ML20205F649

Rev 1 to GE-NE-B13-01935-02, Jet Pump Assembly Welds Flaw Evaluation Handbook for Vermont Yankee

ML20205F649
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Site:	Vermont Yankee File:NorthStar Vermont Yankee icon.png
Issue date:	07/31/1998
From:	Caine T, Kaul M, Mallard L GENERAL ELECTRIC CO.
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ML20205F577	List: ML20205F574 ML20205F649 ML20205F663
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GE-NE-B13-01935, GE-NE-B13-01935-02, GE-NE-B13-1935, GE-NE-B13-1935-2, NUDOCS 9904060344
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GENuclearEnergy TECHNICAL SERVICES GE-NE-B13-01935-02 GE Nuclear Energy Revision 1 175 Curtner Avenue, San Jose, CA 95125 DRF # B13-01935 JET PUMP ASSEMBLY WELDS FLAW EVALUATION HANDBOOK FOR VERMONT YANKEE July 1998 Prepared for Vermont Yankee Nuclear Power Corporation Prepared by GE Nuclear Energy 9904060344 990329

gDR ADOCK 05000271 PDR ,.

GE Nuclear Enero GE-NE-B13-01935-02, Rev.1 l DRF # B13-01935 l 1

I JET PUMP ASSEMBLY WELDS FLAW EVALUATION HANDBOOK FOR l VERMONT YANKEE I

July 1998 Prepared by: d  %

Lerond Mallard, Engineer Structural Mechanics & Materials a

Verified by: WW >

M. K. Kaul, Principal Engineer Structural Mechanics & Materials Reviewed by: b, bw T. A. Caine, Manager Structural Mechanics & Materials i

GENuclect Energy GE-NE-B13-01935-02, Rev.1 DRF # B13-01935 IMPORTANT NOTICE REGARDING 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 Vermont Yankee Nuclear Power Corporation (VYNPC) and GE, and nothing contained in this document shall be construed as changing the contract. The use of this information by anyone other than VYNPC, 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 the completeness, accuracy, or usefulness of information. contained in this document, or that its use may infringe privately owned rights.

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l GE Nuclear Energy GE-NE-B13-01935-02, Rev.1 DRF # B13-01935

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Executive Summary A flaw evaluation, consisting of stress and fracture mechanics analyses of the Vermont Yankee jet pump assembly circumferential welds, was conducted to develop a flaw evaluation handbook. The procedures of BWRVIP-41 were used as a guide in determining the allowable flaw lengths. End-of-cycle ellowable flaw lengths were calculated at all circumferential weld locations. The methodology presented in this report can be used along with consideration of observed IGSCC and evaluation of fatigue crack growth rates to disposition any indications detected during inspections of the jet pumps at Vermont Yankee.

The following tables show a summary of the allowable adjusted flaw lengths for Vermont Yankee. See Figure 1 on page 21 for weld locations. All flaw lengths must be adjusted for IGSCC growth and NDE uncertainty.

Allowable Adjusted Circumferential Flaw Sizes (inchl Weld Flaw Length Flaw Length Weld Designation (Limit Load) (FIV) Description TS-2 19.56 18.75 Thermal Sleeve TS-2a 19.90 16.26 Thermal Sleeve TS-2b 16.17 10.38 Thermal Sleeve RS-1 18.62 8.16 Riser Elbow RS-2 21.43 13.56 Riser Elbow RS-3 21.43 21.43' Riser to Transition Piece RS-4 15.66 13.84 Riser Sleeve RS-5 15.30 11.03 Riser Sleeve j TR-1 37.63 37.63* Transition Piece I IN-1 14.15 11.37 Elbow and Mixer Nozzle MX-1 10.45 3.99 Mixer Flange to Barrel MX-2 11.68 3.35 Barrel to Adapter MX-4 15.48 9.93 Adapter to Flare MX-5a 14.43 2.53 Adapter Sleeve MX-5b 14.43 2.53 Adapter Sleeve DF-1 21.34 18.45 Diffuser Collar to Shell DF-2 30.16 22.21 Diffuser Shell to Tailpipe DF-3 25.74 12.61 Diffuser Tailpipe to Adapter AD-1 22.18 4.47 Adapter Top to Bottom AD-2 21.19 4.00 Adapter Bottom to Support Plate

• FIV does not limit acceptance criterion of as-found flaw size.

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GENuclear Enerv GE-NE-B13-01935-02, Rev.1 DRF # B13-01935 TABLE OF CONTENTS Eate

1. PURPOSE / OBJECTIVE 1 ;

2. METHODS 1 1

3. ASSUMPTIONS 2

4. DESIGN INPUTS 2 4.1. Static Loads 2 4.2. Dynamic Loads S 4.3. Vibration Loads 6 I

5. LOAD COMBINATIONS AND STRESS LEVELS 8 5.1. Load Combinations 8 5.2, Osiculated Stress Levels 8

6. FRACTURE MECHANICS EVALUATION 10 6.1. Limit Load Methodology 10 6.2. Allowable Flaw Length Calculation 13 6.3. Crack Growth Evaluation 14

7. EVALUATION OF INDICATIONS 16 7.1. Procedure for Evaluation of Indications using Ilandbook 16 7.2. Leakage Calculation 16

8.

