ML18040A155

Suppl 1 to GE-NE-B13-01739-04, Nine Mile Unit 1 Shroud

ML18040A155
Person / Time
Site:	Nine Mile Point
Issue date:	02/24/1995
From:	Marisa Herrera, Mehta H GENERAL ELECTRIC CO.
To:
Shared Package
ML17059A702	List: ML17059A701 ML18040A155
References
GE-NE-B13-01739, GE-NE-B13-01739-04, GE-NE-B13-1739, GE-NE-B13-1739-4, NUDOCS 9503100129
Download: ML18040A155 (37)

Text

ATTACHNfNT 3 GE-NE-B13-01739-04 SUPPLEMENT 1 CLASS III DRF B13-01739 SUPPLEMENT 1 TO NINEMILEUNIT 1 SHROUD GENE - B13 - 01739 - 04 FEBRUARY 24, 1995 Prepared by:'N'vt~

M.L. Herrera T.E. Gleason Wzyy Verified by:

H.S. Mehta f

9503100i29 950228 PDR ADOCK 05000220 P

DR

l J

GE.NE-B1341739R svppmMENT 1 CONTENTS

1. INTRODUCTION

2. STIFI'NESS CALCULATIONS Page 2.1 Shroud Stiffness 2.2 Stabilizer StifRess

3. LOADCALCULATIONS 3.1 Preload 3.2 Separation Loads 3.3 Normal Conditions 3.4 Upset Conditions

4. VALIDITYOF SAFETY ANALYSIS

5. CONCLUSIONS 8

9 9

10 12 5.1 Normal Operation 5.2 Upset pressure operation 5.3 Action Following an Upset Event 12 12 12 LIST OF FIGURES Figure 2.1:

Example ofCrack Simulation Figure 2.2a:

ANSYS Mesh Figure 2.2b Shroud Model 13 14 Figure 2.3:

Ring Displacement With Cracks on Ring Side ofWeld 15 Figure 2.4:

Ring Displacement With Cracks on Shell Side ofWeld 16 Figure 2.5:

Ring Displacement WithNo Welds 17

GF NE.B13417394 supprzmm 1

LIST OF TABLES Page Table 2.1:

Core Shroud Conditions Evaluated Table 2.2:

Shroud Stiffness Results Table 3.1:

Compressive Loads at Welds H6B and H7 Table 3.2:

Estimated Pressures for Upset Conditions

t

GE-NE-B1341739<

sUpprzxmrr 1

I. INTRODUCTION This document is a supplement to the Nine MileUnit 1 (NMP-1) Shroud Repair Hardware Stress Analysis. The evaluation in this supplement is in addition to the referenced stress analysis and is consistent with the results or conclusions.

The tie rod loads calculated in this evaluation are bounded by the loads used in the reference analysis.

The purpose ofthis supplement is to further evaluate the possibility ofweld crack separation during normal and upset operation.

The original stress analysis show there was no weld crack separation in the lower shroud welds (H6B &H7) during normal and upset operating conditions, but the spring constants oftwo stru'ctural components (top guide ring and conical support) were determined to be non conservative.

This evaluation supplements the original stress report with an independent ANSYS analysis ofthe top guide support ring for two 360'hrough wall crack conditions (H2 &, H3). This evaluation also includes the stiffnesses ofthe shroud conical support and the tie rod lower support assembly.

The stiffnesses in the original stress report remain valid since they yield the most conservative results for stress evaluation.

2. STIFFNESS CALCULATIONS 2.1 Shroud Stiffness This section presents the evaluation ofthe NMP-1 core shroud stiffness calculation.

The core shroud stiffness, including the conical support, was determined for various conditions.

The conditions considered include through-wall 360'racking simultaneously't H2 and H3. The analysis presented here does not postulate cracking at H8, but covers cracking at all other welds (H1 - H7), IfH8 is assumed to be cracked the load path willbe through the H8 brackets whose stiffness is very high and do not affect the overall shroud stiffness. The shroud conditions considered in this evaluation are summarized in Table 2-1.

Since welds H2 and H3 affect the shroud stiffness, they are specifically cited in Table 2-1.

