusted Adjusted Average Measured Saturated Saturated Reaction Reaction Radius Activity Activity Activity Rate Rate Sample ID Foil ID [cm.] [dps/g] [dps/g] [dps/g] [rpslatom] [rpslatom]

74-2204 Top 158.11 1.08E+05 4.42E+05 4.23E+05 6.45E-17 74-2207 Top-Mid 158.11 9.68E+04 3.96E+05 3.79E+05 5.78E-17 74-2213 Bot-Mid 158.11 1.15E+05 4.70E+05 4.50E+05 6.87E-17 74-2216 Bottom 158.11 1. 15E+05 4.70E+05 4.50E+05 6.87E-17 6.49E-17 74-2202 W-13 158.11 2.08E+06 5.19E+06 5.OOE+06 7.93E-15 74-2200 R-14 158.11 1.98E+06 4.94E+06 4.76E+06 7.55E-15 74-2198 P-18 158.11 2.06E+06 5.14E+06 4.95E+06 7.85E-15 74-2203 W-14 159.11 1.63E+06 4.07E+06 4.71 E+06 7.46E-15 74-2201 R-15 159.11 1.70E+06 4.24E+06 4.91 E+06 7.78E-15 74-2199 P-1 9 159.11 1.85E+06 4.61 E+06 5.34E+06 8.47E-15 7.84E-15 74-2210 Middle 158.11 5.83E+06 7.36E+07 7.03E+07 1.01E-14 1.01E-14 74-2220 Middle 158.35 4.32E+05 7.79E+06 7.79E+06 5.12E-14 4.11E-14 74-2219 Middle 158.35 4.25E+06 7.66E+07 7.66E+07 4.89E-13 4.81 E-13 74-2205 Top 159.11 3.09E+07 1.26E+08 1.22E+08 7.97E-12 74-2208 Top-Mid 159.11 3.14E+07 1.28E+08 1.24E+08 8.10E-12 74-2211 Middle 159.11 2.96E+07 1.21E+08 1.17E+08 7.64E-12 74-2214 Bot-Mid 159.11 2.94E+07 1.20E+08 1.16E+08 7.59E-12 74-2217 Bottom 159.11 2.94E+07 1.20E+08 1.16E+08 7.59E-12 7.78E-12 74-2206 Top 159.11 1.19E+07 4.87E+07 5.69E+07 3.71E-12 74-2209 Top-Mid 159.11 1.18E+07 4.83E+07 5.64E+07 3.68E-12 74-2212 Middle 159.11 1.07E+07 4.38E+07 5.11E+07 3.34E-12 74-2215 Bot-Mid 159.11 1.24E+07 5.07E+07 5.92E+07 3.87E-12 74-2218 Bottom 159.11 1.24E+07 5.07E+07 5.92E+07 3.87E-12 3.69E-12 Notes:

• The average U-238(n,f) reaction rate of 4.1 1E-14 includes the correction of a factor of 0.845 to account for plutonium build-in and an additional factor of 0.950 to account for photo-fission effects in the sensor.

• The average Np-237(nf) reaction rate of 4.8 IE-13 includes the correction of a factor of 0.983 to account for the photo-fission effects in the sensor.

22 Table 12 Measured Sensor Specific Activities and Reaction Rates Capsule T Radially Radially Adjusted Adjusted Average Measured Saturated Saturated Reaction Reaction Radius Activity Activity Activity Rate Rate Sample ID Foil ID [cm.] [dpslg] [dpslg] [dps/g] [rps/atom] [rpslatom]

81-1392 Top 158.11 1.60E+05 3.51 E+05 3.35E+05 5.11E-17 81-1395 Top-Mid 158.11 1.40E+05 3.07E+05 2.93E+05 4.47E-17 81-1402 Bot-Mid 158.11 1.66E+05 3.64E+05 3.48E+05 5.30E-17 81-1415 Bottom 158.11 1.74E+05 3.82E+05 3.64E+05 5.56E-17 5.11E-17 81-3390 S-22 158.11 1.14E+06 3.36E+06 3.19E+06 5.06E-15 81-3392 P-28 158.11 1.27E+06 3.74E+06 3.56E+06 5.64E-15 81-3394 W-21 158.11 1.30E+06 3.83E+06 3.64E+06 5.77E-15 81-3391 S-23 159.11 1.OIE+06 2.97E+06 3.43E+06 5.44E-15 81-3393 P-29 159.11 1.03E+06 3.03E+06 3.50E+06 5.55E-15 81-3395 W-22 159.11 1.10E+06 3.24E+06 3.74E+06 5.92E-15 5.562-15 81-1399 Middle 158.11 8.62E+05 5.25E+07 5.01 E+07 7.17E-15 7.17E-15 81-1388 Middle 158.35 7.41E+05 5.34E+06 5.34E+06 3.51E-14 2.74E-14 81-1389 Middle 158.35 6.09E+06 4.39E+07 4.39E+07 2.80E-13 2.75E-13 81-1390 Top 159.11 3.17E+07 6.96E+07 6.60E+07 4.31 E-12 81-1393 Top-Mid 159.11 3.06E+07 6.72E+07 6.37E+07 4.16E-12 81-1396 Middle 159.11 3.03E+07 6.65E+07 6.31E+07 4.12E-12 81-1400 Bot-Mid 159.11 3.27E+07 7.18E+07 6.81E+07 4.44E-12 81-1403 Bottom 159.11 3.07E+07 6.74E+07 6.39E+07 4.17E-12 4.24E-12 81-1391 Top 159.11 1.21 E+07 2.66E+07 3.06E+07 2.00E-12 81-1394 Top-Mid 159.11 1.13E+07 2.48E+07 2.86E+07 1.87E-12 81-1397 Middle 159.11 1.16E+07 2.55E+07 2.94E+07 1.92E-12 81-1401 Bot-Mid 159.11 1.26E+07 2.77E+07 3.19E+07 2.08E-12 81-1404 Bottom 159.11 1.20E+07 2.63E+07 3.04E+07 1.98E-12 1.97E-12 Notes:

• The average U-238(nf) reaction rate of 2.74E-14 includes the correction of a factor of 0.820 to account for plutonium build-in and an additional factor of 0.955 to account for photo-fission effects in the sensor.

• The average Np-237(nf) reaction rate of 2.75E-13 includes the correction of a factor of 0.984 to account for the photo-fission effects in the sensor.

