ML20137C547

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Rev 1 to GE-NE-B13-01869-02, Internal Core Spray Line Flaw Evaluation for Pilgrim Nuclear Power Station
ML20137C547
Person / Time
Site: Pilgrim
Issue date: 03/18/1997
From: Bothne D, Methta H
GENERAL ELECTRIC CO.
To:
Shared Package
ML20137C531 List:
References
GE-NE-B13-01869, GE-NE-B13-01869-02, GE-NE-B13-1869, GE-NE-B13-1869-2, NUDOCS 9703250038
Download: ML20137C547 (25)
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& GE Nuclear Energy Technical and Modification Services GE-NE-B13-01869-02 Rev i March 1997  ;

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Internal Core Spray Line Flaw Evaluation for ,

1 Pilgrim Nuclear Power Station Prepared by: / NN

'6orothy>(Botline, Engineer Engineering and Licensing Consulting Services Verified by: M

'H. S. Mehia, Project Manager Engineering and Licensing Consulting Services Approved by:

2' $ '

'b H. S. Mehta. Project ManagdI Engineering and Licensing Cnnsulting Services 9703250038 970318 PDR G ADOCK 05000293 4-

, PDR; .

& GE Nuclear Ersergy IMPORTANT NOTICE REG ARDING CONTENTS OF THIS REPORT Please Read Carefully The only undertakings of the General Electric Company (GE) respecting information in this document are contained in the contract between Boston Edison Company (BECo), (Pilgrim) and GE, and nothing contained in this document shall be construed as changing the contract. The use of this information by anyone other than BECo, or for any purpose other than that for which it is intended, is not authorized; and with respect to any )

unauthorized use, GE makes no representation or warranty, express or implied, and assumes no liability as to i the completeness, accuracy, or usefulness of information contained in this document, or that its use may infringe privately owned rights.

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4 Table of Contents

-- --.~ 2 TABLE OF CONTENTS -. - ~ -

_ 3 LIST OF TABLES - - __

l LIST OF FIGURES .. _ -. . _. - .3 i.e iNTRODuCriON _. _ . _ _._._._.~ . 4 ,

I 4 l 2.0 METH0DS. .. _. . . ~

1 3.0 ASSUMPTlGNS .

~ . - - . _ _ . - _ _. . 4

_ - .. - 5 4.0 DESIGN INPUTS _ ~. . -.

-5 4.1 DEADWEIGitT(DW)- .

-S 4.2 DYNAMIC INERTIA - I

. 6 4.3 DYNAMIC ANC110R DisrLACEMENT.. . 1 6

4.4 FLUID DRAG x

. -6 4.5 CORE SPRAY INJECTION LOADING (CSIN)-

... . -7 4.6 TilERMAL LOADS ...

_ 8 5.0 LOAD COMBINATIONS & STRESS LEVELS -

. . . . . -8 5.1 LOADCOMBINATIONS

. .. .9 5.2 CALCULARD SDtESS LEVELS - .

_ - _12 6.0 FRACTURE MECHANICS EVALUATION _

. 12 6.1 LIMIT LOAD METliODOLOGY (CIRCUMFERENTI AL CRACKS) .. ..

-13 6.2 ALLOWABLE FLAW LENGTH CALCULATION (CIRCUMFERENTIAL)..

18 6.3 CRACK GROWTH EVALUATION.. . . . .

.. I S j 6.4 LIMIT LOAD METilODOLOGY & ALLOWABLE VALUE FOR AXtAL CRACKS -

..-.~. .- ...- -. ...~ .. .. - .-. 19 7.0 LEAKAG E EVA LUATION ... . ~.

. .. 19 7.1 LEAK RAE CALCULATION METHODOLOGY .

19 7.2 OVERALL LEAK RATE CALCULATION .

8.0 SU M M A RY & CO NCLUSI O N S.. ~.~.. __. . .~ ..._ ~..~ ... . .-. . . ~ ... . -.. I 9 R E FE R EN C ES . . .. ... . . . ...... . - . . -. ... ..- ... ... -.- ........~..~...-.....-~....~.21 2

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List of Tables TABLE I RELATIVE DYNAMIC ANCllOR MOTIONS 6 TABLE 2 THERMAL DISPLACEMENTS FOR TRANSIENT CONDITIONS FOR THE UPPER CORE SPRAY LINE 8 TABLE 3 THERMAL DISPLACEMENTS FOR TRANSIENT CONDITIONS FOR DIE LOWER CORE SPRAY LINE 8 TABLE 4 PRIMARY MEMBRANE, PRIMARY BENDING AND SECONDARY STRESSES FOR THE LOOP A PIPING 10 TABLE 5 PRIMARY MEMBRANE, PRIMARY BENDING ANDSECONDARY STRESSES FOR THE LOOP D PIPING 11 l

TABLE 6 LOOP A CORE SPRAY LINE ALLOWABLE FLAW LENGTHS FOR ONE CYCLE 14 TABLE 7 LOOP B CORE SPRAY LINE ALLnWABf C FLAW LENGTHS FOI;ONE CYCLE 15 TABLE 8 LOOP A CORE SPRAY LINE ALLOWABLE FLAW LENGTHS FOR TWO CYCLES 16 17 TABLE 9 LOOP B CORE SPRAY LINE ALLOWABLE FLAW LENC':US FOR TWO CYCLES 1

List of Figures 22 I FIGURE I PILGRIM CORE SPRAY "A" LOOP PIPING 23 FIGURE 2 PILGRIM CORE SPRAY "B" LOOP PIPING 24 FIGURE 3 PILGRIM CORE SPRAY NOZZLE 25 FIGURE 4 ANSYS MODEL OF THE PILGRIM CORE SPRAY LINE 27 FIGURE 5 STRESS DISTRIBUTION IN A CRACKED PIPE AT LIMIT LOAD i

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Q GE Nuclear Energy 1.0 Introduction Cracking in the core spray line intemal piping at several Boiling Water Reactor (BWR) plants has been recently observed. Cracking is believed to be i:.tergranular stress corrosion cracking (IGSCC) typically in the vicinity of circumferential welds. In addition, there have been cases of cracking at creviced areas. De availability of a flaw evaluation handbook prior to actual inspection of the line can help reduce any potential outage delay in order to disposition cracking, if any.

