ML20211B090

Analysis Rept Structural Analyses of Main Steam Check & Isolation Valves for Prairie Island,Unit 1

ML20211B090
Person / Time
Site:	Prairie Island
Issue date:	09/15/1973
From:	Elliott A, Wang K, Willey C NUCLEAR SERVICES, INC.
To:
Shared Package
ML20211B062	List: ML20211B056 ML20211B090
References
PI0-01-06, NUDOCS 9908240220
Download: ML20211B090 (112)

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September IL, 1973 PIO-01-06 ANALYSIS RSPORT STRUCTURAL ANALYSES OF MAIN STEAM CHECK & ISOLATION VALVES FOR FRAIRIE ISLAND, UNIT 1 i

v Prepared for Pioneer Service & Engineering Company j

by l Nuclear Services Corporation

, CAMPBEU.. CAUFORNIA j

Approved by:

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September Ib.1973 PIO-01-06 ANALYSIS REPORT ,

STRUCTURAL ANALYSES OF MAIN STEAM CHECK & ISOLATION VALVES FOR PRAIRIE ISLAND, UNIT 1

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Prepared for Pioneer Service & Engineering Company b?

l Nuclear Services Corporation CAMPBELL. CAUFORNIA

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) Nuclear Services Corporation l ABSTRACT I The main steam isolation and check valve, manufactured by Schutte &

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t Koerting Company for Prairie Island Unit 1, were analyzed to verify the ability to withstand the dynamic effects of the closure event

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following a postulated pipe napture. Elastic-plastic finite element techniques were utilized with both static and nonlinear dynamic analyses.

i The valve components examined (the disc, tail link, rock shaft, and valve body seat area) were shown to meet the design criteria. Additional analyses were perfor:ned for the isolation valve to evaluate the effects. 3

x of spurious closures on the ability of the valve to withstand a closure following pipe rupture. A measurable limit was established for the allovable permanent deformation caused by spurious closures.

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1.0 INTRODUCTION

1 2.0 DESIGN ANALYSIS CRITERIA 3 b

3.0 DESCRIPTION

OF CHECK VALVE COMPONENTS AND MATERIALS o

h.0 LOADING CONDITIONS 0 5.0 DESIGN ANALYSIS METHOD 10 I 6.0 RESULTS II 0 I'

7.0 CONCLUSION

S 22 8.0 PIFERENCES 23 APPENDIX A: Analysis of the Disc & Valve Body Seat Area 2h '

APPENDIX B: Analysis of the Tail Link 66 I APPENDIX C: Analysis of the Disc & Tail Link Interaction 86 APPENDIX D: Analysis of the Rock Shaft 92 APPENDIX E: Material Properties 96 I

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1.0 INTRODUCTION

This report, prepared for Pioneer Service and Engineering Company, presents the results of the structural analyses performed by Nuclear Services Corporation on.the Prairie Island Unit 1 main steam check and isolation _

'i. 1 valves for closure following a postulated pipe rupture. The isolation j l

'E valve was also investigated to determine the effects of spurious trips ,

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upon closure following a pipe rupture. j,,]k a

The components analyzed were the rotating assembly, which includes the .

g disc, the tail link, the rock shaft and the valve body seat area. The effects of closure on the valve body proper, attached piping system and I other components of the valve vere not a part of this analysis.

I The primary load conditions imposed on the valve assembly are a result of the kinetic energy developed during valve closure. This closure .

results in a centrifugal force on the tail link and an impact force between the disc and valve body. The kinetic energy and angular velocities of the closure assembly were determined by Nuclear Services Corporation, and are reported in Reference 1. Secondary loading conditions, such as the pressure acting on the disc, asymmetry at impact, and post impact effects, were also considered in the analyses.

For the tail link, disc, and valve body seat area finite element models -

I were develered nnd etwwie elastic-plastic analyses were performed to

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determine load, deflection, and strain relationships. The total energy absorbed by the disc and valve seat as a function of the impact force was I then determined from the combined static load-deflection characteristics

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of the disc and valve seat.

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! I A multi-dimensional, nonlinear dynamic model van developed to determine the distribution of valve: seat loading upon disc impact. It'. served to t

check the assumption of axisymmetric deformation employed in the finite element calculations.

4 A one dimensional, nonlinear dynamic model was used to determine response characteristics, including rebound radial acceleration. A sirdlar model I was used to determine dynamic response of the tail link during angular closure.