SUMMARY

& CONCLUSIONS . 18

9. REFERENCES 19 APPENDIX: Response to NRC Request for Additional Information 24 Note: Changes to Revision 0 are shown by bars "l" in the left margin.

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GE N:: clear Enero GE-NE-B13-01935-02, Rev.1 DRF # B13-01935 .

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1. Purpose / Objective l The objective of this handbook is to document the results of a fracture mechanics evaluation of the Vermont Yankee (VY) jet pump assembly circumferential welds. This evaluation results in the allowable end-of-cycle flaw lengths at all of the jet pump assembly  !

circumferential welds. Figure 1 is a schematic showing the welds of interest with desig, ations in accordance with BWRVIP-41 (Reference 1).

The results presented in the flaw evaluation handbook can be used to disposition indications found in thejet pump assembly at Vermont Yankee.

2. Methods This section presents the methodology and procedure used in performing the jet pump riser weld flaw evaluation. Following are the steps used in the analysis.

1. Review of the reference drawings.

2. Determine the loading and load combinations as suggested in BWRVIP-41.

3. Create a SAP 4G07V (Reference 2) finite element rrodel of the jet pump assembly.

Anchor connection points are the recirculation inlet nozzle, shroud support plate and riser brace.

4. Determine the membrane and bending stresses considering the load combinations.

5. Use the limit load methods of BWRVIP-41 as a guide to determine the allowable flaw lengths. BWRVIP-41 evaluation procedures are used as a guide, since the jet pump is not a part of the reactor pressure boundary.

6. Evaluate IGSCC and fatigue crack growth rate. Calculate IGSCC crack growth for an eighteen month (12,000 hrs) cycle based on a growth rate of 5x10 4inch / hot hour.

Determine if fatigue crack growth rate due to vibration is significant by calculating the stress intensity factor due to flow induced vibration and comparing with threshold stress intensity factor range. If calculated stress intensity factor range is less than threshold, then given crack would have no fatigue crack growth.

7. Leakage rates are calculated as a function of percent of allowable flaw length d'uring normal plant operation per BWRVIP-41 procedures.

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3. Assumptions

1. The jet pump geometry is as described in the reference drawings (Reference 5).

2. The calculations are based on one flaw per riser. However, synergistic effects of multiple flaws in one riser are negligible and would not affect the results of this analysis. TM reason is that even large flaws (180 degrees) do not significantly change the stiffness of the riser and therefore the response to input loadings does not change. Note that neglecting the synergistic effects is conservative because even a small decrease in stiffness due to flaws would reduce loads because ofincreased system compliance.

3. Fatigue due to thermal stresses is negligible due to minimal temperature differentials during normal or transient conditions.

4. The jet pumps are assumed to be in the as-designed configuration because the vibration data on which the fatigue evaluation is based was taken from a new plant in startup. To account for issues such as jet pump fouling and restrainer bracket set screw gap, a conservative value of A.K threshold, 5.0 ksi-Vin, is used in the vibration fatigue evaluation.

4. Design Inputs The design inputs in this evaluation consisted of the geometry of the jet pump and the applied I loads. The geometry of thejet pump was obtained from the drawings listed in Reference 5.

The jet pump riser is 10-inch schedule 40, inlet mixer is 6-inch piping, the diffuser is a frustum with radii of 8 and 14-inches, the adapter is 14-inch piping. The material for the entire assembly is Type 304 stainless steel. A finite element model was developed to determine the stresses from various design loads. Figure 2 shows a line plot of the finite i element model. The SAP 4G07V finite elemen+ program was used to perform the stress j analysis. i 4.1. Static Loads The applied static loads on the jet pump assembly consist of the following: deadweight, hydraulic, fluid drag, loads due to flow induced vibrations, and thermal anchor displacements. Each of these loads are briefly discussed in the following sections.

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c GEN: clear Enero GE-NE-B13-01935-02, Rev.1 DRF # B13-01935

. 4.1.1'. Deadweight (DW)

The deadweight loading consists of the weight of the jet pump and the water enclosed in the jet pump. The stresses for this loading were calculated by applying one 'g' vertical acceleration in the finite element model of the jet pump assembly. For flaw evaluation 4 purposes, the stress from this loading is treated as primary. The designation for this load is-Deadweight: DW 4.1.2. Hydraulic Loads (F1, F2)

The hydraulic loads acting on the jet pump are calculated by summing the fluid momentum and pressure forces in the vertical and horizontal directions. This load definition considers any pressure differences between the annulus and the jet pump. Two hydraulic force values are ;;.51ated and applied to the jet pump. The first value is the longitudinal force in the riser which puts the riser welds in tension. The second value is the net vertical hydraulic load. The net vertical force is predominately caused by the pressure difference between the jet pump and annulus at the slip joint. Because the slip joint can not transmit a vertical load, the vertical load is carried through the riser causing a large bending moment on the riser. The following designations are used:

Normal / Upset Hydraulic Load: F1 Faulted Hydraulic Load: F2 <

For the flaw evaluation purposes, the stresses from the hydranlic loads are treated as primary.

The values of the hydraulic loads were calculated based on VY coreflow drive flow data and their application points are presented in the following table.

