Text

Revised Trojan Reactor Vessel Package Sar
	ML20237B612
Person / Time
Site:	07109271, Trojan File:Portland General Electric icon.png
Issue date:	08/13/1998
From:	; PORTLAND GENERAL ELECTRIC CO.
To:	;
Shared Package
ML20237B610	List: ML20237B612; VPN-040-98, Forwards Suppl Response to NRC 980709 RAI Re Transportation of Trojan Reactor Vessel Package (Rvp) for Disposal.Updated PGE-1076, Trojan Reactor Vessel Package SAR, Encl;
References
	PGE-1076, PGE-1076-02, PGE-1076-2, NUDOCS 9808190106
	Download: ML20237B612 (72)
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L VPN-040-98 August 13,1998

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UPDATE to Trojan Nuclear Plant Safety Analysis Report for Reactor Vessel Package l

(PGE-1076)

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l Instructions for Update The following information is provided as a guide for the insertion of new sheets for changes to l

the " Trojan Nuclear Plant Safety Analysis Report for Reactor Vessel Package," PGE,1076, dated August 13,1998, Books 1 and 2.

Book 1 Remove Insert Cover Page Cover Page List of Effective Pages List of Effective Pages Page 2-19 Page 2-19 Table 2-9 Table 2-9 Book 2 Remove Insert Appendix 2-2 Appendix 2-2 Appendix 2-2B (behind Appendix 2-2)

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I PGE-1076 O

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PORTLAND GENERAL ELECTRIC COMPANY l

I TROJAN REACTOR VESSEL PACKAGE l

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SAFETY ANALYSIS REPORT i

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August 13,1998 i

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l Portland General Electric Company 121 SW Salmon Street Portland, Oregon 97204 j

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r Trojan Reactor Vessel Package - Safety Analysis Report hN l

LIST OF EFFECTIVE PAGES TROJAN REACTOR VESSEL PACKAGE i

SAFETY ANALYSIS REPORT Page Number Revision Table of Contents August 8,1998 j

List of Tables August 8,1998 List of Figures August 8,1998 List of Appendices August 8,1998 List of Effective Pages August 13,1998 Pages 1-1 through 1-20 August 8,1998 Table 1-1 August 8,1998 Table 1-2 August 8,1998 Figure 1-1 (1 and 2) 1 Figure 1-2 1

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Figure 1-3 1

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Figure 1-3A (1) 1 Figure 1-3A (2) 0 Figure 1-4 (1 through 8) 1 Figure 1-5 (1 and 2) 1 Figure 1-5 (3) 0 Figure 1-5 (4 and 5) 1 Figure 1-6 (1 through 8) 1 Page 2-1 August 8,1998 Page 2-2 March 31,1997 Pages 2-3 through 2-18 August 8,1998 Page 2-19 August 13,1998 Pages 2-20 through 2-61 August 8,1998 Table 2-1 August 8,1998 Table 2-2 March 31,1997 Table 2-3 August 8,1998 Table 2-4 March 31,1997 Table 2-5 August 8,1998 Table 2-6 March 31,1997 I

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1 August 13,1998

Trojan Reactor Vessel Package - Safety Analysis Report

.h Table 2-7 August 8,1998 Table 2-8 March 31,1997 Table 2-9 August 13,1998 Table 2-10 March 31,1997 l

Table 2-11 August 8,1998 Table 2-12 March 31,1997

. Table 2-13 March 31,1997 Table 2-14 March 31,1997 l

Table 2-15 March 31,1997

. Table 2-16 March 31,1997 Table 2-17 March 31,1997 Table 2-18 March 31,1997 Figure 2-1 March 31,1997 Figure 2-2 March 31,1997 Figure 2-3 N/A Figure 2-4 N/A Figure 2-5 N/A Figure 2-6 N/A

. Figure 2-7

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l Q Pages 3-1 through 3-6 March 31,1997 -

Pages 3-7 through 3-10 August 8,1998 l

. Table 3-1 August 8,1998 Table 3-2 March 31,1997

' Table 3-3 August 8,1998 Table 3-4 March 31,1997 Page 4-1 March 31,1997 Pages 4-2 through 4-6 August 8,1998 Page 5-1 March 31,1997 Pages 5-2 through 5-11 August 8,1998

-Table 5-1 August 8,1998 Table 5-2 August 8,1998 Table 5-3 March 31,1997 Table 5-4 March 31,1997 Table 5-5 August 8,1998 Table 5-6 March 31,1997 L

' Table 5-7 August 8,1998 Table 5-8 August 8,1998 2

August 13,1998 L-_____-__

O Troian Reactor Vessel Package - Safety Analysis Report l

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- Table 5-9 August 8,1998

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Table 5-10 August 8,1998 Table 5-11 August 8,1998 Table 5-12 August 8,1998 Figure 5-1 N/A Figure 5-2 N/A Figure 5-3 N/A Page 6-1 March 31,1997 Pages 7-1 through 7-7 August 8,1998 Page 7-8 March 31,1997 Figure 7-1 '

N/A Figure 7-2 N/A Figure 7-3 N/A Page 8-1 August 8,1998 Page 8-2 March 31,1997 Pages 8-3 through 8-11 August 8,1998

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Table 8-1 August 8,1998 I

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August 13,1998 s

O Trojan Reactor Vessel Package - Safety Analysis Retsort LO l

barge subjected to a transverse acceleration. The RVP is restrained by the cradles and straps.

l One side of the cradle support structure is in compression and the other side is in tension. Loads are transferred through two beams in the cradle support structure and bolsters to the barge.

l A detailed tiedown analysis was submitted to the National Cargo Bureau and US Coast Guard for review and has been approved.

l The'RVP to transporter tiedown system was designed to meet the requirements of the l

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March 1993 draft of ANSI N14.2 (Reference 2-10). The transporter to barge tiedown system l

was designed to ANSI N14.24 - 1985 (Reference 2-5). The allowable stresses in the transporter l

j to barge tiedown system will be based on AISC standards (American Institute of Steel l

Construction Manual,9'h Edition), which are more conservative than the, corresponding l

allowables in N14.24 - 1985.

Analyses have been performed to demonstrate that RVP shell and nozzle stresses resulting from tiedown imposed loads remain within required allowables. The tiedown system design accelerations are summarized in Table 2-8. The analyses are contained in Appendix 2-2. The stresses in the RVP shell due to the transport induced loads reacted by the tiedown. system are calculated using a static force applied to the CG of the package having a vertical' component of

' 2g, a horizontal component along the direction in which the vehicle travels of 10g, and a horizontal component in the transverse direction of Sg. The stresses produced by these loadings are calculated in a detailed finite element analysis. The results of the finite element analysis are summarized in Table 2-9 and are shown to be within allowable limits, based on Code values for

. RVP material yield strengths.