Structural analyses were also performed on the rock shaft and the shaft portion of the disc where the tail link is attached. Both of these areas I were examined when the rotating assembly was subjected to the peak centri-l fugal force.

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2.0 DESIM ANALYSIS CRITERIA The design criteria as defined for the main steam check and isolation valves ,

I following a pipe break is established to insure that the valve disc, body %k

j seat area, tail link and rock shaft survive the closure and perform the [~

intended function of preventing further steam flow. The valve is considered to meet the design criteria if, following a pipe break:

1. The maximum material strain level within the components is less than 50 percent, of the ultimate strain; and

2. The resulting centerline offset between the disc and valve body is less than the specified disc to seat overlap of 0.625 inches.

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3.0 DESCRIPTION

OF CHECK VALVE COMPONENTS AND MATERIALS Y

7he main steam check and isolation valves for Prairie Island Unit 1 are supplied by Schutte & Koerting Company (S&K). The general configuration of these valves and their principal components (valve body, disc, tail 4

link and rock shaft) are shown in Figure 3-1. Closure of the isolation -

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valve is initiated by a release mechanism which forces the disc into the [

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s 1 flow, thus allowing the steam to accelerate the disc towards its seated l., .

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Table 3-1 identifies each of the components which vere analyzed and the [r drawing number ; defining the appropriate design. The check and isolation  ;

valves are identical except for the tail link. The tail link for the { i

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check valve is slightly stronger as a result of the additional thickness i a

for the limit stop (Figure 3-1). In addition. Table 3-1 identifies the ,

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> material of each component. The material properties are defined in Figures f E-1 to E-5 of Appendix E. Except for the sent velds, material properties from lot tests at room temperatures and properties at the operating tem-

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perature for the disc vere supplied by Pioneer Service & Engineering Company ,3 (References 2 and 3). All other properties were obtained by Nuclear  ;

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Services' Corporation from the references noted in the figures. As can $

I be seen from the stress-strain curves, the materials utilized for the

..np valve components are ductile and can absorb large amounts of energy by plastic deformation.

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S&K COMPONENT DRAWING NO. MATERIAL MATERIAL PROPERTIES  !

Check & Isolation 70-XA-16 N/A N/A ,

i Valve Disc Body h10 CB Figure E-1 ,

73-E-49 E-306-Disc Seat CL-15 Figure E-2 Tail Link-Isolation 70-E-lk" A216 I Tail Link-Check 70-E-85* WCB Figure E-3 l

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k.0 LOADING CONDITIONS The primary load conditions imposed on the valve assembly are a direct result of the kinetic energy which is developed during valve closure.

Energy levels and associated angular velocities are given in Reference 1. U; l

In comparing the check and isolation valves, it can be seen from ,

Reference 1 that the highest energy levels and associated angular velocities are experienced by the isolation valve with the check valve ]

in the full open position.. These vorse case levels were utilized in the analyses for the isolation valve and are identified in Tabl'e h-1. The ,

design pipe break level includes a 12% safety factor per Reference h.

G g Resolution of the energy generated by the closure results in two primary '

loading cases: (1) impact force between the disc and the valve body 3 seat area and, (2) centrifugal force in the tail link. In addition, a i number of secondary load conditions are present, including centrifugal reaction at the rock shaft, centrifugal reaction at the disc shaft attach-;,

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ment, restraint reaction on the link due to restricted are travel after ,

impact, pressure on the disc, and rebound forces.

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. TABLE h-1 i

KINETIC ENERGY & ANGULAR VELOCITY -

DEVELOPED BY THE ISOLATION. ,,

VALVE DURING CLOSURE I 1 Input. l sj

. Angular Energy,E s Dise Velocity, 6

Angular to (103 . ;j .. .

, Operating Condition Travel.0 (Rad /See) (in.-lbs)' '

Pipe Break 80 93 9 1.20 1

Design Pipe Break' 80 99.6 1.35 I Spurious Trip, 130%

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50 DESIGN ANALYSIS METHODS 51 Analysis Method Sunnnavy The analyses were performed in the following steps.