He objective of this report is to document the results of the flaw evaluation of the core spray intemal piping at Pilgrim Nuclear Power Station. He outcome of flaw evaluation is a set of allowable flaw lengths at key locations in the core spray internal piping system. The evaluation also includes leak rate calculations for postulated through-wall indications which can be used in an evaluation to determine if the conclusions predicted by current LOCA analyses remain valid with the postulated leakage. De overall package constitutes a flaw handbook that could be used to disposition indications that might be detected during the inspection of core spray internal piping.

2.0 Methods his section presents the methodology and procedure used in performing the core spray line analysis.

Following are the steps used in the analysis.

1. Create finite element models of the core spray lines using the ANSYS [ Reference 1] computer program.

Determine the forces and moments resulting from the loadings identified in Section 4.

2. Combine the forces and moments according to the load combinations specified in BWRVIP Document

[ Reference 2], and calculate the stresses on the piping, using the combined forces and moments, and the cross sectional properties of the piping.

3. Determine the applied stresses at several key locations in the piping system and use the limit load methods of Paragraph IWB 3640,Section XI, ASME Code [ References 3 and 4] as a guide to determine the allowable flaw lengths. He rules of 1989 Edition of Section XI are used as a guide in determining the allowable flaw lengths.
4. Perform leak rate evaluations.

3.0 Assumptions His Section describes the assumptions made in the methodology of the analysis.

l. De piping system geometry is as described in the referenced drawings [ Reference 5]. The dimension tolerances specified on the reference drawings are such that any variations within those values will have insignificant impact on the calculated stress values. It was also judged that any deviations between thel built geometry and the geometry indicated in the reference drawings would not be significant in terms stress analysis and the allowable flaw calculations. Any discrepancies between drawings were assumed be the most limiting Setween the two.
2. He seismic response spectra and anchor displacements at the core spray nozzles and at the shroud attachment points are as determined in the referenced report and design record file (References 6].
3. nny other assumptions are stated in the body of the report.

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O GE Nuclear Energy 4.0 Design inputs ne internal core spray piping is 5-inch schedule 40 and the material is Type 304 stainless steel . Figures I and 2 show the schematic of both loops of the intemal core spray lines. Figure 3 shows the details of the core spray nozzle. Each of the welds are also identified in these figures. Three finite element models, consisting of upper and lower loops of the internal core spray piping and one branch of the core spray spargers, were developed to determine the stresses from various design loads. Figure 4 provides a line plot showing the elements of one of the core spray piping finite element models. Figure 5 provides a line plot showing the elements of the core i spray sparger finite element model.

ne design inputs in this evaluation consisted of: (1) the geometries of the intemal core spray lines and

! spargers, and (2) the applied loads. He geometries of the internal core spray lines were obtained from the j drawings listed in Reference 5. The applied loads on the core spray line consist of the following: deadweight, seismic inertia, seismic anchor displacements, fluid drag, loads due to flow initiation and thermal (and internal pressure) anchor displacements. Each of these loads is briefly discussed next.

4.1 Dead Weight (DW) ne deadweight loading consists of the weight of the core spray pipe and the weight of the entrapped water.

f 1 The metal weight was determined as 14.57 lb/ft (9.11 lb/ft for sparger) and the weight of the entrapped water l as e o Ib/ft (4.28 lb/ft for sparger). Weight of the nozzles was also included for the sparger. He stresses for this loading were calculated by applying a 1.0g vertical acceleration in the finite element models. For the flaw evaluation purposes, the stress from this loading is treated as a primary stress.

1 4.2 Dynamic inertia De dynamic inertia loading consists of horizontal and vertical inertia forces acting on the entire core spray line due to seismic and hydrodynamic excitation of the RPV and the core shroud. He locations where the l

1. seismic excitation is imparted to the core spray line are the vessel nozzle, the support brackets and the points where it is attached to the shroud.

For the purpose of specifying load combinations, the following designadons are used:

Operating Basis Earthquake inertia - X Direction : OBEIX Operating Basis Earthquake Inertia - Y Direction: OBEIY Operating Basis Earthquake Inertia Vertical' - Z Direction: OBEIZ Safe Shutdown ( or Design Basis) Earthquake - X Direction: SSEIX Safe Shutdown ( or Design Basis) Earthquake - Y Direction: SSEIY Safe Shutdown ( or Design Basis) Earthquake Vertical- Z Direction: SSE!Z For the flaw evaluation purposes, the stresses from the seismic inertia loading are treated as primary stresses.

' These acronyms have been changed from the list in Reference 2 to correspond with the coordinate system used for the finite element model shown in Figures 4 and 5.

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& GE NuclearEnergy 4.3 Dynamic Anchor Displacement Dynamic anchor displacements are applied to the attachment points of the core spray lines at the RPV and the shroud. Table I lists the calculated values of the relative dynamic anchor motions. ).

Table 1 Relative Dynamic Anchor Motions j i

OBE SSE t

Horiz Horiz j 0.037 in. 0.055 in  ;

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The analysis assumed that a reasonable extent of metal ligament is remaining at each of the circumferential -

welds in the shroud. For flaw evaluation purposes, the stresses from the seismic anchor displacement loading l are treated as secondary. The load case designations used are the following j

Operating Basis Earthquake Displacement - X Direction: OBEDX Opc.ating Basis Earthquake Displacement - Y Direction: OBEDY Safe Shutdown (or Design Basis) Earthquake Displacement - X Direction: SSEDX-Safe Shutdown (or Design Basis) Earthquake Displacement - Y Direction: SSEDY I here is also an anchor displacement due to Main Steam Line Break (MSLB) pressure lifting off the top of a f cracked shroud. When SSE occurs simultaneously with MSLB, deflection is increased since the lifted shroud is free to rotate under the SSE loading. This load case is called LOCAPD, and it occurs at the same time as DRG2 l (see below). The vertical anchor motions for LOCAPD were obtained from the Pilgrim Shroud Repair  :

Hardware Stress Analysis Supplement, March 95 [Refemace 6]

4.4 Fluid Drag De drag loads consist of the forces resulting from the fluid flow past the core spray line, ne flow in the f

annulus region during the normal operation exerts some downward drag force on the core spray piping. De magnitude of this loading was determined to be approximately 5.73 lb/ft or 0.688 lb/in, based on a conservative value of 5 ft/second for the fluid velocity in the vessel annulus region. During the upset condition, core spray operation is assumed (no feedwater flow) and, therefore, the drag loads are insignificant. '