Welds H2 and H3 include included the 0.6" filletwelds. The detail shroud drawings show a fullpenetration weld between the shroud cylinders and ring with a 0.63 filleton the ring side. Fillet welds with a 0.6 inch legs are used for conservatism.

Case 1 postulates a crack in through the weld at the ring and assumes the ring pivots about the weld material

T1

GE-NE B13%17394 SUPPLEMENT 1 attached to the shell.

Case 2 postulates the crack in the shroud cylinders above the weld and takes credit for the weld material attached to the ring.

In Case 3 is a conservative bounding case that assumes there are no filletwelds. Allthree cases take no credit for the 8 top guide alignment brackets, the 4 top guide support pads or the 4 core spray line support brackets.

Allofthese items are welded to the top guide support ring and to the shroud shell sections which adds to the stiffness ofthe ring section.

Case 3 is evaluated for information only as a bounding condition and is not considered a realistic condition. The existence ofCase 3 means the shroud was fabricated in severe violation ofthe design requirements and there is no evidence to support such a condition.

Afourth case is included for comparison with cases 1 through 3. Case 4 is the as built condition. There are no cracks and the filletwelds are included in the model.

Table 2-1 Core Shroud Conditions Evaluated Case Details Cracks at H2 and H3 on ring side Cracks at H2 and H3 onshell side H2, H3 Failed, No welds considered No failed welds, with fillets Analysis The analysis ofthe core shroud axial stiffness was comprised ofi) developing a finite element mode, ii) applying loading to the top ofthe core'shroud, iii)Determining the stiffness using the resulting vertical displacement and applied loads.

Finite Element Analysis Afinite element model was developed using the ANSYS computer program (Reference 2-1). The model was comprised oftwo-dimensional axisymmetric quadrilateral elements (STIF 42). The core shroud flange and top guide support ring were included in the model

'long with the shroud cylinders and the support cone.

GE.NE.B13417394 supprzMEm 1

The cracks were modeled at the H2 and H3 locations by including two sets ofcoincident nodes.

One set ofnodes corresponded to the lower row ofelements, and the second set corresponded to the upper row ofelements.

The assumed contact point between the upper and lower portion ofthe shroud at each crack location was modeled by giving two elements a common node.

This modeling is illustrated in Figure 2-1.

Shroud Shell Contact Point-Common Node Coincident Independent MNodes Along Crack Top Guide Support Ring Figure 2-1 Example ofCrack Simulation Figure 2-2a shows the finite element model and Figure 2-2b shows the detail for the H2/H3 weld configuration. The actual crack assumptions are more clearly seen in Figure 2-3 (Case 1), Figure 2-4 (Case 2) and Figure 2-5 (Case 3).

Applied Load The loading was applied as nodal forces in the downward vertical direction at the top of the shroud flange. Atotal of 11 nodes are on this surface and 1000 lbs (per radian) was

'pplied at each node.

Thus, the total force is 27'(11000) = 69,115 lbs.

dl I

N A

GE-NE-813417394 svppU.MENT 1 Stiffness Calculations The stiffness ofthe shroud was calculated by dividing the total applied force by the average nodal vertical displacement at the shroud Qange (F=kh). Table 2-2 summarizes the results ofthe finite element cases and resulting shroud stiffness (Ks).

Figures 2-3 through 2-5 shows the exaggerated displacement which demonstrate the behavior ofthe cracked shroud and show more clearly the crack assumptions.

Case Table 2 Shroud Stiffness Results Average Vertical Displacement (in.)

0.01022 0.01092 0.02556 0.001006 Ks, Stiffness (F/h) lb/in 6.76x10 6.33xlo 2.70x10 68.7x10 2.2 Stabilizer Stiffness The original stress report did not include the stiffness oftie rod lower support assembly in the calculation ofthe overall assembly stiffness.

This supplement includes the lower support assembly stiffness.

The stiffness ofthe lower support is analyzed as a series of springs including the toggle bolts, the lower support hook section and the lower support base section.

Each component is divided into sections that are analyzed as spring elements with the K= AE/Lrelationship.

The toggle bolts have the lowest stiffness in this series of springs controls the overall stiffness.

The net lower support stiffness, including the toggle bolts (Kls) is calculated at 13 x 10 Ibfin.