Hydraulic Loads Service Level Riser Radial Load Net Vertical Load (ib.) = Fx (lb) = Fy Normal / Upset (F1) 11392 4804 Faulted (F2) 11392 4804 4.1.3. Fluid Drag and Acoustic Loads The drag loads consist of the forces resulting from the fluid flowing in the annulus region past the jet pump. The flow F the annulus region during normal operation exerts some downward drag force on the jc pump. A postulated recirculation line break LOCA (suction side) subjects the jet pump to a drag force load in a tangential direction relative to vessel center line. A previous TRACG analysis (Reference 7) performed for a similar size plant l calculated the worst case flow velocities past the jet pump assembly. These velocities were adjusted for VY unique geometry. The worst case velocities correspond to a suction side 1

3 L.

GE Nuclear Energy GE-NE-B13-01935-02, Rev.1 DRF # B13-01935 recirc'ulation line break LOCA. Other breaks do not affect the jet pump nearly as severely due to the proximity to or other lines break size. The velocities, used in the calculation of the drag forces, corresp,nd to the jet pumps nearest to the suction nozzle. The following designation is used:

Drag Loads During Normal Or: ration: DRG1 Drag Loads During LOCA Condition: DRG2 1

For the purposes of flaw evaluation, the stresses from the circumferential drag loads are treated as primary and the radial drag (due to vortex shedding) loads are negligible due to their low frequency (jet pump fundamental frequency being more than 3 times greater than vortex shedding frequency, thus requiring no evaluation).

Drag Loads Service Level Total Tangential Load (ib.) Total Tangential Load (lb.)

on Two Diffusers = Fz on Riser = Fz Normal 578 372 Faulted 9521 2023 During the LOCA, Acoustic loads (ACI & AC2) precede the flow induced drag loads (DRG2). ACI & AC2 are short duration shock loads and the flow induced drag (DRG2) load follows acoustic load in time. Therefore, the time phasing of the two components of this load are considered during Faulted load combination.

The ACl loads are calculated based on acceleration of the annulus fluid producing a tangential force on the diffusers and the risers. These loads, varying with distance from the source (Recirculation outlet nozzle break location), are calculated from the basic principles (Reference 11). The AC2 load is an added hydraulic force due to a pressure spike in the recirculation inlet (Riser) based on a generically postulated values per Reference 7. For the purpose of flaw evaluation, these ACl & AC2 loads are considered primary. Total values of these loads are given in the Table below:

Acoustic Loads Load Total Tangential Load (lb) Total Tangential Load (lb) Total Radial Load (Ib) at Designation on Two Diffusers = Fz on Riser = Fz weld RS1 loca' tion = Fx ACI 93646 23800 -

AC2 -- --

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GENucle:r Energy GE-NE-B13-01935-02, Rev.1 DRF # B13-01935

. 4.1.4: Thermal Loads The three anchor points of the jet pump (the recirculation inlet nozzle, riser braces on the vessel, and the shroud support plate) grow vertically and horizontally at different rates due to differences in the materials (Iow alloy steel for the vessel, versus stainless steel for the jet pump). Also, these displacements are expected to vary during certain transients due to the differences in temperatures between the vessel and the shroud. The loads produced by these thermal anchor displacements and thermal expansion are treated as secondary. The following thermal displacements are considered:

Displacements during Normal Operation: NOD Displacements during Loss of Feed water Pump: LFWPD The following table shows the values of the displacement with respect to the recirculation inlet nozzle safe end.

1 Thermal Anchor Displacements Load Case Riser Brace Riser Brace Recirc Inlet Recirc Inlet Vertical "Y" Horizontal Vertical "Y" Horizontal (inch) "X" (inch) (inch) "X" (inch)

NOD 0.657 -0.062 0.297 -0.062 LFWPD 0.593 -0.057 0.269 -0.057 4.2. Dynamic Loads The applied dynamic loads on thejet pump assembly consist of seismic inertia only.

In order to obtain a realistic dynamic response from the jet pump model, the inclusion of hydrodynamic mass is mandatory. Hydrodynamic mass is also termed " apparent mass" and

" virtual mass". Physically, the hydrodynamic mass effect comes from the force which an accelerating solid object immersed in a fluid must impart to the fluid in order to cause fluid acceleration. Since this force is associated with solid body acceleration, the term

" hydrodynamic mass" is used to describe the effective mass.

4.2.1. Seismic Inertia l The seismic inertia loading consists of horizontal and vertical inertia forces acting on the jet pump due to seismic excitation of the RPV. The locations where the seismic excitation is imparted to thejet pump are the vessel recirculation inlet nozzle, the shroud support plate and the riser brace. An enveloped spectra was used which bounds the input accelerations at various attachment points as evaluated in the seismic analysis for Vermont Yankee Core Spray Flaw Evaluation (Reference 6). The horizontal spectra were applied in the radial and l

tangential directions and separately from the vertical spectra. The resultant response for j seismic loading was taken as the maximum of the two horizontal, plus the vertical response l spectra analyses results added by absolute sum. The following designations are used: j l

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i GE Nuchar Energy GE-NE-B13-0193S-02, Rev.1 DRF # B13-01935 l