23 Table 13 Measured Sensor Specific Activities and Reaction Rates Capsule S Radially Radially Adjusted Adjusted Average Measured Saturated Saturated Reaction Reaction Radius Activity Activity Activity Rate Rate Sample ID Foil ID [cm.] [dpslg] [dpslg] [dpslg] [rps/atom] [rpslatom]

93-3163 Top 158.11 2.06E+05 3.06E+05 2.92E+05 4.45E-17 93-3166 Top-Mid 158.11 1.82E+05 2.70E+05 2.58E+05 3.93E-17 93-3172 Bot-Mid 158.11 1.98E+05 2.94E+05 2.81 E+05 4.28E-17 93-3175 Bottom 158.11 2.18E+05 3.24E+05 3.09E+05 4.71E-17 4.34E-17 93-4326 P-31 158.11 1.62E+06 2.93E+06 2.79E+06 4.42E-15 4.42E-15 93-3169 Middle 158.11 8.51 E+06 4.27E+07 4.06E+07 5.81 E-15 5.81E-15 93-3159 Middle 158.35 1.40E+06 4.63E+06 4.63E+06 3.04E-14 2.19E-14 93-3160 Middle 158.35 1.11E+07 3.67E+07 3.67E+07 2.34E-13 2.30E-13 93-3161 Top 159.11 3.55E+07 5.27E+07 5.05E+07 3.29E-12 93-3164 Top-Mid 159.11 3.71E+07 5.51E+07 5.28E+07 3.44E-12 93-3167 Middle 159.11 3.39E+07 5.03E+07 4.82E+07 3.15E-12 93-3170 Bot-Mid 159.11 3.60E+07 5.35E+07 5.12E+07 3.34E-12 93-3173 Bottom 159.11 3.45E+07 5.12E+07 4.91E+07 3.20E-12 3.29E-12 93-3162 Top 159.11 1.43E+07 2.12E+07 2.47E+07 1.61E-12 93-3165 Top-Mid 159.11 1.37E+07 2.03E+07 2.37E+07 1.54E-12 93-3168 Middle 159.11 1.31 E+07 1.95E+07 2.26E+07 1.48E-12 93-3171 Bot-Mid 159.11 1.45E+07 2.15E+07 2.50E+07 1.63E-12 1 93-3174 Bottom 159.11 1.35E+07 2.00E+07 2.33E+07 1.52E-12 1.56E-12 Notes:

• The average U-238(n,f) reaction rate of 2.19E-14 includes the correction of a factor of 0.755 to account for plutonium build-in and an additional factor of 0.953 to account for photo-fission effects in the sensor.

'The average Np-237(nf) reaction rate of 2.30E-13 includes the correction of a factor of 0.983 to account for the photo-fission effects in the sensor.

24 Table 14

~7~) Summary of Sensor Reaction Rates from Capsules V, R, T, and S Measured Reaction Rate [rms/nudeusl Sensor Reaction Capsule V Capsule R Capsule T Capsule S Cu-63r(n,a)Co-60 6.82e-17 6.49e-17 5.1 1e-17 4.34e-17 Fe-54(n,p)Mn-54 7.55e-15 7.84e-15 5.56e-15 4.42e-15 Ni-58(n,p)Co-58 8.90e-15 1.0le-14 7.17e-15 5.81e-15 U-238(n,f)Cs-137 Cd Covered 3.91e-14 4.1le-14 2.74e-14 2.19e-14 Np-237(n,f)Cs-137 Cd Covered Rejected 4.8le-13 2.75e-13 2.30e-13 Co-59(n,y,) Co-60 None 7.78e-12 4.24e-12 3.29e-12 Co-59(ny) Co-60 Cd Covered None 3.69e-12 1.97e-12 1.56e-12

)

i)

25 33.4 LEAST SQUARES EVALUATION OF SENSOR SETS

)

Least squares adjustment methods provide the capability of combining the measurement data with the neutron transport calculation resulting in a Best Estimate neutron energy spectrum with associated uncertainties. Best Estimates for key exposure parameters such as W(E> 1.0 MeV) or dpa/s along with their uncertainties are then easily obtained from the adjusted spectrum. In general, the least squares methods, as applied to surveillance capsule dosimetry evaluations, act to reconcile the measured sensor reaction rate data, dosimetry reaction cross-sections, and the calculated neutron energy spectrum within their respective uncertainties. For example, g

relates a set of measured reaction rates, R,, to a single neutron spectrum, ýg, through the multigroup dosimeter reaction cross-section, aig, each with an uncertainty S. The primary objective of the least squares evaluation is to produce unbiased estimates of the neutron exposure parameters at the location of the measurement.

5 For the least squares evaluation of the R. E. Ginna surveillance capsule dosimetry, The FERRET code [ ]

was employed to combine the results of the plant specific neutron transport calculations and sensor set reaction rate measurements to determine best estimate values of exposure parameters (<(E > 1.0 MeV) and 2 dpa) along with associated uncertainties for the three in-vessel capsules withdrawn to date.

The application of the least squares methodology requires the following input:

1 - The calculated neutron energy spectrum and associated uncertainties at the measurement location.

2 - The measured reaction rates and associated uncertainty for each sensor contained in the multiple foil set.

3 - The energy dependent dosimetry reaction cross-sections and associated uncertainties for each sensor contained in the multiple foil sensor set.

For the R. E. Ginna application, the calculated neutron spectrum was obtained from the results of plant specific neutron transport calculations described in Section 3.2 of this report. The sensor reaction rates were derived from the measured specific activities using the procedures described in Section 3.3. The dosimetry reaction cross-sections and uncertainties were obtained from the SNLRML dosimetry cross-section libraryl' 1 . The SNLRML library is an evaluated dosimetry reaction cross-section compilation recommended for use in LWR evaluations by ASTM Standard El018, "Application of ASTM Evaluated Cross-Section Data File, Matrix E 706 (liB)".

26 The uncertainties associated with the measured reaction rates, dosimetry cross-sections, and calculated neutron spectrum were input to the least squares procedure in the form of variances and covariances. The assignment of the input uncertainties followed the guidance provided in ASTM Standard E 944, "Application of Neutron Spectrum Adjustment Methods in Reactor Surveillance.

The following provides a summary of the uncertainties associated with the least squares evaluation of the R. E. Ginni'surveillance capsule sensor sets:

Reaction Rate Uncertainties The overall uncertainty associated with the measured reaction rates includes components due to the basic measurement process, the irradiation history corrections, and the corrections for competing reactions. A high level of accuracy in the reaction rate determinations is assured by utilizing laboratory procedures that conform to the ASTM National Consensus Standards for reaction rate determinations for each sensor type.

After combining all of these uncertainty components, the sensor reaction rates derived from the counting and data evaluation procedures were assigned the following net uncertainties for input to the least squares evaluation:

Reaction Uncertainty Cu 63(n, )Co60 5%

Fe"(n,p)Mne 5%

Ni58(n,p)Co58 5%

U238(nf)Cs137 10%

Np237(nf)Cs13 10%

Co59(n,Y)Co60 5%

Dosimetry Cross-Section Uncertainties The reaction rate cross-sections used in the least squares evaluations were taken from the SNLRML library. This data library provides reaction cross-sections and associated uncertainties, including covariances, for 66 dosimetry sensors in common use. Both cross-sections and uncertainties are provided in a fine multigroup structure for use in least squares adjustment applications. These cross-sections were compiled from the most recent cross-section evaluations and they have been tested with respect to their accuracy and consistency for least squares evaluations. Further, the library has been empirically tested for use in fission spectra determination as well as in the fluence and energy characterization of 14 MeV neutron sources. Detailed discussions of the contents of the SNLRML library along with the evaluation process for each of the sensors is provided in Reference 16.