A. A finite element analysis of the dise and valve seat area assuming static conditions and axisymmetry was performed. Impact load, deflection andistrain relationships were developed frcm .the results of this analysis. By combining the load deflection characteristics j of the valve and disc and equating input kinetic energy to strc.in' '

energy, a-relationship between impact load and kinetic energy was ,

determined.

f B. A nonlinear, single degree of freedom analysis was performed for the disc to valve impact. The purpose was to determine the potential for separation between the disc and valve following initial impact., ,

A separation would result in an additional energy input which was'- (

S not considered'in the previous static energy analysis. An additional ,

3 objective was to determine the post impact rebound velocities so that the effect of the rebound centrifugal force on the tail link. J could be determined.

.fi C. A nonlinear, multi degree of freedom model of the disc to valve  ?

impact was utilized to confirm the assumption of axisymmetry of

'i impact loads utilized in the finite element analysis.

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D. The dise pin attachment to the tail link and the rock shaft were analyzed when the disc van subjected to the maximum centrifugal force.

! E. A single degree of freed m , nonlinear dynamic analysis was performed on the tail link with the disc mass during the angular acceleration [

'l J prior.to impact. The purpose was to determine the influence 'of' the dynamic time dependent centrifugal force upon the tail link i-g .$.~

3 reaction load in emparison to a static centrifugal force. .

F. A finite element analysis of the isolation tail link was performed.

1 Both the effects of the centrifugal force prior to impact and the effect of the disc restraint upon the free are travel of the tail I

link following impact vere e.amined.

52 Static h, Elastic-plastic analyses methode vere utilized since the available clastic I

1 strain energy.is insufficient to balance the input kinetic energy and the b l

The finite element computer applied centrifugal forces are above yield.

program MARC-CDC (Reference 5), which has the capr.bility to perform elastic-plastic analysis, was used in the stress analysen. A portion of j

the theoretical basis of this computer program is described in a paper  !

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by Marcal and King, Reference 6. 4 g

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governed by the Prandtl-Reuss stress-strain relations. Isotropic strain @

.,.-m-hardening effects are taken into account, and the analysis follows an M I incremental procedure. Initially, the program performs an elastic AnM Mt-I ,

analysis to determine the element with the maximum strain. The loads are subsequently scaled so that this element is at incipient yiele The Nm y,B Y[7, U$

loads are then increased in specified increments. With each increment, vy, l%

results are obtained which show the changt of plastic deformation.within

- the structure. The method utilizes small deflection theory. g i

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In describing the structure, the cross section is broken into quadri- 1

7,3 l lateral and triangular constant strain elements. The element type used Ji to analyze the disc and valve body was the axisymmetric, quadrilateral e*

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1 Q1 The finite' element models for the disc, valve body, and tail link;are ' l shown on Figures A.1-1, A.2-l', and B.1-1, respectively. Appendices A }k ] '

i .,I and B describe this portion of the analysis in greater detail.

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- Dynamic nonlinear analyses techniques were utilized to assess the symmetry A.

of disc impact, rebound characteristics and effect of the time dependent radial acceleration upon the tail link during closure. All dynamic analyses were conducted with the aid of the computer program PIPERUP y, e

(Reference 7). The PIPERUP cceputer program performs nonlinear, elastic-

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plastic analysis of three dimensional systems subjected to dynamic time history forcing functions. The program computes and outputs restraint spring reaction forces and system deflections as a function of time.

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PIPD1UP is an adaptation of the finite element method. The three dimen-

sional system is mathematically modeled as an assembly of weightless structural members connecting discrete nodal points. Weight of the system including the distributed weight and concentrated weights is

'E lumped at selected model mass points. An incremental procedure is ,

I used to account for the nonlinear effects of plastic deformation of the system and restraints.

Stress-strain characteristics of the members which connect node points are idealized by three linear segments. The first portion represents i linear and perfectly elastic behavior, the second represents linear' S

= strain hardening, with the third portion representing perfectly plastic behavior. In situations where stress reversal and unloading occur., an ,

isotropic strain hardening model is used; that is, unloading is always e along the. elastic line. (Figure 5-1).

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two subelements which are defined to have a sum stiffness equal'to the strain hardening stiffness of:the system. At the second transition the r 1

process is repeated leaving a single subelement with a very small n stiffness.

Restraint springs are modeled in PIP D UP vith an 11tial gap and a tri-linear stiffness curve. Again, the first stiffness represents linear a

elastic behavior, the second stiffness models linear strain hardening, with the third stiffness modeling perfectly plastic behavior. (Figure 5-2).