. Operating Basis Earthquake Inertia: OBEI  :

Safe Shutdown (or Design Basis) Earthquake Inertia: SSEI OBE HorizoWal (X & Z) Response Soectra

! (Reference 6 Enveloce of Nodes 57 and 5) -

Period (Sec.) Acceleration (g) Period (Sec) Acceleration (g) Period (Sec) Acceleration (g) 0.002 0.850 0.050 0.850 1.00 0.850 OBE Vertical m Resoonse Soectra (Reference 6 Enveione ofNodes 57 and 5)

Period (Sec.) Acceleration (g) Period (Sec) Acceleration (g) Period (3ec) . Accebration (g) 0.002 0.305 0.050 0.305 1.00 0.305 SSE Horizontal (X & Z) Resoonse Soectra l (Reference 6 Enveloce of Nodes 57 and 5)

Period (Sec.) Accelerution (g) l Period (Sec) Acceleration (g) Period (Sec) Acceleration (g) 0.002 1.00 1 0.050 1.00 1.00  ! .00 SSE Vertical m Resoons: Soectra (Reference 6 Enveloce ofNodes 57 and 51  ;

Period (Sec.) Acceleration (g) Period (Sec) Acceleration (g) Period (Sec) Acceleration (g)

_ 0.002 0.305 0.050 0.305 1.00 0.305 4.3. Vibration Loads The flow induced vibration (FIV) loads are caused by turbulent flow in the piping exciting the naturai frequencies of the jet pump assembly. The method of calculating the vibration  ;

stress from the test data can be summarized as follows: '

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'l. Review the start up vibration data for VY(Reference 8) to determine the primary

structural modes ofinterest for the jet pump.

2. Using a finite element model of the jet pump, determine the, naturm frequencies, inode shapes, and modal stresses of all stmetural modes ofinici.est.

3. Normalize the modsl stresses such that they are equal to the measured strain data  ;

observed during start up testing.

4. Determine the normalized modal stress at each weld locaticn of interest in the jet.  !

pump assembly for each mode. The FIV stress is calculated by summing of 6 -

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GENuclear Enero GE-NE-B13-01935-02, Rev.1 DRF # B13-01935 l

i stresses for the individual modes. The numerical values of the peak to peak FIV are tabulated below.

Flow Induced Vibration Stress (FIV)

Weld FIV Stress

[ psi]

TS-2 303 TS-2a 379 i TS-2b 622 )

RS-1 832

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RS-2 429 RS-3 163 RS-4 416 RS-5 574 TR-1 50 IN-1 346 MX-1 1499 VJ-2 1798 MX-4 556 MX-5a 2141 l MX-5b 2141 DF-1 182 DF-2 222 DF-3 528 AD-1 1540 AL-2 1677 The following designation is used:

FlowInduced Vibration Stresses: FIV 7 ,

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GE Nuclear Energy GE-NE-B13-01935-02, Rev.1 DRF H B13-01935

5. Load Combinations and Stress Levels This section describes the manner in which the various loads were combined for the purpm of obtaining stress levels for the flaw evaluation. The limiting stress levels at the welds are then summarized.

5.1. Load Combinations The flaw evaluation methodology to be used makes the distinction oetween primary and seconda y stresses by specifying different safety factors. The flaw evaluation methodology (Reference 1) also 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. Load combinations are consistent with BWRVIP-41.

The following set ofload combinations were considered for the evaluation of normal / upset condition:

(1) DW(P) +Fl(P) + FIV(P) + DRG1(P) + NOD (S)

(2) DW(P) +Fl(P) + FIV(P) + DRGl(P) + LFWPD(S)

(3) DW(P) +Fl(P) + FIV(P) + DRGl(P) + OBEl(P) + NOD (S)

Note that the letter in the parenthesis indicates whether a load is primary or secondary as d: fined by the ASME Code. The set ofload combinations used for the Emergency / Faulted conditions are the following:

(4) DW(P) +F2(P) + SSEI(P) + AC1(P) + NOD (S)

(5) DW(P) +F2(P) + SSEI(P) + AC2(P) + NOD (S)

(6) DW(P) +F2(P) + SSEI(P) + DRG2(P) + NOD (S) 5.2. CalculatedStress.2evels The forces and moments at various nodes in the model for all of the load sources were calculated using the SAP 4G07V finite element code (Reference 2). These forces and moments were then combined to obtain the total forces and moments for a given load combination. The absolute values of the loads are summed, except for the deadweight and the hydraulic load, which are summed algebraically, then added absolutely. Thus, 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 of expansion stress. As described later, the flaw evaluation methodology uses the primary membrane (Pm), primary bending (Pb) and the  !

expansion stress (Pe).

l The calculated values of Pm, Pb, and Pe stress levels at the circumferential weld locations are summarized in the following two tables for the governing load combinations. ,

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Summary of Calculated Stress in Governing Normal / Upset Load Con bination Weld ID P. P. P, l (Figure 1) (psi) (psi) (psi)

TS-2 772 2489 327 l TS-2a 772 2405 334 TS-2b 1021 2575 586 RS-1 1021 1195 683 RS-2 351 803 359 RS-3 381 321 56 l RS-4 360 1330 790 RS-5 360 1496 953 TR-1 136 68 39 IN-1 114 659 320

_ NIX-1 55 1282 520 MX-2 ,39 1743 1868 MX-4 5 581 739 l MX-5a 11 1524 1045 MX-5b 11 1524 1045 l

DF-1 10 209 451

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DF-2 30 571- 1350 l DF-3 57 1062 1918 AD-1 68 1872 2280 l AD-2 72 2039 2386 Note: The third load case was the governing ;oad combination for Normal / Upset Conditions.