For sensors included in the R. E. Ginna surveillance capsules, the following uncertainties in the fission spectrum averaged cross-sections are provided in the SNLRML documentation package.

/

27 Reaction Uncertainty Cu 63(n,a)Co° 4.08-4.16%

FeM(n,p)Mnl 3.05-3.11%

Ni51(np)Co58 4.49-4.56%

Ui 8(n,f)Cs137 0.54-0.64%

Np 37(nf)CsI 37 10.32-10.97%

Co 59(n,y)Co6° 0.79-3.59%

These tabulated ranges provide an indication of the dosimetry cross-section uncertainties associated with the sensor sets used in LWR irradiations.

Calculated Neutron Spectrum The neutron spectrum input to the least squares adjustment procedure was obtained directly from the results of plant specific transport calculations for each surveillance capsule location. The spectrum at each location was input in an absolute sense (rather than as simply a relative spectral shape). Therefore, within the constraints of the assigned uncertainties, the calculated data were treated equally with the measurements.

While the uncertainties associated with the reaction rates were obtained from the measurement procedures and counting benchmarks and the dosimetry cross-section uncertainties were supplied directly with the SNLRML library, the uncertainty matrix for the calculated spectrum was constructed from the following relationship:

Mg.,g =Rý +Rg*Rg.

• P, 8 where R. specifies an overall fractional normalization uncertainty and the fractional uncertainties R., and RPspecify additional random groupwise uncertainties that are correlated with a correlation matrix given by:

P919 = [1 -1"1/45g' +,9*e-H where 2

H (g _ g,)

2 2r The first term in the correlation matrix equation specifies purely random uncertainties, while the second term describes the short range correlations over a group range y (0 specifies the strength of the latter term).

The value of 8 is 1.0 when g = g' and 0.0 otherwise.

28

-x The set of parameters defining the input covariance matrix for the R. E. Ginna calculated spectra was as

• 9 follows:

Flux Normalization Uncertainty (RJ) 15%

Flux Group Uncertainties (Rg, Re.)

(E > 0.0055 MeV) 15%

(0.68 eV < E < 0.0055 MeV) 29%

(E < 0.68 eV) 52%

Short Range Correlation (0)

(E > 0.0055 MeV) 0.9 (0.68 eV < E < 0.0055 MeV) 0.5 (E < 0.68 eV) 0.5 Flux Group Correlation Range (y)

(E > 0.0055 MeV) 6 (0.68 eV < E < 0.0055 MeV) 3 (E < 0.68 eV) 2 Results of the least squares evaluation of the four sensor sets withdrawn from the R. E. Ginna reactor are provided in Tables 15 and 16. In Table 15, measured, calculated, and best estimate sensor reaction rates are given for Capsules V, R, T, and S. The improvement in the fit of the adjusted spectra to the measurements is evident for all four capsule data sets. Prior to the application of the adjustment procedure M/C ratios for individual foil reactions ranged from 0.75 to 1.32, while after the adjustment M/BE ratios ranged from 0.91 to 1.12. Thus, demonstrating a significant improvement in the data fits.

In Table 16, the calculated and best estimate exposure rates and integrated exposures of Capsules V, R, T, and S are given. Data are provided in terms of both fluence (E > 1.0 MeV) and iron atom displacements.

)

29 Table 15 I Comparison of Measured, Calculated, and Best Estimate Reaction Rates at the Surveillance Capsule Center Surveillance Capsule V Reaction Rate frps/aom]

Best Reaction Measured Calculated Estimate M/C MIBE Cu'3(n,c)Co60 6.82e-17 6.69e-17 6.54e-17 1.02 1.04 FeW4(n,p)Mn5 4 7.55e-15 8.14e-15 7.34e-15 0.93 1.03 Ni58(n,p)Co58 8.90e-15 1.14e-14 9.74e-15 0.78 0.91 23 U 8(n,f)Cs' 37 Cd 3.91e-14 4.37e-14 3.77e-14 0.89 1.04 Np237 (n,f)CsI 37 Cd Co 59(n,y)Co 60 CoS9( n,)Co6" Cd Surveillance Capsule R Reaction Rate Irps/atomi Best Reaction Measured Calculated Estimate M/C M/BE Cu63(na)Co6 6.49e-17 6.41e-17 6.39e-17 1.01 1.02 Fe54 (n,p)Mn54 7.84e-15 7.79e-15 7.74e-15 1.01 1.01 NiSS(n,p)Co6s 1.0le-14 1.09e-14 1.06e-14 0.93 0.95 u238 (n,f)Cs137 Cd 4.11e-14 4.18e-14 4.22e-14 0.98 0.97 Npý1 (n,)Cs1 37 Cd 4.81e-13 3.64e-13 4.28e-13 1.32 1.12 Co 59(ny)Co60 7.78e-12 9.22e-12 7.84e-12 0.84 0.99 Co59(n,y)Co6° Cd 3.69e-12 3.66e-12 3.68e-12 1.01 1.00 Surveillance Capsule T Reaction Rate [rps/atom]

Best Reaction Measured Calculated Estimate M/C M/BE Cu 63(n,cz)Co 60 5.11e-17 5.09e-17 5.05e-17 1.00 1.01 Fe54(n,p)Mns4 5.56e-15 5.5le-15 5.51e-15 1.01 1.01 Ni5S(n,p)Co58 7.17e-15 7.58e-15 7.46e-15 0.95 0.96 U3 8 (n f)Cs137 Cd 2.74e-14 2.70e-14 2.76e-14 1.01 0.99 Np 23 7(n,fCs137 Cd 2.75e-13 2.14e-13 2.49e-13 1.29 1.10 Co5 9(n,y)Co 60 4.24e-12 5.Ole-12 4.27e-12 0.85 0.99 Co 59(nny)Co6" Cd 1.97e-12 1.89e-12 1.96e-12 1.04 1.01 i

30 Table 15 (continued)

Comparison of Measured, Calculated, and Best Estimate Reaction Rates at the Surveillance Capsule Center Surveillance Capsule S Reaction Rate [rps/aom]

Best Reaction Measured Calculated Estimate M/C M/BE Cu63(n, a)Co60 4.34e-17 4.25e-17 4.21e-17 1.02 1.03 Fe5 4(n,p)M n5 4 4.42e-15 4.66e-15 4.45e-15 0.95 0.99 Ni5s(n,p)Coss 5.8le-15 6.43e-15 6.05e-15 0.90 0.96 U0 8 (nm)Cs137 Cd 2.19e-14 2.33e-14 2.24e-14 0.94 0.98 NpM7 (n,f)Cs 137 Cd 2.30e-13 1.88e-13 2.06e-13 1.22 1.12 Co59(nY)Co60 3.29e-12 4.36e-12 3.32e-12 0.75 0.99 Co59(n,-y)Co6o Cd 1.56e-12 1.70e-12 1.56e-12 0.92 1.00

)

31 Table 16 Comparison of Calculated and Best Estimate Exposure Parameters at the Surveillance Capsule Center Time Averaged Exposure Rates

• b(E > 1.0 MeV) [n/cm2-s!

