ML15322A091

Calculation No. 1001527.303, Revision 0, Feedwater, Water Level Instrument, and Core Dp Nozzle Fracture Mechanics Evaluation for Hatch, Unit 1 and Unit 2 Pressure-Temperature Limit Curve Development

ML15322A091
Person / Time
Site:	Hatch, 05000360
Issue date:	12/30/2011
From:	Griesbach T J, Qin M, Sommerville D V Structural Integrity Associates
To:	Office of Nuclear Reactor Regulation
Shared Package
ML15322A088	List: ML15322A089 ML15322A090 ML15322A091 ML15322A092 NL-15-2034, Calculation No. 1001527.303, Revision 0, Feedwater, Water Level Instrument, and Core Dp Nozzle Fracture Mechanics Evaluation for Hatch, Unit 1 and Unit 2 Pressure-Temperature Limit Curve Development NL-15-2034, Calculation No. 1001527.304, Revision 2, Hatch Unit 1, P-T Curve Calculation for 38 and 49.3 EFPY NL-15-2034, Calculation No. 1001527.305, Revision 2, Hatch Unit 2 P-T Curve Calculation for 37 and 50.1 EFPY NL-15-2034, E.I. Hatch, Units 1 and 2 - Enclosure 4, Non-Proprietary Version of the Epri/Bwrvip Information Requested in RAI 1 NL-15-2034, E.I. Hatch, Units 1 and 2 - Response to Request for Additional Information Regarding Application for Amendment to Technical Specifications for Relocation of Pressure and Temperature (P-T) Curves to the Pressure and Temperature Limits Repo
References
NL-15-2034 1001527.303, Rev. 0
Download: ML15322A091 (30)

Text

V Structural Integrity Associates, Inc.Y File No.: 1001527.303 Project No.: 1001527 CALCULATION PACKAGE Quality Program: [] Nuclear [] Commercial PROJECT NAME: Plant Hatch Unit l&2 P-T Curve Evaluation CONTRACT NO.: P0: SNG10018845, Rev. 0 CONTRACT:

19862, Rev. 0 CALCULATION TITLE: Feedwater, Water Level Instrument, and Core DP Nozzle Fracture Mechanics Evaluation for Hatch Unit 1 dLIu UIII L r I~uI Lullli '.uL w JVW1Up1L~11 Document Affected Project Manager Preparer(s)

&Revision Pages Revision Description Approval Checker(s)

Signature

& Date Signatures

& Date 01 -30 Initial Issue Responsible Eni~ineer D. V. Sommerville D. V. Sommerville 12/30/2011 12/30/2011 Responsible Verifiers M. Qin*12/30/2011 T. J. Griesbach 12/30/2011 Page 1 of 30 F0306-01R1

$jmSbiwbra kIturity Associates, Inc.=Table of Contents 1.0 OBJECTIVE

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

4 2.0 METHODOLOGY.............................................................................

4 2.1 Unit and Nozzle Specific Methodology Overview ...............................

6 2.1.1 Fee dwater Nozzle .....................................................................

6 2.1.2 Water Level Instrument Nozzle ......................................................

7 2.1.3 Core Differential Pressure Nozzle...................................................

8 2.2 2-D FEM Correction Factor ........................................................

8 2.3 Boundary Integral Equation / Influence Function Methodology

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

10 2.4 WRC Bulletin 175 Methodology

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

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

11 2.5 ASME XI, G-22 14.3 Methodology for Radial Thermal Gradients

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

12 3.0 DESIGN INPUTS...................................

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

12 4.0 ASSUMPTIONS............................................................................

13 5.0 CALCULATIONS

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

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

14 5.1 Hatch Uniti1........................................................................

14 5.1.1 Feedwater Nozzle ....................................................................

14 5.1.2 Water Level Instrument Nozzle......

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

15 5.1.3 Core Differential Pressure Nozzle .................................................

15 5.2 Hatch Unit 2........:................................................................

16 5.2.1 Feedwater Nozzle ....................................................................

16 5.2.2 Water Level Instrument Nozzle.............................................

.........

17 5.3 Justification for Linear Scaling of Thermal Stress Intensity Factor Solutions 17 6.0

SUMMARY

OF RESULTS...............................................................

18

7.0 REFERENCES

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

18 File No.: 1001527.303 Page 2 of 30 Revision:

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$jsbcrm~raI tor~ly Associates, Inc.=List of Tables Table 1 : Hatch Unit 1 FW Nozzle Blend Radius Path Stress Distributions

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

22 Table 2: Hatch Unit 2 FW Nozzle Blend Radius Path Stress Distributions...................

23 Table 3: Hatch Unit 1 Core DP Nozzle 1000 psi Internal Pressure Load Case Path Stress Distribution................................................................

24 Table 4: Hatch Unit 1 FW Nozzle K 1 p and KIT using the BIB/IF Methodology...............

25 Table 5: Hatch Unit 2 FW Nozzle K 1 p and KIT using the BIB/IF Methodology................

25 Table 6: Hatch Unit 1 Core DP Nozzle K 1 p using the BIB/IF Methodology...................

25 Table 7: Hatch Unit 1 FW Nozzle K 1 p using the WRC Bulletin 175 Methodology...........

25 Table 8: Hatch Unit 1 Core DP Nozzle K 1 p using the WRC Bulletin 175 Methodology.....25 Table 9: Hatch Unit 2 FW Nozzle K 1 p using the WRC Bulletin 175 Methodology...........

26 Table 1.0: Summary of Nozzle K 1 t Results Using the BIB/IF and G-221 14.3 Methodologies.................................

i.............................

26 Table 11 : Maximum Stress Intensity Factor for WLI Nozzle Considering Thermal Load Cases......................................................

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

26 Table 12: Summary of Nozzle Stress Intensity Factors .........................................

27 List of Figures Figure 1. Typical Nozzle Corner Crack Stress Extraction Path Orientation...................

28 Figure 2. WRC Bulletin 175, Figure A5-i, Bstimates of Stress Intensity Factors for Flaws at a Nozzle Corner ..................................................

........28 Figure 3. Hatch Unit 1 FW Nozzle Blend Radius Path Stress Distributions...................

29 Figure 4. Hatch Unit 2 FW Nozzle Blend Radius Path Stress Distributions...................

29 Figure 5. Hatch Unit 1 Core DP Nozzle 1000 psi Internal Pressure Load Case Path Stress Distribution................................................................

30 Figure 6. Plant C WLI Nozzle KIT for Three Thermal Transients

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

30 File No.: 1001527.303 Page 3 of 30 Revision:

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1.0 OBJECTIVE

The objective of this calculation is to calculate the Mode I stress intensity factors, KI, for the Edwin I.Hatch Unit 1 and Unit 2 Feedwater (FW) nozzles and Water Level Instrument (WLI) nozzles, and the Unit 1 Core Differential Pressure (DP) nozzle, necessary for creation of the Pressure-Temperature (P-T)limit curves for the pressure test (Curve A), core not critical (Curve B) and core critical (Curve C)conditions for the Hatch Unit 1 and Unit 2 reactor pressure vessels (RPV). Both internal pressure and thermal transient load cases are considered.

2.0 METHODOLOGY

Consistent with the Structural Integrity Associates, Inc. (SI) Boiling Water Reactor (BWR) P-T Curve Licensing Topical Report (LTR) [1], the FW nozzle is normally taken as the limiting component in the non-beltline region of the RPV. This assumption is made because: 1. The geometric discontinuity caused by the nozzle penetration in the RPV shell causes a stress concentration which results in larger pressure induced stresses than would be calculated in the shell regions of the RPV, 2. The FW nozzle experiences more severe thermal transients than most of the other nozzles because of the feedwater injection temperature which causes larger thermal stresses than are experienced in the shell regions of the RPV, 3. Although some other nozzles can experience thermal transients which would cause thermal stresses larger than those calculated for the shell regions of the RPV and some nozzles are larger diameter than the FW nozzle, which could result in a slightly larger KIp, the combined stresses from the applied thermal and pressure loads are considered to bound all other non-beltline discontinuities.

The Hatch Unit 1 and Unit 2 Adjusted Reference Temperature calculations

[2] identify the WLI nozzle as contained within the beltline region of both the Unit 1 and Unit 2 RPV. Consequently, the effects of these nozzles must be considered in the beltline P-T curve development.

These nozzles will cause a stress intensification in the beltline shells. Further, since the beltline region experiences a reduction in toughness caused by neutron irradiation, it is not obvious whether the FW nozzle will bound the WLI Nozzle throughout the life of Hatch Unit 1 and Unit 2. Consequently, the effects of the WLI Nozzle on the beltline P-T curves must be specifically considered.

The SI P-T Curve LTR [1] addresses the bottom head penetrations by conservatively applying a stress concentration factor (SCF) of 3.0 for a hole in a flat plate to the pressure induced membrane stress in the bottom head shell and using the thermal stress intensity factor solution given in ASME XI, Non-mandatory Appendix G, Paragraph G-22 14.3 [8]; discussed below. Review of the Hatch Unit 1 general assembly drawing [30] shows that the Core DP nozzle exists in the thinnest section of the bottom head.Based on prior experience from a similar bottom head design, application of the conservative SCF=3.0 methodology to the Core DP nozzle penetration will result in a bottom head P-T curve which controls the entire RPV. Consequently, a detailed evaluation of the Core DP nozzle is performed to remove excess conservatism.

File No.: 1001527.303 Page 4 of 30 Revision:

0 F0306-01R1 Consistent with 10OCFR50 Appendix G [3] the RPV P-T curves are applicable for normal operation and all anticipated operating occurrences.

Consequently, all Level A and Level B (Normal and Upset)operating events defined on the RPV, FW nozzle, WLI nozzle, and Core DP nozzle thermal cycle diagrams (TCD) [4, 5, 6] are considered in selecting bounding thermal and pressure conditions for preparing P-T curves.The SI P-T Curve LTR [1 ] identifies acceptable methodologies for calculating applied pressure and thermal stress intensity factors for postulated nozzle corner flaws. These methodologies include: Pressure Load Case: 1. Welding Research Council Bulletin 175 [7]2. Boundary Integral Equation / Influence Function (BIE/IF) [1]*Thermal Transient Load Case: 1. BIB/IF [1]2. ASME XI, Non-mandatory Appendix G, Paragraph G-2214.3 [8]The WRC Bulletin 175 [7] methodology for calculating a KI for an internal pressure load case is convenient to apply since no nozzle specific finite element analysis (FEA) is necessary.

The only inputs required are the nozzle and vessel geometry and the hoop stress calculated for the vessel shell, remote from discontinuities.

The BIB/IF methodology is applicable to any load case provided that a third order polynomial curve fit to the applicable stress distribution is available.

The appropriate path for a postulated nozzle corner crack has a path origin located at the peak stress location in the blend radius for the pressure load case and it is oriented parallel to a 45° line through the nozzle, as shown in Figure 1. This methodology requires detailed stress distributions through the nozzle blend radius which are typically obtained from a plant specific FEA.The flat plate thermal stress intensity factor solution given in the American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel (B&PV) Code Section XI, Non-mandatory Appendix G, Paragraph G-22 14.3 [8] is considered to be a conservative, simplified methodology for obtaining the KIT at the nozzle corner path resulting from a heat-up/cool-down transient of constant rate (i.e. 100 °F/hr), when the wall thickness is taken asthe path length along the 450 path identified in Figure 1. This approach is considered to be conservative because: 1. Geometric discontinuities do not intensify thermal stresses in a manner similar to stresses from mechanical loading; thus, fracture mechanics solutions which inherently consider the nozzle geometry are not necessarily required, 2. Thermal stresses increase as section thickness increases because the differential thermal strain increases with thickness.

Consequently, the practice of taking a wall thickness determined by the path length of the 450 path results in a thicker wall.File No.: 1001527.303 Page 5 of 30 Revision:

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Integrit Associates, Inc: SI has prepared a complementary P-T Curve LTR [9] which provides simplified methods for calculating the K~p and KIT for the BWR WLI nozzles. This method is convenient since the necessary stress intensity factors can be calculated using simple geometry and material properties data without the need for a plant specific FEA.The methods identified above are used in this calculation to calculate stress intensity factors for the Hatch Unit 1 and Unit 2 FW nozzles and WLI nozzles, and the Unit 1 Core DP nozzle. The sections below identify the general methodology applied for each nozzle at each unit.2.1 Unit and Nozzle Specific Methodology Overview The specific methodologies used to calculate the Kjp and KIT for the FW nozzle, the WLI nozzle, and the Core DP nozzle are discussed separately below.2.1.1 Feedwater Nozzle Various FW nozzle fracture mechanics analyses have previously been performed for Hatch Unit 1 [10]and Unit 2 [11] to satisfy the requirements of NUREG-0619

[ 12] regarding FW nozzle rapid thermal cycling caused by leakage past the thermal sleeve seals. The load cases necessary for development of P-T curves and those required to address NUREG-06 19 are similar. Consequently, the previous evaluations performed for Hatch Units 1 and 2 will be utilized, as appropriate, for the present evaluation.

The following methodology is used: Internal Pressure Load Case 1. Pressure stress distributions reported in References

[10, 11] for Hatch Unit 1 and Unit 2 are taken for the 1000 psig internal pressure load case.2. Recognizing that the Reference

[10, 11] evaluations were performed using a 2-D axi-symmetric finite element model (FEM) and that it is known that the stress intensification caused by the nozzle geometry is under predicted in a 2-D axi-symmetric representation of the nozzle, a correction factor must be applied to the stresses obtained from the 2-D axi-symmetric FEM. The internal pressure load case stresses, in the nozzle blend radius region, are corrected using the methodology presented in Reference

[13]. Plant specific dimensions are used to calculate the correction factor.3. The BIE/IF methodology presented in the SI P-T Curve LTR [1] is used to calculate K~p by fitting a third order polynomial equation to the path stress distribution for each plant specific pressure load case. The resulting K 1 p can be linearly scaled to determine the K 1 p for various RPV internal pressures.

4. 'The methodology given in WRC Bulletin 175 [7] is used to obtain an independent confirmation of the K 1 p calculated for each unit. The resulting K 1 p can be linearly scaled to determine the K~p for various RPV internal pressures.

File No.: 1001527.303 Page 6 of 30 Revision:

0 F0306-01RI iuru hItud Associates, Inc." Thermal Transient Load Case 1. Thermal shock load case path stress distributions reported in References

[10, 11] for Hatch Unit 1 and Unit 2 are taken. The FW nozzle thermal shock is the most severe Level A/B thermal transient for the FW nozzle; thus, the load case considered in References

[10, 11] is appropriate for P-T curve calculations.