During a postulated double-ended break, the drag loads on the core spray line were determined to be 0.5 psi ~l downward at the upper portion of the core spray line and 5 psi downward at the lower portion of the core spray line ne drag loads are treated as primary loads for the flaw evaluation purposes and are designated as l follows:

Drag Load During Normal Op: ration: DRG1 Drag Loads During LOCA Conditions: DRG2 4.5 Core Spray Injection Loading (CSIN)

Two types orloads result when the core spray flow is initiated: internal pressure and the axial loads due to .

flow. During normal operation, the pressure differential between the inside and the outside of the core spray 6

& GE Nuclear Energy line is essentially negligible. Durmg core spray injection, a bounding internal pressure value of 68.8 psi was used (68.8 psi is the pressure drop across the core spray lines corresponding to the test condition with the highest flow listed,4500 gpm, on Core Spray System Process Diagram).

The axial load due to flow, which is a function of flow velocity , was calculated to be 19.3 psi. (equivalent pressure) This was added to the 68.8 psi due to pressure difference to yield 224 psi total (equivalent pressure).

The membrane stress due to this internal pressure was calculated using strength of materials formulas. Also, reaction forces and mcments due to flow out of the sparger nonles were applied at the nonle locations. The force calculation is similar to the calculation of axial force in the noule. Moment is based on the force times the applicable moment arm.

Stresses due to water hammer loads are insignificant since the core spray inlet valve ramps open over a period of time up(.n system actuation.

4.6 Thermal Loads ne two anchor points of the internal core spray line (the core spray nozzle and the brackets on the vessel at one end and the shroud attachment points at the other end) expand vertically and horizontally at different rates due to differences in the materials

  • thermal expansion coefficients (low allow steel for the vessel versus stainless steel for the shroud). Also, these displacements are expected to vary during certain transients due to the differences in temperatures between the vessel and the shroud. He loads produced by these thermal anchor disp'scements and thermal expansion are treated as secondary. He intemal pressure in the vessel also produces vertical and horizontal anchor motion at the nozzle and brackets , This displacement was included along with the thennat anchor displacements. The following thermal load cases need to be considered:

Thermal displacements during Normal Steady State operation : NOD Thermal displacements during Loss of Feedwater Pump transient: LFWPD Thermal displacements during LOCA: LOCAD The LOCA thermal displacements may consist of several sub-cases. One case occurs when the core spray is l just initiated following the LOCA event. Another sub-case may be several hours following the LOCA event.  :

ne only difference between the various LOCA sub-cases would be the assumed temperatures for the vessel, ,

the shroud, the shroud support brackets and the core spray piping. An intermediate case , for which the '

shroud, the support brackets and the core spray line have reached final temperature, but vessel temperature has not changed, resulted in the greatest difference in displacements and was determined to be bounding.

Pilgrim core support structure does not include shroud support legs (stilts). The support consists of a cylinder madc of nickel-chrome-iron (Alloy 600). The RPV, shroud, and cylinder heights are used to determine the thermal displacements. The calculated values of differential thermal displacements for the various transient conditions are shown in Tables 2 and 3.

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4 Table 2 Thermal Displacements for Transient Conditions For The Upper Core Spray Line Operating Temperatures (*F) RPV Displacements (in.) * *

  • Pipe Condition / RPV Support Shroud
  • Pressure Nozzle RPV Shroud '" Relative Temp.

Cylinder (psi) Safe End Horiz. Horiz Vert. (*F)

Transient 534 1000 0.550 0.468 0.414 -0.076 522 NOD $22 522 300 432 1100 0.308 0.269 0.318 0.243 300 j LFWPD 300 281 35 0.483 0.399 0.178 -0.697 173 f LOCAD $22 281 l-Table 3 Thermal Displacements for Transient Conditions for The Lower Core Spray Line Temperatures (*F) RPV Displacements (in.) * " Pipe Operating '

Shroud' Pressure Nozzle RPV Shroud " Relative Temp.

j Condition / RPV Support Cylinder (psi) Safe End Horiz. Horiz Vert. (*F)

Transient 534 1000 0.550 0.468 0.414 -0.119 522 NOD $22 522 432 1100 0.308 0.269 0.318 0.210 300 LFWPD 300 300 281 35 0.483 0.399 0.178 -0.715 173 j LOCAD 522 281

  • The Shroud temperature is assumed to be the average between the Annulus temperature and the Core temperature.

" Shroud vertical displacement - RPV vertical displacement

'" LOCAD pipe temp is assumed the average of annulus fluid temperature and tne temperature of the fluid inside the pipe.

He temperatures and pressures stated in the above table are derived from the information contained in the RPV and nozzle thermal cycle drawings [ References 8 and 9].

5.0 LOAD COMBINATIONS & STRESS LEVELS ,

his section describes the manner in which the various loads were combined for the purpose of obtaining stress levels for flaw evaluation. He limiting stress levels in three critical areas are then summarized.

5.1 Load Combinations ne flaw evaluation methodology used in this analysis (similar to that in Section XI in the ASME Code), "

makes the distinction between the normal / upset (Level A/B) condition loads, for which the factor of safety is 2.77, and the emergency / faulted (Level C/D) condition loads, for which the safety factor is 1.39.

De following load combinations were considered for normal / upset conditions for the core spray line:

(1) DW(P) + DRGf(P) + NOD (S)

(2) DW(P) + CSIN(P)

(3) DW(P) + DRGl(P) + LFWPD(S)

(4) DW (P) + DRGl(P) + OBElX(P) + OBEIZ(P) + NOD (S) + OBEDX (S) 1 4

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(5) DW(P) + DRGl(P) + OBElY(P) + OBElZ(P) + NOD (S) + OBEDY(S) i 1

Note that the letter in the parenthesis indicates whether a load is primary (P) or secondary (S) as defined by the l ASME Code. He set ofload combinations used for the Emergency / Faulted conditions consist of the following for the core spray line:

(6) DW(P) + DRG2(P) + SSElX(P) + SSEIZ (P) + NOD (S) + SSEDX(S) + LOCAPD(S) l (7) DW(P) + DRG2(P) + SSElY(P) + SSE!Z (P) + NOD (S) + SSEDY(S) + LOCAPD(S)

(8) DW(P) + CSIN(P) + SSElX(P) + SSElZ(P) + LOCAD(S) + SSE"X(S)

(9) DW(P) + CSIN(P) + SSElY(P) + SSE!Z(P) + LOCAD(S) + SSEDY(S) ne LOCAD loads need not be included in the emergency / faulted combinations 6 and 7 since these i displacement-controlled loadings develop much later in time when the drag loads due to LOCA have j

, decreased to an insignificant level. On the other hand, LOCAPD is added to Faulted corr.binations 6 and 7 l since it occurs at the same time as DRG2.