Best Calculated Estimate Uncertainty BE/C Capsule V 1.32e+11 1.13e+11 7% 0.86 Capsule R l.26e+11 1.30e+11 6% 1.03 Capsule T 7.79e+10 8.1le+10 6% 1.04 Capsule S 6.78e+10 6.62e+10 6% 0.98 Iron Atom Displacements Idpa/s]

Best Calculated Estimate Uncertainty BE/C Capsule V 2.40e-10 2.05e-10 9% 0.86 Capsule R 2.30e-10 2.37e-10 7% 1.03 Capsule T 1.36e-10 1.40e-10 7% 1.03 Capsule S 1.19e- 10 1.16e- 10 7% 0.97 Integrated Capsule Exposure 0 (E> 1.0 MeV) [n/cm?

Best Calculated Estimate Uncertainty BE/C Capsule V 5.87e+18 5.03e+18 7% 0.86 Capsule R 1.02e+19 1.05e+19 6% 1.03 Capsule T 1.69e+19 1.76c+19 6% 1.04 Capsule S 3.64e+19 3.55e+19 6% 0.98 Iron Atom Displacements [dpal Best Calculated Estimate Uncertainty BE/C Capsule V 1.07e-02 9.15e-03 9% 0.86 Capsule R 1.85e-02 1.9le-02 7% 1.03 Capsule T 2.94e-02 3.04e-02 7% 1.03 Capsule S 6.38"-02 6.22e-02 7% 0.97

32

.1) 3.5 COMPARISON OF MEASUREMENTS AND CALCULATIONS In this section, comparisons of the measurement results from the four surveillance capsules withdrawn to date with the corresponding analytical predictions at the measurement locations are provided. These comparisons are given on two levels. In the first instance, calculations of individual sensor reaction rates are compared directly with the corresponding values obtained from the measured specific activities. In the second case calculations of fast neutron exposure rates in terms of ý(E > 1.0 MeV) and dpa/s are compared with the best estimate results obtained from the least squares evaluation of the three capsule dosimetry results. These two levels of comparison yield consistent and similar results with all measurement to calculation comparisons falling within the 20% limits specified as the acceptance criteria in Regulatory Guide 1.190.

In the case of the direct comparison of measured and calculated sensor reaction rates, the M/C comparisons for fast neutron reactions range from 0.78-1.32 for the 19 samples included in the data set. In the comparisons of best estimate and calculated fast neutron exposure parameters, the corresponding BE/C comparisons range from 0.85 -1.04 for the four surveillance capsules withdrawn to date.

Based on these comparisons, it is concluded that the data comparisons validate the use of the calculated fast neutron exposures provided in Section 3.2 of this report for use in the assessment of the condition of the materials comprising the beltline region of the R. E. Ginna reactor pressure vessel.

)

33 Table 17 Comparison of Measured and Calculated Neutron Sensor Reaction Rates For In-Vessel Surveillance Capsules V, R, T, and S M/C Ratio Capsule Cu-63(na) Fe-54(n,p) Ni-58(n,p) U-238(n,o Np-237(n,f Average  % std dev V 1.02 0.93 0.78 0.89 0.91 10.9 R 1.01 1.01 0.93 0.98 1.32 1.05 14.8 T 1.00 1.01 0.95 1.01 1.29 1.05 12.7 S 1.02 0.95 0.90 0.94 1.22 1.01 12.7 Average 1.01 0.97 0.89 0.96 1.28 1.01 13.3

% std dev 0.8 4.2 8.4 5.5 3.9 Note: The average and % std dev values in bold face type represent the average and standard deviation of the entire 19 sample threshold foil data set.

Table 18 Comparison of Best Estimate and Calculated Fast Neutron Exposure Rates 9 For In-Vessel Surveillance Capsules V, R, T, and S BE/C Ratio Capsule Neutron Fluence Iron Atom Displacements V 0.86 0.86 R 1.03 1.03 T 1.04 1.03 S 0.98 0.97 Average 0.98 0.97

% std dev 8.7 8.6

34 4 CRITERIA FOR ALLOWABLE PRESSURE-TEMPERATURE RELATIONSHIPS 4.1 OVERALL APPROACH The ASMEqpproach for calculating the allowable limit curves for various heatup and cooldown rates specifies that the total stress intensity factor, Ki, for the combined thermal and pressure stresses at any time during heatup or cooldown cannot be greater than the reference stress intensity factor, KI, for the metal temperature at that time. K10 is obtained from the reference fracture toughness curve, defined in Code Case N-640, "Alternative Reference Fracture Toughness for Development of PT Limit Curves for Section XI"' 3 " 4] of the ASMIE Appendix G to Section XI. The K10 curve is given by the following equation:

Ki, = 3 3 .2 +20. 7 3 4 *e[°°2(T-RT*T)] (1) where, K,, = reference stress intensity factor as a function of the metal temperature T and the metal reference nil-ductility temperature RTNDT This Ki, curve is based on the lower bound of static critical K, values measured as a function of temperature on specimens of SA-533 Grade B Classl, SA-508-1, SA-508-2, SA-508-3 steel.

4.2 METHODOLOGY FOR PRESSURE-TEMPERATURE LIMIT CURVE DEVELOPMENT The governing equation for the heatup-cooldown analysis is defined in Appendix G of the ASME Code as follows:

C* Ki. + Kit < Ki (2) where, K = stress intensity factor caused by membrane (pressure) stress Kit = stress intensity factor caused by the thermal gradients KIc = function of temperature relative to the RTNDT of the material C = 2.0 for Level A and Level B service limits C = 1.5 for hydrostatic and leak test conditions during which the reactor core is not critical

)

35 For membrane tension, the corresponding K1 for the postulated defect is:

Kim= M.x (pRit) (3) where, Mm for an inside surface flaw is given by:

Mm = 1.85 for t < 2, Mm = 0.9264I- for 2*%Ft <3.464,5 Mm = 3.21 for .ft > 3.464 Similarly, Mm for an outside surface flaw is given by:

Mm = 1.77 for rt < 2, Mm = 0.893 ft- for 2 fl- < 3.464, Mm = 3.09 for t- > 3.464 and p = internal pressure, Ri = vessel inner radius, and t = vessel wall thickness.

For bending stress, the corresponding K, for the postulated defect is:

• Krb = Mb

• Maximum Stress, where Mb is two-thirds of Mm The maximum K, produced by radial thermal gradient for the postulated inside surface defect of G-2120 is Ku = 0.953x10-3 x CR x t 2-, where CR is the cooldown rate in 'F/hr., or for a postulated outside surface defect, Kt = 0.753x10"3 x HU x t", where HU is the heatup rate in OF/hr.