The stress le"els in the preceding tables were used in the allowable flaw evaluations as

! described in the next section.

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GE Nuclear Enero GE-NE-B13-01935-02, Rev.1 DRF # B13-01935

.Sumniary of Calculated Stress in Governing Emergency / Faulted Load Combination Weld ID P. P, P, (Figure 1) (psi) (psi) (psi)

TS-2 777 6110 327 TS-2a 777 5785 334 TS-2b 1027 5845 586 RS-1 1848 809 683 RS-2 361 2186 359 RS-3 388 2365 56 RS-4 370 7406 790 RS-5 369 7745 953 TR-1 139 184 39 IN-1 121 1619 320 '

MX-1 58 7026 520 MX-2 41 3685 1868 MX-4 5 1099 739 -

MX-5a 11 2207 1045 MX-5b 2207 1945

) 11 DF-1 10 288 451 DF-2 30 1354 1350 DF-3 57 3436 1918 -

AD-1 68 5799 2280 AD-2 72 6562 2386 Note: The fourth load case was the governing load combination for Emergency / Faulted Conditions, except RS-1 where the fifth load case governed.

The calculations presented are for one flaw per riser and do not inc ..de coupling effects of multiple flaws. An analysis was performed to study the effects of multiple flaws.

Circumferential f;aws of up to 180 degrees have little effect on the stiffness of the system and would not change the results of this analysis.

6. Fracture Mechanics Evaluation The limit load methodology, as specified in BWRVIP-41, was used in calculating the allowable flaw lengths. This methodology is first described followed by the results of allowable flaw evaluations.

6.1. LimitLoadMethodology Circumferential Flaws Consider a circumferential crack of length,1 = 2Rcx and constant depth, d. In order to determine the point at which limit load is achieved, it is necessary to apply the equations of i

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GE Nudear Eneru GE-NE-B13-01935-02, Rev.1 DRF # B13-01935 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 Figure 3 where the maximum stress is the flow stress of the material, cf. Equilibrium oflongitudinal forces and moments about the axis gives the following equations:

i p = [(n- ad/t) - (P,/cr)n]/2 (1) ]

Pi = (2cf /n)(2 sin p - d/t sin a) (2) l Where, t = pipe thickness, inches a = crack half-angle as shown in Figure 3 d = crack depth, inches R = pipe ra Uus, inches p = angle *.iat defines the location of the neutral axis Z = weld type factor i

P,= piping expansion stress, psi

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P = primary membr"e stress, psi P, = primary bending stress, psi Pi = failure bend!ng stress, psi of = 3Sm, flow stress, psi Sm = allowable stress, psi The safety factor (SF) is then incorporated as follows:

Pi = Z*SF (P, + P,+ P,% - P. O)

The P, and P, are primary stresses. P, is a secondary stress and includes stresses from ali displacement-controlled loadings such as thennal expansion, seismic anchor motion, etc. All three quantities are calculated from the analysis of applied loading. The safety factor value is 2.77 for normal / upset conditions and 1.39 for emergency / faulted conditions. The crack angle (2a) is the value for which equation 2 is equal to equation 3.

Z Factor The test data considered by the ASME Code 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 welds produced by a process using the flux had lower toughness. To account for the reduced toughness of the flux welds (as compared to non-flux welds) the Section XI procedures (Reference 3) preacribe a penalty factor, called a 'Z' factor. The 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 weld-base metal interface. The expressions for the value of Z factor in Appendix C are given as the following:

Z =

1.15 [1 + 0.013(OD-4)] for SMAW

=

1.30 [1 + 0.010(OD-4)] for SAW I1

GENuclear Energy GE-NE-B13-01935-02, Rev.1 DRFM B13-01935 where OD is the nominal pipe size (NPS) in inch-s. The procedures of Paragraph IWB-3640 require the use'of OD = 24 for pipe sizes less than 24-inches. This approach is very conservative and, therefore, the use of actual NPS (OD=10 inches) was made in calculating the 'Z' factor. This approach is considered reasonable and has been accepted by Section XI

[ Item # 96-169, ISI-95-32, Mechanical Engineering Magazine September 1996, p 128]. The welding process used was a combination of shielded metal arc type (SMAW) or submerged are welds (SAW) and gas tungsten arc weld type (GTAW). Since a non-flux process was specified for at least part of the weld, it must be assumed that the welds are flux welds (SAW). The Z-factor is thus:

Z io.;,,, = 1.30 [1 + 0.010(10-4)] = 1.38 i

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l GE Nudear Energy GE-NE-B13-01935-02, Rev.1 DRF # B13-Ol935 6.2. Allowable Flaw Length Calculation The stresses from the table in the preceding section were utilized to determine the acceptable ,

through-wall flaw lengths. The acceptable flaw size was determined by requiring a safety I factor on stress. The flow stress was taken as 3S,, (S,,, = 16.9 ksi for Type 304 stainless steel at 550*F. As specified in Reference 1, 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 end-of-cycle allowable flaw lengths are tabulated in the following table.