De loads applicable for the sparger are: deadweight, seismic inertia, and core spray initiar,on. The core spray initiation also produces sparger nozzle loads that need to be included with the other lost. He fluid drag forces act in the upward direction, and are less than the weight loading and therefore can be conservatively neglected. j i

ne following load combinations were considered for normal / upset conditions for the sparger and the reaction forces at the sparger brackets:

l (1) DW(P) + OBElX(P) + 0BE!Z(P) l (2) DW(P) + OBElY(P) + OBElz(P)

(3) DW(P)+CSIN(P)

The load combinations evaluated for the emergency / faulted conditions for the sparger are the following:

(4) DW(P) + CSIN(P) + SSElX(P) + SSEIZ(P) + LOCAD(S)

(5) DW(P) + CSIN(P) + SSElY(P) + SSElZ(P) + LOCAD(Si ne CSIN loads include the hydraulic load at the sparger nozzles.

5.2 Calculated Stress Levels ne force and moments at various nodes in the model for all of the load sources were calculated using the ANSYS finite element code (Reference 1]. These forces and moments were then combined to obtain the total forces and moments for a given load combination. Rus, for each load combination and each node, a set of forces and moments were obtained. Furthermore, within each set, the forces and moments from the displacement-controlled loadings were tabulated separately for the calculation the P, stress. As described later, the flaw evaluation methodology uses the primary membrane (P ), primary bending (P ) and the expansion stress (P ).

15 weld locations were considered for the purpose of allowable flaw evaluations (see Figures I,2 and 3). The calculated values of P., P., and P,, stress levels at these seven locations are summarized in Tables 4 and 5 for the goveming normal / upset and emergency / faulted condition load combinations. De sparger stress is given at the weld between the sparger arm and the sparger T-Box. The stress levels in Tables 4 and 5 are used in the allowable flaw eva.aations as described in the next sect;ca.

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, 1 O GE Nuclear Energy Table 4. Primary Membrane, Primary Bending and Secondary Stresses for the Loop A Piping Weld Location Element # Gov. Load P,, P. P.

Weld, (Figures I and 2) (Figure 4) Type Combination (Psi) (Psi) (Psi) 2 910.0 682.6 0.0 Pla(14-A-13F) 78 non-flux 7 129.0 2315.6 5774.8 Qux 2 910.0 752.1 0.0 Plb(14-A 14F) 77 non-flux 7 128.1 2209.0 5687.2 f flux 2 910.0 786.1 0.0 l Plc(14-A-15F) 75 non-flux 7 126.2 20052 5506.1 l flux 2 292.7 24.9 0.0 l SWl3 A(14-A-16F) 71 non-flux 9 324.5 39.8 968.2 flux 5 37.3 1313.8 3201.6^

P3(PWil AN/S) 70 non-flux 9 422.6 1333.3 10221.4 flux i 2 377.9 174.9 0.0 P4a(PW10AN/S) 21 non-flux 9 395.5 291.0 7639.7 flux 2 401.3 254.4 0.0 P4b(PW9AN/S) 20 non-flux 9 415.8 442.0 7143.6 flux 2 424.7 53.2 0.0 ,

P5(FW8AN,SW8AS) 12 non-flux 9 450.4 163.9 5825.1 l flux 2 425.6 58.7 0.0 .l P6(FW7AN,SW7AS) 1I non-flux 9 451.7 169.5 6280.1  ;

flux 2 427.4 73.5 0.0 non-flux P7(SW6AN/S) 10 9 454.4 211.4 7197.3 l flux I 2 429.0 88.8 0.0 P4c(PW5AN/S) 5 non-Dux 9 456.7 267.6 8005.2 flux 2 375.8 154.7 0.0 P4d(PW4AN/S) 4 non-Dux 9 381.6 373.8 9136.1 flux 2 729.1 386.6 0.0 P8b(SWlAN/S) I non-flux 9 737.9 696.5 10881.2 flux 2 729.1 386.6 0.0 P8a(SW2AN/S) I non-Gux  !

9 737.9 696.5 10881.2 flux T*812 P2(SW12A) non-flux t=812 Dux -

297 243 0 Sparger 39(Fig 5) non-Dux 3 4 335 305 733 Dux

~ Non-flux weld category also mcludes base metal.

    • 7his value is calculated but not used in the flaw evaluation for this non-Dux or base metal case.

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Q GE Nuclear Energy Table 5. Primary Membrane, Primary Bending andSecondary Stresses for the Loop B Piping Weld Location Element # Weld Gov. Load P. P. P.

(Figures I and 2) (Figure 4) Type Combination (Psi) (Psi) (Psi) 78 2 909.4 787.6 0.0 Pla(14-B-13F) non-flux 7 129.8 2311.0 5723.1 flux 77 2 909.4 753.4 0.0 Pib(14-B-14F) non-flux 7 128.9 2211.2 5637.3 flux 75 2 909.4 683.8 0.0 l Plc(14-B-15F) non-flux 7 127.0 2007.5 5459.8 flux 71 2 292.5 25.0 0.0 l SW13 B(14-B-16F) non-flux 9 325.0 39.8 988.6 flux 70 5 37.8 1332.4 3106.3~

P3(PWl1BN/S) non-flux 9 424.0 1359.1 9860.0 4

flux 21 2 377.1 169.4 0.0 P4a(PW10AN/S) non-flux 9 397.1 290.9 7098.7 flux 20 2 * ^ ! .2 251.4 0.0 P4b(PW9BN/S) nor.-flux 9 415.9 426.9 6818.1 flux 12 2 424.6 49.9 0.0 P5(FW8BN.SW8BS) non-flux