The through-wall temperature difference associated with the maximum thermal K, can be determined from Fig. G-2214-1. The temperature at any radial distance from the vessel surface can be determined from Fig.

G-2214-2 for the maximum thermal K1 .

(a) The maximum thermal K, relationship and the temperature relationship in Fig. G-2214-1 are applicable only for the conditions given in G-2214.3(a)(1) and (2).

(b) Alternatively, the K, for radial thermal gradient can be calculated for any thermal stress distribution and at any specified time during cooldown for a 1/4-thickness inside surface defect using the relationship:

Kit = (1.0359Co + 0.6322C, + 0.4753C2 + 0.3855C 3) * (4)

)

36 or similarly, Krr during heatup for a 1/4-thickness outside surface defect using the relationship:

Kit = (1.043Co + 0.630Ci + 0.481C2 + 0.401C3) * (5) where the coefficients Co, C1, C2 and C3 are determined from the thermal stress distribution at any specified time during the heatup or cooldown using the form:

a'(x) = Co+ Ci(x / a) +C2(x/ a)2 + C3(x/ a)3 (6) and x is a variable that represents the radial distance from the appropriate (i.e., inside or outside) surface to any point on the crack front and a is the maximum crack depth.

Note, that equations 3, 4 and 5 were implemented in the OPERLIM computer code, which is the program used to generate the pressure-temperature (P-T) limit curves. No other changes were made to the OPERLIM computer code with regard to P-T calculation methodology. Therefore, the P-T curve methodology is unchanged from that described in WCAP-14040, "Methodology used to Develop Cold Overpressure Mitigating System Setpoints and RCS Heatup and Cooldown Limit Curves"'21 Section 2.6 (equations 2.6.2-4 and 2.6.3-1) with the exceptions just described above.

At any time during the heatup or cooldown transient, Ki, is determined by the metal temperature at the tip of a postulated flaw at the I/4T and 3/4T location, the appropriate value for RTNDT, and the reference

- fracture toughness curve. The thermal stresses resulting from the temperature gradients through the vessel wall are calculated and then the corresponding (thermal) stress intensity factors, Kit, for the reference flaw are computed. From Equation 2, the pressure stress intensity factors are obtained and, from these, the allowable pressures are calculated.

For the calculation of the allowable pressure versus coolant temperature during cooldown, the reference flaw of Appendix G to the ASME Code is assumed to exist at the inside of the vessel wall. During cooldown, the controlling location of the flaw is always at the inside of the wall because the thermal gradients produce tensile stresses at the inside, which increase with increasing cooldown rates. Allowable pressure-temperature relations are generated for both steady-state and finite cooldown rate situations. From these relations, composite limit curves are constructed for each cooldown rate of interest.

The use of the composite curve in the cooldown analysis is necessary because control of the cooldown procedure is based on the measurement of reactor coolant temperature, whereas the limiting pressure is actually dependent on the material temperature at the tip of the assumed flaw. During cooldown, the I/4T vessel location is at a higher temperature than the fluid adjacent to the vessel inner diameter. This condition, of course, is not true for the steady-state situation. It follows that, at any given reactor coolant temperature, the AT (temperature) developed during cooldown results in a higher value of KI( at the 1/4T location for finite cooldown rates than for steady-state operation. Furthermore, if conditions exist so that the increase in K10 exceeds Kit, the calculated allowable pressure during cooldown will be greater than the steady-state value.

1

37 The above procedures are needed because there is no direct control on temperature at the 1/4T location and, therefore, allowable pressures may unknowingly be violated if the rate of cooling is decreased at various intervals along a cooldown ramp. The use of the composite curve eliminates this problem and ensures conservative operation of the system for the entire cooldown period.

Three separate calculations are required to determine the limit curves for finite heatup rates. As is done in the cooldown analysis, allowable pressure-temperature relationships are developed for steady-state conditions as well as finite heatup rate conditions assuming the presence of a l/4T defect at the inside of the wall. The heatup results in compressive stresses at the inside surface that alleviate the tensile stresses produced by internal pressure. The metal temperature at the crack tip lags the coolant temperature; therefore, the K1 , for the l/4T crack during heatup is lower than the Ki, for the 1/4T crack during steady-state conditions at the same coolant temperature. During heatup, especially at the end of the transient, conditions may exist so that the effects of compressive thermal stresses and lower K1 cvalues do not offset each other, and the pressure-temperature curve based on steady-state conditions no longer represents a lower bound of all similar curves for finite heatup rates when the l/4T flaw is considered. Therefore, both cases have to be analyzed in order to ensure that at any coolant temperature the lower value of the allowable pressure calculated for steady-state and finite heatup rates is obtained.

The second portion of the heatup analysis concerns the calculation of the pressure-temperature limitations for the case in which a 1/4T flaw located at the l/4T location from the outside surface is assumed. Unlike the situation at the vessel inside surface, the thermal gradients established at the outside surface during heatup produce stresses which are tensile in nature and therefore tend to reinforce any pressure stresses present. These thermal stresses are dependent on both the rate of heatup and the time (or coolant

• ) temperature) along the heatup ramp. Since the thermal stresses at the outside are tensile and increase with increasing heatup rates, each heatup rate must be analyzed on an individual basis.

Following the generation of pressure-temperature curves for both the steady-state and finite heatup rate situations, the final limit curves are produced by constructing a composite curve based on a point-by-point comparison of the steady-state and finite heatup rate data. At any given temperature, the allowable pressure is taken to be the lesser of the three values taken from the curves under consideration. The use of the composite curve is necessary to set conservative heatup limitations because it is possible for conditions to exist wherein, over the course of the heatup ramp, the controlling condition switches from the inside to the outside, and the pressure limit must at all times be based on analysis of the most critical criterion.

4.3 CLOSURE HEAD/VESSEL FLANGE REQUIREMENTS 10 CFR Part 50, Appendix Gfs1 addresses the metal temperature of the closure head flange and vessel flange regions. This rule states that the metal temperature of the closure flange regions must exceed the material unirradiated RTNDT by at least 120'F for normal operation when the pressure exceeds 20 percent of the preservice hydrostatic test pressure (3106 psi), which is 621 psig for R.E. Ginna. The limiting unirradiated RTNDT of -52*F occurs in the vessel flange of the R.E. Ginna reactor vessel, so the minimum allowable temperature of this region is 68°F at pressures greater than 621 psig. This limit is shown in Figures 6-1 and 6-2 wherever applicable.

)

38 5 CALCULATION OF ADJUSTED REFERENCE TEMPERATURE From Regulatory Guide 1.99, Revision 2, the adjusted reference temperature (ART) for each material in the beltline region is given by the following expression:

ART = Initial RTNDT + ARTNDT + Margin (7)

Initial RTNDT is the reference temperature for the unirradiated material as defined in paragraph NB-2331 of Section IEl of the ASME Boiler and Pressure Vessel Code[' 71. If measured values of initial RTNDT for the material in question are not available, generic mean values for that class of material may be used if there are sufficient test results to establish a mean and standard deviation for the class.