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m u Two s.'rgle degree of_ freedom-systems were modeled. For one, the disc and a portion of the tail link was the mass, and the tail link " radial" stiff-ness was the spring. The time dependent radial acceleration due to angular t

i velocity was inputed for the maximum energy condition. In the other system the dice impact was modeled, thus the spring included the effective stiffness of the disc and valve combined. From this analysis, rebound velocitie's and peak loads were determined. e b

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j. Q. v-6.0 RESULTS The results of the analyses are given in Appendices A through D. Appendix  ::e

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bO A summarizes the analyses performed on the dice and valve body seat. area. Y f .

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E Appendix B includes finite element and dynamic analyses of the tail. link.

g yQ 9,4 Appendix C and D contain analyces of the pin attachment area of the a@c, E dice and the rock shaft. +

rl KIN I , The total maximum strains at several representative operating conditions  ; y<

fe

!E given in neference 1 for the dise, t 11 11ck, valve boar seat area, and I3 y%

rock shaft are identified in Table 6-1. As can be seen, all trips produce ,f[g[

.s .

some plastic deformation. Design pipe break, the worst condition,Jresults in maximum strains which are all below the allovables as ohown in Table I 6-2. (See Appendix A.1, A.2 and B.1),

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d I The tail link when subjected to the maximum centrifugal force of 1h8.9 kips Nf.L

$a -

I 6 which corresponds to an input energy of 1 35 x 10 in.-lbs extends:0 52 inches as shown in Figure B.1-6.

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.f' This is less than the allovable of .625 inch. %

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D, 3

I The results in Appendix A.3 and B.2 show that the isolation valve vill meet the design critieria for a design pipe break, assuming the measured n

gj 69 I maximum permanent deformation at the center of the disc following any  %

yp 4

combination of spurious closures is less than 0.19 inches. In addition, f E

3 the valve vill meet the criteria if three or less verse case spurious p$ y,;

.3.je trips,asdefinedinTableb-1,haveoccurredprevioustoadesignpipe I. break closure.

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l Analyses of the dise pin attachment to the tail link as shown in Appendix C indicaten that it was not yielded due to radial acceleratis.2 prior to i

5 impact. #

s in The analysis of the rock shaft as shown in Appendix D shoved that the j f- stress did not reach yield during the worst case angular acceleration.

,/

The results of the one dimensional non-linear dynamic analysis in Figures l

A.h.1-3 and A.k.1-6 show +, hat the disc impacts the valve body, moves back slightly and stabilizes but does not separate from the valve after initial impact. This confirms the static energy analysis assumption that separ-ation vould not occur as a result of the rebound energy.

Other results of the dynamic analysis sununarized in Table A.h-1 show the

'I low rebound angular velocities of the disc after impact. As can be seen

• rom Figure B.1-k the angular velocity produces a lov centrifugal force in the tail link. The strain force results in Figure B.1-3 for the tail link show that the rebound centrifugal force would not produce any addi- -

tional plastic dercmation.

i The results of the multidimensional, non-linear dynamic analysis of the disc impact indicated that the compliance between the disc and the valve T, r:

s:

l seat was sufficient to distribute the peak loads reasonably uniform ch l around the circumference. This can be seen by examining the peak response

.; E 3

of the eight springs representing the valve body in Figures A.k.2-2 to

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NuclearIServices Corporation A.h.2-6. This confirmed the validity of the assumption of axisynnetry loading utilized in the finite element analyses of the disc and valve body.

Centrifugal. forces from static analysis were utilized to determine plastic strains in the tail link. However, this is conservative since as shown by the dynamic analyses results in Figure B.2-3, the peak dynamic reaction is 21% less than the maximum centrifugal force predicted from static

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analysis as shown in Figure E.1 h.