' 9 450.3 213.0 3534.4 flux TE[/W7BN,SW7BS) 11 2 425.5 48.8 0.0 non-flux 9 451.6 194.0 3873.3 flux 10 2 427.3 52.0 0.0 P7(SW6BN/S) non-flux 9 454.3 170.9 4566.2 flux 5 2 433.4 95.9 0.0 P4c(PW5BN/S) non-flux 9 463.3 287.6 6979.5 flux 4 2 375.5 166.5 0.0 P4d(PW4BN/S) non-flux 9 382.1 396.0 7858.0 flux I 2 728.6 415.9 0.0 P8b(SWlBN/S) non-flux 9 738.6 734.9 9346.5 flux I 2 728.6 415.9 0.0 P8a(SW2BN/S) non-flux 9 738.6 734.9 9346.5 flux t=812 P2(SW12B) 3 297 243 0 Sparger 39(F;g 5) non-flux Dux 4 335 305 733 Non-flux weld category also includes base metal.

    • This value is calculated but not used in the flaw evaluation for this non-flux or base metal case.

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Q GE Nuclear Energy 6.0 FRACTURE MECHANICS EVALUATION 1 i

he limit load methodology was used in calculating the tillowable flaw lengths. His methodology is first l

described followed by the results of allowable flaw evaluations.

6.1 Limit Load Methodology (Circumferential Cracks) ,

Consider a circumferential crack of length, l = 2Ra and constant depth, d. In order to determine the point at l whkh limit load is achieved, it is necessary to apply the equations of equilibrium assuming that the cracked section behaves like a hinge. For this condition, the assumed stress state at the cracked section is as shown in l Figure 6 where the maximum stress is the flow stress of the material, eg. Equilibrium oflongitudinal forces )

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and moments about the axis gives the following equations: l p = [(n ad/t)-(P,/o )n]/2 (1) r Pb' = (2c/n)(2 sin p - d/t sin a) (2)

Where t = pipe thickness, inches a = crack half-angle as shown in Figure 3 p = angle that defines the location of the neutral axis Z= Weld type factor P,= Piping expansion stress P. = Primary membrane stress P.,= Primary bending stress  ;

P = Failure bending stress ASME Sect XI IWB-3640 and Appendix C limit the applicability of the limit load equations (Equations I and l

2) to d/t less than 0.75 for base metal and non-flux welds, and 0.6 for flux welds. These limitations are not based on structural mechanics considerations, but were intended to preclude operation of a pressure boundary component with a leaking crack and the fluid spilling out during normal plant operation (e.g., see discussion in Reference 4). He core spray line in the RPV annulus region is not a part of the reactor pressure boundary and some leakage during its operation is acceptable. Rus, the use of Equations I and 2 for assumed through-wall flaw geometries is acceptable.

De safety factor is then incorporated as follows:

(3)

P[ = Z'SF (P. + P + P/SF)- P.

P, and P are primary stresses. P,is secondary stress and includes stresses from all displacement controlled loadings such as thermal expansion and dynamic anchor motion. All three quantities are calculated from the analysis of applied loading. De safety factor value is 2.77 for normal / upset conditions and 1.39 for emergency / faulted conditions.

I Factor The test data considered by the ASME Code in developing the flaw cvaluation procedure (Appendix C,Section XI) indicated that the welds produced by a process without using a flux had fracture toughness as good or better than the base metal. However, the flux welds had lower toughness. To account for the reduc toughness of the flux welds (as compared to non-flux welds) the Section XI procedures prescribe a penalty factor, called a 'Z' factor. ne examples of flux welds are submerged arc welds (SAW) and shielded metal arc welds (SMAW). Gas metal-arc welds (GMAW) and gas tungsten-arc welds (GTAW) are examples of non-flux welds. Figure IWB-3641 1 may be used to define the weld base metal interface. The expressions for the value of the Z factor in Appendix C of Section XI are given as follows:

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~w GE Nuclear Energy Z =

1.15 {l + 0.0l3(OD-4)] for SMAW

=

1.30 [1 + 0.010(OD-4)] for SAW where OD is the nominal pipe size (NPS) in inches. He procedures of Appendix C recommend the use of OD

= 24 for pipe sizes less than 24 inches. His approach is very conservative and, therefore, the use of actual .

l NPS was made in calculating the Z' factor. This approach is considered reasonable as recent discussions in I

the Section XI Code Working Group on Pipe Flaw Evaluation indicate that for small diameter pipes, such as '

the 5 inch diameter core spray piping, the Z-factor may be close to or less than 1.0.

The flux-type welding process used in the field was shielded metal arc type (SMAW). Herefore, the Z-factors are-  !

(sparger)

Zw = 1.15[1 + 0.013(4-4)] = 1.15 (core spray line)

Zs a = 1.15[1 + 0.013(5-4)] = 1.165 Zu a = 1.15[1 + 0.013(12-4)] = 1.27 (T-box)

Z i o.a = 1.15[1 + 0.013(NPS,-4)} = 1.24 (nermal Sleeve)

Zo .a = 1.15[1 + 0.0l?(NPS,-4)} = 1.18 (Collar) where N'"',, 10 in - 0.1875 (thickness) = 9.8 NPS, = 6.4 in - 0.156 (thickness) = 6.2

'If the indication is located in the non Oux weld or base metal, Z is assumed as 1.0 and the P, stresses are not used in the calculation, consistent with Section XI Appendix C guidelines.

6.2 Allowable Flaw Length Calculation (Circumferential) \

i The stresses from the table in the preceding section were used to determine the acceptable end-of-cycle through wall flaw lengths. The acceptable flaw size was de'.crmined by requiring a safety factor. The flow stress was taken as 3S (S. = 16.9 ksi for Type 304 stainless steel at 55CT). As specified in Reference 2, a safety factors of 2.77 for the normal / upset conditions and 1.39 for the emergency / faulted conditions, respectively, were used. The calculated values of the allowable flaw leneths for the core spray piping are tabulated, and shown in Tables 5 and 6. The allowable flaw lengthe are valid for multiple welds and do not include NDE uncertainty. The NDE uncertainties should be accounted for prior to comparison to the allowable flaw sizes. For welds P5, P6 and P7 the allowable flaw sizes are conservatively calculated using the OD of the 5 inch piping. For the thicker slip joint section, it is recommended that the allowable flaw angle is used along with the corresponding OD at the flaw location. He allv able flaw sizes listed for the non-flux welds can also be used for flaws found in the base metal of the piping.