ARTNDT is the mean value of the adjustment in reference temperature caused by irradiation and should be calculated as follows:

ARTNDT = CF

• f(o. 28 .-o, loogf) (8)

To calculate ARTNDT at any depth (e.g., at 1/4Tor 3/4T), the following formula must first be used to attenuate the fluence at the specific depth.

dpthx) = fff

• e (..24x) (9) where x inches (vessel beltline thickness is 6.5 inches) is the depth into the vessel wall measured from the

- vessel clad/base metal interface. The resultant fluence is then placed in Equation 8 to calculate the ARTNDT at the specific depth.

The Westinghouse Radiation Engineering and Analysis Group evaluated the vessel fluence projections in Section 3 of this report. The evaluation used the ENDF/B-VI scattering cross-section data set. This is consistent with methods presented in WCAP-14040-NP-A, "Methodology Used to Develop Cold Overpressure Mitigating System Setpoints and RCS Heatup and Cooldown Limit Curves". Tables 6 and 7 contain the calculated vessel surface fluences values at various azimuthal locations, including the longitudinal location for the nozzle shell forging to intermediate shell forging girth weld. Tables 19 and 20 contain the 1/4T and 3/4T calculated fluences and fluence factors, per the Regulatory Guide 1.99, Revision 2, used to calculate the ART values for all beltline materials in the R.E. Ginna reactor vessel.

9

39 TABLE 19

') Summary of the Vessel Surface, 1/4T and 3/4T Fluence Values used for the Generation of the 52 EFPY Heatup/Cooldown Curves Material Surface 1/4T 3/4T 2

(n/cm ,E > 1.0 MeV) 2 (n/cm ,E > 1.0 MeV) (n/cm 2 ,E > 1.0 MeV)

Inter. & Lower Shell Forgings and the Inter, to Lower Shell 4.85 x 10'9 3.28 x 10'9 1.51 x 1019 Girth Weld (Peak Fluence)

Nozzle Shell Forging and the 1.92 x 1018 1.30 x 10"8 5.96 x 1017 Nozzle to Inter. Shell Girth Weld TABLE 20 Summary of the Calculated Fluence Factors used for the Generation of the 52 EFPY Heatup and Cooldown Curves Material 1/4T Fluence 1/4T FF 3/4T Fluence 3/4T FF 2 2

) (n/cm ,E > 1.0 MeV) (n/cm ,E > 1.0 MeV)

Inter. & Lower Shell Forgings 3.28 x 1019 1.31 1.51 x 10'9 1.11 and the Inter. to Lower Shell Girth Weld Nozzle Shell Forging and the 1.30 x 1018 0.47 5.96 x 1017 0.32 Nozzle to Inter. Shell Girth Weld Margin is calculated as, M = 2 Ja? I + eA . The standard deviation for the initial RTNDT margin term, is a, 00 F when the initial RTNroT is a measured value, and 17°F when a generic value is available. The standard deviation for the ARTNDT margin term, cya, is 17F for plates or forgings, and 8.5°F for plates or forgings when surveillance data is used. For welds, aA is equal to 28'F when surveillance capsule data is not used, and is 141F (half the value) when credible surveillance capsule data is used. aA need not exceed 0.5 times the mean value of ARTNT.

40 Contained in Tables 21 and 22 are the calculations of the 52 EFPYART values used for generation of the

) heatup and cooldown curves.

TABLE 21 Calculation of the ART Values for the 1/4T Location @ 52 EFPY Material RG 1.99 R2 CF FF IRTNDT(a) ARTNDT Margin ART<)

Method (OF) (OF) (OF) (OF) (OF)

Nozzle Shell Forging 123PI18(a) IPosition 1.1 44.0 0.47 30 20.7 20.7 71 Inter. Shell Forging 125S255 Position 1.1 44.0 1.31 20 57.6 34 112 Position 2.1 16.6 1.31 20 21.7 34(c) 76 Lower Shell Forging 125P666 Position 1.1 31.0 1.31 40 40.6 34 115 Position 2.1 28.3 1.31 40 37.1 17 94 Nozzle to Intermediate Shell Position 1.1 167.6 0.47 10 78.8 56 145 Girth Weld (Heat # 71249) Position 2.1 180.8 0.47 10 85.0 56(c) 151 Intermediate Shell to Lower Shell Position 1.1 170.4 1.31 -4.8 223.2 56 274

) Girth Weld (Heat # 61782) Position 2.1 161.9 1.31 -4.8 212.1 4 8 .3 (d) 256 Notes:

(a) Initial RTTDT values are measured values.

(b) ART = Initial RTNDT + ARTNDT + Margin (OF)

(c) Surveillance Data is not credible, thus the full oa is used in calculating the margin term.

(d) Based on Additional tests by B&W and documented in the Ginna PTLR.. .Used per the request of RGE.

41 TABLE 22

) Calculation of the ART Values for the 3/4T Location @ 52 EFPY Material RG 1.99 R2 CF FF IRTNa) ARTNDT Margin ART(b)

Method (OF) (OF) (OF) (OF) (OF)

Nozzle Shell forging 123P118(a) Position 1.1 44.0 0.32 30 14.1 14.1 58 Inter. Shell Forging 125S255 Position 1.1 44.0 1.11 20 48.8 34 103 Position 2.1 16.6 1.11 20 18.4 34(c) 72 Lower Shell Forging 125P666 Position 1.1 31.0 1.11 40 34.4 34 108 Position 2.1 28.3 1.11 40 31.4 17 89 Nozzle to Intermediate Shell Position 1.1 167.6 0.32 10 53.6 53.6 117 Girth Weld (Heat # 71249) Position 2.1 180.8 0.32 10 57.9 56(c) 124 Intermediate Shell to Lower Shell Position 1.1 170.4 1.11 -4.8 189.1 56 240 Girth Weld (Heat # 61782) Position 2.1 161.9 1.11 -4.8 179.7 4 8 .3(d) 223 Notes:

(a) Initial RTNT values are measured values.

(b) ART = Initial RTNDT + ARTNDT + Margin (OF)

(c) Surveillance Data is not credible, thus the full 0 a is used in calculating the margin term.

(d) Based on Additional tests by B&W and documented in the Ginna PTLR... Used per the request of RGE.

)

42 The upper to intermediate shell forging to lower shell forging girth weld has the highest overall ART.

However, since Code Case N-641 allows for less restrictive methodology to be used when the highest ART comes from a girth weld, then the highest non-girth weld ART must be identified. The intermediate shell forging has the highest non-girth weld ART.

Contained in Table 23 is a summary of the limiting ARTs to be used in the generation of the R.E. Ginna reactor vessl heatup and cooldown curves. It has been determined based on the relief allowed via ASME Code Case N-641 (Circ. Flaw Methodology) and the magnitude of the girth weld ART values as compared to the intermediate shell forging ART values, that the intermediate shell forging ART values would produce a more conservative pressure temperature curve at the lower temperature portion of the curves. Thus, composite curves were created from two computer runs: 1) The highest "circ. flaw" ART with the "Circ.