End-of-Cycle Allowable Flaw Lengths Based on Outside [ inches]

Weld Flaw Length Weld Designatior. (Limit Load) Description TS-2 19.56

• Thermal Sleeve TS-2a 19.90 Thermal Sleeve TS-2b 16.17 Thermal Sleeve RS-1 18.62 Riser Elbow ,

RS-2 21.43 Riser Elbow RS-3 21.43 Riser to Transition Piece j RS-4 15.66 Riser Sleeve RS-5 15.30 Riser Sleeve TR-1 37.63 Transition Piece IN-1 14.15 Elbow and Mixer Nozzle MX-1 10.45 Mixer Flange to Barrel MX-2 11.68 Barrel to Adapter MX-4 15.48 Adapter to Flare MX-5a 14.43 Adapter Sleeve MX-5b 14.43 Adapter Sleeve DF-1 21.34 Diffuser Collar to Shell DF-2 30.16 Diffuser Shell to Tailpipe DF-3 25.74 Diffuser Tailpipe to Adapter AD-1 22.18 Adapter Top to Bottom AD-2 21.19 Adapter Bottom to Support Plate l

These allowable values are the end of cycle values and they do not consider the crack growth E due to IGSCC or fatigue. The crack growth is discussed in Section 6.3. The crack growth due to these mechanisms must be added to the existing crack length at each end to determine whether the projected end of cy cle length is acceptable.

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. 6.3. ' Crack Growth Evaluation j Prior crack growth analyses performed for BWR shroud and cere spray line indications have I

used an IGSCC crack growth rate of 5x10-5 inch / hot hour. This crack grow th rate translates into a crack length increase per eighteen month cycle of(12,000 hrs x 5x10-5) or 0.6 inch at j each end of an indication. Thus, the projected length, if of any indication whose current l length at the time of inspection is, Ip, would be (lp + 0.6 x 2) inches. A factor of 2 in the

! preceding parenthesis is to account for the growth at each end of the indication.'  !

' In addition to IGSCC growth, fatigue growth due to FIV is evaluated for the indications. The expected fatigue growth is a strong function of the crack size and cannot be determined until the crack is characterized. With a characterized crack, the stress intensity factor (AK) can be computed and compared to the threshold stress intensity factor (AKth). The AKth is the value l at which fatigue crack growth for high cycle stress becomes significant for high cycle events i

! and must be considered. At values below AKth, fatigue growth can be neglected. For 304 l stainless steel, Reference 12 reports a AKth value of 5.5 ksidin, at R ratios of < 0.5. A l threshold value of 5.0 ksiVin is used here to cover vibration loading and R ratio uncertainties.

The stress intensity factor calculation method described in Reference 13 was used to calculate  !

l- the~ AK values. The calculated FIV threshold crack size for all weld locations are tabulated below. In the cases where the values are less than the limit load method allowable, fatigue l

crack growth would begin prior to reaching the limit load crack sires; otherwise, fatigue l crack growth need not be considered.  !

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GE Nudear Enerv GE-NE-B13-01935-02, Rev.1 l DRF # B13-01935 Flow induced Vibration Stress (FIV)

Weld FIV Stress FIV Flaw Size

[ psi) [in]

TS-2 303 18.75 TS-2a 379 16.26 TS-2b 622 10.38 i RS 832 8.16 RS-2 429 13.56 RS-3 163

• RS-4 416 13.84 RS-5 574 11.03 TR-1 50

• IN-1 346 11.37 '

MX-1 1499 3.99 MX-; 1798 3.35 MX-4 556 9.93 MX-5a 2141 2.53 MX-5b 2141 2.53 DF-1 182 18.45 DF-2 222 22.21 DF-3 528 12.61 AD-1 1540 4.47 AD-2 1677 4.00

• FIV does not lima acceptance criterion of as-found flaw size Thermal expansions due primarily to system start up and shutdown were also considered in the evaluation. The AK value v.a calculated using the methods of Reference 13 and the expansion stress (P,) from Section 5.2. The AK value for thermal expansion is less than 20 ksi (in)3. This corresponds to about a 0.000050 inch per cycle based on the data in Reference

14. The fatigue crack growth due to thermal expansion stress cycling is negligible due to the limited number of cycles. Thus, total crack growth due to system thermal transients is insignificant compared with IGSCC cracs growth.

The flaws at Vermont Yankee are evaluated in the next section along with values of flaw length conesponding to threshold stress intensity, AK..

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GE Nuclear Energy GE-NE'-B13-Ol935-02, Rev.1 DRF #,B13-01935

7. Evaluation ofIndications 7.1. Procedurefor Evaluation ofIndications using Handbook Any indications at HAZ of the welds in Figure I can be evaluated by use of this handbook in the following manner.

With a crack growth of 0.6 inches at each of both ends of the indications, total actual flaw size at the end of cycle (EOC) including NDE uncertainty values of 0.38" for the UT ,

inspection (Reference 10) are as shown in the Table below.