For the T-box cover plate (weld P2-SW12A/B), shear stress due to CSIN was calculated, and shear flow stress of 1.5S, was used. Equations are the same as for the other welds.

The first set of allowable crack lengths given in Tables 6,7,8 and 9 do not include the projected crack growth during the next operating cycle. Essentially, with a crack growth of 1.6 inches cased on the re-inspection at the end of one 24 month fuel cycle and 3.2 inches based on the re-mspection at the end of two 24 month fuel cycles , the allowable crack lengths are reduced by 1.6 inches (one cycle) or 3.2 inches (tw o cycle). The allowable flaw lengths which include the projected crack growth during the next operating cycle are tabulated in the bottom half of Tables 6 and 7 and for the next two operating cycles are tabulated in the bottom half of Table 8 and 9. The basis for the 1.6 inch or 3.2 inch crack growth is discussed in the next subsection.

13

& GE NuclearEnergy Teb!: 6 Loop A Core Spray Line A!lowable Flaw Lengths for One Cyle Weld non flux or base metal flux Location Angle Length Angle Length (Figure 1) (deg.) (in.) (deg.) (in.)

l 230.4 20.1 190.1 (6.6 Pla(14-A-13F) 231.1 20.2 191.7 (6.7 Plb(14-A-14F)

Total 232.8 20.3 195.0 17.0 Pic(14-A-15F)

Allowable SWl3A(14-A-16F) 283.9 26.6 266.4 25.0 P2(SWI2A) 228.3 22.5 219.3 2I.6 Efective Crack w/o P3(PWIIAN/S) 245.2 Il.9 172.4 8.4 Crack 270.2 '13.1 196.9 9.6 P4a(PW10AN/S)

Growth 265.4 12.9 198.6 9.6 P4b(PW9AN/S)

Added P5(FW8AN.SW8AS) 273.0 13.3 2l0.4 10.2 272.7 13.2 206.9 10.0 P6(FN7AN.SW7AS)

P7(SW6AN/S) 271.9 13.2 199.9 9.7 j 271.0 13.2 194.0 9.4 P4c(PW5AN/S) 271.2 13.2 186.8 9.1 P4d(PW4AN/S) 246.5 13.7 169.6 9.4 P8a(SW2AN/S) 246.5 13.7 169.6 9.4 F8b(SWIAN/S)

Sparger 267.3 9.3 260.5 9.1 212.0 18.5 171.8 15.0  ;

Pla(14-A-13F)

Plb(14-A-14F) 212.8 18.6 (73.4 15.t 214.4 18.7 176.6 15.4 Total Plc(14-A-15F) 266.8 25.0 249.3 23.4 Allowable SWl3A(14-A-16F) '

212.0 20.9 203.1 20.0 Efective P2(SW!2A) 212.2 10.3 139.4 6.8 .

Crack with P3(PWilAN/S)

)

One Cycle P4a(PWl0AN/S) 237.2 Il.5 164.0 8.0 ofCrack P4b(PW9AN/S) 232.5 lI.3 165.6 8.0 Growth P5(FW8AN.SW8AS) 240.0 lI.7 (77.5 8.6 239.7 I I .6 (74.0 8.4 P6(FW7AN.SW7AS) 238.9 11.6 167.0 8.1 P7(SWBAN/S) 238.1 11.6 161.0 7.8 P4c(PWSAN/S) 238.3 11.6 153.8 7.5 P4d(PW4AN/S) 217.7 12.1 140.9 7.8 P8a(SW2AN/S) 217.7 12.1 140.9 7.8 P8b(SWlAN/S)

Spareer 221.5 7.7 214.6 7.5 . ,

14

Q GE NuclearEnergy Table 7 Loop B Core Spray Line Allowable Flaw Lengths for One Cycle Weld non-flux or base metal flux Location Angle Length Angle Length (Figure 1) (deg.) (in.) (deg.) (in.)

230.3 20.1 190.5 16.6 Pla(14-B-13F) 231.1 20.2 192.0 16.8 P1b(14-B-14F) 232.8  :.J.3 195.3 17.0 Total Plc(14-B-15F)

Allowable SW13B(14 B-16F) 283.9 26.6 266.0 25.0 J

212.0 20.9 203.1 20.0 Efective P2(SW12B)

Crack w/o P3(PWilBN/S) 244.6 Il.9 174.2 8.5 270.5 13.1 200.6 9.7 Crack P4a(PWl0AN/S)

Growth P4b(PW9BN/S) 26:.6 12.9 20l.0 9.8 Added P5(FW8BN.SW8BS) 273.1 13.3 229.3 Il.1 273.1 13.3 226.4 l 1.0 P6(FW7BN.SW7BS) 272.9 13.2 220.5 10.7 P7(SW6BN/S) 270.5 13.1 200.6 9.7 l P4c(PWSBN/S)

P4d(PW4BN/S) 270.7 13.1 194.7 9.5 f 245.7 13.7 177.8 9.9 P8a(SW2BN/S) 245.7 13.7 177.8 9.9 PBb(SWIBN/S)

Sparger 267.3 9.3 260.5 9.1  :

212.0 18.5 IT2.2 (5.0 }

Pla(14-B-13F) 212.8 18.6 173.7 15.2 Plb(14-B-14F) 214.4 18.7 176.9 15.4 Total Plc(14-B-15F) 266.8 25.0 248.9 23.4 Allowable SW13B(14-B-16F) 195.8 19.3 186.8 18.4 Efective P2(SW12B) 21I.7 10.3 141.2 6.9 Crack with P3(PWIIBN/S) 237.5 l1.5 167.6 8.1 One Cycle P4a(PW10AN/S) 232.6 11.3 168.0 8.2 ofCrack P4b(PW9BN/S) 240.2 I l.7 196.4 9.5 Growth P5(FW8BN.SW8BS) 240.2 11.7 193.4 9.4 P6(FW7BN.SW7BS) 239.9  ! ! .6 187.5 9.1 P7(SW6BN/S) 237.5 11.5 167.6 8.1 P4c(PWSBN/S) 237.7 11.5 161.7 7.9 P4d(PW4BN/S) 216.9 12.1 149.1 8.3 P8a(SW2BN/S) 216.9 12.1 149.1 8.3 P8b(SWIBN/S)

Sparger 221.5 7.7 214.6 7.5 15

O GE Nuclear Energy Table 8 Loop A Core Spray Line Allowable Flaw Lengths for Two Cycles Weld non-flux or base metal flux Location Angle Length Angle Length (Figure 1) (deg.) (in.) (deg.) (in.)