The total strain reported in Table 6.1 for the tail link is based on the conservative assumption that the tail link is constrained to deflect the same amount as the center line axial displacement of the disc. With respect to this assumption, two analyses were performed. The displacement-strain results in Figure B.1-9 are from the first case with the tail link initially stress free which corresponds to a spurious trip. The '

displacement-strain results in Figure B.1-11 are from the second case with the tail link initially sub.jected to a strain state corresponding to the centrifugal force from a design pipe break trip. For both cases, maximum strain levels were greater than the strain levels from the centri-w fugal force only case. '

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I TABLE 6-2 j

SUMMARY

OF RESULTS FOR DESIGN PIPE BREAX Maximum Allevable Percent of Component Strain A11ovable Strain Total Strain Disc 8.9% 9 0% 99% l Disc Face Weld b.2% 10.0% , h2%

Tail Link 9 5% 12.0% 79%

I Valve Body 7 3% 12.0% '61%

Valve Seat Weld 8.3 % 20.5 h1 %

Rock Shaft * -

I l

• less than yield 1 I

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Nuclear Services Corporation 70 CONCLUSIONS n For the check valve, the analyses demonstrated that the comiponents ex-

, amined'(the disc, valve seat, tail link and M ck shaft) meet the spec-

l ified design criteria for a postulated pipe rupture. j For the isolation valve, the analyses d monstrates that the components .,

examined meet the specified design c.riteria for a postulated pipe rupture, ,

i assuming that the measured maximum permanent deformation at the center of the valve disc following any combination of spurious closuree'is less ~"

than 0.19 inch. Furthemore, the components will meet the specified I criteria if three or less vorst case spurious trips have occured previous to a design pipe break closure.

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8. 0 REFERENCES

1. Nuclear Services Corporation, " Analysis Report: Maximum Energy '

of Disc Impact Following Pipe Rupture Main Steam Check and ,

Isolation Values for Kevaunee Unit 1", KEW-01-01, July 30,1973, Revised September 6, 1973.

2. Personal Con:munication from A. J. Elliott, Nuclear Services.

' Corporation, to C. Didier, Pioneer Service & Engineering Company, on July _ 27, 1973, concerning lot test data for valve materials. i' I ,

3.

Personal Communication from C. Didier, Pioneer Service & Engineering Company, to G. R. Randall, Nuclear Service Corporation on August 28, 1973 I h. Personal Communication from C. Didier, Pioneer Service & Engineering Company, to G. R. Randall, Nuclear Service Corporation on' July 30, l 1973.

g 5 Marc Analysis Corporation, MARC-CDC: Non-linear Finite Element M Analysis Program", 1971.

6. Marcal, V. P. and King, I. P. , " Elastic-Plastic Analysis of Two- ,

Dimensional Stress Systems by the Pinite Element Method,"iInt. J. )

Mech. Sci., Vol 9, pp lh3-155, 1967.

{

7. Nuclear Services Corporation, "PIPERUP: A Computer Program for Analysis of Piping Systems Subject to Pipe Rupture Loads".

8. Personal Communication from C. Didier, Pioneer Service &' Engineering 3 Company, to G. R. Randall, Nuclear Service Corporation on July 28,1973.

e g 3-9 " Steels' for Elevated Temperature Service," United States Steel Corp., Pittsbur6 h, Pa.

-I 10. Aerospace Structural Metals Handbook,1973 Publication, Mechan-s ical Properties Data Center, Belfour Stulen, Inc. AF ML-TR-68-115

11. De' signer's Guide for Welded Construction The Lincoln Electric Company, Cleveland, Ohio, No. 1100.1. ;O

~

12. Personal Communication from A. J. Elliott, Nuclear Services Corpora-tion, to R. Sabertino, Shutte & Koerting Company on July 27,1973. I

13. " Steel:for Nuclear Applications", United States Stee.1 Corporation, Pittsburgh, Pennsylvania, 1967. '

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APPENDIX A Analysis of Dine and Valve Body Seat Area k

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TABLE OF CONTEXTS en A-1 Static, Axisymmetric Analysis of Disc g .i s

26 hh A Aximyummetric Analysis of Valve Body Seat 36 A Static, Axisynenetric Analysis of Dise and Valve h2 Body Interaction I Ah Dynamic. Axisyunnetric Analysis of Dise and Valve Body Interaction 50 in I i [.

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d; Nuclear Services Corporation I A .1 Dise ' Static N The finite element computer code MARC - CDC, which incorporates' elastic- j-I plastic

analysis with strain hardening, was used for the static stress JA

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analysis of the disc. The model (Figure A.1-1) was constructed; from 296 63' axicymmetrie quadrilateral and triangular ring elemente to sinkil' ate the  : ,.

dise geometry. Analytical accuracy in the area of the impact face veld was enhanced by the use of a very fine mesh with the appropriate material f3 d.

properties. .