An alternative method (Appendix 2-2B) was also used to calculate the stresses induced on the RVP shell and nozzles by the longitudinal restraint system. The results of the evaluation compare.well with the finite element analysis and also demonstrate that nozzle stresses under a 10g longitudinal load to the RVP do not exceed material yield strengths. The evaluation compares the calculated stresses to the Code minimum yield strength (S = 50 ksi) and to the y

actual yield strengths for the RVP materials. The actual material yield strengths, taken from test reports, exceed 60 ksi. Thus, a comparison to the actual yield strengths shows there is an even greater margin to yield.

l 2-19 August 13,1998

Trolan Reactor l'essel Packaze - Safety Analysis Retwrt h

nV Table 2-9 Summary of Margins of Tiedown Induced Stresses i

Section Condition Location Calculated Allowable

% of Value Value Allowable (ksi)

(ksi)

Value l

2.5.2 Vertical Restraint Head Flange

- Maximum Stress 29.6

'S, = 3 8 78 l

2.5.2 Longitudinal Restraint Nozzle Transition l

- Maximum Stress 43.5 S, = 50 87 l

2.5.2 Lateral Restraint Vessel Wall l

- Maximum Stress 48.5 S = 50 97 l

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A pendix 2-2 P

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PORTLAND GENERAL ELECTRIC COMPANY TROJAN NUCLEAR PLANT APPENDIX 2-2

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LONGITUDINAL RESTRAINT AND j

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TIEDOWN STRUCTURAL ANALYSIS PORTLAND i

GENERAL ELECTRIC COMPANY August 11,1998

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.io.uofDME-gG8 394-REVISION 3 tone,us. esme-i u - e u __

suPruen DocuesENT STAMP

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gFinal Submittal.

/ p g(

2 0 Work may proceed. Sutmt Anal document l

3 0 Revue and resubmit. Work mey pmesed prepgred g ;y i

sutgect to incorporation of changes Indicoled.

4 0 Revue and resutml Work may not EdeHeh 8(UdHICha 6oR not required. Work may proceed 6 0 Other Checked &

i Permlesion to proceed does not constitute Approved By:

acceptance or approval of design detalle, f

calculadone, analyses, test rnethode, or 7

meteriale developed or selected by the supplier Norman Lace and does not relieve suppeler from full compliance with contractual obilgatione o INFORyb,LY PProved By:

Prepwedb b"M#

Robert Morgan 8-Mti oeie 5 %%M, concu, red byh%9, oeie m,.. w -

r APPENDIX 2-2 Table of Centents For LONGITUDINAL RESTRAINT & TIEDOWN STRUCTURAL ANALYSIS

' q l V

l. OBJECTIVE................................................................................................................1

11. DESCRIPTION OF LONGITUDINAL RESTRAINT & TIEDOWN DESIGN................1 i

I l l. AS S U M P TI O N S.......................................................................................

IV. R E F E R E N C E S........................................................................................

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V. ANA LYTICA L M ETH 0 D S.......................................................................................... 2 A. Package Equilibrium Analysis and Vessel Wall StructuralIntegrity Analysis 2

I B. Longitudinal Restrsint Stress Analysis 3

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C. RPV Local Stresses Due to Cradle and Tiedown Strap 3

VI. ANALYTICAL CALCU LATIONS & EVALUATIONS................................................ 4 A. Package Equilibrium Analysis and Vessel Wall Structural Integrity Analysis -

5 B. Longliudinal Restraint Stress Asalysis 10 O

C. RPV Flange Local Stresses Iaduced by Tiedowns 34 Vl l. C O N C L U S l O N S....................................................................................................

l FIGURES 2-2-1 General Arrangement of Package, Cradles and Tiedown System 2 2 2 Longitudinal Restraint Total pages in this appendix =48 l

Revision Record:

I' Rev.1, initial issue; Rev. 2, intemal revision (not issued); Rev. 3, revised pages relating to paragraphs VB and VIB, Longitudinal Restraint Analysis. Note: Pages with revised contents (other than page number

~ changes) show "Rev. 3" at the top. Pages that were just renumbered show their original revision number at the top, e.g., Rev.1.

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dwpf crojan tydnsar2 doc i

m Appendix 2-2 Rev.1 LONGITUDINAL RESTRAINT & TIEDOWN STRUCTURAL ANALYSIS L

OBJECTIVE To demonstrate that the RPV package assembly's longitudinal restraint and tiedown systems do not compromise the package structural integrity.

D. DESCRIPTION OF LONGITUDINAL RESTRAINT & TIEDOWN DESIGN

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l During transpon, the RPV package rests on two structural steel cradles, as shown in Figure 2-2-1 of this Appendix. One cradle is positioned under the RPV head flange, and the other one is positioned near the bottom head of the vessel, under the 5" primary shielding jacket that surrounds the RPV. The RPV package is secured to the transporter with eight rigid struts. These struts, two per connection point, are pinned between brackets on the two package tiedown straps and similar brackets connected to the transporter The cradleind cradle support beams, along with the rigid struts, will resist the inertia ib:ces associated with the vertical acceleration as defined by ANSI N14.2 (1.5 g's) and the maximum barge trans ~

verse load of 1.6 g's.

The longitudinal restraint utilizes a clamping assembly, that is built into the cradle support steel, to clamp two of the RPV outlet nozzles so they can not move in the longitudinal direc-tion (see Figure 2-2-2). The structural steel restraint system is designed to withstand a 1.5 g loading times the package weight (1020 tons). Since the RPV nozzles are an integral part of C

the package, the two nozzles used as the longitudinal restraint must withstand a load of 1020 l

tons times a 10 g Load Amplification Factor (LAF), per 10 CFR 71.45. The included finite element analysis demonstrates that the nozzles are adequate for this service condition. The tie rod clamp design of the longitudinal restraint clamping assembly only restrains the pack-age in the longitudinal direction, so the 10 CFR 71 venical and transverse loads of 5 and 2 g's are not applicable to the tiedown system.

BL ASSUMPTIONS

1. The two nozzles are analyzed to the 10 CFR 7! 45 requirement for withstanding a design LAF of 10 g's, in the longitudinal direction. The restraint system can not load the noz-zles in either the transverse or vertical directions, so these LAF's are not considered.

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2. Weight of the package is 1020 tons, including the RPV, reactor internals, internal grout, primary shielding jacket assembly, spot shielding. and impact limiters.

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Appendix 2-2 Rev.1 IV. REFERENCES

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Provide a listing of all applicable references, e.g., applicable codes and standards, and com-puter programs.

1. AShE Boiler & Pressure Vessel Code,Section VIII,(1992).

2. ASME Boiler & Pressure Vessel Code,Section II, Part D (1992).

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3. AShE Boiler & Pressure Vessel Code,Section IX (1992).

4. Pactec Impact Limiter Analysis Calculation, number ED-017.1, dated 3/3/97.

5. 10 CFR 71 - Packaging & Transportation of Radioa'etive Material.

6. ' ANSI N14.2 - 3/5/93 drafi, Proposed American National Standard for Tiedowns for Truck Transport ofRadioactive Materials

7. General Finite Element Program, COSMOS \\M, version 1.71, 1994.

8. Weight and Center of Gravity calculations for the Trojan RVAIR project, TLG task j

number 201-03.