The stiffness ofthe upper support, tie rod and C-spring (Ktr) is calculated at 514,000 lb/in. The combined stiffness ofa single stabilizer assembly including the lower support is Kst.

1/Kst = 1/Ktr+ 1/Kls 1/Kst = 1/514,000+ 1/13 x 10 Kst = 494,450 lb/in The stiffness ofthe four stabilizer assemblies, Kcc= 4 x 494,450 lbfin= 1.978 x 10 lb/in

GE-NE B13417394 SUPPLEMENT 1

3. LOADCALCULATIONS 3.1 Preload The preload is calculated for 4 different conditions ofthe top guide support ring resulting in four different spring constants.

At normal thermal conditions, the unrestricted thermal expansion ofthe shroud, dJ., exceeds the stabilizer by 0.155 in. The four different conditions ofthe top guide support ring are summarized below.

Case

1. Welds H2 and H3 have a 360'hrough wall crack on the ring side ofthe filletweld. (Ks = 6.76 x 10'bfin)

Case 2. Welds H2 and H3 have a 360'hrough wall crack on the shroud shell side ofthe filletweld. (Ks=6.33 x10 lb/in)

Case 3. Welds H2 and H3 have a 360'hrough wall cracks and there is no fillet weld. (Ks =2.70 x 10 lb/in)

Case 4. Welds H2 and H3 are not cracked.

(Ks = 68.7 x 10'lbfin)

The preload for each condition is calculated as shown below and Table 3.1 shows the total compressive load on welds H6B and H7. This preload calculation conservatively neglects the total mechanical preload of 12,000 lb applied by tightening the tie rod nuts.

P = [(Ks x Kcc) /(Ks+ Kcc)] x dL, The net compressive load at welds H6B and H7 is a combination ofthe thermal preload and the weight, less buoyancy, ofthe internals.

The net weight ofthe internals acting against the reactor pressure is calculated in the referenced stress report as 174,910 lb.

~"

GE-NE-B13417394 svpprzMENT 1 TABLE3.1 COMPRESSIVE LOADATWELDS H6B ANDH7 CASE NUMBER THERMAL PRELOAD, LB 237 188 233,596 176,954 298,010 NET WEIGHT, LB 174 910 174,910 174,910 174,910 TOTAL LOAD,LB 412,098 408,506 351,864 472,920 3.2 Separation Loads The loads tending to separate the lower shroud welds are applied by the steady state pressure drop across the core plate and the shroud head.

Seismic loads are not included in this crack separation analysis. The separation loads during normal operation with 105%

flow and the design basis upset conditions at 100% flow are calculated below. The NMP-1 design basis is 100% core flow and 105% core flow is used to provide margin in the analysis.

The liftload is the sum ofthe LQ"s times the areas for each event.

The applicable areas are from the referenced stress report.

Effective Core Plate Area = Acs = 12,346 in Effective Shroud Head Area =Ash = 24,328 in The liftload = Acs x pcs + Ash x irish 3.3 Normal Conditions During normal operation at 105% core flow, the core support pressure drop, pcs, is 15.9 psi and the shroud head pressure drop, leash, is 5.9 psi. The calculated liftload is 339,836 lb.

The results show there is no crack separation for all the cases considered.

The compressive thermal preload plus weight ofthe internals exceeds the 339,836 lb load tending to separate the welds for all load cases.

f A

GE-NF 813%17394 SUPPLEMENT 1 With welds H2 and H3 not cracked, the ring stifRess is calculated at 68.7 x 10'b/in and the thermal preload is calculated to be 298,101 lb. The load on each stabilizer assembly is 74,503 lb which is less than the bounding load of79,670 lb used for the original structural analysis.

3.4 Upset Conditions At design bases upset pressure with 100% core flow, the core support pressure drop, pcs, is 16.9 psi and the shroud head pressure drop, dZsh, is 8.3 psi. The calculated lift load is 410,569 lb.

There willbe no separation for Case 1 since the compressive load is 412,098 lb and a separation load is 410,569 lb.