When path stress distributions at multiple times are presented in References

[10, 11 ] the bounding distribution is selected for the current evaluation.

2. The BIE/IF methodology presented in the SI P-T Curve LTR [1] is used to calculate KIT by fitting a third order polynomial equation to the path stress distribution for each plant specific thermal shock load case. The resulting KIT can be linearly scaled to determine the KIT for various shock amplitudes.

3. A KIT for a uniform 100 °F/hr and 200 "F/hr heat-up/cool-down transient is calculated using the equation given for a radial thermal gradient in ASME XI, Appendix G, Paragraph G-22 14.3 [8]4. Results from various plant specific evaluations are presented to support the use of G-22 14.3 [8]for calculating KIT for 100 "F/hr and 200 0 F/hr heat-up/cool-down transients.

2.1.2 Water

Level Instrument Nozzle Simplified methods for calculating the K~p and KIT for the WLI nozzles in General Electric designed BWRs are given in Reference

[9], which is a companion LTR to the P-T curve LTR [1]. The K~p and KIT terms are calculated using Equations (8-1) and (8-2) of Reference

[9], which are repeated below, for convenience, as Eq. (1) and Eq. (2): K 1 Pesue .94~ ~1000Q psig internal pressure (1)KIRa,np =

++/-t,)]- 20.715, 100 °F/hr cooldown transient (2)Where: R is the inside radius of the pressure vessel, in tv is the wall thickness of the pressure vessel, in tn is the thickness of the WLI nozzle insert near the postulated crack location, in oa is the coefficient of thermal expansion at the highest temperature in the transient, in/in/°F The units of Ki in Eq. (1) and Eq. (2) are ksi-in°5.File No.: 1001527.303 Page 7 of 30 Revision:

0 F0306-01R1 jS b-c orl Aitii~ ssociates, In.=2.1.3 Core Differential Pressure Nozzle The following methodology is used: Internal Pressure Load Case: 1. The methodology given in WRC Bulletin 175 [7] is used to calculate the K 1 p. The resulting K 1 p can be linearly scaled to determine the K~p for various RPV internal pressures.

2. Since a 2-D axi-symmetric finite element analysis of a Core DP nozzle for a plant with a similar design has previously been performed for development of P-T curves [28], and since the dimensions of the Core DP nozzle modeled [29], in the vicinity of the bottom head penetration, are identical to Hatch Unit 1 [30], the results of the previous Core DP nozzle evaluation may be used to obtain an independent benchmark of the K 1 p obtained using the WRC Bulletin 175 [7]methodology.

3. A path stress distribution, in the vicinity of the Core DP nozzle, caused by a 1000 psig internal pressure load case, is taken from Reference

[28]. The location and orientation of the path are consistent with that used for nozzle evaluations.

4. The BIE/IF methodology presented in the SI P-T Curve LTR [1] is used to calculate K 1 p by fitting a third order polynomial equation to the path stress distribution for the plant specific pressure load case.Thermal Transient Load Case: 1. Similar to the methodology given in the P-T Curve LTR [1], a KIT for a uniform 100 °F/hr and 200 0 F/hr heat-up/cool-down transient is calculated using the equation given for a radial thermal gradient in ASME XI, Appendix G, Paragraph G-22 14.3 [8]2. Results from various plant specific evaluations are presented to support the use of G-2214.3 [8]for calculating KIT- for 100 0 F/hr and 200 °F/hr heat-up/cool-down transients.

2.2 2-D FEM Correction Factor When a cylindrical nozzle intersection with a cylindrical pressure vessel is modeled using a 2-D axi-symmetric simplification, the geometry is approximated as a nozzle intersection with a spherical shell.This simplification results in a non-conservative treatment of the geometric discontinuity at the blend radius region as well as a resulting reduction in the far field membrane stress in the shell caused by the approximation of the pressure vessel as a sphere rather than a cylinder.

Consequently, stress results obtained from a 2-D axi-symmetric FEM, for mechanical loads such as pressure, must be corrected before they can be used in subsequent analyses.

Sommerville and Walter [13] describe a methodology for correcting stress distributions obtained from the nozzle blend radius region of 2-D axi-symmetric models. The correction factor is given in Reference

[13] as: File No.: 1001527.303 Page 8 of 30 Revision:

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~jSbircraI laturif Associatus, Inc: SCF3DHo CF = 2. 3DHo (3)SCF 2 D)_Hoop Where: SCF 3 DHoop is the stress concentration factor (SCF) defined with respect to the the hoop stress direction, for the 3D geometry.SCF 2 DHoop is the stress concentration factor for the 2D geometry.The expected SCF for the 3-D geometry can be estimated using the following equation for a circular hole in a pressurized cylinder [14]: For 0132, K(, (/) = 2.5899 +0.8002.13+/-+4.0112.132

_-1.8235-/83

+ 0.3751 ./34 (4a)~j.( 2)f 2 " (4b)Where: Kt(I3) is the SCF defined with respect to the far field hoop stress, PR/t.v is the Poisson's ratio of the material, assumed equal to 0.3.r is the radius of the nozzle bore, in R is the inside radius of the pressure vessel, in tv is the wall thickness of the pressure vessel, in Kt(t3) is used for the SCF3D)_Hoop term in Eq. (3). The SCF2DHoop can be calculated from the results of the 2-D axi-symmetric FEM by calculating the SCF using the following equation: 2 ., O'totl,max (5)SCF 2-H°°P = P. R Where: P is the RPV internal pressure for the pressure load case, psi O'total,max is the largest total hoop stress in the blend radius region, psi Inserting Eq. (5) and Eq. (4a, 4b) into Eq. (3) gives the correction factor which can be used to uniformly scale the pressure load case path stress distribution from the 2-D axi-symmetric FEM. This path stress distribution can then be fit with the equation for a 3 rd order polynomial and used with the BIE/IF solution described in Reference

[1], and discussed below, to obtain Kjp.File No.: 1001527.303 Page 9 of 30 Revision:

0 F0306-01R1 mIhtugril Associates, Inc.=2.3 Boundary Integral Equation / Influence Function Methodology The following discussion is excerpted from the SI P-T curve LTR [1]. Note that the equation and reference numbers in the excerpt below refer to Reference

[1]: The stress intensity factors for the feedwater nozzle may be calculated using the results of a detailed finite element model of the nozzle. In some cases, such results may already be available from the governing design basis stress report for the fee dwater nozzle. The details of the finite element process are not included here," rather, the extraction of the. appropriate finite element results and their use in developing P-T limit curves is discussed.

For a path through the limiting nozzle inner blend radius corner, as shown in Figure 2-7, the thermal and pressure hoop stress distributions should be extracted from the finite element model. Each of the stress distributions should befit with a third-order polynomial that reasonably fits the calculated stresses in the region of interest.The thermal stress intensity factor, K 1 t, is computed based on either of the nozzle corner solutions shown in Figure 2 -8 for a postulated 1/4t (based on the section thickness) axial defect, as follows." K 1 , = [ 0.723 C 0 , + 0.55 1 ~1)C 1 , + 0.462~Z C 2 , + 0.4 0 8 j J C 3 ,] (2.5.3-3 a)K 1 , V~[O.06C~

+0.53 C 1 + .4 4~ i~j C, +/-.39~jj ~ 3~j(2.5.3-3b) where: Kit the thermal stress intensity factor for the limiting normal/upset transient (psi inc)a = 1/4tpostulatedflaw depth (inches)t = thickness of the cross-section through the limiting nozzle inner blend radius corner, as Shown in Figure 2- 7.Co,, C 1 ,., C 2 ,, C 3 , thermal stress polynomial coefficients based on fits to finite element analysis.File No.: 1001527.303 Page 10 of 30 Revision:

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$bSl wc~ Ilte! grit Associates, Inc.=Equation 2.5.3-3 a is based upon a Boundary Integral Equation / Influence Function (BIE/IF) solution developed for a quarter crack in an infinite quarter space.Equation 2.5. 3-3b is the average of the BIE/IF solutions developed for a quarter crack in an infinite quarter space and a semicircular crack in an infinite half space.These equations have been investigated by the NRC and Oak Ridge National Laboratory and shown to be acceptable for use in calculating the applied stress intensity factor for a corner cracked nozzle [16]. Although both solutions given above are evaluated in Reference

[16], it is acknowledged by the original authors of these formulations, in the basis work used to develop the approach [171, that the twvo formulations differ very little and in fact provide KI values which differ only by approximately 5%. This can be seen by review of the coefficients used in each equation above. Consequently, either Equation 2.5.3-3 a or 2.5.3-3b may be used for any nozzle configuration in a BWR.The BIE/IF solution introduced above is applicable to any load case and any BWR nozzle; thus, it is applied both to the internal pressure and thermal transient loads considered for P-T curve development.

2.4 WRC Bulletin 175 Methodology An alternative solution for determining the applied pressure stress intensity factor, Kip, is to use the method given in Appendix 5 of WRC Bulletin .175 [7], where: K 1~ zFar~.P l~ (6 Where: K 1 p is the applied pressure stress intensity factor, psi ic P is the operating pressure, psi a is the 1/4t postulated flaw depth, inches R is the vessel inner radius, inchesis the thickness of the vessel shell, inches F(a/r,) is the shape factor given in Figure 2, where rn ri + 0.29rc, ri is the actual inner radius of nozzle, inches rc is the nozzle blend radius, inches A functional form of F(a/rn) is given in Reference

[15, Pg. 11.1-17] as:= .42 .0 .46 (7)Where, Eq. (7) is applicable for 0.,07 < a/rn< 0.9.File No.: 1001527.303 Page 11 of 30 Revision:

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$jtSinwwrul I~l~itedt Associates, Inc.*2.5 ASME XI, G-2214.3 Methodology for Radial Thermal Gradients ASME XI, Non-mandatory Appendix G, Paragraph G-22 14.3 [8] gives the following, simplified method for calculating KIT for a radial thermal gradient: Kt =-0.953.10-3**CR. tv 2 5 (8)Where: CR is the cooldown rate, 0 F/hr.tv is the thickness of the vessel shell, inches Paragraph G-22 14.3 [8] states that Eq. (8) will yield conservative results if used for cool-down rates greater than 100 °F/hr.This methodology is used to calculate the KIT in the nozzle blend radius region by using the path length along the 450 path shown in Figure 1 as the shell thickness, tv.3.0 DESIGN INPUTS The following design inputs are used for this evaluation:

• Previous NUREG-0619 evaluations:

References

[10, 11]* Core DP Nozzle evaluation:

Reference

[28]* Thermal transient definitions for Level A/B: References

[4, 5, 6]* WLI Nozzle dimensions and material:

Reference

[16]Hatch Unit 1 : Nozzle insert material:

Inconel, SB- 166 RPV inside radius: 110.375 inches Nozzle insert thickness:

0.28 1 inches Vessel shell thickness:

5.375 inches Note." Some drawings show a shell thickness of 5.875 inches;" however, the minimum dimension given in the general arrangement drawing is used for this evaluation.

Hatch Unit 2: Nozzle insert material:

Inconel, SB- 166 RPV inside radius: 110.375 inches Nozzle insert thickness:

0.66 1 inches Vessel shell thickness:

5.375 inches File No.: 1001527.303 Page 12 of 30 Revision:

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$j~biicfriru Iutugdty Associates, Inc.e* FW Nozzle dimensions:

Hatch Unit 1: Nozzle bore diameter: Nozzle blend radius: RPV inside radius: Vessel shell thickness:

Hatch Unit 2: Nozzle bore diameter: Nozzle blend radius: RPV inside radius: Vessel shell thickness:

• Core DP Nozzle dimensions:

Hatch Unit 1: Vessel shell thickness:

Bottom head radius: Nozzle inside radius:* Material Properties:

Hatch Unit 1 and Unit 2 WLI Nozzle Insert: Coefficient of thermal expansion:

Poisson's ratio: Reference

[ 17]6.7 inches 2.5 inches 110.375 inches 5.375 inches 6.5 inches 5.0 inches 110.375 inches 5.375 inches Reference

[30]3.188 inches 110.5 inches 1.250 inches, see assumption 3.Reference

[ 18]7.7x106 in/ir/°F at 550 0 F 0.3 (assumed)4.0 ASSUMPTIONS The following assumptions are used in this calculation and supported by data presented below: 1. The thermal stress intensity factor, KIT, for a 200 °F/hr thermal transient can be conservatively calculated using Eq. (8) above for vessel shells and forged vessel nozzles, where the shell thickness in the equation is taken as the vessel shell thickness when evaluating the shell, and the nozzle blend radius path length when evaluating forged nozzles. Adequacy of this assumption is demonstrated in Section 5.3 below.2. The thermal stress intensity factor, KIT, for a 200 °F/hr thermal transient can be conservatively calculated using Eq. (2) above for a WLI nozzle by scaling the KIT obtained from Eq. (2) by 2.Similarly higher heat-up/cool-down rates can be addressed by appropriate scaling factors.Adequacy of this assumption is demonstrated in Section 5.3 below.File No.: 1001527.303 Revision:

0 Page 13 of 30 F0306-01RI

$b~ml inc.=3. The stress concentration effect of the Core DP penetration in the bottom head is conservatively addressed by treating the penetration as a nozzle in which the radius considered in the WRC Bulletin 175 pressure stress intensity methodology is taken as the radius of the hole in the shell rather than the ID of the Core DP penetration.

This assumption is validated in Section 5.1.3 below by comparing the KIP obtained using this approach with the KIP obtained using the BIB/IF methodology.

5.0 CALCULATIONS

The calculations for the FW nozzles, WLI nozzles, and Core DP nozzle are presented for each unit, separately, below.5.1 Hatch Unit 1 The calculations for the FW, WLI, and Core DP nozzles are presented in separate sections.5.1.1 Feedwater Nozzle Table 1 presents a tabulation of the path stress distribution taken from the Unit 1 plant specific FW nozzle FEM, in the blend radius region of the nozzle. Both the 1000 psig internal pressure and 450 0 F thermal shock load case path stress distributions are taken from Reference

[i10a]. The thermal transient load case path stress distribution is fit with a 3 rd order polynomial equation.