9. Burns and Roe calculation number; i

2030 02-23-001

10. MathCAD 6.0 (for Windows) Computer Program

11. Attachment L to Specification No. TD-16.

V. ANALYTICAL METHODS To demonstrate that the RPV package cradle assemblies and tiedown system do not com-promise the pa' kage integrity and calculations were performed to examine the effects on the c

package integrity and stability. The analysis is bounded by a loading condition of 10, 5, and 2 g's acting simultaneously at the CG of the package. A diagram of this condition is shown in this appendix's'Section VI A calculations. All supporting systems are considered infinitely rigid. The specific areas analyzed are as follows:

A. Package Equilibrium Analysis and Vessel Wall Structuralintegrity Analysis-To examine a worst case transportation loading condition's effect on the integrity of the RPV shell, the 10 CFR 71.45 loads of 10, 5, and 2 g's, are applied to the package's center of gravity simultaneously. In this analysis, the supporting systems are considered to be infinitely rigid and the reaction forces are calculated through an equilibrium analysis, which is shown diagrammatically in the calculations. To achieve a package equilibrium condition, the RPV

,O d wpruojantvdn,sar doc 2

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cradles, tiedown struts, and straps will resist these loads and achieve overall equilibrium. The Primary Membrane Stress in the vessel shell is then calculated and compared to the RPV ma-terial yield stress for acceptance. This is based on the Maximum Shear Stress theory for fail-h3 ure (Tresca).

In addition, an analysis is included that derives a multiplication factor for these loads that can be used to factor up the original finite element design calculations to show what the stresses would be in various components under the simultaneous 10, 5, 2 g loading event.

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B. Longitudinal Restraint Stress Analysis-4 The longitudinal restraint consists of a clamping assembly, that is built into the cradle support steel, to clamp two of the RPV outlet nozzles so they can not move in the longitudinal direc-tion (see Figure 2-2 2 of this appendix). The two RPV outlet nozzles utilized irtthis ar-rangement have been analyzed for the LAF loading specified in 10 CFR 71.45, i.e.,10, 5, and 2 g's. These two nozzles have been analyzed for 1020 tons times a LAF of10 (g's) in the longitudinal direction. The vertical and transverse accelerations can not be transferred to the l

vessel nozzles or shell, since the assembly is not restrained in those directions, i.e., the tierods only restrain the package along tierods' axis.

The analytical methodology for the evaluation of the local conditions at the longitudinal re-straint uses the ASME B&PV Code,Section VIII, Division 2, with application of the compu-tational rules of Appendix 4 (Alternative Methods By Detailed Stress Analysis). The calcu-lations for this assessment were performed using the Finite Element Method (Reference 7).

The allowable stress limit for the nozzle material is based on 10 CFR 71.45, which states that for the specified loading conditions, the stress generated in the base materiaf shall not exceed O

its yield strength. Thus, the nozzle walls and the surrounding vessel shell that constitute this l

integral package attachment, shall have resultant stresses that are less than the material yield stress.

C. RPV Local Stresses Due to Cradle and Tiedown Strap -

The RPV head end flange has been analyzed for tiie local stresses caused by the bearing of both the cradle and the steel tiedown strap against'the package, i.e., against the flange or the steel shielding jacket. A finite element model was used to determine the local stresses due to the loading from the steel tiedown strap on the RPV head end flange, and by similitude on the shield jacket. The finite element calculation run used a load induced from the strap as being equal to the smits package load of 1020 tons, amplified by a 2 g LAF, which is the vertical 4

component called out in 10 CFR 71.45 for integral package attachments, even though the j

straps and cradles are not integral attachments. This load was then compared to the resultant j

from the simultaneous 10,5,2 g loads calculated in Section VI A and a multiplication factor

{

of 2.741 was calculated for application to the finite element results for the simultaneous 10, j

5,2 g resultant load. The results from this factoring are shown in the results table. Based on l

the calculations for the head end flange, an assessment of the nresses in the 5" shield area can also be made and this is also shown in the table.

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Appendix 2-2 Rev.1 For the stresses induced by the cradle in the head end flange and the shielding jacket, a finite element analysis was performed for the cradle / shield interaction. This analysis was done for O

the cradle nearest the lower reactor head, since the highest reaction load occurs here. By similitude these results also apply to the head end flange, with the allowable values changed for the RPV flange material. Since the head end flange area that the cradle contacts is thicker than the shi !d area where the finite element case was mn, and since the resultant loads at the t

head end flange are less than those at the other cradle, the actual stresses in the flange area would be lower than the calculated values. These finite element loads were then compared to the resultants from the simultaneous 10,5,2 g loads calculated in Section VI A and a multi-plication factor of 3.488 was calculated for application to the finite element results for the simultaneous 10,5,2 g resultant load. The results from this factoring are shown in the re-sults table.

The allowable stress limit for the nozzle material is based on 10 CFR 71.45, which states that for the specified loading conditions, the stress generated in the base material shall not exceed its yield strength.

VI. ANALYTICAL CALCULATIONS & EVALUATIONS 1

This section of the appendix contains the actual calculations and assessments relating to each of the " analytical methods" previously discussed. Each of the analyses stans with its own ti-tle page, and they are ordered the same as presented in discussion of the " analytical methods" used. The results of all of these calculations have been summarized and interpreted in Sec-tion VII of this appendix.

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Appendix 2 2 Rev.1 A. Package Equilibrium Analysis and Vessel Wall StructuralIntegrity Analysis -

r Vessel wall structural integrity evaluation under simultaneously acting collision loads:

Determination of the structural wall integrity under the collision loads is based on following assumptions:

Supporting systems, which are pan of the transportation concept are considered a-infinitely rigid (this assumption is applicable for this evaluation only).

b-Collision loads are :

Longitudinal Direction : log Transverse Direction : Sg Vertical Direction : 2g c-Allloads are acting simultaneously d-General Primary Membrane Stress Intensity is determined at vessel maximum resultant bending moment.

(For graphicalinterpretations see 4

Moment Determination:

attached figure).

R = 190.375 L :274.25 WR :10410' Weight of the vessel (Ib) 2 a : 174.75 b : 99.5 CGlocation (in)

F1 :WR 10 Ft : WR 5 Fv : WR2 M1 :R Fi 1 Mt a d Ft Mr :N Fv Maximum moments from individual 2

L L

components Fl = 2.04 10' (lbs)

Ml = 9.709 10' (in-lbs)

Ft = 1.02 10' Mt = 6 467 10' Fv = 4.0810' Mv = 2.587 10' Resultant maximum moment:

2 M = 1.389 10' (in lbs)

M= (M1 - Mv)2, pg Vessel section properties:

Do :190 375 wt : 8.5 Di : Do - 2 wt RPV properties calculated based on most critical parameters - no l

1:L Do'- Di' I = 2.013 10' 64 meluded rveN.s*Ruco 5

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Appendix 2-2 Rev.1 Vessel Primary Membrane Stress Intensity O

e ob = M 2-ob - 6 571 10' (psi)

General Membrane Stress Intensity 1

Using the Maximum Shear Stress Theory for failure (Tresca) 1/2 ob < 1/2 cy where ey is the material yield stress for -20 to 100 0F i

l For SA-533 GR.B CL 1, o y = 50.0 (ksi)

I ey =50000.