{

Flaw Sizes for any Arbitrary Weld After one Cycle Indication Indication IGSCC NDE Adjusted Allowable Allowable Location Measured Growth untertainty Size Flaw' Size Flaw Size Weld Length (Inches) (Limit Load) (FIV)

TS-2 A" 1.2" 0.38" A + 1.58 19.56 18.75 RS-1 B" 1.2" 0.38" B + 1.58 18.62 8.16 MX-1 C" 1.2" 0.38" C + 1.58 10.45 3.99 The NDE uncertainty values are based on upcoming revision to BWRVIP-03 per Reference

10. The NDE uncertainty here is maximum (for 60 deg. shear probe) out of two methods i.e.

60 deg. shear probe or 45 deg probe. If 45 deg shear probe is used instead, then adjusted size een be further reduced by 0.068",i.e., by 68 mils. If the adjusted flaw lengths after one cycle are less than the allowable flaw sizes, then the existing flaws are acceptable for continued '

operation without any modifications. l 7.2. Leakage Calculation Leakage from postulated through-wall flaws with length equal to the allowable end of cycle (EOC) flaw size are calculated in this section. The leakage rate through an adication was estimated assuming incompressible Bernoulli flow through the crack:

Q = CAg2g,AP /p (5) where, Q = Leakage C= flow coefficient A= arca p= mass density of fluid at 520 F AP = pressure difference acess the pipe .

A AP value of 123 psi was used based on VY specific process data. This is the AP design value for the pressure difference between the annulus and jet pump riser. This value is bounding for normal operating conditions.

16

GE Nuclear Energy GE-NE-B13-Oi935-02, Rev.1 DRF # B13-01935 Leak' rate from the through-wall indications in the riser can be estimated using the preceding j equation with the value of flow coefficient, C, 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 ]

l

]

i to calculate a conservative value c ' crack opening displacement, S, and assume the crack j

opening configuration to be like a r :tangular slot with one side being the crack length,2a, l

and the other side as the crack operiing displacement. The opening displacement is j

calculated using S = 4sl/E (Reference EPRI Report NP-2472, Vol. 2, D-2) where 1 is one half

)

the crack length (allowables calculated in section 6.2), s is the applied streu, and E is i Young's modulus. If the calculated crack openings are less than 10 mils, then a conservative crack opening of 10 mils was used. The crack opening area is then simply:

A = 2a (S) (6)

The table below shows the maximum leakage rates for flaw sizes of 25%,50% and 100% of the limit load allowable flaw lengths calculated in Section 6.2. Leakage for any other percentage of allowable flaw size length can be linearly interpolated. Acceptance criterion for leakage values is outside the scope of this handbook.

)

Leakage versus Percentage of EOC Allowable Flaw Length for Normal Operation Weld Leak Rate (gpm) @ % of allowable length Location 25 % 50 % 100 %

TS-2 23.5 47.0 94.0 TS-2a 23.9 47.8 95.6 TS-2b 19.4 38.9 77.7 RS-1 22.4 44.7 89.5 RS-2 25.7 51.5 103.0 RS-3 25.7 51.5 103.0 RS-4 18.8 37.6 75.2 RS-5 18.4 36.7 73.5 TR-1 45.2 90.4 180.8 IN-1 17.0 34.0 68.0 MX-1 12.6 25.1 50.2 MX-2 14.0 28.1 56.1 MX-4 18.6 37.2 74.4 MX-5a 17.3 34.7 69.3 MX-5b 17.3 34.7 69.3 DF-1 25.6 51.3 102.5 DF-2 36.2 72.5 144.9 DF-3 30.9 61.8 123.7 AD-1 26.6 53.3 106.6 AD-2 25.4 50.9 101.8 17

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GENuclear Energy GE-NE-B13-01935-02, Rev.1 DRF # B13-01935

8. S'ummary & Conclusions 1

A flaw evaluation, consisting of stress and fracture mechanics analyses of the Vermont Yankee jet pump circumferential welds was conducted to develop a flaw evaluation handbook. The procedures of BWRVIP-41, were used as a guide in determining the allowable flaw lengths. End-of-cycle allowable flaw lengths were calculated at all circumferential weld locations. The results presented in this report can be used along with l

consideration ofIGSCC crack growth to disposition any indications detected during future inspections of thejet pumps at Vermont Yankee.

End-of-Cycle Allowable Flaw Lengths Based on Outside Diameter Weld Flaw Length Flaw Length Weld Designation (inch) (FIV) Description TS-2 19.56 18.75 Thennal Sleeve TS-2a 19.90 16.26 Thermal Sleeve j TS-2b 16.17 10.38 Thermal Sleeve RS-1 18.62 8.16 Riser Elbow RS-2 21.43 13.56 Riser Elbow RS-3 21.43 21.43* Riser to Transition Piece RS-4 15.66 13.84 Riser Sleeve RS-5 15.30 11.03 Riser Sleeve TR-1 37.63 37.63* Transition Piece IN-1 14.15 11.37 Elbow and Mixer Nozzle MX-1 10.45 3.99 Mixer Flange to Barrel MX-2 11.68 3.35 Barrelto Adapter MX-4 15.48 9.93 Adapter to Flare MX-5a 14.43 2.53 Adapter Sleeve MX-5b 14.43 2.53 Adapter Sleeve DF-1 21.34 18.45 Diffuser Collar to Shell DF-2 30.16 22.21 Diffuser Shell to Tailpipe  !