There willbe no separation for Case 2 ifcredit is taken for the 12,000 lb mechanical preload applied by the tie rod nuts. However, ifone excludes the mechanical preload, the maximum difference between the separation load and the thermal preload plus dead weight is 2063 lb. The maximum crack separation with this load is 0.001 inch. Separation is calculated be dividing the excess load by the spring constant ofthe four stabilizer assemblies (1.978 x 10 Ibfin) This separation is acceptable and there is no need for any inspections or other operator actions followingthe upset pressure events.

Crack separation is predicted for the Case 3 conditions. The thermal preload plus weight ofthe internals result in a 351,864 lb compressive load with a 410,569 lb load tending to separate the welds. The separation load exceeds the preload by 58,705 Ib, and the maximum separation is 0.03 inches. It should be emphasized that this case is overly conservative since it assumes that there are no filletwelds.

The resulting shroud stifRess is much higher and the thermal preload willbe higher if credit is taken for the 8 top guide alignment brackets, the 4 top guide support pads and the 4 core spray pipe brackets.

These items are welded to the top guide ring and to the shroud shell section and willincrease the shroud stifRess at the top ring section.

The upset pressure drops across the core support and the shroud head are also calculated on the high side. Table 3.2 show a comparison ofthe design basis upset pressures and the calculated pressures for several upset events.

Similar upset pressures are not available for.,

the 105% flow conditions, but the table gives an idea ofthe margins involved in the pressure calculations.

10

GE-NE-B134 17394 SUPPLEMENT 1

TABLE3.2 ESTIMATED PRESSURES FOR UPSET CONDITIONS COMPONENT Shroud Head Upper Shroud Core Plate Shroud Support NORMAL 100%

FLOW 5.5 4.7 14.5 19.2 UPSET DESIGN BASIS 8.3 7.1 16.9 21.6 RECIRC.

FLOW RUNOUT 5.8 5.1 15.2 20.3 PR. REG.

FAILURE 7.5 6.7 14.5 21.2 4.

VALIDITYOF SAFETY ANALYSIS The safety evaluation (GE-NE-B13-01739-5) was reviewed for areas in which a reduction in shroud stiffness would alter the results and/or conclusions.

The following is the result of this review:

1.

Part B.2.1:

Safety Design Basis (Conformance)

The maximum temporary vertical displacement willincrease slightly (-0.02 inch). This negligible increase will not alter the conclusion that the ECCS piping willperform its intended safety function. The design margin for this event was originally approximately 250%.

2.

Part B.3.1:

Flow Partition (Conformance)

The conclusion that this design verifies acceptable leakage through the flow partition resulting from weld separation during accident and transient events that meet the normal operations is still valid.

The maximum potential weld separation predicted for upset pressure conditions has no impact on the consequences of the previously analyzed NMP1 FSAR anticipated operational transients.

The anticipated operational events which result in limiting safety conditions for MCPR, LHGR, and Over-Pressure do not lead to weld separation during the event.

This is the result of constant or decreasing shroud pressure differences during these limiting events.

The consequences of anticipated operational events which result in weld separation, and therefore, shroud leakage are less severe than when separation does not occur.

This is the result of reduced core moderation, because of the reduced core flow, leading to lower core power generation.

Therefore, the consequences of anticipated operational transients are unchanged by potential weld separation.

11

I'

OE-NE-B13417394 supprzMENT 1 Item 1 above is the only area in which a non-conservative result is reported. Allother areas ofthe Safety Evaluation either have no impact or use refer to pre-loads which bound the lower pre-load values calculated by a reduced shroud stiflness.

There are no changes to any ofthe conclusions ofthe safety evaluation.

5. CONCLUSIONS 5.1. Normal Operation The results show there willbe no crack separation for all cases for normal operating conditions at 105% core flow.

5.2. Upset Pressure Operations The results show there willbe'no crack separation for Case 1 and crack separation is not expected for Case 2.

The results show up to 0.03 inch crack separation for the Case 3, but this condition is not a true representation ofthe shroud configuration.

5.3 Action Following an Upset Event The Nine MilePoint Unit 1 Shroud Repair Hardware Stress Analysis show the stabilizer and shroud stresses remain below the elastic limitduring.all normal and upset events.

The loads evaluated in this analysis are bounded by the loads evaluated in original stress analysis.