.ob < ey 1 ob = 3.28510 1 ey = 2.5 10' (psi) l 3

2 2

Therefore, the structuralintegrity of the vesselshellis satisfied.

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Appendix 2 -2 Rev.1 I

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a 4

b O

i Fav Fbi

.Suppp t.B j

FI

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support A Fbt F 3 Ft C.

) k Rs!

Fbv

' ' M.V.

Mt' -

1 r L,a,b... cG location dimensions R........ RPV radius Mt,MI,Mv.. Moments due to Transverse, Longitudinal and Vertical Loads LONGITUDINAL EQUILIBRIUM TRANSVERSE EQUILIBRIUM Fal y

O S

'S FI

\\Fta,b l

'. -l A k a

Na,b Rs Fbl VERTICAL EQUILIBRIUM i

Fv J L J L Fav Fbv COLLISION LOAD EQUILIBRIUM 7

l

1 Appendix 2 2 f

Rev.1 O

Factorization of the Finite Elements Results, Using the Free Body Diagram.

lU The factorization of the Finite Element stress results derived for the transponation loads 1.5.1.6 t

& 1.5 g (loads considered acting independently), can be used to determine the results from the application of the 2,5 & 10 g (loads acting simultaneously). This factorization requires additional calculations that are based on the transverse plane equilibrium. The transverse input load becomes a resultant vector out of 2 g (venical amplification) and 5 g (transverse amplification). In order to obtain the resultant vector, the SRSS method is appropriate. Using this load, new equilibrium is constructed and the resultant reaction input for saddle F.E. and tiedown F.E. is calculated. The ratio of the new versus originalinput values, gives the required coefficient. These coefficients are used for factorization of the Stress Intensities derived for transpon loads based on application of simultaneous 10,5 and 2 g LAF's.

WR = 2.04 10' M: WR*S M = 6 49910' Transverse Load L

l l

N : WR

2 -

N = 2.6 10' Venical Load L

j hM - N = 7 10' 2

2 bTnew hTnew = 1.071 10' Load / reaction required for the sin 37 35 -." -

saddle contact stress F.E. model ISO input.

cp = hTnew cos#37.35 *,. N cp = 1.112 10' Load / reaction required for the i

380-strap F.E. model input l

E = 5.558 10' 2

S :.*E.

I S = 5,644 10' Individual strap loads ( not used c,,/10 *-\\

in the vessel design) 2

\\

180, Factorization Coemeients:

hTong 3.0910' csaddle :

csaddle - 3 46-Cradle F.E. Coemcient hTong RRv = 4 0810'

P estrap :RRv cstrap = 2.724 Strap F.E. Coemclent These factors are used in other pans of the appendix calculations to obtain results for other appropriate t

LAF conditions that were analyzed.

TYDN.SAR McD OV

4 Appendix 2 - 2 Rev.1 O

l l

Factorization of the Finite Element Results I

l r NN 10'deg NFnew O

cp 37.35 deg i

-- - - - y j

~

S N= 2(WR)a/L l

=

M= 5(WR)a/L l

9 O

l l

i

Appendix 2 -2 Rev. 3 q(g B. Longitudinal Restraint Stress Analysis -

Nozzle Local Stress Evaluation Engineering Methodology:

The nozzle local stress derivations used the Finite Element technique. Similarly, as was done for the analyses oflocal stresses for other RPV support and vessel itself, this method is the most reliable and most accurate for determination of the local stresses.

The Outlet Nozzle Finite Element model was generated by eight nodded brick element - SOLID.

This element is commonly used for modeling thick structures and bodies with complex geometry.

The element has three translatory degrees of freedom for each node. The longitudinal restraint components have been modeled to represent the actual geometry of the main restraint components, e.g., the nozzle, the surrounding vessel shell plate, and the reactor head flange. Because of geometrical and loading symmetry, the restraint system can be modeled as one half of the system,.

i.e., one nozzle and the associated vessel shell. The model utilized the following boundary -

conditions:

1. The reactor vessel head flange is a fixed condition due to its massiveness and rigidity.

2. On the remaining three boundaries, the shell is free to grow in the longitudinal direction, but is restrained in the circumferential and radial directions.

These boundary conditions and the load application are shown in the figure that follows this Q

discussion of boundary conditions and load application.

O The load is applied over the nozzles contact surface at the point where the locking mechanism clamps over the nozzles. The detail three dimensional sketch of the mechanism is shown in the attached drawing. The blocking mechanism is designed to amplification of 1.5 g. The nozzles uresses, however are checked for load amplification of 10 g. The distribution of the applied load between the two contact surfaces on the nozzle is as shown immediately following:

Determination of the N.ozzle Pressure loading due to the RVP weight 10G Amplification.

The load is based on dimensions as shown in the calculation shown below. The pressure

is distributed over two locations of the outlet nozzles, i.e., over the pipe connection portion and over the nozzle main body portion. (Two outlet nozzles are resisting the longitudinal loading effect). Two half moon blocks of the restraining mechanism are clamped over the nozzle circumference. Also, there were some slight dimensional deviations between "as modeled" versus "as built" nozzle dimensions. The deviatio'ns are shown below and were found to contribute to a conservative calculation. The analyzed geometry is more conservative, thus the model geometry of the finite element is not effected.-

I 10 i

TYNDNsAR2 MCD 4

Appendix 2-2 Rev. 3 i

O Restraining Concept with Boundary Conditions and Pressure Load Application

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ux=uy=0 l

pt.

E,_4.z -

uz/0

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uz/0 U

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uz=0 i

G v

1 11

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Appendix 2 -2 ~

Rev. 3

- n The pressure is assumed to be distributed based on the relative stiffness of the two nozzle sections V

(top and bottom), where it is applied, based on the bending stiffness for a cantilever beam. Thus, the stiffness is de6ned as shown in the following sketch:

i Top load S

i G

li B

B

/

i

\\

./

\\ Bottom load in i

3 kt = (3Eli)/h 3

kb = (3Elb)/le From drawings of the nozzle, it was found that I - 11.625 inches and I ~ 11.5 inches. Since 3E t

b 3

is a constant, and 1:3~lb, then k / k "I / I -

t b

t b O

Outside / inside pipe end designed diameter:

Outside pipe end "as built" diameter:

Do2.= + 10 + 24 Do2 = 34.125 (in)

Do2foud = 34.06 Do2foud = 34.12 (in)

Di2 = 5 + 24 Di2 = 29 (in)

Outside / inside large end designed diameter:

Outside large end "as built" diameter:

Dol = 2 + 412 Dol =50 (in)

Do2foud = 50.0 (in)

Dit = 5 + 24 Dit =29 (in)

The pressure distribution for the finite element application is determined based on the stiffness, or rather the moment ofinertia. The derivations are shown below:

11.=

-(Do!' - Dil")

11 = 2.721 10 (ind) 8 1

i 12 = 1(Do2' - Di2')

12 = 3.185 10' (ind) 64

()

12 TYNDNsAR2 MCD u________________

j

i Appendix 2 -2 Rev. 3 C)

E =0.117 (Pressure is distributed over top and bottom of the nozzle is based on this ratio) f~

11-Total longitudinal load of the RVP is as follows:

WI.= 1020 2000 WT = 2.04 10' (ib)

G = 10 7

Flong a WT G Flong = 2.04 10 (Ib)

Bearing surface total:

al = 8 a2.= 6 (in)

Design bearing lengths A.= (al Dol + a2 Do2) 2 A =1.21 10' (in2)

Two nozzles are resisting Flong p :=

p = 1.687 10' (psi)

A Distribution based on Moment ofinertia / stiffness pb : p - p b pb = 1.489 10 (psi) 4 Il 3

pt p.

pt = 1.974 10 (psi)

Check:

p = pt + pb p = 1.687 10" (psi)

Ok Pressures are proportional to the not stiffness (or Moment of inertia) of pipe or heavy nozzle sections

.....j.......

pt = 1.974 10' gy g /////a l

i pb = 1.489 10" j

i l

l 13 l

TYNDNsAR2.MCD l

i

Appendix 2 -2 Rev. 3 The blocking mechanism is capable of transferring longitudinal shear load only. The transverse shear direction is not restrained, since this degree of freedom, is resisted only by bending stiffness of the tie rods. This mechanism is also capable of shear transfer. The bending couple, represented by the longitudinal load vector and the longitudinal shear vector at the blocking mechanism, is resisted - equalized - by the reaction at the saddle and the strut assembly. There is minimum or no clearance at the restraining mechanism, since the mechanism consists of two adjustable halves, which are tightly assembled against the outlet nozzles. The loading is described in detail in the load equilibrium section of the report.

The description of the Finite Element model and supporting load calculations for the model are presented next.

Design Input Variables:

Material:

SA-533 Gr. B Class I and SA-508 C12 Yield Stress for -20 to 100 d 50.0(ksi)

Young's Modulus:

27.8 x 106 (p !)

Min. Tensile Stress:

80.0(ksi)

Poisons Ratio:

0.3 Allowable value:

Sy = 50.0 (ksi)

Load over the nozzle bearing surface was applied to the model.

Wr = 1020 200010 Wr = 2.04 10' Amplified Load by 10 g

= 9.0 Nozzle width (in)

D = 4 12 + 2 Nozzle diameter (in) A = s D A = 450 Area - since two nozzles:

i l

2 At = A 2 At = 900 (in )

Bearing pressure:

p=

p = 2.267 10' (psi) i Finite Element stress computer analysis post processing utilized a search routine to find the maximum component finite element stresses in the modeled structure. These values were then factored up to account for a slight deviation in the applied load versus the ' ctua!

a

-load. This factoring (increased by 6.6%) is explained in the discussion that follows the table. The following table shows these values (in bold) at the indicated elements.

14 TYNDNsAR2 MCD I

1

Appendix 2 -2 Rev. 3 O

Maximum stress components in modeled structure, as obtained by FE search:

Element I.D.

ex cy az Txy Tyz txz Number 1817 1109

-1258

-13005

-1418

-5064 8603 1944

-613

-3059

-12472

-809

-10873 5074 2399 12152 9434 13645 7227

-239 116 2521 18122 14604 28569 994

-16310

-517 2948 18335 22493 22386

-947

-21000 957 Similarly, a search was performed to determine the elements with the highest Principle Stresses and the highest Stress Intensity. These represent the highest combination of stresses within the stmeture.

Element I.D. No.

Pl P2 P3 S.I.

1944 6045

-1827

-20403 26451 2948 43490 18431 1286 42204 The maximum calculated component and Principle Stresses, and the Stress Intensity shown in the above table were all found to be below the allowable yicld limit of 50 (ksi), as required by 10 CFR 71.45(b). Note that the above values have been increased by 6.6% over the actual F.E.

output values. This increase is due to a slight deviation in the computer generated area ofload

/]

application on the nozzle. The reason for this increase is more fully described on the next page.

V It should be noted that the maximum stress intensity shown above, closely approximates the -

effective von Mises stress that should be directly compared to the uniaxial material yield strength.

The ASME Code uses the Tresca approach (principle stress differences) as the basis for comparison to allowable stresses determined from yield and ultimate strength. Therefore, the RPV outlet nozzles' stresses induced by the longitudinal package loading, are below the allowable limit and acceptable.

Evaluation of the Load Application Surface Deviation.

l Design load surface per one nozzle:

C Ades.= 604.75 (in2)

Finite element surface over which the pressure was applied:

Do2 = 34.125 m2 = 6 (in)

Atdes = Do2-m2 Atdes = 204.75 (in2)

I Dol =50.

ml.=8 (in)

Abdes : Dol ml Abdes =400 (in2)

TYNDNsAR2.se 15

Appendix 2 -2 Rev. 3 i

Note: Since the finite element mesh generator created element sizes which did not allowed to model the pressure surface exactly, the error coefficient is calculated herein in order to l

1 correct the small stress result deviations.

Large nozzle contact dimensions deviations Amic = 8 - 6.84 Amic = 1.16 (in)

@ zero degrees with respect to global coordinate system Amts = 8 - 7.67 Amis =0.33 (in)

@ ninety degrease F.E. Large diameter nozzle area:

)

AbFE = Dol 8 -

Dol-(Amic - Amts) + Dol Amts AbFE = 362.75 (in2)

Ades = Abdes + Atdes Ader =604.75 (in2) l AFE = AbFE + Atdes AFE = 567.5 (in2)

Ades,

e AFE i

1 Therefore, all stress results shallbe factored upwards by 6.6% in order to offset the FE medellng i

deviation.

69 16 TYNDNSAR1MCD i

Appendix 2 -2 Rev. 3 O

Design Pressure Contact Surface:

Nozzle Center Location "c" l

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Appendix 2 2 Rev.1 1

O C. RPV Flange Local Stresses Induced by Tie Down -

U Engineering Analyses and Methodology:

This evaluation was carried out in a similar way as the analyses for the local stresses under the saddle.

The main tool for the local stress determination was Finite Element technique. The RPV head and flange geometry was modeled as a 1/4 model. The symmetry of the structure and simultaneously syrnmetry of the loading was satisfied. The local stress solution does not require modeling of the whole structure, since the local stresses has only local character thus diminish within the 3/ beta distance (usually 20" from the discontinuity, for the RPV geometry). The model was constructed out of SOLID element with 20 nodes per element. The reason was to increased the accuracy of stress distribution across the vessel wall. The boundary conditions and modeling details are given in the attached figures.