230.4 20.1 190.1 16.6 Pla(14-A-13F) 231.1 20.2 191.7 16.7 Plb(14-A-141) 232.8 20.3 195.0 17.0 Total Pic(14-A 15F)

Allowable SWl3A(14-A-16F) 283.9 26.6 266.4 25.0 228.3 22.5 219.3 21.6 Efective P2(SW12A) 245.2 11.9 172.4 8.4 Crack w/o P3(PWIIAN/S)

Crack P4a(PWl0AN/S) 270.2 13.1 196.9 9.6

' Grovth 265.4 12.9 198.6 9.6 P4b(PW9AN/S) 273.0 13.3 210.4 10.2 Added P3(FW8AN.SW8AS)

P6(Fn7AN,SW7AS) 272.7 13.2 206.9 10.0 271.9 13.2 199.9 9.7 P7(SW6AN/S) 271.0 13.2 194.0 9.4 P4c(PWJAN/S) 271.2 13.2 186.8 9.1 P4d(PW4AN/S) 246.5 13.7 169.6 9.4 P8a(SW2AN/S) 246.5 13.7 169.6 9.4 P8b(SWIAN/S)

Sparger 267.3 9.3 260.5 9.1 193.7 16.9 153.5 13.4 Pla(14-A-13F) 194.5 17.0 155.0 83.5 Pib(14-A-14F) 196.1 17.1 158.2 13.8 Total Pic(14-A-13F) 249.7 23.4 232.2 21.8 Allowable SW13A(14-A-16F) 195.8 19.3 186.8 18.4 Efective P2(SW12A) 179.2 8.7 106.6 5.2 Crack with P3(PWilAN/S) 204.2 9.9 131.2 6.4 Two Cycles P4a(PWIDAN/S) 199.6 9.7 132.5 6.4 ofCrack P4b(PW9AN/S) 207.2 10.1 144.5 7.0 Growth PS(FW8AN.SW8AS)

P6(FW7AN,SW7AS) 206.6 10.0 140.9 6.8 205.9 10.0 134.0 6.5 P7(SW8AN/S) 205.3 10.0 128.0 6.2 P4c(PWJAN/S) 205.4 10.0 12i.0 5.9 P4d(PW4AN/S) 188.9 10.5 112.0 6.2 P8a(SW2AN/S) 188.9 10.5 112.0 6.2 P8b(SWIAN/S)

Sparger 175.5 6.1 168.8 5.9 16

I .

O GE Nuclear Energy Table 9 Loop B Core Spray Line Allowable Flaw Lengths for Two Cycles Weld non-flux or base metal flux Location Angle Length Angle Length (Figure 1) (deg.) (in.) (deg.) (in.)

230.3 20.1 190.5 16.6 Pla(14 B-13F) 231.1 20.2 192.0 16.8 P1b(14-B-14F) 232.8 20.3 195.3 17.0 Total Pic(14-B-15F)

Allowable SW13B(14-B-16F) 283.9 26.6 266.0 25.0 P2(SWI2B) 212.0 20.9 203.1 20.0 Efective Crack w/o P3(PWIIBN/S) 244.6 11.9 174.2 8.5 Crack P4a(PW10AN/S) 270.5 13.1 200.6 9.7 Growth 265.6 12.9 201.0 9.8 P4b(PW9BN/S)

Added P5(FW8BN.SW8BS) 273.1 13.3 229.3 11.1 P6(FW7BN.SH7BS) 273.1 13.3 226.4 11.0 272.9 13.2 220.5 10.7 P7(SW6BN/S) 270.5 13.1 200.6 9.7 P4c(PWJBN/S) ,

P4d(PW4BN/S) 270.7 13.1 194.7 9.5 245.7 13.7 177.8 9.9 PBa(SW2BN/S) 245.7 13.7 177.8 9.9 l PBb(SWIBN/S)

Sparger 267.3 9.3 260.5 9.1 193.7 16.9 153.8 13.4 l Pla(14-B-13F) 194.5 17.0 155.4 13.6 Pib(14-B-14F) 196.0 17.1 158.5 13.8 Total Pic(14-B ISF)

SW13B(14-B-16F) 249.7 23.4 231.9 21.8 Allowable 179.6 17.7 (70.5 16.8 Efective P2(SWl2B)

P3(PWilBN/S) 178.8 8.7 108.4 5.3 Crack with 204.4 9.9 134.5 6.5 \

Two Cycles P4a(PW10AN/S) 199.7 9.7 135.2 6.6  ;

ofCrack P4b(PW9BN/S) 207.3 10.1 163.3 7.9 \

Growth P5(FW8BN.SW8BS) 207.3 10.1 160.5 7.8  !

P6(FW7BN.SW7BS) 206.8 10.0 154.5 7.5 P7(SW6BN/S) 204.4 9.9 134.5 6.5 P4c(PWJBN/S) 204.6 9.9 128.9 6.3 P4d(PW4BN/S)

P8a(SW2BN/S) 182.2 10.5 120.3 6.7 P8b(SWIBN/S) 182.2 10.5 120.3 6.7 Sparger 175.5 6.1 168.8 5.9 17

9 h GE Nuclear Energy 6.3 Crack Growth Evaluation Prior crack growth analyses performed for BWR shroud indications have conservatively used a crack growth rate of 5x10-5 inch / hot hour, ne stresses induced in the core spray line are very low, as evidenced by the stress results presented in the previous section. Rose stress results also conservatively include the effects of seismic and core spray injection loads, which are not typically present. Therefore, the applied stress intencity factor is low, and the corresponding crack growth rate would be significantly below the upper bound value of 5x10 inch / hot hour used here.