In the previous evaluation

[10a], the pressure load case path stress distribution was corrected using a different methodology than utilized for the present calculation; therefore, this correction factor is removed before the correction factor calculated using the methodology described in Section 2.2 is applied. The corrected pressure path stress distribution is fit with a 3 rd order polynomial equation.The Hath Unit 1 FW nozzle correction factor is calculated below: r = 6.7 in FW nozzle bore radius R = 110.38 in RPV radius adjacent to FW nozzle tv= 5.38 in RPV shell thickness adjacent to FW nozzle 3=- 0.177 -Eq. (4)Kt 2.85 -SCF for hole in cylinder, Eq. (4)SCF 2 0= 2.38 -SCF from 2-D axi-symmetric FEA, Eq. (5)CF = 2.39 -Correction factor, Eq. (3)Figure 3 is a plot of the pressure and thermal shock path stress distribution in the Hatch Unit 1 FW nozzle blend radius region with the polynomial curve fit equations and correlation coefficients shown.Table 4 summarizes the polynomial coefficients for each load case and presents the Kip and KIT for the Hatch Unit 1 FW nozzle.Table 7 presents the K 1 p calculated using the WRC Bulletin 175 methodology

[7].File No.: 1001527.303 Page 14 of 30 Revision:

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$b~o~grl Associates, Inc.=Recognizing that the BIB/IF methodology has been shown to provide conservative estimates of the stress intensity factors for both pressure and thermal load cases [19, 20] and that both the WRC Bulletin 175[7] and BIE/IF [1] methodologies are accepted methods for calculating the KI from the pressure load case, the lower KI is used for this evaluation.

Reference

[10Ob] identifies that the extended power uprate conditions result in an increase in the RPV dome pressure of 50 psig (1000 psig to 1050 psig) and an increase in the FW fluid temperature of 6 °F (450 °F to 456 'F). The increase in dome pressure is accommodated by linear scaling during preparation of the P-T curves in a subsequent calculation.

The increase in the FW fluid temperature is a change of less than 1.5%. The work documented in References

[19, 20] shows that the BIB/IF methodology is significantly conservative (on the order of 30-50% when compared to the RMS K 1 calculated from a 3-D finite element fracture mechanics evaluation) for the 1/4 thickness flaws required for P-T curves;consequently, the KIT calculated for the 450 °F shock is not further increased in this evaluation to account for the small increase in FW temperature.

5.1.2 Water

Level Instrument Nozzle Using Eq. (1) and (2) and considering Hatch Unit 1 WLI nozzle dimensions and material properties, the K 1 p and KIT are calculated below: tv= 5.38 in RPV shell thickness adjacent to WLI nozzle R = 110.38 in RPV radius adjacent to WLI nozzle tn= 0.28 1 in WLI nozzle thickness cx 7.70E-06 in/in/OF Thermal expansion coefficient for nozzle material, Inconel, 550 'F K = 71.6 ksi-in 0.5 1000 psi pressure stress intensity factor, Eq. (1)KIT = 17.4 ksi-in 0.5 100 'F/hr cool-down transient stress intensity factor, Eq. (2)KIT= 34.8ksi-n 0 5 200 'F/hr cool-down transient stress intensity factor, Eq. (2) scaled by a KXT = 34.8ksi'n°'S factor of 2 5.1.3 Core Difjferential Pressure Nozzle The K 1 p calculated using the WRC Bulletin 175 methodology

[7] is shown in Table 8.The pressure path stress distribution is extracted from Reference

[28]. The stress distribution is tabulated in Table 3 and plotted in Figure 5. Because of the orientation of the path and the configuration of the nozzle, which contains a tube which penetrates the RPV, the path stress distribution exhibits a lower stress at the path origin than slightly inboard along the path. Consequently, the 3 rd order polynomial equation is fit to the stress distribution, after omitting the first point along the curve. The curve fit and equation are shown in Figure 5. The resulting polynomial coefficients and K 1 p obtained using the BIB/IF methodology are listed in Table 6.File No.: 1001527.303 Page 15 of 30 Revision:

0 F0306-01iRI

$j~bvg~aunIlatgrily ssoca~tes, Inc.'The K~p calculated using the BIE/IF methodology

[1] benchmarks well against the K 1 p calculated using the WRC Bulletin 175 methodology

[7]. Recognizing that the BIE/IF methodology has been shown to provide conservative estimates of the stress intensity factors for both pressure and thermal load cases[ 19, 20] and that both the WRC Bulletin 175 [7] and BIE/IF [1 ] methodologies are accepted methods for calculating the K 1 from the pressure load case, the lower KI is used for this evaluation.

The KIT term is calculated using Eq. (8) and is presented in Table 12 for heat-up/cool-down rates of 100 0 F/hr and 200 0 F/hr. Section 5.3 provides justification for using the ASME XI, Non-mandatory Appendix G, Paragraph G-2214.3 [8] methodology for heat-up/cool-down rates greater than 100 0 F/hr.5.2 Hatch Unit 2 The calculations for the FW and WLJ nozzles are presented in separate sections.5.2.1 Feedwater Nozzle Table 2 presents a tabulation of the path stress distribution taken from the Unit 2 plant specific FW nozzle FEM, in the blend radius region of the nozzle. Both the internal pressure and thermal shock load case path stress distributions are taken from Reference

[11 la]. The thermal transient load case path stress distribution is fit with a 3r order polynomial equation.

In the previous evaluation

[1 la], the pressure load case path stress distribution was corrected using a different methodology than utilized for the present calculation; therefore, this correction factor is removed before the correction factor calculated using the methodology described inSectioh 2.2 is applied. The corrected pressure path stress distribution is fit with a 3 rd order polynomial equation.The Hath Unit 2 FW nozzle correction factor is calculated below: r = 6.5 in FW nozzle bore radius R = 110.38 in RPV radius adjacent to FW nozzle t-- 5.38 in R!PV shell thickness adjacent to FW nozzle 13= 0.172 -Eq. (4)Kt 2.84 -SCF for hole in cylinder, Eq. (4)SCF 2 D = 2.05 -SCF from 2-D axi-symmetric FEA, Eq. (5)CF = 2.77 -Correction factor, Eq. (3)Figure 4 is a plot of the pressure and thermal shock path stress distribution in the Hatch Unit 2 FW nozzle blend radius region with the polynomial curve fit equations and correlation coefficients shown.Table 5 summarizes the polynomial coefficients for each load case and presents the K 1 p and KIT for the Hatch Unit 2 FW nozzle.Table 9 presents the K 1 p calculated using the WRC BUlletin 175 [7] methodology.

File No.: 1001527.303 Page 16 of 30 Revision:

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hi/t egd/ty Associateus, Inc." Recognizing that the BIE/IF methodology has been shown to provide conservative estimates of the stress intensity factors [19, 20] and that both the WRC Bulletin 175 [7] and BIE/IF [1] methodologies are accepted methods for calculating the KI from the pressure load case, the lower K 1 is used for this evaluation.

Reference

[1 lc] identifies that the extended power uprate conditions result in an increase in the RPV dome pressure of 50 psig (1000 psig to 1050 psig) and an increase in the FW fluid temperature of 6 0 F (delta T from 450 0 F to 456 0 F). The increase in dome pressure is accommodated by linear scaling during preparation of the P-T curves in a subsequent calculation.

The increase in the FW fluid temperature is a change of less than 1.5%. The work documented in References

[19, 20] shows that the BIE/IF methodology is significantly conservative (on the order of 3 0-50% when compared to the RMS K 1 calculated from a 3-D finite element fracture mechanics evaluation) for the 1/4 thickness flaws required for P-T curves; consequently, the KIT calculated for the 450 0 F shock is not further increased in this evaluation to account for the small increase in FW temperature.

5.2.2 Water

Level Instfrument Nozzle Using Eq. (1) and (2) and considering Hatch Unit 2 WLI nozzle dimensions and material properties, the K~p and KIT are calculated below: tv= 5.38 in RPV shell thickness adjacent to W/LI nozzle R = 110.38 in RPV radius 'adjacent to WLI nozzle tn= 0.66 1 in WLI nozzle thickness ct= 7.70E-06 in/in/°F Thermal expansion coefficient for nozzle material, Inconel, 550 0 F Kw= 80.0 ksi-in°'s 1000 psi pressure stress intensity factor, Eq. (1)KI 1 = 19.9 ksi-in°'s 100 °F/hr cooldown transient stress intensity factor, Eq. (2)200 °F/hr cooldown transient stress intensity factor, Eq. (2) scaled by a factor KIT = 39.9 ksi-in°'s of 2.0 5.3 Justification for Linear Scaling of Thermal Stress Intensity Factor Solutions Data is presented in this section which supports the Assumptions 1 and 2 identified in Section 4.0.Table 10 summarizes the KIT calculated for four nozzles from three separate BWRs. Both a 100 0 F/hr and a 300 0 F/hr cool-down rate were evaluated.

Both the BIE/IF methodology

[1] and the methodology given in ASME XI, Non-mandatory Appendix G, Paragraph G-22 14.3 [8] were used to calculate KIT.The results confirm that the simplified method given in Paragraph G-22 14.3 [8] may be used to obtain a KIT. Further, as suggested in WRC 175 [7] and G-2214.3 [8] use of this methodology is expected to be increasingly conservative for cool-down rates larger than 100 0 F/hr. This trend is seen in the two KIT results Obtained for Plant .C at 100 0 F/hr and 300 0 F/hr, where it is shown that the KIT calculated using the flat plate solution exceeds that calculated using the BIB/IF methodology by a ratio of 2.34 for a 100 0 F/hr ramp rate and 2.71 for ai 300 0 F/hr ramp rate. Since the flat plate solution given in the ASME Code[8] is obtained from a quasi-steady state temperature distribution through the wall thickness it is anticipated that the solution becomes increasingly conservative for faster ramp rates since there is File No.: 1001527.303 Page 17 of 30 Revision:

0 F0306-01R1

$SbwouraI hIterf Associates, Inc." insufficient time for a quasi-steady state temperature distribution to develop through-wall since the vessel cools from approximately 550 0 F to 100 °F; for faster ramp rates the temperature ramp ends before a quasi-steady state thermal distribution can develop.Table 11 presents the KIT calculated using the BIE/IF methodology and using the simplified methodology given in the WLI Nozzle LTR [9] which was developed for a 100 0 F/hr cool-down transient.

Considering that the stress analysis and fracture mechanics methodologies are both linear elastic methods, the results should be scalable.

It is recognized that at faster ramp rates the vessel will not be able to develop the quasi-steady state temperature distribution possible for slower ramp rates;therefore, the methodology given in Reference

[9] is expected to yield conservative results for ramp rates faster than 100 °F/hr. The results in Table 11 and Figure 6 show a trend consistent with expectations.

Consequently, the methodology given in Reference

[9] for the 100 °F/hr ramp rate may be used to obtain a conservative KIT value for higher ramp rates.6.0

SUMMARY

OF RESULTS Table 12 summarizes the KI values calculated for the FW, WLI, and Core DP nozzles at Hatch Unit 1 and Unit 2 for the internal pressure load case, a 450 0 F thermal shock, a 100 °F/hr, and a 200 °F/hr cool-down transient.

These values will be used in subsequent calculations to prepare P-T curves for both Hatch Units.

7.0 REFERENCES

1. Sommerville, D.V., "Pressure-Temperature Limits Report Methodology for Boiling Water Reactors," SIR-05-044, Rev. 1, June 2011.2. Adjusted Reference Temperature Calculations:

a. Sommerville, D.V., "Hatch Unit 1 RPV Material Summary and ART Calculation," SI Calculation No. 1001527.301, Rev. 0.b. Sommerville, D.V., "Hatch Unit 1 RPV Material Summary and ART Calculation," SI Calculation No. 1001527.302, Rev. 0.3. Title 10 Code of Federal Regulations Part 50, Appendix G, "Fracture Toughness Requirements." 4. Reactor Pressure Vessel Thermal Cycle Diagrams: a. SNOC Dwg. S 15025, GE Dwg. 729E762, "Reactor Thermal Cycles," SI File No. GPCO-31 Q-209.b. SNOC Dwg. S-41615, GE Dwg. 761E246, Sht. 1, "Reactor Vessel Thermal Cycles (Including Black Start)," SI File No. 1001527.211.

c. SNOC Dwg. S-416 16, GE Dwg. 761E246, Sht. 2, "Reactor Vessel Thermal Cycles (Including Black Start)," SI File No. 1001527.211.

File No.: 1001527.303 Page 18 of 30 Revision:

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hItugdl Associates, Inc." 5. Feedwater Nozzle Thermal Cycle Diagrams: a. GE Dwg. 135B9990, "Nozzle Thermal Cycles (Feedwater)," SI File No. 1001527.211.

b. SNOC Dwg. $26421, GE Dwg. 158B8369, Sht. 4, Rev. 2, "Nozzle Thermal Cycles -Including Black Start (Feedwater

-Normal & Upset Conditions)," SI File No. 1001527.211.

c. SNOC Dwg. S26422, GE Dwg. 158B8369, Sht. 5, Rev. 2, "Nozzle Thermal Cycles -Including Black Start (Feedwater-Emergency

& Fault Conditions)," SI File No. 1001527.211.

6. Water Level Instrument Nozzle Thermal Cycle Diagrams: a. GE Dwg. 135B9990, Sht. 7, Rev. 0, "Nozzle Thermal Cycles (Instrumentation

& Core Diff.Press & Liquid Control)," SI File No. 1001527.211.

b. SNOC Dwg. S26426, GE Dwg. 1 58B 8369, Sht. 9, Rev. 2, "Nozzle Thermal Cycles -Including Black Start (Instrumentation

& Core Diff. Press & Liquid C," SI File No. 1001527.211.

7. PVRC Recommendations on Toughness Requirements for Ferritic Materials, WRC Bulletin 175, August 1972.8. American Society of Mechanical Engineers, Boiler and Pressure Vessel Code,Section XI, Rules for Inservice Inspection of Nuclear Power Plant Components, Non-mandatory Appendix G, "Fracture Toughness Criteria for Protection Against Failure," 2001 Ed. through 2003 Addenda.9. Sommerville, D. V., "Linear Elastic Fracture Mechanics Evaluation of General Electric Boiling Water Reactor Water Level Instrument Nozzles for Pressure-Temperature Curve Evaluations," SI Report 0900876.401, Rev. 0, June 2011.10. Hatch Unit 1 NUREG-0619 Evaluations:

a. Liffengren, D. J., et al., "Edwin I. Hatch Nuclear Power Station, Unit 1 Feedwater Nozzle Fracture Mechanics Analysis to Show Compliance with N UREG-06 19," NEDE-30238, DRF-137-0010, August 1983, General Electric Company. SI File No. 1001527.210.