Load application, representing the strap tie down pressure / forces, are derived and shown next.

Design Input Variables:

Material:

SA 533 GR B Cl 1 Allowable values:

Yield Stress for -20 to 100 cF: 50.0 (ksi)

Sy = 50.0 (ksi)

Min. Tensile Stress:

80.0 (ksi)

Young's Modulus:

29.2 x 106 (psi)

Poisons Ratio:

0.3 Loading

The loading is represented by discrete point load vectors at the contact of the strap against the flange i

RPV head. From the equilibrium evaluation of the forces, most critical loading was selected for RPV local stress evaluation Discrete Load:

i = 42 Numbers of nodal points at the flange contact WR :1020 2000 2 WR = 4.08 10' Load is amplified by 2 g RRv = 4.0810' Total vertical load at the tie down (worst vertical load) (lb) -

j Please note that total load was considered at the tie down.

l This is conservative and for check oflocal stresses in the RPV j

head is acceptable, as long as the limits were not exceeded.

This load application was not necessary. (Actual vertical reaction 6

Rr = 1.3 x 10 x 2 Rr = 2.6 x 106 = RRv Rsiy :

Rsiy = 9.714 10' (lb) l TYDN,SAR M*D 34

Appendix 2-2 Rev.1 I

1 r"

From the finite element analysis and the attached plots, the Maximum Stress Intensities are found to be

\\

as follows l'or the independently applied LAF's of 1.6,1.5, & l.5 g:

3830 (psi) for the RPV flange head (dome area) local stress induced by the tiedown strap, 8670 (psi) for the RPV wall at the bottom end head stress induced from the cradle, and 13900 (psi) for the 5" shield, contact stresses induced by the cradle, 10800 (psi) for the RPV head flange (flange area) local stress induced by the tiedown strap, j

These values are based on the maximum values (instead of the average values that the cod(

uses). Therefore these results are very conservative.

l 1

To obtain the results for the simultaneous application of the 10,5, and 2 g LAF's, the a' bove values j

have to be factored using the factors calculated in section VI A:

J 3.83 2.741 = 10.498 (psi) for the RPV flange head (dome wall) local stress induced by the tiedown strap 8 67 3.488 = 30.241 (psi) for the RPV wall at the bottom end head stress induced from the cradle 13.9 3.488 = 48 483 (psi) for the RPV wall under the 5" shield, contact stresses induced by the cradle

)

10 81741 = 29.603 (psi) for the RPV hejd flange (flange area) local strus induced by tiedown strap Db The allowable for the above stress calculations is the material yield strength, Sy = 50.0 (ksi) for the SA 533 Gr. B, Cl 1 material for -20 to 100 of.

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Appendix 2-2 Rev.1 q

Cradle Shield Interaction - Finite Element Solution for Local Stresses in the RPV She Q

and 5" Shield at the Cradle Contact.

Finite Element input for the local stress determination consists ofdetermination of:he discrete load l

vectors applied at the contact between the cradle and the RPV Shield. From the equilibrium I

calculations, the "N" reaction at the cradle contact, was found to be:

N = 3.0910' (Ib)..... Reaction is intensified by factor 1.6 g.

The load is distributed over six nodal points, so that the discrete value at the individual nodal point is:

i =6 Ni :

Ni = 5.15 10' The load reaction are spread over sweep angle of 37.350 Since the design allows the conserv-ative interpretation of the reaction distribution, the individual load vectors are derived as follows:

7.351 Nw = 1.649 10 Niy : Ni stn 8

\\2 180 (lb)

N Niz :Ni ces 1,-

Niz = 4.879 10' (lb)

\\2 180 Q}

These vectors are applied to the Finite Element model - File: CRACONT.* The boundary

/

conditions and the modelis given descriptively in the attached figures. The summary of the maximum Stress Intensity and the component stresses for the extremities are in attached table.

Description of the Finite Element Model for the 5" Shield & RPV Shell Saddle Contact Stress.

The F.E. model utilizes geometrical as well as loading symmetry in modeling. Only 1/4 of the 5" shield jacket and RPV shell with cylindrical head, is modeled. This model generates local Stress Intensities and tlie stress field under the 1.6 g amplified total RPV weight load at the shield and cradle contact. The loading is derived as a point load at Nodal Points, where the cradle is in contact with the 5" shield. Calculation of the individual load vectors is given in previous section.

The boundary conditions are apparent from the attached plots. Namely one end of the shell is restrained in longitudinal direction ("x") and the other boundaries reflect the geometrical symmetry. It has to be stressed that the model derives local stresses Pl + Pb only. The stresses at the vicinity of the boundaries can not be considered accurate, since the loading here is not of a symmetrical character. The contact stresses for design purposes are correct, since the model extends far enough from the contact discontinuity ( 3/ beta). F.E. utilized SOLID eighth nodded element. Analysis is classified as Linear - Static. For the design purposes, this modelis l

considered satisfactory - there is no need for greater accuracy of stress distribution across the shell wall, since the stress levels are well below allowable limits and the loading is conservative.

TYDN,,sAR McD tO 37 l

I

Appendix 2-2 Rev.I Stress Components as Derived by F.E. at the RPV Shield /Shell under the Cradle

Contact:

y

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Average stress components

-6450.

-7420.

-2034.

-30.5 101.4 1570.

Using the Mohr Circle Theorem for determination of the Principal Stresses and evaluating the actual stress component magnitudes (bold values), the Principal Stresses are solved as follows:

el :.6450 e2 m.7420 c3 :.2034 (psi)

Stress Intensities are defined as follows:

SI =cl - c2 St = 970 Note: The Stress Intensities SI, S2, & S3 are absolute S2 s e2 - e3 S2 = -5.386 10' S3 = e3 - of S3 = 4 41610' Local Membrane plus Bending P1 + Pb Stress Intensity - max.- Pi + Pb = 5.4 (ksi)

Allowable Limit is Sy = 38.0 (ksi) for the S A 516 Gr. 70 shield material for -20 to 100 oF (allowable value is higher for the shell).

To obtain the results for the simultaneous application of the 10,5, and 2 g LAF's, the above values

' have to be factored esing the factors calculated in section VI A:

5.4 3 488 = 18.835 (psi) for the 5" shield jacket contact stresses induced by the cradle (the allowable is still Sy = 38.0 (ksi).