Since there is no plastic deformations, the shroud willreturn to the same condition as prior to the event.

Crack separation postulated to occur during an upset event is temporary and willclose followingthe event.

The tie rod assembly remain tight and there is no need for an immediate inspections ofthe tie rod assembly.

Reference G.J. DeSalvo and R.W. Gorman, ANSYS Engineering Analysis System User's Manual, Swanson Analysis Systems, Inc., Houston, PA, Revision 4.4a, May 1, 1989.

I

ANSYS

4. 4Al FEB 15 1995 16:49:08 PREP7 ELENENTS TYPE NUN ZV 1

+DIST=8.719

• XF

=91.496 4YF 31.32 MIND=2 ZV

=1 DIST=38.198 XF

=91. 25 YF

=34.725 Point 1

Figure 2.2a:

ANSYS Mesh

'1

ANSYS 4,4A1 FEB 20 1995 14:40:13 POST1 ELEMENTS TYPE NUN ZV

=1 DIST=144.114 XF

=95.75 YF

=-61. 562 Nine Nile Point 1

Figure 2.2b Shroud Model

I

I 1$

I+

I I lg I'

I I

I'>I I 1 I

I r~>T~~

I ~~44$V7k PrT I

.tl-++t.~

PIrly I TTTg(

Lpga

)pal $'$g.j I+

II IN IIII fl ff I'L fff II ff ANSYS

4. 4Al.

FEB 15 1995 15:50:53 POST1 STRESS STEP 1

ITER=l SY (AVG)

S GLOBAL DHX =D.01889 SHN =-1021 SHNB "1954 SHX =498.97 SHXB=949.389 ZV

=1

+DIST 8.991

• XF

=91. 469

• YF

=31.19 "1D21 "851.765

-682.923

-514.081 gag 345 '39 "176.398

~IIIIr'P.

-7.556 161.286 330.128 498.97 t,

I I I

,'r I; IIII

, I III e

1 Figure 2.3:

Ring Displacement With Cracks on Ring Side ofWeld MIND=2 ZV 1

DIST=38. 198 XF

=91. 25 YF 34.725 ggg

-1021 gag

-851.765

-682.923 161 '86 330.128 498.97

L

pT>g~

.) t

't

~I ~

) \\/

II IIIl H

I IIr ll ANSYS

4. 4Al FEB 15 1995 16:24:23 POST1 STRESS STEP=1 ITER=1 SY (AVG)

S GLOBAL DNX =0.01527 SNN "788.146 SMNB -1115 SNX =618.038 SNXB 652.781 ZV

=1

+DIST 11.813

• XF

-91. 443 4YF

=30.279

-788.146

-631.903 "475.661

-319 '18 "163.175

-6 '33 149.31 305.552 461.795 618.038 NHP 1 s

."4 2

F'g 2 4'ing Displacement With Cracks on Shell Side ofWeld WIND 2 ZV

=1 DIST=38.198 Xf

=91.25 YF 34.725

-788.146 "631.903

-475,661 305.552 461.795 618.038

F M

I IP I I'I ANSYS

4. 4Al FEB 15 1995

.16:38."59 POST1 STRESS STEP 1

ITER=1 SV (AVG)

S GLOBAL DN){ =0.036456 SHN

-1206 SNNB -1261 SMX =1033 SMNB=1087 4X

)

I I

I

)

1

) I 6

3 4X II Figure 2.5:

Ring Displacement With No Welds ZV

=1

+ DIST=11. 199

• NF

-91. 34?

~YF

=31.013,

~gg 1206 "957 '23

-708.369

-459.616 gag

-210.862 37.892 286.646 535.4 784.154 1033 MIND=2 ZV 1

DIST=38. 198 NF

91. 25 YF

=34.725 "1206

-95?.123

-708.369 535.4

?84.154 1033

f I

r

+

w-"

~

~

~ 7 ~

Subject Number

.'Originator

~GE Nuclear Energy Engineering Calculation Sheet Sheet of Oate

, Oate NEO@87 (REV. 1-93)

Ct

Subject Number F

8/ - l' a

C.

, Originator 1

Verifier' Nuclear Energy ngineering Calculation Sheet Sheet of i Date Date C

NE0487 (REV. 1-93)