Flaw Methodology (ASME Code Case N-64 1), and 2) The highest "Axial Flaw" ART with the Axial Flaw (1996 ASME Code) Methodology. The limiting pressures and temperature were taken from both sets of curves to create the most limiting PT curves overall. These limiting curves will be presented in Section 6.

TABLE 23 Summary of the Limiting ART Values Used in the Generation of the R.E. Ginna Heatup/Cooldown Curves 1/4 T Limiting ART 3/4 T Limiting ART Intermediate to Lower Shell Girth Weld (Limiting Circ. FlawMaterial) 256 1 223 Intermediate Shell Forging (LimitingAxialFlaw Material) 112 103

).

43 6 HEATUP AND COOLDOWN PRESSURE-TEMPERATURE LIMIT CURVES Pressure-temperature limit curves for normal heatup and cooldown of the primary reactor coolant system have been calculated for the pressure and temperature in the reactor vessel beltline region using the methods discussed in Sections 4.0 and 5.0 of this report. This approved methodology is also presented in WCAP-14040-NP-A, Revision 2 with exception to those items discussed in Section 1 of this report.

Figure 3 presents the limiting heatup curves without margins for possible instrumentation errors using heatup rates of 60 and 100'F/hr applicable for the first 52 EFPY. These curves were generated using a combination of the 1996 ASME Code Section XI, Appendix G with the limiting axial flaw ARTs and the ASME Code Case N-641 with the limiting circ. flaw ARTs. The limiting pressures and temperatures between the two cases produce a bounding heatup curves for the given rates. Figure 4 presents the limiting cooldown curves without margins for possible instrumentation errors using cooldown rates of 0, 20, 40, 60 and 100*F/hr applicable for 52 EFPY. Again, these curves were generated using a combination ofthel996 ASME Code Section XI, Appendix G with the limiting axial flaw ARTs and the ASME Code Case N-641 with the limiting circ. flaw ARTs. Allowable combinations of temperature and pressure for specific temperature change rates are below and to the right of the limit line shown in Figures 3 and 4. This is in addition to other criteria, which must be met before the reactor is made critical, as discussed below in the following paragraphs.

The reactor must not be made critical until pressure-temperature combinations are to the right of the criticality limit line shown in Figure 3. The straight-line portion of the criticality limit is at the minimum permissible temperature for the 2485 psig inservice hydrostatic test as required by Appendix G to 10 CFR Part 50. The governing equation for the hydrostatic test is defined in Code Case N-641131(approved in February 1999) as follows:

1.5 K, < K1,

where, Ki is the stress intensity factor covered by membrane (pressure) stress, Ki, = 33.2 + 20.734 e[r02 CrO-RTr)],

T is the minimum permissible metal temperature, and RTNDT is the metal reference nil-ductility temperature.

The criticality limit curve specifies pressure-temperature limits for core operation to provide. additional margin during actual power production as specified in Reference 5. The pressure-temperature limits for core operation (except for low power physics tests) are that the reactor vessel must be at a temperature equal to or higher than the minimum temperature required for the inservice hydrostatic test, and at least 40°F higher than the minimum permissible temperature in the corresponding pressure-temperature curve for heatup and cooldown calculated as described in Section 4.0 of this report. For the heatup and cooldown curves without margins for instrumentation errors, the minimum temperatures for the in service hydrostatic leak tests for the R.E. Ginna reactor vessel at 52 EFPY is 162'F, respectively. The vertical line drawn from

44 these points on the pressure-temperature curve, intersecting a curve 407F higher than the pressure-temperature limit curve constitutes the limit for core operation for the reactor vessel.

Figures 3 and 4 define all of the above limits for ensuring prevention of nonductile failure for the R.E.

Ginna reactor vessel for 52 EFPY. The data points used for the heatup and cooldown pressure-temperature limit curves shown in Figures 3 and 4 are presented in Tables 24 and 25.

)

)

45 MATERIAL PROPERTY BASIS LIMITING MATERIAL: INTER- to LOWER SHELL FORGING GIRTH WELD and INTER- SHELL FORGING LIMITING ART VALUES AT 52 EFPY: 1/4T, 256-F (Circ FlawART), 112°F (Axial FlawART) 3/4T, 223°F (Circ FlawART), 103°F (Axial FlawART) 2500 2250 2000 1750 0 1500 0

U)

(n a 1250 10

= 1000 U

'U 750 500 250 I

0 0 50 100 150 200 250 300 350 400 450 500 550 Moderator Temperature (Deg. F)

Figure 3 R.E. Ginna Reactor Coolant System Heatup Limitations (Heatup Rates of 60 &

100 0F/hr) Applicable for the First 52 EFPY (Without Margins for Instrumentation Errors) Using 1996 App.G Methodology

)

46 MATERIAL PROPERTY BASIS LIMITING MATERIAL: INTER. to LOWER SHELL FORGING GIRTH WELD and INTER- SHELL FORGING LIMITING ART VALUES AT 52 EFPY: 1/4T, 256-F (Circ FlawART), 112°F (Axial FlawART) 3/4T, 223°F (Circ Flaw AR7), 1031F (Axial FlawART) 2500 2250 2000 1750 1500 1250

1000 750 500 250 0

0 50 100 150 200 250 300 350 400 450 500 550 Moderator Temperature (Deg. F)

Figure 4 R.E. Ginna Reactor Coolant System Cooldown Limitations (Cooldown Rates up to 100 0F/hr) Applicable for the First 52 EFPY (Without Margins for Instrumentation Errors) Using 1996 App.G Methodology 1

47 TABLE 24

)/ 52 EFPY Heatup Curve Data Points Using 1996 App. G & ASME Code Case N-641 (without Uncertainties for Instrumentation Errors) 60 Heatup 60 Critical Limit* 100 Heatup 100 Critical Limit Leak Test Limit T (01) P (psig) T (OF) P (psig) I T (OF) P (psig)I T (OF) P (psig) T (OF) P (psig) 60 0 162 0 60 0 162 0 112 1500 60 621 162 621 60 621 162 621 143 2000 65 621 162 621 65 621 162 621 162 2485 68 621 162 861 68 621 162 844 68 861 162 861 68 844 162 844 70 861 162 865 70 844 162 844 75 865 162 875 75 844 162 844 80 875 162 889 80 844 162 848 85 889 162 908 85 848 162 856 90 908 162 930 90 856 162 868 95 930 162 955 95 868 162 883 100 955 162 984 100 883 162 902 105 984 162 1017 105 902 162 924 110 1017 162 1054 110 924 162 950 115 1054 162 1095 115 950 162 980 120 1095 165 1141 120 980 165 1014 125 1141 170 1192 125 1014 170 1052 130 1192 175 1248 130 1052 175 1094 135 1248 180 1310 135 1094 180 1142 140 1310 185 1378 140 1142 185 1195 145 1378 190 1445 145 1195 190 1254 150 1445 195 1460 150 1254 195 1319 155 1460 200 1476 155 1319 200 1349 160 1476 205 1495 160 1349 205 1362 165 1495 210 1515 165 1362 210 1377 170 1515 215 1537 170 1377 215 1394 175 1537 220 1561 175 1394 220 1412 180 1561 225 1589 180 1412 225 1433 185 1589 230 1619 185 1433 230 1456 190 1619 235 1652 190 1456 235 1481 195 1652 240 1689 195 1481 240 1510 200 1689 245 1730 200 1510 245 1541 205 1730 250 1775 205 1541 250 1576 210 1775 255 1824 210 1576 255 1614 215 1824 260 1879 215 1614 260 1657