I L

l 38 Tve~_s" uco V

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I NT E Y O

T B N D I

SE SS EU R A T C S

L Q L

+E bPH

+S lP DN FO A T D O A L E PH 1

=c L

S S

E R

T S

O n

i L

G 0 0 0 G G G G G G G 0 0 0 0 G G G G G G G

+

v + + + + + + G G G E E E E E E E E G G G 3 5 6 8 G 2 3 5 G

8 4 0 6 3 9 5 8 3 G 1

s

.8 8

n3 3 3 2 2 7 3 G 1

1 1

e t

1 I

I 1

n DA EH EGNA LF VPR EHT X

N I

S O

E IT ISNETN I

S SERTSbP+

l PFO TOLP 1

=c L

S S

E R

T S

O n

i L

\\

\\\\\\\\\\\\\\

I e= c.,w w't 7*.

n d

4 4 4 3 3 3 3 3 3 38 8 8 8 8 8 8 8 8 8 88 8 8 S 8 s 8 8 S G 8 8

+ + t + + + + + * + 8 E E E E E E E E E EG 9 6 3 7 3 8 9 9 s 8.

3 2 1

3 s3 s 7 5 2g s

,.....g ng 3 3 7 6 4 3 2g 1

g e

1 N

t n

I 1

DNA D

A E.S H D VA PO R,L DT LC.

EAT I

H N S O E C H E T

L ND I

A NR OC I

TE UI I

BT I

RR TE SD ID N U

SSL EL RE TH SS 1

=c t

S S

E R

T S

I L

>w

~

kI i

\\

i f'

i' i

Appendix 2-2

+

Rev.1 VB. CONCLUSIONS The calculations discussed in this appendix, and their results, show that the tiedown system for the Trojan RPV package meets all applicable requirements and is fit for-sersice for this application.

I The following table summarizes the results of these analyses for the components and loading l

conditions described herein:

j i

l S

O e

e O

a t___

uV d tgl r

s n sf l K

oir o

d o m.d e

Vn a

a c a c e

P a ni c o

,a e

nt nt l

r e led n el h et R

rb wo wt e r f

t h n i

c s a

a e r io ol e

e d

t n e e

R n, tr S

A fi ns T

M n qioui r o r o i

t c

o ios al u ud t

o t

ot u

o st st t

a t d

i n al t

d me l

eh ot en e n n c e tal b s imd t

o et a v v

i t

n L

E it aS e luil l

u e a si r cirl s n s n f

ed e

s o f

d U

R a ee a u nat nat d

l e m u

ieet r E

l

.hi oc d n d n a

r li lcyei l e

hl c h s vc a cn a cn e r t

i wa r

l s

uTs l

r s e r

a o a e s bVb sl rh vh y e el S

c t

s e t

r a t

c s

f ga o ga o sf ime weed y l

a s t r

P d e g S op S o p u es R

m mi n el m el m ui is gloh s e r c

s c

eRl k s t

l t

st e u

e a

t si l

s n

el a o z

hl o hl o c

a a a ya er e r ai um oeo ie r

a a

S S gfb s Tcc Tc c Cvb Bl h ot r

f t f s co Nh w hh I

t t t S

Y

.C n -

P F A B

B CdA mii L

v s e C dA L P E

I I

nI oa e un A A A R V V

V VaV F a r oi VaV I

nI r

t l

C d p N

A l

l l

l E

C0 C0 C0 C0 C0 o

1 L

L i

B r-i

,2

,t 00 2

s B

B B

72 B

A BS k r

r r

r i. o

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r.

r s rf R

AT i

s rf s rf s. o

. o

. o i

5 s

k Gl k Gl U

WI kGl k Gl kGo kGl 2

)

M 0

aF

)

f

)

T OI =

3i 0.3 iaF 0.3 iaF 0.3 i *aF 0.6 l 0.3 i 'F a

r*

,0 3 03 03 r "

03 r 81 a) r e0 55 0 55 e0 iF 03 LL S5 5 e0 55 e

e 35 r 55 0

C L

x= Am1 =

t t

- a0 t

a0

- a 0

- a0

- e

- a 0 Am1 = Am

= Am1

= At 0

= Am U

A 1

1 a

R

%S,S o

,S o

,S m1 S(

0

,S o

S(

( I t

I I

i t

t t

1 t

T S

8 s

=

s9 s

s i

s4 s

s N

S

- n s

e e

e k

T mI r0 tr tr W

L e s 3 e

e S1 S

S t

s e

e U

Me 3 ms ms m=

mi mi O

s s

r S

yt =

u s u

s, u y uk uk D

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I imn i ime.

me mi ix6 x2 y

m3 m1 a e t s3 iue s3 s

E R

s s

n 0

i4 i

s x e e

a n a tr te e

x r

at a

a t

T Pb MS n MS n MI i M2 M3 r r e

o o

n s 9

0 t

k t

x g

e e

x g

T 2

g g

- 2

- 2 s u a

a s u s u s u nt d nt d nt d nt d N

oign c

c oign oign oign k

k I

t a

n a

a n

n n

A 0 o p

p 0 o 0 o 0 o 2l e 2l e 2 l e

2l e R NS s

t t

0 s

s s

GD D gr a

a 0 gr 1

0 e 1

0 gr 0 gr 0 e 1

s s

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1 A E v

n n

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m1 m1 m1 m1 s

S S

S EO U s

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E r u ad r uad D L f o r a 0

0 f or a f or a f or a R

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l l

l i gG gG a a Sl 1

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t a

lt lul it a

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0 e mn i e e mn Ris d i e e mn r

r r

N i v 1

C 1 C Risd v Risd v Risd v i e I

D la d

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nt R at u Rtsf Rtut u R e s sf r i r

t l s

t s

t

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sl lo S

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r f

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al 1

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h r

t e

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l 2

f e l

ip 's 7

7 7

4 1

S s

c s 9

2 8

5 3

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n e.

8 8

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c i i t r

1 1

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8 L P E

nI 5

d r

1 1

1 I

n C

a ii P.hw I

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1 N

f g

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o t

n A

n l

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l l

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2 0

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t 9

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6 3

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r

i. o i.-

o h n R

AT s r f s r r 8

9 3

6 6

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3 4

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e k

W.

kG l

)

f

)

kGo G ois f

a T

OI 03 r

a) 8 1 LL 55 e0 35 iF e n n

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=

a n

e*

g oi t

r a0 U

A A m1 = At 0 ll 1,M.

a k

e a f

3 4

6 7

7

,S 0

R S(

o

,S X hr.

0 7

S( m1 c

c

(

I 5

1 1

t 6

0 a n s T

=

a i

W.

8 pi e

5 5

9 1

T s

t "v

S nd s

,i s

s s

s ote n N

S e

e

' 4,.

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r d c T

t W

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t e n 3

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6 7

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4 0

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5 S

uk uk e u e 1

1 2

D E

m3 m5 r

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r

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8 8

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e a 2 d y s Y'

9 7

1 4

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nl e

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4 9

4 x

g x

g pn s T

8 2

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9 1

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pa e

- 2

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t n a t

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l A

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m1 S S 1

1 s

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1 E O U os ns o s n s p Vs u

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e n t

Y D L f o r a f o r a i o e n nr e R

t t

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1 1

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c a

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p a

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h ON O N

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a o u e c

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2 O n s e u 1

8 e

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s u s l

m.