GE Proprietary Information.

b. Bothne, D., "Power Uprate Evaluation Report for Edwin I. Hatch Unit 1, Feedwater Nozzle NUREG-06 19 Fracture Mechanics Analysis for Extended Power Uprate Conditions," GE-NE-B13-01869-065-01, July 1997, General Electric Company. SI File No. 1001527.210.

GE Proprietary Information.

File No.: 1001527.303 Page 19 of 30 Revision:

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it Associates, 11. Hatch Unit 2 NUREG-0619 Evaluations:

a. Liffengren, D. J., et al., "Edwin I. Hatch Nuclear Power Station, Unit 2 Feedwater Nozzle Fracture Mechanics Analysis to Show Compliance with NUREG-0619," NEDC-30256, DRF-137-0010, August 1983, General Electric Company. SI File No. 1001527.210.

GE Proprietary Information.

b. Stevens, G. L., "Updated Feedwater Nozzle Fracture Mechanics Analysis for Edwin I. Hatch Nuclear Power Station Unit 2," GE-NE-523-95-0991, Rev. 0, DRF B 13-01524, September 1991, General Electric Company. SI File No. 1001527.210.

c. Bothne, D., "Power Uprate Evaluation Report for Edwin I. Hatch Unit 2, Feedwater Nozzle NUREG-0619 Fracture Mechanics Analysis for Extended Power Uprate Conditions," GE-NE-B13-01869-065-02, July 1997, General Electric Company. SI File No. 1001527.210.

GE Proprietary Information.

12. BWR Feedwater Nozzle and Control Rod Drive Return Line Nozzle Cracking, NUREG-06 19, November 1980, Nuclear Regulatory Commission.

13. Sommerville, D., Walter, M., "An Investigation into the Effects of Modeling Cylindrical Nozzle to Cylindrical Vessel Intersections Using 2D Axisymmetric Finite Element models and a Proposed Method for Correcting the Results," ASME PVP201 1-57001, Proceedings of the 2011 ASME Pressure Vessel and Piping Division Conference.

14. Pilkey, W.D., Pilkey, D.F., Peterson's Stress Concentration Factors, 3 rd. Ed., John Wiley & Sons, 2008.15. Zahoor, A., "Ductile Fracture Handbook," EPRI Report NP-6301, Volume 3. January 1991.16. Water Level Instrument Nozzle Drawings: a. SNOC Sketch 1-BE-i, Rev. 0, "Nil, N12, and N16 Instrumentation Nozzle Detail," SI File No. 1001527.208.

b. SNOC Sketch 2-BE-2, Rev. 1, "2N12 and 2N16 Instrumentation Nozzle Detail," SI File No.1001527.209.

17. Feedwater Nozzle Drawings: a. SNOC Dwg. SX18921, "Reactor Vessel Feedwater Nozzle As Built," SI File No.100 1527.208 b. SNOC Sketch 2-BF-4, Rev. 1, "2N4 Nozzle Assembly (Feedwater)," SI File No. GPCO-31Q-208.18. American Society of Mechanical Engineers, Boiler and Pressure Vessel Code,Section II, Part D, Materials, 2001 Ed. through 2003 Addenda.19. Yin, S., Bass, B. R., Stevens, G. L., Stress and Fracture Mechanics Analyses of Boiling Water Reactor and Pressurized Water Reactor Pressure Vessel Nozzles, ORNL/TM--2010/246, December 2010.File No.: 1001527.303 Page 20 of 30 Revision:

0 F0306-O01R1 rai lte gril Associates, Inc;" 20. Sommerville, D.V., Qin, M., Houston, E., "An Investigation of the Adequacy of a Simplified Boundary Integral Equation / Influence Function Equation Linear Elastic Fracture Mechanics Solution for Nozzle Corner Cracking," ASME PVP201 1-57742, Proceedings of the 2011 ASME Pressure Vessel and Piping Division Conference.

21. SI Calculation Package 1100445.302, Rev. 0.22. SI Calculation Package 1100445.303, Rev. 0.23. SI Calculation Package NPPD-13Q-302, Rev. 1.24. SI Calculation Package 1000847.302, Rev. 0.25. SI Calculation Package 1000847.303, Rev. 2.26. SI Calculation Package 1100151.302, Rev. 0.27. S1 Calculation Package 1100151.303, Rev. 0.28. SI Calculation Package 1100445.304, Rev. 0.29. Combustion Engineering Drawing 232-242, Nozzle Details, SI File No. 1100445.204.

30. Core DP Nozzle Drawings: a. SNOC Dwg. S-15227A, Combustion Engineering Drawing 234-244-5, Nozzle Details for 218" I.D. BWR, SI File No. 1001527.208.

b. SNOC Sketch 1-BE-2, Rev. 1, "N10 Standby Liquid Control & Core Differential Pressure Nozzle Detail," SI File No. 100 1527.208.c. SNOC Dwg. S15523, Combustion Engineering Drawing 234-270, Rev. 3, "General.Arrangement Elevation for 218" ID BWR," SI File No. 100 1527.208.File No.: 1001527.303 Revision:

0 Page 21 of3O0 F0306-01R1

$Sbwcbuu hlgl Assadate .°Table 1: Hatch Unit 1 FW Nozzle Blend Radius Path Stress Distributions.

1000 psig Internal Pressure Load Case Thermal Load Case 450 "F shock Path Distance, °h) Oh(2' 3 o, Gb, in. pipsi psi psi 0.000 f.l }.1 24461 58463 Di Di 0.075 ]} 24052 57486 II }I 0.225 1 /] 23283 55649 i }J-0.400 ft }1 22437 53626 1}0.600 It Di 21527 51452 U1 Di 0.850 U /] 20466 48916 [ }.i.1.150 }1 19262 46037 Di }1 1.500 J} 18081 43215 ft }1 1.960 1.1 }1 16848 40268 Ii 1.2.476 ft }1 15206 36344 Di Di 3.052 I.I ]J 13724 32802 Ui Di 3.628 II }i 12372 29571 Di4.204 II 11110 26554 IL Di 4.779 Di }) 9898 23657 lit Di 5.355 Di Di 8692 20774 sD 5.93 1 Di Di 7439 17780 ft }6.507 Di 1) 6073 14515 UJ Di 6.811 Di }i 5384 12868 ft Di Notes: 1. From Reference [10a], with CF=1.6557.

2. From Reference

[l0a], with CF=1.6557 removed.3. From Reference [10a], with CF as given in PVP201 1-5700 1 [13].File No.: 1001527.30, Revision: 0 3 Page 22 of 30 This page contains GEH PROPRIETARY INFORMATION which has been redacted F0306-01IRI mnssa t whlt A~ s A,.Table 2: Hatch Unit 2 FW Nozzle Blend Radius Path Stress Distributions. 1000 psig Internal Pressure Load Case Thermal Load Case 450 OF shock Path in. pipsi psi psi 0.000 U1 } 21010 58242 L1 }IL.0.075 LI }) 20651 57247 Li IL 0.225 LI }i 19970 55359 H I 0.400 I( }L 19213 53261 fL IL 0.600 Li II 18403 51015 Li IL 0.850 Li /) 17440 48345 LI II.1.150 IL 16452 45607 Li IL 1.500 LI IL 15206 42153 Li IL 1.954 Li IL 14233 39455 Li IL 2.693 LI IL 11909 33013 Li }L 3.284 LI II 10634 29479 {L Li 3.874 Zl H9468 26245 IL 4.465 Li II 8365 23188 {L }L 5.056 Li I}. 7289 20206 LI IL 5.647 LiII} 6194 17170 Li IL 6.247 Li IL 5029 13941 Li IL 6.838 LI II! 3718 10307 Li IL 7.127 Li IL 3040 8427 LI IL Notes: 1. From Reference 2. From Reference 3. From Reference[1 la], with CF=1.5987. [1 la], with CF=1.5987 removed.[1 la], with CF as given in PVP201 1-5700 1 [13].File No.: 1001527.303 Revision: 0 Page 23 of 30 This page contains GEH PROPRIETARY INFORMATION I which has been redacted F0306-01RI $b~jSlcrral hIt Writ Associates, Inc.=Table 3: Hatch Unit 1 Core DP Nozzle 1000 psi Internal Pressure Load Case Path Stress Distribution [281.Path Distance, ah, in. psi 0.000 26186 0.225 31610 0.450 30525 0.675 28449 0.901 26377 1.126 23518 1.351 22272 1.576 21189 1.801 20222 2.026 19351 2.251 18559 2.476 17831 2.702 17156 2.927 16522 3.152 15916 3.377 15327 3.602 14746 3.827 14169 4.052 13594 4.277 12838 4.503 12045 File No.: 1001527.303 Revision: 0 Page 24 of 30 F0306-01R1 8 q trailafugri Associates, mnc: Table 4: Hatch Unit 1 FW Nozzle K 1 p and KIT using the BIE/IF Methodology. Pressure JThermal A0 58242 51488 Al -11913 -15980'A2 1484.2 2002.4 A3 -105.18 -127.21 K 1 81.1 65.3 ksi-in°'5 Table 5: Hatch Unit 2 FW Nozzle K 1 p and KIT using the BIE/IF Methodology. Pressure II Thermal A0 58136 48077 Al -12560 -28076 A2 1522.5 4948.5 A3 -103.53 -598.8 K 1 81.3 46.8' ksi-in°'5 Table 6: Hatch Unit 1 Core DP Nozzle Kwp using the BIE/IF Methodology. Pressure A0 35123 Al -12874 A2 3144.7 A3 ,316.07 KI 38.9 ksi-inO.S Table 7: Hatch Unit 1 FW Nozzle K 1 p using the WRC Bulletin 175 Methodology. F(a/rn) 1.61 -PR 1/t(7ta)0 5 47.49 ksi-in°'5 KI 76.6 ksi-in°'s Table 8: Hatch Unit 1 Core DP Nozzle K~p using the WRC Bulletin 175 Methodology. F(a/rn) j0.99 -PRi/t(ira)° 5 32.58 ksi-in°'5 K______ I 32.3 ksi-in°'s File No.: 1001527.303 Page 25 of 30 Revision: 0 F0306-01R1

Assaciatus, Irnc=Table 9: Hatch Unit 2 FW Nozzle K~p using the WRC Bulletin 175 Methodology. F(a/r,) 1.62 -PRi/t~ta)0 5 48.58 ksi-in°5 K~p78.9 ksi-in 0'5 Table 10: Summary of Nozzle KIT Results Using the BIE/IF and G-22114.3 Methodologies. NozzleKIKT Plant ID Tp()Path Length, (BIE/IF Solution), (Flat Plate Soln.), Ratio Reference Te 2 in ksi-in°'5 ksi-in°'5 Plant A (3) FW 7.60 11.1 15.2 1.37 [21], [22], [23]Plant B FW 7.73 7.0 15.8 2.26 [24],[25]Plant B RI 9.29 9.4 25.1 2.67 [25]Plant C (1) FW 8.66 23.3 63.1 2.71 [26]Plant C FW 8.66 9.0 21.0 2.34 [27]1.2.3.300 *F/hr cooldown transient. FW is Feedwater nozzle, RI is Recirculation Inlet nozzle.Path length taken as 205 *RPV shell thickness to estimate the 45 degree path length.Table 11: Maximum Stress Intensity Factor for WLI Nozzle Considering Thermal Load Cases."ksi'in°0 S Ratio of Simplified Method to.*Load Case BIEiIF Simplified. ... ..BIE/IF Methiod 1271 Method Shutdown with 100 °F/hr 39.2 37.1 0.95 Shutdown with 300 0 F/hr 60.5 111.3 1.84 Turbine By-Pass & SRV Blow-down 55.1 1. Where Reference [27] gives tv=5.98 inches, tn=0.795 inches, xc=9.75E-6 in~in/°F File No.: 1001527.303 Revision: 0 Page 26 of 30 F0306-01RI $jj'sinww l Interity Associates, Inc.Table 12: Summary of Nozzle Stress Intensity Factors.Pressure (1,3)(1000 psig)Thermal (1,4)(450 TF shock)Thermal (1,5)(100 °F/hr)(200 °F/hr)...Nozzle j Unit 1 Unit 2 Unit 1 Unit 2 J Unit 1 Unit 2 Unit o1 Unit 2 FW J 76.6 78.9 65.3 46.8 j 11.5 12.9 j 23.1 25.8 WLI j 71.6 80.0 n/a n/a 17.4 19.9 34.8 39.9 Core DP 32.3 n/a n/a n/a 1.73 n/a 3.46 n/a 1. KI in units of ksi-in°5.2. 200 °F/hr results are scaled from 100 °F/hr assuming response is linear.3. Pressure load case results are obtained using WRC 175 methodology [7].4. Thermal shock results are obtained using BIE/IF methodology [1].5. Thermal ramp results are obtained using ASMIE XI, Non-mandatory Appendix G, Paragraph G-2214.4 methodology [81.File No.: 1001527.303 Revision: 0 Page 27 of 30 F0306-01RI 'jjuwtowau lategrify Associates, Inc." Nozzle Blend Radius Path N Nozzle Corner Crack Location Figure 1. Typical Nozzle Corner Crack Stress Extraction Path Orientation. 2 A" o t (3.1 0.2 0.), 0.4 0.5' (3,0 0.1 RA1IO OF-CRACIK SIZE T'O MDI,, OR NO)ZZ'L[ RAOiUS lilyn 0,5 0.9 Figure 2. 'WRC Bulletin 175, Figure A5-i, Estimates of Stress Intensity Factors for Flaws at a Nozzle Corner [71.File No.: 1001527.303 Revision: 0 Page 28 of 30 F0306-01R1 Figure 3. Hatch Unit 1 FW Nozzle Blend Radius Path Stress Distributions [10b].Figure 4. Hatch Unit 2 FW Nozzle Blend Radius Path Stress Distributions [lla].File No.: 1001527.302 Revision: 0 3 Page 29 of 30 SThis page contains GEH PROPRIETARY INFORMATION which has been redacted F0306-01R 1 350000 -Ivy= -316.07x3 + 3144.7x 2 -12874x + 35123 25 0 ....... I ... ..R = 0.9964 ...* .1 50 0 ...... ....... .. ..... .0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 Nozzle Path Distance, in ,,,,OOO00 psig Pressure Note: The first data point in the series is excluded for the curve fit.Figure 5. Hatch Unit 1 Core DP Nozzle 1000 psi Internal Pressure Load Case Path Stress Distribution. 70................................... ........ ... .... .. ........um Krr= 6 6.5 M Pa'rnm (t0.5 60 ~~~~~MaximumKIA= 60.6 MPa im (15_5.1 ksh ............. 50 40 ,#2o-10--.-Shutdown at 55.56 degree C/hr-e-Shutdown at 166.67 degree C/;nr o Turbne By-Pa. & SRVM Biowdown 6000 8000 10000 12000 14000 16000 18000 20000 Thne (sec)0 2000 4000 Figure 6. Plant C WLI Nozzle KIT for Three Thermal Transients 1271.File No.: 1001527.303 Revision: 0 Page 30 of 30 F0306-0I1R V Structural Integrity Associates, Inc.Y File No.: 1001527.303 Project No.: 1001527 CALCULATION PACKAGE Quality Program: [] Nuclear [] Commercial PROJECT NAME: Plant Hatch Unit l&2 P-T Curve Evaluation CONTRACT NO.: P0: SNG10018845, Rev. 0 CONTRACT: 19862, Rev. 0 CALCULATION TITLE: Feedwater, Water Level Instrument, and Core DP Nozzle Fracture Mechanics Evaluation for Hatch Unit 1 dLIu UIII L r I~uI Lullli '.uL w JVW1Up1L~11 Document Affected Project Manager Preparer(s) &Revision Pages Revision Description Approval Checker(s) Signature & Date Signatures & Date 01 -30 Initial Issue Responsible Eni~ineer D. V. Sommerville D. V. Sommerville 12/30/2011 12/30/2011 Responsible Verifiers M. Qin*12/30/2011 T. J. Griesbach 12/30/2011 Page 1 of 30 F0306-01R1 $jmSbiwbra kIturity Associates, Inc.=Table of Contents 1.0 OBJECTIVE ................................................................................... 4 2.0 METHODOLOGY............................................................................. 4 2.1 Unit and Nozzle Specific Methodology Overview ............................... 6 2.1.1 Fee dwater Nozzle ..................................................................... 6 2.1.2 Water Level Instrument Nozzle ...................................................... 7 2.1.3 Core Differential Pressure Nozzle................................................... 8 2.2 2-D FEM Correction Factor ........................................................ 8 2.3 Boundary Integral Equation / Influence Function Methodology ............... 10 2.4 WRC Bulletin 175 Methodology .............. .................................. 11 2.5 ASME XI, G-22 14.3 Methodology for Radial Thermal Gradients ............ 12 3.0 DESIGN INPUTS...................................