DF-3 25.74 12.61 DiffuserTailpipe to Adapter _ l AD-1 22.18 4.47 Adapter Top to Bottom '

AD-'2 21.19 4.00 Adapter Bottom to Support Plate

• FIV does not limit acceptance criterion of as-found flaw size.

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GE Nudear Enero GE-NE-B13-01935-02, Rev.1 DRF # B13-01935 1

9. References

[1] BWR Jet Pump Assembly Inspection and Flaw Evaluation 'iuidelines, (BWRVIP-41), EPRI Report No. TR-108728, October 1997.

[2] SAP 4G07V Users Manual, NEDO 109W, Addenda 1,199(.

[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 and Appendix C.

[4] Ranganath, S. and Mehta, H. S., " Engineering Methods for the Assessment of Ductde Fracture Margin in Nuclear Power Plant Piping," Elastic Plastic Fracture: Second Symposium, Volume II - Fracture Resistance Curves and Engineering Applications, ASTM STP 803, C.F. SI<ih and J. P. Gudas, Eds., American Society for Testing and Materials,1983, pp. II-309 - II-330.

[5] Vermont Yankee Jet Pump Drawing No. 730E438G7, G8 Rev. 9 Vermont Yankee Jet Pump Riser Drawing No. 730E770 G3, G4 Rev. 3 Vermont Yankes Jet Pump Elbow Drawing No. I17Cl475P1 Rev.1 Vermont Yankee Reactor Vessel Drawing No. 919D294 Rev. 8 Vermont Yankee Riser Brace Drawing No. I17C4614G1 Rev. 0 Vermont Yankee Reactor Assembly Drawing No.104R940 Rev.10

[6] DRF # B13-01805, Report Number B13-01805-66, Vermont Yankee Core Spray Flaw Evaluation, September 1996.

[7] DRF # A71-00014,Index C-13, TRACG Acoustic Loads Analysis.

[8] Vermont Yankee Vibration Test Result" - Transmittal letter from E.J. Romesberg to W.J Neal, ERJ-96-74, June 12,1974.

[9] Vermont Yankee Reactor Thermal Cycles Drawing No. 729E762 Rev. O, Nozzle Thermal Cycles Drawing No. 761E708 Rev.1.

[10] " Review of Peach Bottom jet pump riser VT and UT inspections", Letter from Greg Selby of EP.RI NDE center to Tom Hinkle of PECO Energy Company, dated Octoter 28,1997.

[11] " Fluid Acceleration Forces on Jet Pumps and Risers", by Frederick Moody, General Electric Company, February 1998, DRF # B13-01915, Section E.

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I GENuclexr Enero GE-NE-B13-01935-02, Rev.1 i- DRF# B13-01935

[12]' Barsom, J.M., Rolfe, S.T., Fracture and Fatigue Control in Structures, Second -

Edition, Prentice-Hall, Inc.

'[13] Zahoor, A., Ductile. Fracture Handbook, Prepared. for Novetech Corporation and l EPRI, EPRI Report Number NP-6301-D.

[14] Hale, D.A., Yuen, J., and Gerber, T., Fatigue Crack Growth in Piping and RPV Steels ;

In Simulated BWR Water Environment, GEAP-24098, January 1978. 1

[15] Vermont Yankee Final Safety Analysis Report, FSAR Rev.14, November 1997.

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GENuclearEnergy GE-NE-B13-01935-02, Rev.1 DRF M B13-01935 E

I vEnMourvaux E JET PUMP A$$EMBLY 4

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13 13 13 13 13 13 13 13 Il !!' '

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22

GE Nucleer Enero GE-NE-B13-01935-02, Rev.1 DRF # B13-01935 l

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I Nominal Stress 1 in the Uncracked I Section of Pipe Crack Length = 2Ra p,,p, 4 6 Flow Stress, e,

, 4__+ j i

4- >

I d 4- I I 4- _ _ _ - _ . _ . I al4-/ 4- I A i / 4- I

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Neutral Axis '  !

p, 4 4_

Stress Distribution in P. = Applied Membrane Stress in Uncracked Section the Cracked Section at P. = Applied Bending Stress in Uncracked Section the Point of Collapse Figure 3. Stress Distribution in a Cracked Pipe at the Point of Collapse 23

GE Nuclear Enero GE-NE-B!3-01935-02, Rev.1 DRF # B13-01935 AP'PENDIX: RESPONSE TO NRC REQUEST FOR ADDITIONAL INFORMATION REGARDING JET PUMP RISER WELD INSPECTIONS AT VERMONT YANKEE l

. l 24

r t vamos r vnue secu:..m Powne cowoimios

, , Docket No. 50-271 RVY 99-43 l

l 1

1 Attachment 3 i

Vermont Yankee Nuclear Power Station l Jet Pump Riser Circumferential Weld Inspections and Flaw Evaluation Projected Two-Cycle Flaw Length and Leakage Rate Calculations l

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