• Data Points in Bold are the limiting pressures and temperature from the Circ. Flaw Run 9

48 TABLE 24 - (Continued)

) 52 EFPY Heatup Curve Data Points Using 1996 App. G& ASME Code Case N-641 (without Uncertainties for Instnumentation Errors) 60 Heatup 60 Critical Limit* 100 Heatup 100 Critical Limit Leak Test Limit T (OF) P (psig) T (OF) P (psig) T (OF) P (psig) T CF) P (psig) T (IF) P (psig) 220 1879 265 1929 220 1657 265 1704 225 1929 270 1980 225 1704 270 1756 230 1980 275 2036 230 1756 275 1814 235 2036 280 2099 235 .1814 280 1877 240 2099 285 2168 240 1877 285 1947 245 2168 290 2244 245 1947 290 2025 250 2244, 295 2328 250 2025 295 2111 255 2328 300 2421 255 2111 300 2205 260 2421 260 2205 305 2310 265 2310 310 2425 270 2425 Data Points in Bold are the limiting pressures and temperature from the Circ. Flaw Run i)

• )

49 TABLE 25 52 EFPY Cooldown Curve Data Points Using 1996 App. G &ASME Code Case N-641 (without Uncertainties for Instrumentation Errors)

Steady State f 20°F/hr. 40OF/hr. I 60*F/hr. I 100°F/hr.

T (OF) I P (psig) IT (OF) I P (psig) IT (OF) I P (psig) IT (OF) I P (psig) IT (OF) I P (psig) 60 0 60 0 60 0 60 0 60 0 60 621 60 621 60 621 60 621 60 621 65 621 65 621 65 621 65 621 65 621 68 621 68 621 68 621 68 621 68 621 68 870 68 851 68 831 68 812 68 776 70 877 70 858 70 839 70 821 70 785 75 897 75 879 75 861 75 843 75 810 80 918 80 901 80 885 80 869 80 838 85 942 85 926 85 911 85 897 85 869 90 969 90 954 90 941 90 928 90 904 95 998 95 985 95 973 95 962 95 942 100 1030 100 1019 100 1009 100 1000 100 985 105 1066 105 1057 105 1049 105 1042 105 1032 110 1105 110 1098 110 1093 110 1088 110 1084 115 1149 115 1144 115 1141 115 1140 115 1142 120 1197 120 1195 120 1195 120 1197 120 1193 125 1250 125 1250 125 1250 125 1250 125 1199 130 1309 130 1309 130 1309 130 1309 130 1207 135 1374 135 1374 135 1374 135 1341 135 1216 140 1446 140 1446 140 1412 140 1351 140 1226

) 145 1525 145 1482 145 1422 145 1361 145 1237 150 1552 150 1493 150 1433 150 1372 150 1250 155 1563 155 1505 155 1445 155 1385 155 1264 160 1576 160 1518 160 1459 160 1399 160 1279 165 1590 165 1532 165 1474 165 1415 165 1296 170 1605 170 1548 170 1490 170 1433 170 1316 175 1622 175 1566 175 1509 175 1452 175 1337 180 1641 180 1585 180 1530 180 1474 180 1361 185 1662 185 1607 185 1553 185 1498 185 1388 190 1685 190 1631 190 1578 190 1524 190 1418 195 1710 195 1658 195 1606 195 1554 195 1451 200 1738 200 1687 200 1637 200 1587 200 1488 205 1769 205 1720 205 1671 205 1623 205 1528 210 1803 210 1756 210 1709 210 1663 210 1574 215 1841 215 1796 215 1752 215 1708 215 1624 220 1883 220 1840 220 1798 220 1757 220 1679 225 1929 225 1889 225 1850 225 1812 225 1741 230 1980 230 1943 230 1907 230 1873 230 1810 235 2036 235 2003 235 1970 235 1940 235 1886 240 2099 240 2069 240 2040 240 2014 240 1970 245 2168 245 2142 245 2118 245 2097 245 2063 250 2244 250 2222 250 2204 250 2188 250 2166 255 2328 255 2312 255 2298 255 2289 255 2281 260 2421 260 2410 260 2403 260 2400 260 2400

• Data Points in Bold are the limiting pressures and temperature from the Circ. Flaw Run

50 7 REFERENCES

)

1. Regulatory Guide 1.99, Revision 2, "Radiation Embrittlement of Reactor Vessel Materials," U.S.

Nuclear Regulatory Commission, May 1988.

2. WCAP-14040-NP-A, Revision 2, "Methodology used to Develop Cold Overpressure Mitigating system Setpoints and RCS Heatup and Cooldown Limit Curves", J.D. Andrachek, et. al., January 1996.

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[Sub Reference 1: ASME Code Case N-640, "AlternativeReference FractureToughnessfor Development of P-T Limit Curvesfor Section X, Division 1 ", February26, 1999.]

[Sub Reference 2: ASME Boiler andPressure Vessel Code, Case N-588, "Attenuationto Reference Flaw OrientationofAppendix Gfor CircumferentialWelds in Reactor Vessels",Section X, Division 1, Approved December 12, 1997.]

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5. Code of Federal Regulations, 10 CFR Part 50, Appendix Q "Fracture Toughness Requirements,"

U.S. Nuclear Regulatory Commission, Washington, D.C., Federal Register, Volume 60, No. 243, dated December 19, 1995.

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Lippincott, April 1994.

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) Discrete Ordinates Neutron/Photon Transport Code System," August 1996.

51

14. RSIC Data Library Collection DLC-185, "BUGLE-96, Coupled 47 Neutron, 20 Gamma-Ray Group Cross-Section Library Derived from ENDF/B-VI for LWR Shielding and Pressure Vessel Dosimetry Applications," March 1996.

15. A. Schmittroth, "FERRET Data Analysis Code," HEDL-TME-79-40, Hanford Engineering Development Laboratory, Richland, WA, September 1979.

16. RSIC Data Library Collection DLC-178, "SNLRML Recommended Dosimetry Cross-Section Compendium," July 1994.

17. 1989 Section III, Division 1 of the ASME Boiler and Pressure Vessel Code, Paragraph NB-233 1, "Material for Vessels."

)