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

12 4.0 ASSUMPTIONS............................................................................ 13 5.0 CALCULATIONS .....................

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

14 5.1 Hatch Uniti1........................................................................ 14 5.1.1 Feedwater Nozzle .................................................................... 14 5.1.2 Water Level Instrument Nozzle...... .................................. 15 5.1.3 Core Differential Pressure Nozzle ................................................. 15 5.2 Hatch Unit 2........:................................................................ 16 5.2.1 Feedwater Nozzle .................................................................... 16 5.2.2 Water Level Instrument Nozzle............................................. ......... 17 5.3 Justification for Linear Scaling of Thermal Stress Intensity Factor Solutions 17 6.0

SUMMARY

OF RESULTS............................................................... 18

7.0 REFERENCES

............................................................................. 18 File No.: 1001527.303 Page 2 of 30 Revision: 0 F0306-01RI $jsbcrm~raI tor~ly Associates, Inc.=List of Tables Table 1 : Hatch Unit 1 FW Nozzle Blend Radius Path Stress Distributions ................... 22 Table 2: Hatch Unit 2 FW Nozzle Blend Radius Path Stress Distributions................... 23 Table 3: Hatch Unit 1 Core DP Nozzle 1000 psi Internal Pressure Load Case Path Stress Distribution................................................................ 24 Table 4: Hatch Unit 1 FW Nozzle K 1 p and KIT using the BIB/IF Methodology............... 25 Table 5: Hatch Unit 2 FW Nozzle K 1 p and KIT using the BIB/IF Methodology................ 25 Table 6: Hatch Unit 1 Core DP Nozzle K 1 p using the BIB/IF Methodology................... 25 Table 7: Hatch Unit 1 FW Nozzle K 1 p using the WRC Bulletin 175 Methodology........... 25 Table 8: Hatch Unit 1 Core DP Nozzle K 1 p using the WRC Bulletin 175 Methodology.....25 Table 9: Hatch Unit 2 FW Nozzle K 1 p using the WRC Bulletin 175 Methodology........... 26 Table 1.0: Summary of Nozzle K 1 t Results Using the BIB/IF and G-221 14.3 Methodologies................................. i............................. 26 Table 11 : Maximum Stress Intensity Factor for WLI Nozzle Considering Thermal Load Cases......................................................

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

26 Table 12: Summary of Nozzle Stress Intensity Factors ......................................... 27 List of Figures Figure 1. Typical Nozzle Corner Crack Stress Extraction Path Orientation................... 28 Figure 2. WRC Bulletin 175, Figure A5-i, Bstimates of Stress Intensity Factors for Flaws at a Nozzle Corner .................................................. ........28 Figure 3. Hatch Unit 1 FW Nozzle Blend Radius Path Stress Distributions................... 29 Figure 4. Hatch Unit 2 FW Nozzle Blend Radius Path Stress Distributions................... 29 Figure 5. Hatch Unit 1 Core DP Nozzle 1000 psi Internal Pressure Load Case Path Stress Distribution................................................................ 30 Figure 6. Plant C WLI Nozzle KIT for Three Thermal Transients .............................. 30 File No.: 1001527.303 Page 3 of 30 Revision: 0 F0306-01RI

1.0 OBJECTIVE

The objective of this calculation is to calculate the Mode I stress intensity factors, KI, for the Edwin I.Hatch Unit 1 and Unit 2 Feedwater (FW) nozzles and Water Level Instrument (WLI) nozzles, and the Unit 1 Core Differential Pressure (DP) nozzle, necessary for creation of the Pressure-Temperature (P-T)limit curves for the pressure test (Curve A), core not critical (Curve B) and core critical (Curve C)conditions for the Hatch Unit 1 and Unit 2 reactor pressure vessels (RPV). Both internal pressure and thermal transient load cases are considered.

2.0 METHODOLOGY

Consistent with the Structural Integrity Associates, Inc. (SI) Boiling Water Reactor (BWR) P-T Curve Licensing Topical Report (LTR) [1], the FW nozzle is normally taken as the limiting component in the non-beltline region of the RPV. This assumption is made because: 1. The geometric discontinuity caused by the nozzle penetration in the RPV shell causes a stress concentration which results in larger pressure induced stresses than would be calculated in the shell regions of the RPV, 2. The FW nozzle experiences more severe thermal transients than most of the other nozzles because of the feedwater injection temperature which causes larger thermal stresses than are experienced in the shell regions of the RPV, 3. Although some other nozzles can experience thermal transients which would cause thermal stresses larger than those calculated for the shell regions of the RPV and some nozzles are larger diameter than the FW nozzle, which could result in a slightly larger KIp, the combined stresses from the applied thermal and pressure loads are considered to bound all other non-beltline discontinuities. The Hatch Unit 1 and Unit 2 Adjusted Reference Temperature calculations [2] identify the WLI nozzle as contained within the beltline region of both the Unit 1 and Unit 2 RPV. Consequently, the effects of these nozzles must be considered in the beltline P-T curve development. These nozzles will cause a stress intensification in the beltline shells. Further, since the beltline region experiences a reduction in toughness caused by neutron irradiation, it is not obvious whether the FW nozzle will bound the WLI Nozzle throughout the life of Hatch Unit 1 and Unit 2. Consequently, the effects of the WLI Nozzle on the beltline P-T curves must be specifically considered. The SI P-T Curve LTR [1] addresses the bottom head penetrations by conservatively applying a stress concentration factor (SCF) of 3.0 for a hole in a flat plate to the pressure induced membrane stress in the bottom head shell and using the thermal stress intensity factor solution given in ASME XI, Non-mandatory Appendix G, Paragraph G-22 14.3 [8]; discussed below. Review of the Hatch Unit 1 general assembly drawing [30] shows that the Core DP nozzle exists in the thinnest section of the bottom head.Based on prior experience from a similar bottom head design, application of the conservative SCF=3.0 methodology to the Core DP nozzle penetration will result in a bottom head P-T curve which controls the entire RPV. Consequently, a detailed evaluation of the Core DP nozzle is performed to remove excess conservatism. File No.: 1001527.303 Page 4 of 30 Revision: 0 F0306-01R1 Consistent with 10OCFR50 Appendix G [3] the RPV P-T curves are applicable for normal operation and all anticipated operating occurrences. Consequently, all Level A and Level B (Normal and Upset)operating events defined on the RPV, FW nozzle, WLI nozzle, and Core DP nozzle thermal cycle diagrams (TCD) [4, 5, 6] are considered in selecting bounding thermal and pressure conditions for preparing P-T curves.The SI P-T Curve LTR [1 ] identifies acceptable methodologies for calculating applied pressure and thermal stress intensity factors for postulated nozzle corner flaws. These methodologies include: Pressure Load Case: 1. Welding Research Council Bulletin 175 [7]2. Boundary Integral Equation / Influence Function (BIE/IF) [1]*Thermal Transient Load Case: 1. BIB/IF [1]2. ASME XI, Non-mandatory Appendix G, Paragraph G-2214.3 [8]The WRC Bulletin 175 [7] methodology for calculating a KI for an internal pressure load case is convenient to apply since no nozzle specific finite element analysis (FEA) is necessary. The only inputs required are the nozzle and vessel geometry and the hoop stress calculated for the vessel shell, remote from discontinuities. The BIB/IF methodology is applicable to any load case provided that a third order polynomial curve fit to the applicable stress distribution is available. The appropriate path for a postulated nozzle corner crack has a path origin located at the peak stress location in the blend radius for the pressure load case and it is oriented parallel to a 45° line through the nozzle, as shown in Figure 1. This methodology requires detailed stress distributions through the nozzle blend radius which are typically obtained from a plant specific FEA.The flat plate thermal stress intensity factor solution given in the American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel (B&PV) Code Section XI, Non-mandatory Appendix G, Paragraph G-22 14.3 [8] is considered to be a conservative, simplified methodology for obtaining the KIT at the nozzle corner path resulting from a heat-up/cool-down transient of constant rate (i.e. 100 °F/hr), when the wall thickness is taken asthe path length along the 450 path identified in Figure 1. This approach is considered to be conservative because: 1. Geometric discontinuities do not intensify thermal stresses in a manner similar to stresses from mechanical loading; thus, fracture mechanics solutions which inherently consider the nozzle geometry are not necessarily required, 2. Thermal stresses increase as section thickness increases because the differential thermal strain increases with thickness. Consequently, the practice of taking a wall thickness determined by the path length of the 450 path results in a thicker wall.File No.: 1001527.303 Page 5 of 30 Revision: 0 F0306-01IRI

Integrit Associates, Inc: SI has prepared a complementary P-T Curve LTR [9] which provides simplified methods for calculating the K~p and KIT for the BWR WLI nozzles. This method is convenient since the necessary stress intensity factors can be calculated using simple geometry and material properties data without the need for a plant specific FEA.The methods identified above are used in this calculation to calculate stress intensity factors for the Hatch Unit 1 and Unit 2 FW nozzles and WLI nozzles, and the Unit 1 Core DP nozzle. The sections below identify the general methodology applied for each nozzle at each unit.2.1 Unit and Nozzle Specific Methodology Overview The specific methodologies used to calculate the Kjp and KIT for the FW nozzle, the WLI nozzle, and the Core DP nozzle are discussed separately below.2.1.1 Feedwater Nozzle Various FW nozzle fracture mechanics analyses have previously been performed for Hatch Unit 1 [10]and Unit 2 [11] to satisfy the requirements of NUREG-0619 [ 12] regarding FW nozzle rapid thermal cycling caused by leakage past the thermal sleeve seals. The load cases necessary for development of P-T curves and those required to address NUREG-06 19 are similar. Consequently, the previous evaluations performed for Hatch Units 1 and 2 will be utilized, as appropriate, for the present evaluation. The following methodology is used: Internal Pressure Load Case 1. Pressure stress distributions reported in References [10, 11] for Hatch Unit 1 and Unit 2 are taken for the 1000 psig internal pressure load case.2. Recognizing that the Reference [10, 11] evaluations were performed using a 2-D axi-symmetric finite element model (FEM) and that it is known that the stress intensification caused by the nozzle geometry is under predicted in a 2-D axi-symmetric representation of the nozzle, a correction factor must be applied to the stresses obtained from the 2-D axi-symmetric FEM. The internal pressure load case stresses, in the nozzle blend radius region, are corrected using the methodology presented in Reference [13]. Plant specific dimensions are used to calculate the correction factor.3. The BIE/IF methodology presented in the SI P-T Curve LTR [1] is used to calculate K~p by fitting a third order polynomial equation to the path stress distribution for each plant specific pressure load case. The resulting K 1 p can be linearly scaled to determine the K 1 p for various RPV internal pressures.

4. 'The methodology given in WRC Bulletin 175 [7] is used to obtain an independent confirmation of the K 1 p calculated for each unit. The resulting K 1 p can be linearly scaled to determine the K~p for various RPV internal pressures.

File No.: 1001527.303 Page 6 of 30 Revision: 0 F0306-01RI iuru hItud Associates, Inc." Thermal Transient Load Case 1. Thermal shock load case path stress distributions reported in References [10, 11] for Hatch Unit 1 and Unit 2 are taken. The FW nozzle thermal shock is the most severe Level A/B thermal transient for the FW nozzle; thus, the load case considered in References [10, 11] is appropriate for P-T curve calculations. When path stress distributions at multiple times are presented in References [10, 11 ] the bounding distribution is selected for the current evaluation.

2. The BIE/IF methodology presented in the SI P-T Curve LTR [1] is used to calculate KIT by fitting a third order polynomial equation to the path stress distribution for each plant specific thermal shock load case. The resulting KIT can be linearly scaled to determine the KIT for various shock amplitudes.

3. A KIT for a uniform 100 °F/hr and 200 "F/hr heat-up/cool-down transient is calculated using the equation given for a radial thermal gradient in ASME XI, Appendix G, Paragraph G-22 14.3 [8]4. Results from various plant specific evaluations are presented to support the use of G-22 14.3 [8]for calculating KIT for 100 "F/hr and 200 0 F/hr heat-up/cool-down transients.