Pre-operational testing of BWR internals has demonstrated that high cycle fatigue resulting from flow induced vibration is not a concern for the core spray piping. Additionally, low cycle fatigue caused by assumed thermal transients which could be potentially imposed by cold fluid injections through the feedwater spargers located directly above the core spray lines have been found to be insignificant. Herefore, fatigue crack propagation ofindications in the core spray lines is concluded to be negligible, and is not considered to be a further contributor to the crack growth values discussed here.

Rus, a conservative crack growth rate of 5> 10 in/ hot hr can be used in the flaw evaluations. This crack growth rate. translates into a crack length increase of(8000 hr. per year)(2 years) (5x10) or 0.8 inch at each end of an indication assuming a 24-month fuel cycle or 1.6 inch at each end of an indication assuming two 24

-month fuel cycles. Thus, the projected length afte: one cycle, /,, of any indication whose current length at the time ofinspection is l ,p would be (1, +0.8x2) inches. A factor of 2 in the preceding parenthesis is to account j

for the growth at each end of the indication, i 6.4 Limit Load Methodology & Allowable Value for Axial Cracks 1

An approach similar to that outlined in Appendix C, ASME Section XI Code was followed in calculating the '

allowable axial crack length. It was assumed that the crack is throughwall. The allowable crack length,1,is given by the following equation:

I l=V2.48Rt([3S /(SF x o n )f-l) where:

R = Radius of the pipe t = thickness of the pipe SF = Safety Factor = 3.0 on = Circumferential Stress ne circumferential stress is essentially due to internal pressure.

The allowable flaw length for the nominal core spray pipe diameter of 5 inches was cciculated as 30.3 inches.

The allowable flaw length for the 10 inch OD thermal sleeve was calculated as 13.9 inches.

18

f...

& GE Nuclear Energy l

7.0 LEAKAGE EVALUATION 7.1 Leak Rate Calculation Methodology l

The leakage from the core spray line into the RPV annulus could come from a number of sources such as through the gap between the sleeve and the nozzle ID, and through the presence of any through-wall cracks in l

the piping. The leakage rate was estimated assuming incompressible Bernoulli flow through a hole:

i I

i Q = CA,/2g,AP / p (4) where Q= leakage C= flow coefficient A= area p= mass density of fluid AP = pressure difference across the pipe A AP value of 44 psi based on Reference 7 was used. This is the uoper bound value of steady state pressure l during the core spray operation for this plant.

Leak rate from the through-wall indications in the core spray line can be estimated using the preceding equation with the value of flow coefficient, C, conservatively assumed as 1.0. A key input needed is the crack opening area, A.

The approach used in this evaluation to calculate the value of A, was to assume a conservative value of crack opening displacement,6, and assume the crack opening configuration to be like a rectangular slot with one side being the crack length,2a, and the other side as the crack opening displacement. A value of 0.01 inch was assumed for 6. Linear clastic fracture mechanics calculations indicated that this assumed value of 5 is conservative for crack lengths up to half of the circumference. The crack opening area is then simply; ,

A= 2a (S) (5) 7.2 Overall Leak Rate Calculation The leak rates from any indications would be a function of the detected number and lengths of the indications which will be known only after an examination of the internal core spray piping has been conducted. To facilitate this calculation after the examination results are determined, leak rate per inch of crack length is provided herein. This leak rate was calculated as 2.52 gpm per inch of crack length.

As an example of overall leakrate calculations, consider a case where the overall length (including the prcjected crack growth during the next inspection interval) of the indications to be 15 inches, then the overall leakage is (15 x 2.52) or 37.8 gpm.

8.0

SUMMARY

& CONCLUSIONS l

A flaw evaluation, consisting of stress and limit load analyses of the internal core spray piping and spargers at

Pilgrim was performed to develop a flaw disposition handbook. The methodology outlined in the BWRVIP l

l 19

O GE NuclearEnergy

\

i core spray line ! & E guidelines document was followed in this evaluation. Allowable flaw lengths were calculated at several critical locations and leak rate calculation results are presented. It is seen the core spray line is fairly flaw tolerant. At the worst location, on the upper line near the T-Box connection (P3 weld), the minimum allowable flaw size for one cycle is 139' of circumference or 106*of circumference for two cycles.

The methodology presented in this report can be used to disposition any indications detected during future inspections of the intemal core spray lines and spargers.

I 20

.. l 1

O GE Nuclear Energy l  !

i l J

REFERENCES

[1] ANSYS Engineering Analysis System User's Manual, Revision 4.4, Swanson Analysis Systems, Inc.,

Houston, PA, May 1,1989

[2].

Caine,T.A., et al, BWR Core Spray Intemals inspection and Flaw Evaluation Guidelines, GE-NE.

B13-01805-21

[3] ASME Boiler and pressure Vessel Code,Section XI, Rules for In-Service Inspection of Nuclear

}

Power Plant Components, American Society of Mechanical Engineers,1989 Edition, Paragraph IWB 3640.

[4] Ranganath, S. and Mehta, H. S " Engineering Methods for the Assessment of Ductile Fracture Margin in Nuclear Power Plant Piping," Elastic plastic Fracture: Second Symposium, Volume 11 -

Fracture Resistance Curves and Engineering Applications, ASTM STP 803, C.F. Shih and J.P Gudas, Eds., American Society for Testing and Materials,1983, pp.11-309 330 l 1

[5] Pilgrim Drawings: GE Drawings 919D932,919D923,719E427, VPF Drawings 1979-216-6,197-257-5,2426-47-2,2426-069-00 i,

[6] Pilgrim Shroud and Shroud Repair Hardware Stress Analysis , March 95, GENE-B11-00617 j

[7] Core Spray Process Diagram, GE Drawing No.161F322, Rev 2

[8] Nozzle Thermal Cycle Diagrams, GE Drawing No.136B1944, Rev 0

, [9] RPV nermal Cycle Diagram, GE Drawing No. 730E491, Rev 0 l

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& GE Nuclear Energy 1

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N Pilgrin core spray sparger Analysis Figure 4 Ansys Model of the Pilgrim Core Spray Line Sparger 26

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$ GE Nuclear Energy l

NominalStrees in the unemcked Section of Pipe Crack t.ength = 2Ra pg 4 4- Flow Stmos,c, 4%

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