1 4

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s a

n it l

t

/ n w

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9 3

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i I

s di 1

1 2

2 2

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CS e g e l

v D wt l

E1 n

r E

c e ie e t

e e

e D N al a

n j

h s o h h o

s O Vt d s

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P a

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k or r

t C Rc c 5 s o

h t

N1.

2 3

d

)

_ (

c e

n

~

m 3

. e(

d e

R l

=

=-

E j

e P

T c

Y t

- o

$T P

.r

(

5

. O R

N

)e E.

E A

T c

R. (

R T

o O

E P

=

.\\

S N. )

e(

A R

T

=

= _ _

. )o

(

, =

n

. L_

E

)

. e(

L D

A i

R C

.. )

t e

i T

f

. (

0 P

S i

NI

)

P A

e Y

R T i

~

TE

~

. )e E

R

(

a t

U l

I r,'

5

)

5 t

EL e

f E PS r

5 RE w

O V

.)

_ p)b A

o(

s TC s

E

~

n" R

x

.. i

)e

(

fs v_,,6 m,

e

@g-

)

8 8 8 i

f i4 p.p a

I I s

)

l I

j e

7l

}'

,,f

.k '

s )

)

i,

. (

..)

. L_

o

(

3=

~

=

7=:~ _

.w

)

e

=

)

=

)

= __. - =_

O z_

. c_

T

)

i 6-

. ~G

/

t,. 6

n. d

,.=

Appendix 2-2 Rev. 3 O

1

(

i

. [I.D'.'

[.

nt,..,,,

O M

Notes:

1. RPV nozzles provide shear key function only in longitudinal direction

2. Shear key restraint beams transmit shear load to cradle support steel Figure 2-2-2 l

Typical Longitudinal Restraint Concept Typical of(2) RPV Outlet Nozzles O

48 Last page of App. 2-2

.i n,-.-

c_

O

e O

Appendix 2-2B e

' O

llr ar,b.ah.

2

' *" QA RECORD WHEN COMPLETED

gl",,*.

7.

N. o.'

c I

- TROJAN NUCLEAR PLANT /ISFSI CALCULATION COVER SHEET Sheet / Cont'd on Sheet f.,

l Titie' 2kWLDfnd/MfAsf/ odd /AsVJ//A4'7145 /f 4 lans7W#d/bf/2Aff 64 //Ai1147bPUffS2*V

~

Trojan Nuclear Plant /ISFSI Calculation No. #-74/

I Structure

4 Supersedes Calculation No.

System A#

Quality-Related (vet / No Component AAft*&P Mff//

References (PMR/bPMR, SPEER, MR, PSC, etc.)

A V4/# ###//##

Has Been Changed by Affected identify Change or Revision has been Responsible Document Vehicle: (MR, DPMR, DCP, Deferred by (identify Memo, Supervisor /Date No.

PCF, SPEER, PSC, etc.)

CTL, etc.)

(Deferrels Only)

Nk*

b p.

Calculation Objective

\\

$$$ SW$$ A*

Revision Description Rev.

No.

Preparer Date Verified By Date Approve $ By

. Date o

vMmx

ski $.#. L %I@hhh M b.h>ht l

Page.1 of 1 O

U TPP 18-9 Revision 3 Page 13 of 13

E?r aAtds l

Laser No WA I

f.". O" **

A

- - QA RECORD WHEN COMPLETED ***

Q l

DESIGN REVIEW REPORT l.

Design Document (Number, Revision, and Title):

TC-761. Rev.0. " Independent Evaluation of 10CFR71.4510 a Longitudinal Load on Reactor Vessel."

l 11.

Method of verification:

i E

Design Review l

l O

Altemate Calculation l

i O

Qualification Testing til.

Scope of review:

Reviewed entire calculation for methodoloov. assumptions. mathematical accuraev. results. and conclusions. These were found to be acoropriate for the intended function; i

O Continued? Oyes E No IV. Attachments:

Basic Checklist (Attachment 2)

O Yes E No Calculation Checklist (Attachment 3)

E Yes O No l

Modification Checklist (Attachment 4)

O Yes E No Qualification test (s)

O Yes E No NPEP 200-11 Page 1 of 2 Revision 4 Page 9 of 15

.r\\

U t

e n..

4 Design Document (Number, Revision, and Title):

l TC-761. Rev.0 " Independent Evaluation of 10CFR71. 4510 o Longitudinal. Load on Reactor Vessel" V.

Findings:

None 1

Continued? Oyes E No Design verifier Karl K. Gross d

.( 2 d[12,[7 Print

/

' Sign Date VI.

Resolution:

nb Continued? Oyes O No Design verifier Print

/

Sign Date Vll. Approval Sls/53 Responsible Manager J Wu 4,11<.L GN' Print

/

j3ign Date NPEP 200-11 Page 2 of 2 Revision 4 Page 10 of 15 m

l i

lL__-___-____-_______

l l

iA 4.4

"*QA RECORD WHEN COMPLETED"*

T--

C "4;

OG CALCULATION VERIFICATION CHQ% L

)

M.

1 Calculation No. TC-761 Verified By K.K. Gross Date 8/12/98 I

Title Independent Evaluation of 10CFR71.4510 a Longitudinal Load on R.V.

l Yes No N/A l

1.

Are analytical methods appropriate?

E O

O l

2.

Can a person technically qualified in the subject E

O O

l review and understand the calculation without recourse to the originator?

{

l 3.

Is the calculation mathematically accurate?

.E O

O 4.

Are the calculation objectives and design inputs E

O O

l

)

(including source) clearly defined?

5.

Do calculational parameters comply with design E

O O

{

criteria /dimenslaps?

l n

6.

Does the calculation reference or list major E

O O

l

()

equation sources?

7.

Is input data appropriate?

E O

O 8.

Are assumptions properly referenced and correct?

E O

O 9.

Was an applicable and validated computer O

O E

program used?

j 10.

If applicable, were program error notices reviewed?

O O

E I

11.

If an unverified program was used,is an attemate

~J O

E I

check calculation completed?

~

12.

Do the output results seem reascnable?

E O

O NPEP 200-11 Page l'of 1 Revision 4

}(

Page 13 of 15 lE

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O Calculation TC-761 Rev.O Page 10 of10 References 1.

RVAIR SAR Appendix 2-2, Burns & Roe Calculation " Longitudinal Restraint and Tiedown St:uctural Analysis", dated August 11,1998 2.

PGE Vendor Drawing 6478-MI A(2)-1-4 (Chicago Bridge and Iron General Arrangement Drawing Cora,ct 68-3780) 3.

PGE Vendor Drawing 6478-MI A(2)-68-2(Chicago Bridge and Iron Outlet Nozzle Assembly Drawing Contract 68-3780) 4.

PGE Vendor Drawing 6478-M1 A(2)-12(Westinghouse 685J837) 5.

Welding Research Council Bulletin WRC107, " Local Stresses in Spherical and Cylindrical Shells due to External Loadings," Wichman, Hopper, Mershon, March 1979 6.-

Chicago Bridge and Iron, Cor. tract 68-3780, Material Test Report Package

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