2.1.2 Water

Level Instrument Nozzle Simplified methods for calculating the K~p and KIT for the WLI nozzles in General Electric designed BWRs are given in Reference [9], which is a companion LTR to the P-T curve LTR [1]. The K~p and KIT terms are calculated using Equations (8-1) and (8-2) of Reference [9], which are repeated below, for convenience, as Eq. (1) and Eq. (2): K 1 Pesue .94~ ~1000Q psig internal pressure (1)KIRa,np = ++/-t,)]- 20.715, 100 °F/hr cooldown transient (2)Where: R is the inside radius of the pressure vessel, in tv is the wall thickness of the pressure vessel, in tn is the thickness of the WLI nozzle insert near the postulated crack location, in oa is the coefficient of thermal expansion at the highest temperature in the transient, in/in/°F The units of Ki in Eq. (1) and Eq. (2) are ksi-in°5.File No.: 1001527.303 Page 7 of 30 Revision: 0 F0306-01R1 jS b-c orl Aitii~ ssociates, In.=2.1.3 Core Differential Pressure Nozzle The following methodology is used: Internal Pressure Load Case: 1. The methodology given in WRC Bulletin 175 [7] is used to calculate the K 1 p. The resulting K 1 p can be linearly scaled to determine the K~p for various RPV internal pressures.

2. Since a 2-D axi-symmetric finite element analysis of a Core DP nozzle for a plant with a similar design has previously been performed for development of P-T curves [28], and since the dimensions of the Core DP nozzle modeled [29], in the vicinity of the bottom head penetration, are identical to Hatch Unit 1 [30], the results of the previous Core DP nozzle evaluation may be used to obtain an independent benchmark of the K 1 p obtained using the WRC Bulletin 175 [7]methodology.

3. A path stress distribution, in the vicinity of the Core DP nozzle, caused by a 1000 psig internal pressure load case, is taken from Reference

[28]. The location and orientation of the path are consistent with that used for nozzle evaluations.

4. The BIE/IF methodology presented in the SI P-T Curve LTR [1] is used to calculate K 1 p by fitting a third order polynomial equation to the path stress distribution for the plant specific pressure load case.Thermal Transient Load Case: 1. Similar to the methodology given in the P-T Curve LTR [1], a KIT for a uniform 100 °F/hr and 200 0 F/hr heat-up/cool-down transient is calculated using the equation given for a radial thermal gradient in ASME XI, Appendix G, Paragraph G-22 14.3 [8]2. Results from various plant specific evaluations are presented to support the use of G-2214.3 [8]for calculating KIT- for 100 0 F/hr and 200 °F/hr heat-up/cool-down transients.

2.2 2-D FEM Correction Factor When a cylindrical nozzle intersection with a cylindrical pressure vessel is modeled using a 2-D axi-symmetric simplification, the geometry is approximated as a nozzle intersection with a spherical shell.This simplification results in a non-conservative treatment of the geometric discontinuity at the blend radius region as well as a resulting reduction in the far field membrane stress in the shell caused by the approximation of the pressure vessel as a sphere rather than a cylinder. Consequently, stress results obtained from a 2-D axi-symmetric FEM, for mechanical loads such as pressure, must be corrected before they can be used in subsequent analyses. Sommerville and Walter [13] describe a methodology for correcting stress distributions obtained from the nozzle blend radius region of 2-D axi-symmetric models. The correction factor is given in Reference [13] as: File No.: 1001527.303 Page 8 of 30 Revision: 0 F0306-01RI ~jSbircraI laturif Associatus, Inc: SCF3DHo CF = 2. 3DHo (3)SCF 2 D)_Hoop Where: SCF 3 DHoop is the stress concentration factor (SCF) defined with respect to the the hoop stress direction, for the 3D geometry.SCF 2 DHoop is the stress concentration factor for the 2D geometry.The expected SCF for the 3-D geometry can be estimated using the following equation for a circular hole in a pressurized cylinder [14]: For 0132, K(, (/) = 2.5899 +0.8002.13+/-+4.0112.132 _-1.8235-/83 + 0.3751 ./34 (4a)~j.( 2)f 2 " (4b)Where: Kt(I3) is the SCF defined with respect to the far field hoop stress, PR/t.v is the Poisson's ratio of the material, assumed equal to 0.3.r is the radius of the nozzle bore, in R is the inside radius of the pressure vessel, in tv is the wall thickness of the pressure vessel, in Kt(t3) is used for the SCF3D)_Hoop term in Eq. (3). The SCF2DHoop can be calculated from the results of the 2-D axi-symmetric FEM by calculating the SCF using the following equation: 2 ., O'totl,max (5)SCF 2-H°°P = P. R Where: P is the RPV internal pressure for the pressure load case, psi O'total,max is the largest total hoop stress in the blend radius region, psi Inserting Eq. (5) and Eq. (4a, 4b) into Eq. (3) gives the correction factor which can be used to uniformly scale the pressure load case path stress distribution from the 2-D axi-symmetric FEM. This path stress distribution can then be fit with the equation for a 3 rd order polynomial and used with the BIE/IF solution described in Reference [1], and discussed below, to obtain Kjp.File No.: 1001527.303 Page 9 of 30 Revision: 0 F0306-01R1 mIhtugril Associates, Inc.=2.3 Boundary Integral Equation / Influence Function Methodology The following discussion is excerpted from the SI P-T curve LTR [1]. Note that the equation and reference numbers in the excerpt below refer to Reference [1]: The stress intensity factors for the feedwater nozzle may be calculated using the results of a detailed finite element model of the nozzle. In some cases, such results may already be available from the governing design basis stress report for the fee dwater nozzle. The details of the finite element process are not included here," rather, the extraction of the. appropriate finite element results and their use in developing P-T limit curves is discussed. For a path through the limiting nozzle inner blend radius corner, as shown in Figure 2-7, the thermal and pressure hoop stress distributions should be extracted from the finite element model. Each of the stress distributions should befit with a third-order polynomial that reasonably fits the calculated stresses in the region of interest.The thermal stress intensity factor, K 1 t, is computed based on either of the nozzle corner solutions shown in Figure 2 -8 for a postulated 1/4t (based on the section thickness) axial defect, as follows." K 1 , = [ 0.723 C 0 , + 0.55 1 ~1)C 1 , + 0.462~Z C 2 , + 0.4 0 8 j J C 3 ,] (2.5.3-3 a)K 1 , V~[O.06C~ +0.53 C 1 + .4 4~ i~j C, +/-.39~jj ~ 3~j(2.5.3-3b) where: Kit the thermal stress intensity factor for the limiting normal/upset transient (psi inc)a = 1/4tpostulatedflaw depth (inches)t = thickness of the cross-section through the limiting nozzle inner blend radius corner, as Shown in Figure 2- 7.Co,, C 1 ,., C 2 ,, C 3 , thermal stress polynomial coefficients based on fits to finite element analysis.File No.: 1001527.303 Page 10 of 30 Revision: 0 F0306-01R1 $bSl wc~ Ilte! grit Associates, Inc.=Equation 2.5.3-3 a is based upon a Boundary Integral Equation / Influence Function (BIE/IF) solution developed for a quarter crack in an infinite quarter space.Equation 2.5. 3-3b is the average of the BIE/IF solutions developed for a quarter crack in an infinite quarter space and a semicircular crack in an infinite half space.These equations have been investigated by the NRC and Oak Ridge National Laboratory and shown to be acceptable for use in calculating the applied stress intensity factor for a corner cracked nozzle [16]. Although both solutions given above are evaluated in Reference [16], it is acknowledged by the original authors of these formulations, in the basis work used to develop the approach [171, that the twvo formulations differ very little and in fact provide KI values which differ only by approximately 5%. This can be seen by review of the coefficients used in each equation above. Consequently, either Equation 2.5.3-3 a or 2.5.3-3b may be used for any nozzle configuration in a BWR.The BIE/IF solution introduced above is applicable to any load case and any BWR nozzle; thus, it is applied both to the internal pressure and thermal transient loads considered for P-T curve development. 2.4 WRC Bulletin 175 Methodology An alternative solution for determining the applied pressure stress intensity factor, Kip, is to use the method given in Appendix 5 of WRC Bulletin .175 [7], where: K 1~ zFar~.P l~ (6 Where: K 1 p is the applied pressure stress intensity factor, psi ic P is the operating pressure, psi a is the 1/4t postulated flaw depth, inches R is the vessel inner radius, inchesis the thickness of the vessel shell, inches F(a/r,) is the shape factor given in Figure 2, where rn ri + 0.29rc, ri is the actual inner radius of nozzle, inches rc is the nozzle blend radius, inches A functional form of F(a/rn) is given in Reference [15, Pg. 11.1-17] as:= .42 .0 .46 (7)Where, Eq. (7) is applicable for 0.,07 < a/rn< 0.9.File No.: 1001527.303 Page 11 of 30 Revision: 0 F0306-0lRl $jtSinwwrul I~l~itedt Associates, Inc.*2.5 ASME XI, G-2214.3 Methodology for Radial Thermal Gradients ASME XI, Non-mandatory Appendix G, Paragraph G-22 14.3 [8] gives the following, simplified method for calculating KIT for a radial thermal gradient: Kt =-0.953.10-3**CR. tv 2 5 (8)Where: CR is the cooldown rate, 0 F/hr.tv is the thickness of the vessel shell, inches Paragraph G-22 14.3 [8] states that Eq. (8) will yield conservative results if used for cool-down rates greater than 100 °F/hr.This methodology is used to calculate the KIT in the nozzle blend radius region by using the path length along the 450 path shown in Figure 1 as the shell thickness, tv.3.0 DESIGN INPUTS The following design inputs are used for this evaluation:

• Previous NUREG-0619 evaluations:

References [10, 11]* Core DP Nozzle evaluation: Reference [28]* Thermal transient definitions for Level A/B: References [4, 5, 6]* WLI Nozzle dimensions and material: Reference [16]Hatch Unit 1 : Nozzle insert material: Inconel, SB- 166 RPV inside radius: 110.375 inches Nozzle insert thickness: 0.28 1 inches Vessel shell thickness: 5.375 inches Note." Some drawings show a shell thickness of 5.875 inches;" however, the minimum dimension given in the general arrangement drawing is used for this evaluation. Hatch Unit 2: Nozzle insert material: Inconel, SB- 166 RPV inside radius: 110.375 inches Nozzle insert thickness: 0.66 1 inches Vessel shell thickness: 5.375 inches File No.: 1001527.303 Page 12 of 30 Revision: 0 F0306-01IRI $j~biicfriru Iutugdty Associates, Inc.e* FW Nozzle dimensions: Hatch Unit 1: Nozzle bore diameter: Nozzle blend radius: RPV inside radius: Vessel shell thickness: Hatch Unit 2: Nozzle bore diameter: Nozzle blend radius: RPV inside radius: Vessel shell thickness:

• Core DP Nozzle dimensions:

Hatch Unit 1: Vessel shell thickness: Bottom head radius: Nozzle inside radius:* Material Properties: Hatch Unit 1 and Unit 2 WLI Nozzle Insert: Coefficient of thermal expansion: Poisson's ratio: Reference [ 17]6.7 inches 2.5 inches 110.375 inches 5.375 inches 6.5 inches 5.0 inches 110.375 inches 5.375 inches Reference [30]3.188 inches 110.5 inches 1.250 inches, see assumption 3.Reference [ 18]7.7x106 in/ir/°F at 550 0 F 0.3 (assumed)4.0 ASSUMPTIONS The following assumptions are used in this calculation and supported by data presented below: 1. The thermal stress intensity factor, KIT, for a 200 °F/hr thermal transient can be conservatively calculated using Eq. (8) above for vessel shells and forged vessel nozzles, where the shell thickness in the equation is taken as the vessel shell thickness when evaluating the shell, and the nozzle blend radius path length when evaluating forged nozzles. Adequacy of this assumption is demonstrated in Section 5.3 below.2. The thermal stress intensity factor, KIT, for a 200 °F/hr thermal transient can be conservatively calculated using Eq. (2) above for a WLI nozzle by scaling the KIT obtained from Eq. (2) by 2.Similarly higher heat-up/cool-down rates can be addressed by appropriate scaling factors.Adequacy of this assumption is demonstrated in Section 5.3 below.File No.: 1001527.303 Revision: 0 Page 13 of 30 F0306-01RI $b~ml inc.=3. The stress concentration effect of the Core DP penetration in the bottom head is conservatively addressed by treating the penetration as a nozzle in which the radius considered in the WRC Bulletin 175 pressure stress intensity methodology is taken as the radius of the hole in the shell rather than the ID of the Core DP penetration. This assumption is validated in Section 5.1.3 below by comparing the KIP obtained using this approach with the KIP obtained using the BIB/IF methodology.

5.0 CALCULATIONS

The calculations for the FW nozzles, WLI nozzles, and Core DP nozzle are presented for each unit, separately, below.5.1 Hatch Unit 1 The calculations for the FW, WLI, and Core DP nozzles are presented in separate sections.5.1.1 Feedwater Nozzle Table 1 presents a tabulation of the path stress distribution taken from the Unit 1 plant specific FW nozzle FEM, in the blend radius region of the nozzle. Both the 1000 psig internal pressure and 450 0 F thermal shock load case path stress distributions are taken from Reference [i10a]. The thermal transient load case path stress distribution is fit with a 3 rd order polynomial equation. In the previous evaluation [10a], the pressure load case path stress distribution was corrected using a different methodology than utilized for the present calculation; therefore, this correction factor is removed before the correction factor calculated using the methodology described in Section 2.2 is applied. The corrected pressure path stress distribution is fit with a 3 rd order polynomial equation.The Hath Unit 1 FW nozzle correction factor is calculated below: r = 6.7 in FW nozzle bore radius R = 110.38 in RPV radius adjacent to FW nozzle tv= 5.38 in RPV shell thickness adjacent to FW nozzle 3=- 0.177 -Eq. (4)Kt 2.85 -SCF for hole in cylinder, Eq. (4)SCF 2 0= 2.38 -SCF from 2-D axi-symmetric FEA, Eq. (5)CF = 2.39 -Correction factor, Eq. (3)Figure 3 is a plot of the pressure and thermal shock path stress distribution in the Hatch Unit 1 FW nozzle blend radius region with the polynomial curve fit equations and correlation coefficients shown.Table 4 summarizes the polynomial coefficients for each load case and presents the Kip and KIT for the Hatch Unit 1 FW nozzle.Table 7 presents the K 1 p calculated using the WRC Bulletin 175 methodology [7].File No.: 1001527.303 Page 14 of 30 Revision: 0 F0306-0 IRI $b~o~grl Associates, Inc.=Recognizing that the BIB/IF methodology has been shown to provide conservative estimates of the stress intensity factors for both pressure and thermal load cases [19, 20] and that both the WRC Bulletin 175[7] and BIE/IF [1] methodologies are accepted methods for calculating the KI from the pressure load case, the lower KI is used for this evaluation. Reference [10Ob] identifies that the extended power uprate conditions result in an increase in the RPV dome pressure of 50 psig (1000 psig to 1050 psig) and an increase in the FW fluid temperature of 6 °F (450 °F to 456 'F). The increase in dome pressure is accommodated by linear scaling during preparation of the P-T curves in a subsequent calculation. The increase in the FW fluid temperature is a change of less than 1.5%. The work documented in References [19, 20] shows that the BIB/IF methodology is significantly conservative (on the order of 30-50% when compared to the RMS K 1 calculated from a 3-D finite element fracture mechanics evaluation) for the 1/4 thickness flaws required for P-T curves;consequently, the KIT calculated for the 450 °F shock is not further increased in this evaluation to account for the small increase in FW temperature.

5.1.2 Water

Level Instrument Nozzle Using Eq. (1) and (2) and considering Hatch Unit 1 WLI nozzle dimensions and material properties, the K 1 p and KIT are calculated below: tv= 5.38 in RPV shell thickness adjacent to WLI nozzle R = 110.38 in RPV radius adjacent to WLI nozzle tn= 0.28 1 in WLI nozzle thickness cx 7.70E-06 in/in/OF Thermal expansion coefficient for nozzle material, Inconel, 550 'F K = 71.6 ksi-in 0.5 1000 psi pressure stress intensity factor, Eq. (1)KIT = 17.4 ksi-in 0.5 100 'F/hr cool-down transient stress intensity factor, Eq. (2)KIT= 34.8ksi-n 0 5 200 'F/hr cool-down transient stress intensity factor, Eq. (2) scaled by a KXT = 34.8ksi'n°'S factor of 2 5.1.3 Core Difjferential Pressure Nozzle The K 1 p calculated using the WRC Bulletin 175 methodology [7] is shown in Table 8.The pressure path stress distribution is extracted from Reference [28]. The stress distribution is tabulated in Table 3 and plotted in Figure 5. Because of the orientation of the path and the configuration of the nozzle, which contains a tube which penetrates the RPV, the path stress distribution exhibits a lower stress at the path origin than slightly inboard along the path. Consequently, the 3 rd order polynomial equation is fit to the stress distribution, after omitting the first point along the curve. The curve fit and equation are shown in Figure 5. The resulting polynomial coefficients and K 1 p obtained using the BIB/IF methodology are listed in Table 6.File No.: 1001527.303 Page 15 of 30 Revision: 0 F0306-01iRI $j~bvg~aunIlatgrily ssoca~tes, Inc.'The K~p calculated using the BIE/IF methodology [1] benchmarks well against the K 1 p calculated using the WRC Bulletin 175 methodology [7]. Recognizing that the BIE/IF methodology has been shown to provide conservative estimates of the stress intensity factors for both pressure and thermal load cases[ 19, 20] and that both the WRC Bulletin 175 [7] and BIE/IF [1 ] methodologies are accepted methods for calculating the K 1 from the pressure load case, the lower KI is used for this evaluation. The KIT term is calculated using Eq. (8) and is presented in Table 12 for heat-up/cool-down rates of 100 0 F/hr and 200 0 F/hr. Section 5.3 provides justification for using the ASME XI, Non-mandatory Appendix G, Paragraph G-2214.3 [8] methodology for heat-up/cool-down rates greater than 100 0 F/hr.5.2 Hatch Unit 2 The calculations for the FW and WLJ nozzles are presented in separate sections.5.2.1 Feedwater Nozzle Table 2 presents a tabulation of the path stress distribution taken from the Unit 2 plant specific FW nozzle FEM, in the blend radius region of the nozzle. Both the internal pressure and thermal shock load case path stress distributions are taken from Reference [11 la]. The thermal transient load case path stress distribution is fit with a 3r order polynomial equation. In the previous evaluation [1 la], the pressure load case path stress distribution was corrected using a different methodology than utilized for the present calculation; therefore, this correction factor is removed before the correction factor calculated using the methodology described inSectioh 2.2 is applied. The corrected pressure path stress distribution is fit with a 3 rd order polynomial equation.The Hath Unit 2 FW nozzle correction factor is calculated below: r = 6.5 in FW nozzle bore radius R = 110.38 in RPV radius adjacent to FW nozzle t-- 5.38 in R!PV shell thickness adjacent to FW nozzle 13= 0.172 -Eq. (4)Kt 2.84 -SCF for hole in cylinder, Eq. (4)SCF 2 D = 2.05 -SCF from 2-D axi-symmetric FEA, Eq. (5)CF = 2.77 -Correction factor, Eq. (3)Figure 4 is a plot of the pressure and thermal shock path stress distribution in the Hatch Unit 2 FW nozzle blend radius region with the polynomial curve fit equations and correlation coefficients shown.Table 5 summarizes the polynomial coefficients for each load case and presents the K 1 p and KIT for the Hatch Unit 2 FW nozzle.Table 9 presents the K 1 p calculated using the WRC BUlletin 175 [7] methodology. File No.: 1001527.303 Page 16 of 30 Revision: 0 F0306-01R1

hi/t egd/ty Associateus, Inc." Recognizing that the BIE/IF methodology has been shown to provide conservative estimates of the stress intensity factors [19, 20] and that both the WRC Bulletin 175 [7] and BIE/IF [1] methodologies are accepted methods for calculating the KI from the pressure load case, the lower K 1 is used for this evaluation. Reference [1 lc] identifies that the extended power uprate conditions result in an increase in the RPV dome pressure of 50 psig (1000 psig to 1050 psig) and an increase in the FW fluid temperature of 6 0 F (delta T from 450 0 F to 456 0 F). The increase in dome pressure is accommodated by linear scaling during preparation of the P-T curves in a subsequent calculation. The increase in the FW fluid temperature is a change of less than 1.5%. The work documented in References [19, 20] shows that the BIE/IF methodology is significantly conservative (on the order of 3 0-50% when compared to the RMS K 1 calculated from a 3-D finite element fracture mechanics evaluation) for the 1/4 thickness flaws required for P-T curves; consequently, the KIT calculated for the 450 0 F shock is not further increased in this evaluation to account for the small increase in FW temperature.

5.2.2 Water

Level Instfrument Nozzle Using Eq. (1) and (2) and considering Hatch Unit 2 WLI nozzle dimensions and material properties, the K~p and KIT are calculated below: tv= 5.38 in RPV shell thickness adjacent to W/LI nozzle R = 110.38 in RPV radius 'adjacent to WLI nozzle tn= 0.66 1 in WLI nozzle thickness ct= 7.70E-06 in/in/°F Thermal expansion coefficient for nozzle material, Inconel, 550 0 F Kw= 80.0 ksi-in°'s 1000 psi pressure stress intensity factor, Eq. (1)KI 1 = 19.9 ksi-in°'s 100 °F/hr cooldown transient stress intensity factor, Eq. (2)200 °F/hr cooldown transient stress intensity factor, Eq. (2) scaled by a factor KIT = 39.9 ksi-in°'s of 2.0 5.3 Justification for Linear Scaling of Thermal Stress Intensity Factor Solutions Data is presented in this section which supports the Assumptions 1 and 2 identified in Section 4.0.Table 10 summarizes the KIT calculated for four nozzles from three separate BWRs. Both a 100 0 F/hr and a 300 0 F/hr cool-down rate were evaluated. Both the BIE/IF methodology [1] and the methodology given in ASME XI, Non-mandatory Appendix G, Paragraph G-22 14.3 [8] were used to calculate KIT.The results confirm that the simplified method given in Paragraph G-22 14.3 [8] may be used to obtain a KIT. Further, as suggested in WRC 175 [7] and G-2214.3 [8] use of this methodology is expected to be increasingly conservative for cool-down rates larger than 100 0 F/hr. This trend is seen in the two KIT results Obtained for Plant .C at 100 0 F/hr and 300 0 F/hr, where it is shown that the KIT calculated using the flat plate solution exceeds that calculated using the BIB/IF methodology by a ratio of 2.34 for a 100 0 F/hr ramp rate and 2.71 for ai 300 0 F/hr ramp rate. Since the flat plate solution given in the ASME Code[8] is obtained from a quasi-steady state temperature distribution through the wall thickness it is anticipated that the solution becomes increasingly conservative for faster ramp rates since there is File No.: 1001527.303 Page 17 of 30 Revision: 0 F0306-01R1 $SbwouraI hIterf Associates, Inc." insufficient time for a quasi-steady state temperature distribution to develop through-wall since the vessel cools from approximately 550 0 F to 100 °F; for faster ramp rates the temperature ramp ends before a quasi-steady state thermal distribution can develop.Table 11 presents the KIT calculated using the BIE/IF methodology and using the simplified methodology given in the WLI Nozzle LTR [9] which was developed for a 100 0 F/hr cool-down transient. Considering that the stress analysis and fracture mechanics methodologies are both linear elastic methods, the results should be scalable. It is recognized that at faster ramp rates the vessel will not be able to develop the quasi-steady state temperature distribution possible for slower ramp rates;therefore, the methodology given in Reference [9] is expected to yield conservative results for ramp rates faster than 100 °F/hr. The results in Table 11 and Figure 6 show a trend consistent with expectations. Consequently, the methodology given in Reference [9] for the 100 °F/hr ramp rate may be used to obtain a conservative KIT value for higher ramp rates.6.0

SUMMARY

OF RESULTS Table 12 summarizes the KI values calculated for the FW, WLI, and Core DP nozzles at Hatch Unit 1 and Unit 2 for the internal pressure load case, a 450 0 F thermal shock, a 100 °F/hr, and a 200 °F/hr cool-down transient. These values will be used in subsequent calculations to prepare P-T curves for both Hatch Units.

7.0 REFERENCES

1. Sommerville, D.V., "Pressure-Temperature Limits Report Methodology for Boiling Water Reactors," SIR-05-044, Rev. 1, June 2011.2. Adjusted Reference Temperature Calculations:

a. Sommerville, D.V., "Hatch Unit 1 RPV Material Summary and ART Calculation," SI Calculation No. 1001527.301, Rev. 0.b. Sommerville, D.V., "Hatch Unit 1 RPV Material Summary and ART Calculation," SI Calculation No. 1001527.302, Rev. 0.3. Title 10 Code of Federal Regulations Part 50, Appendix G, "Fracture Toughness Requirements." 4. Reactor Pressure Vessel Thermal Cycle Diagrams: a. SNOC Dwg. S 15025, GE Dwg. 729E762, "Reactor Thermal Cycles," SI File No. GPCO-31 Q-209.b. SNOC Dwg. S-41615, GE Dwg. 761E246, Sht. 1, "Reactor Vessel Thermal Cycles (Including Black Start)," SI File No. 1001527.211.

c. SNOC Dwg. S-416 16, GE Dwg. 761E246, Sht. 2, "Reactor Vessel Thermal Cycles (Including Black Start)," SI File No. 1001527.211.

File No.: 1001527.303 Page 18 of 30 Revision: 0 F0306-01R1

hItugdl Associates, Inc." 5. Feedwater Nozzle Thermal Cycle Diagrams: a. GE Dwg. 135B9990, "Nozzle Thermal Cycles (Feedwater)," SI File No. 1001527.211.

b. SNOC Dwg. $26421, GE Dwg. 158B8369, Sht. 4, Rev. 2, "Nozzle Thermal Cycles -Including Black Start (Feedwater

-Normal & Upset Conditions)," SI File No. 1001527.211.

c. SNOC Dwg. S26422, GE Dwg. 158B8369, Sht. 5, Rev. 2, "Nozzle Thermal Cycles -Including Black Start (Feedwater-Emergency

& Fault Conditions)," SI File No. 1001527.211.

6. Water Level Instrument Nozzle Thermal Cycle Diagrams: a. GE Dwg. 135B9990, Sht. 7, Rev. 0, "Nozzle Thermal Cycles (Instrumentation

& Core Diff.Press & Liquid Control)," SI File No. 1001527.211.

b. SNOC Dwg. S26426, GE Dwg. 1 58B 8369, Sht. 9, Rev. 2, "Nozzle Thermal Cycles -Including Black Start (Instrumentation

& Core Diff. Press & Liquid C," SI File No. 1001527.211.

7. PVRC Recommendations on Toughness Requirements for Ferritic Materials, WRC Bulletin 175, August 1972.8. American Society of Mechanical Engineers, Boiler and Pressure Vessel Code, Section XI, Rules for Inservice Inspection of Nuclear Power Plant Components, Non-mandatory Appendix G, "Fracture Toughness Criteria for Protection Against Failure," 2001 Ed. through 2003 Addenda.9. Sommerville, D. V., "Linear Elastic Fracture Mechanics Evaluation of General Electric Boiling Water Reactor Water Level Instrument Nozzles for Pressure-Temperature Curve Evaluations," SI Report 0900876.401, Rev. 0, June 2011.10. Hatch Unit 1 NUREG-0619 Evaluations:

a. Liffengren, D. J., et al., "Edwin I. Hatch Nuclear Power Station, Unit 1 Feedwater Nozzle Fracture Mechanics Analysis to Show Compliance with N UREG-06 19," NEDE-30238, DRF-137-0010, August 1983, General Electric Company. SI File No. 1001527.210.

GE Proprietary Information.

b. Bothne, D., "Power Uprate Evaluation Report for Edwin I. Hatch Unit 1, Feedwater Nozzle NUREG-06 19 Fracture Mechanics Analysis for Extended Power Uprate Conditions," GE-NE-B13-01869-065-01, July 1997, General Electric Company. SI File No. 1001527.210.

GE Proprietary Information. File No.: 1001527.303 Page 19 of 30 Revision: 0 F0306-O01RI

it Associates, 11. Hatch Unit 2 NUREG-0619 Evaluations:

a. Liffengren, D. J., et al., "Edwin I. Hatch Nuclear Power Station, Unit 2 Feedwater Nozzle Fracture Mechanics Analysis to Show Compliance with NUREG-0619," NEDC-30256, DRF-137-0010, August 1983, General Electric Company. SI File No. 1001527.210.

GE Proprietary Information.

b. Stevens, G. L., "Updated Feedwater Nozzle Fracture Mechanics Analysis for Edwin I. Hatch Nuclear Power Station Unit 2," GE-NE-523-95-0991, Rev. 0, DRF B 13-01524, September 1991, General Electric Company. SI File No. 1001527.210.

c. Bothne, D., "Power Uprate Evaluation Report for Edwin I. Hatch Unit 2, Feedwater Nozzle NUREG-0619 Fracture Mechanics Analysis for Extended Power Uprate Conditions," GE-NE-B13-01869-065-02, July 1997, General Electric Company. SI File No. 1001527.210.

GE Proprietary Information.

12. BWR Feedwater Nozzle and Control Rod Drive Return Line Nozzle Cracking, NUREG-06 19, November 1980, Nuclear Regulatory Commission.

13. Sommerville, D., Walter, M., "An Investigation into the Effects of Modeling Cylindrical Nozzle to Cylindrical Vessel Intersections Using 2D Axisymmetric Finite Element models and a Proposed Method for Correcting the Results," ASME PVP201 1-57001, Proceedings of the 2011 ASME Pressure Vessel and Piping Division Conference.

14. Pilkey, W.D., Pilkey, D.F., Peterson's Stress Concentration Factors, 3 rd. Ed., John Wiley & Sons, 2008.15. Zahoor, A., "Ductile Fracture Handbook," EPRI Report NP-6301, Volume 3. January 1991.16. Water Level Instrument Nozzle Drawings: a. SNOC Sketch 1-BE-i, Rev. 0, "Nil, N12, and N16 Instrumentation Nozzle Detail," SI File No. 1001527.208.

b. SNOC Sketch 2-BE-2, Rev. 1, "2N12 and 2N16 Instrumentation Nozzle Detail," SI File No.1001527.209.

17. Feedwater Nozzle Drawings: a. SNOC Dwg. SX18921, "Reactor Vessel Feedwater Nozzle As Built," SI File No.100 1527.208 b. SNOC Sketch 2-BF-4, Rev. 1, "2N4 Nozzle Assembly (Feedwater)," SI File No. GPCO-31Q-208.18. American Society of Mechanical Engineers, Boiler and Pressure Vessel Code, Section II, Part D, Materials, 2001 Ed. through 2003 Addenda.19. Yin, S., Bass, B. R., Stevens, G. L., Stress and Fracture Mechanics Analyses of Boiling Water Reactor and Pressurized Water Reactor Pressure Vessel Nozzles, ORNL/TM--2010/246, December 2010.File No.: 1001527.303 Page 20 of 30 Revision:

0 F0306-O01R1 rai lte gril Associates, Inc;" 20. Sommerville, D.V., Qin, M., Houston, E., "An Investigation of the Adequacy of a Simplified Boundary Integral Equation / Influence Function Equation Linear Elastic Fracture Mechanics Solution for Nozzle Corner Cracking," ASME PVP201 1-57742, Proceedings of the 2011 ASME Pressure Vessel and Piping Division Conference.

21. SI Calculation Package 1100445.302, Rev. 0.22. SI Calculation Package 1100445.303, Rev. 0.23. SI Calculation Package NPPD-13Q-302, Rev. 1.24. SI Calculation Package 1000847.302, Rev. 0.25. SI Calculation Package 1000847.303, Rev. 2.26. SI Calculation Package 1100151.302, Rev. 0.27. S1 Calculation Package 1100151.303, Rev. 0.28. SI Calculation Package 1100445.304, Rev. 0.29. Combustion Engineering Drawing 232-242, Nozzle Details, SI File No. 1100445.204.

30. Core DP Nozzle Drawings: a. SNOC Dwg. S-15227A, Combustion Engineering Drawing 234-244-5, Nozzle Details for 218" I.D. BWR, SI File No. 1001527.208.

b. SNOC Sketch 1-BE-2, Rev. 1, "N10 Standby Liquid Control & Core Differential Pressure Nozzle Detail," SI File No. 100 1527.208.c. SNOC Dwg. S15523, Combustion Engineering Drawing 234-270, Rev. 3, "General.Arrangement Elevation for 218" ID BWR," SI File No. 100 1527.208.File No.: 1001527.303 Revision:

0 Page 21 of3O0 F0306-01R1 $Sbwcbuu hlgl Assadate .°Table 1: Hatch Unit 1 FW Nozzle Blend Radius Path Stress Distributions. 1000 psig Internal Pressure Load Case Thermal Load Case 450 "F shock Path Distance, °h) Oh(2' 3 o, Gb, in. pipsi psi psi 0.000 f.l }.1 24461 58463 Di Di 0.075 ]} 24052 57486 II }I 0.225 1 /] 23283 55649 i }J-0.400 ft }1 22437 53626 1}0.600 It Di 21527 51452 U1 Di 0.850 U /] 20466 48916 [ }.i.1.150 }1 19262 46037 Di }1 1.500 J} 18081 43215 ft }1 1.960 1.1 }1 16848 40268 Ii 1.2.476 ft }1 15206 36344 Di Di 3.052 I.I ]J 13724 32802 Ui Di 3.628 II }i 12372 29571 Di }}4.204 II 11110 26554 IL Di 4.779 Di }) 9898 23657 lit Di 5.355 Di Di 8692 20774 sD 5.93 1 Di Di 7439 17780 ft }6.507 Di 1) 6073 14515 UJ Di 6.811 Di }i 5384 12868 ft Di Notes: 1. From Reference [10a], with CF=1.6557.

2. From Reference

[l0a], with CF=1.6557 removed.3. From Reference [10a], with CF as given in PVP201 1-5700 1 [13].File No.: 1001527.30, Revision: 0 3 Page 22 of 30 This page contains GEH PROPRIETARY INFORMATION which has been redacted F0306-01IRI mnssa t whlt A~ s A,.Table 2: Hatch Unit 2 FW Nozzle Blend Radius Path Stress Distributions. 1000 psig Internal Pressure Load Case Thermal Load Case 450 OF shock Path in. pipsi psi psi 0.000 U1 } 21010 58242 L1 }IL.0.075 LI }) 20651 57247 Li IL 0.225 LI }i 19970 55359 H I 0.400 I( }L 19213 53261 fL IL 0.600 Li II 18403 51015 Li IL 0.850 Li /) 17440 48345 LI II.1.150 IL 16452 45607 Li IL 1.500 LI IL 15206 42153 Li IL 1.954 Li IL 14233 39455 Li IL 2.693 LI IL 11909 33013 Li }L 3.284 LI II 10634 29479 {L Li 3.874 Zl H9468 26245 IL 4.465 Li II 8365 23188 {L }L 5.056 Li I}. 7289 20206 LI IL 5.647 LiII} 6194 17170 Li IL 6.247 Li IL 5029 13941 Li IL 6.838 LI II! 3718 10307 Li IL 7.127 Li IL 3040 8427 LI IL Notes: 1. From Reference 2. From Reference 3. From Reference[1 la], with CF=1.5987. [1 la], with CF=1.5987 removed.[1 la], with CF as given in PVP201 1-5700 1 [13].File No.: 1001527.303 Revision: 0 Page 23 of 30 This page contains GEH PROPRIETARY INFORMATION I which has been redacted F0306-01RI $b~jSlcrral hIt Writ Associates, Inc.=Table 3: Hatch Unit 1 Core DP Nozzle 1000 psi Internal Pressure Load Case Path Stress Distribution [281.Path Distance, ah, in. psi 0.000 26186 0.225 31610 0.450 30525 0.675 28449 0.901 26377 1.126 23518 1.351 22272 1.576 21189 1.801 20222 2.026 19351 2.251 18559 2.476 17831 2.702 17156 2.927 16522 3.152 15916 3.377 15327 3.602 14746 3.827 14169 4.052 13594 4.277 12838 4.503 12045 File No.: 1001527.303 Revision: 0 Page 24 of 30 F0306-01R1 8 q trailafugri Associates, mnc: Table 4: Hatch Unit 1 FW Nozzle K 1 p and KIT using the BIE/IF Methodology. Pressure JThermal A0 58242 51488 Al -11913 -15980'A2 1484.2 2002.4 A3 -105.18 -127.21 K 1 81.1 65.3 ksi-in°'5 Table 5: Hatch Unit 2 FW Nozzle K 1 p and KIT using the BIE/IF Methodology. Pressure II Thermal A0 58136 48077 Al -12560 -28076 A2 1522.5 4948.5 A3 -103.53 -598.8 K 1 81.3 46.8' ksi-in°'5 Table 6: Hatch Unit 1 Core DP Nozzle Kwp using the BIE/IF Methodology. Pressure A0 35123 Al -12874 A2 3144.7 A3 ,316.07 KI 38.9 ksi-inO.S Table 7: Hatch Unit 1 FW Nozzle K 1 p using the WRC Bulletin 175 Methodology. F(a/rn) 1.61 -PR 1/t(7ta)0 5 47.49 ksi-in°'5 KI 76.6 ksi-in°'s Table 8: Hatch Unit 1 Core DP Nozzle K~p using the WRC Bulletin 175 Methodology. F(a/rn) j0.99 -PRi/t(ira)° 5 32.58 ksi-in°'5 K______ I 32.3 ksi-in°'s File No.: 1001527.303 Page 25 of 30 Revision: 0 F0306-01R1

Assaciatus, Irnc=Table 9: Hatch Unit 2 FW Nozzle K~p using the WRC Bulletin 175 Methodology. F(a/r,) 1.62 -PRi/t~ta)0 5 48.58 ksi-in°5 K~p78.9 ksi-in 0'5 Table 10: Summary of Nozzle KIT Results Using the BIE/IF and G-22114.3 Methodologies. NozzleKIKT Plant ID Tp()Path Length, (BIE/IF Solution), (Flat Plate Soln.), Ratio Reference Te 2 in ksi-in°'5 ksi-in°'5 Plant A (3) FW 7.60 11.1 15.2 1.37 [21], [22], [23]Plant B FW 7.73 7.0 15.8 2.26 [24],[25]Plant B RI 9.29 9.4 25.1 2.67 [25]Plant C (1) FW 8.66 23.3 63.1 2.71 [26]Plant C FW 8.66 9.0 21.0 2.34 [27]1.2.3.300 *F/hr cooldown transient. FW is Feedwater nozzle, RI is Recirculation Inlet nozzle.Path length taken as 205 *RPV shell thickness to estimate the 45 degree path length.Table 11: Maximum Stress Intensity Factor for WLI Nozzle Considering Thermal Load Cases."ksi'in°0 S Ratio of Simplified Method to.*Load Case BIEiIF Simplified. ... ..BIE/IF Methiod 1271 Method Shutdown with 100 °F/hr 39.2 37.1 0.95 Shutdown with 300 0 F/hr 60.5 111.3 1.84 Turbine By-Pass & SRV Blow-down 55.1 1. Where Reference [27] gives tv=5.98 inches, tn=0.795 inches, xc=9.75E-6 in~in/°F File No.: 1001527.303 Revision: 0 Page 26 of 30 F0306-01RI $jj'sinww l Interity Associates, Inc.Table 12: Summary of Nozzle Stress Intensity Factors.Pressure (1,3)(1000 psig)Thermal (1,4)(450 TF shock)Thermal (1,5)(100 °F/hr)(200 °F/hr)...Nozzle j Unit 1 Unit 2 Unit 1 Unit 2 J Unit 1 Unit 2 Unit o1 Unit 2 FW J 76.6 78.9 65.3 46.8 j 11.5 12.9 j 23.1 25.8 WLI j 71.6 80.0 n/a n/a 17.4 19.9 34.8 39.9 Core DP 32.3 n/a n/a n/a 1.73 n/a 3.46 n/a 1. KI in units of ksi-in°5.2. 200 °F/hr results are scaled from 100 °F/hr assuming response is linear.3. Pressure load case results are obtained using WRC 175 methodology [7].4. Thermal shock results are obtained using BIE/IF methodology [1].5. Thermal ramp results are obtained using ASMIE XI, Non-mandatory Appendix G, Paragraph G-2214.4 methodology [81.File No.: 1001527.303 Revision: 0 Page 27 of 30 F0306-01RI 'jjuwtowau lategrify Associates, Inc." Nozzle Blend Radius Path N Nozzle Corner Crack Location Figure 1. Typical Nozzle Corner Crack Stress Extraction Path Orientation. 2 A" o t (3.1 0.2 0.), 0.4 0.5' (3,0 0.1 RA1IO OF-CRACIK SIZE T'O MDI,, OR NO)ZZ'L[ RAOiUS lilyn 0,5 0.9 Figure 2. 'WRC Bulletin 175, Figure A5-i, Estimates of Stress Intensity Factors for Flaws at a Nozzle Corner [71.File No.: 1001527.303 Revision: 0 Page 28 of 30 F0306-01R1 Figure 3. Hatch Unit 1 FW Nozzle Blend Radius Path Stress Distributions [10b].Figure 4. Hatch Unit 2 FW Nozzle Blend Radius Path Stress Distributions [lla].File No.: 1001527.302 Revision: 0 3 Page 29 of 30 SThis page contains GEH PROPRIETARY INFORMATION which has been redacted F0306-01R 1 350000 -Ivy= -316.07x3 + 3144.7x 2 -12874x + 35123 25 0 ....... I ... ..R = 0.9964 ...* .1 50 0 ...... ....... .. ..... .0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 Nozzle Path Distance, in ,,,,OOO00 psig Pressure Note: The first data point in the series is excluded for the curve fit.Figure 5. Hatch Unit 1 Core DP Nozzle 1000 psi Internal Pressure Load Case Path Stress Distribution. 70................................... ........ ... .... .. ........um Krr= 6 6.5 M Pa'rnm (t0.5 60 ~~~~~MaximumKIA= 60.6 MPa im (15_5.1 ksh ............. 50 40 ,#2o-10--.-Shutdown at 55.56 degree C/hr-e-Shutdown at 166.67 degree C/;nr o Turbne By-Pa. & SRVM Biowdown 6000 8000 10000 12000 14000 16000 18000 20000 Thne (sec)0 2000 4000 Figure 6. Plant C WLI Nozzle KIT for Three Thermal Transients 1271.File No.: 1001527.303 Revision: 0 Page 30 of 30 F0306-0I1R}}