I For the; purpose of the disc analysis, the boundary between the dise face g-

.x

.k;

.I weld and thelvalve seat face veld was modeled by fixing one degree of freedom for"a single node (Figure A.1-2). This restraint provided an axisymmetric reaction for the applied loads but allowed the disc to ro- ,

tate freely without developing shear at the contact surface, thus'realis- i .i tically simulating the zero shear and zero tension interface. ,,

The disc inertia loads were simulated by applying an equivalent .axisym- [ ,

1

.I metric pressure as indicated in Figure A.1-2. The pressure distribution corresponded to the actual mass distribution of the dise with an ampli- .,

t, . ;

fication near the center to account for the effect of the tail link mass. 2f The elastic-plastic stress analysis preceded by incrementally; increasing 3

' ,f e H the pressure amplitudes in a manner that perserved the shape.of the dis- L W.

3l tribution. The total " impact" load for a given increment equals the in- W.

1;

-- p-

- g tegrated area of the pressure distribution. For each increment a point 7 '

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I d j on a total force - equivalant deflection curve can be detemined. The e f

strains for each elecent of the dise model can then be related to the.

J total force.

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>Mi The total force applied versus equivalant deflection is abovn in Figure ~ g A.1-3. This curve was generated by equating the sumation of external  %

1 work done at the pressure boundary (equivalant to the strain energy) to ]p d the external work done by the total force moving through a distance equal

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to an equivalant deflection. The above relationships are given by

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- , F gdr Total Force = 2w p .r.dr r

a and }

Total Ihtternal Work I!L Equivalant Deflection a wW Total Force ,

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1 The total energy l absorbed (strain energy) by the dise at a given load-is .,

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equal to the area under that portion of the total force-equivalant deflec-

. ...+

tioncurve.]Thetotalstrainattwolocationsofinterest(faceveld' con- i '

N tact and extreme tensile element) as a function of the total force:is &

v u'"

shown in Figure A.1 14 The sequence of growth of the plastic zone vith increasing. total applied load is described schematically in Figure'A.1-5 (' '

The ,redicted actua1 eef1ection at the center of the dise as a f2nction g g,  ;

of applied total';1oad is given in Figure A.1-6.

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• E As a result of the large deflections encountered and the accompanying ,g,

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load increment nine (2,616 kips). The dashed extension of the curves cK t* ..>,.

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i. . A.2 Valve Body Seat Area - Static P 9

An axisymmetric, elastic-plastic, strain hardening stress analysis utilizing the finite element computer code MARC - CDC was performed to obtain the M

y equivalant static; response of the valve seat upon disc impact. The model d Ig (Figure A.2.1) incorporaten 125 axisymmetrie quadrilateral and triangular g T

ring elenents'to simulate the seat geometry. An increased density mesh s

9 was utilized in the area of the face veld in order to enhance the accuracy p

of localized compressive strain predictions. M 4

a l 1

' The disc impact load was applied to the seat as an axisym:netric annular . N pressure distributed linearly over a radial distance of 0.156 inches, as 5 "

shown in Figure A.2-1. This distribution corresponded to the reaction.

y distribution on the disc as predicted per Section A.l. The distribution t f

van determined by examining the stress state of the elements at the y

boundary of the disc while the disc vns loaded and constrained by the y b undary conditions of Section A.1. '

y 31; l

l h

}

The geometry of the. finite element model of the valve body was truncated $

both upstream and downstream from the valve seat. Fixed boundary conditions j for reacting the impact load were established along line A-B (Figure A.2-1) which lies 3.6 inches dovnstream from the face veld surface. In addition. /

the boundary along C-D, upstream, was fixed. One criterion for selection of these boundaries included consideration of the anticipated plastic zone.

3 Based on the extent of the plastic zone indicated in Figure A.2-h, the  !

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1

. W ,

0 Y 4

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y

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.n I spurious predictions of the plastic strain. An additional boundary selection i

criterion was based on the effects of the reactive restraint provided by the

~

(;

m o

of I large inertia of the valve both upstream and downstream of the valve seat area.

hX These inertias vill tend to prog 1ee a fixed condition during the I high velocity impact of the disc upon the valve seat area. 9 u,

t u h,

W Since the stress'in the valve body in the seat region resulting from an .

es-;

L I internal pressure of 1000 psi was negligible when compared to the impact stresses, it was not included in the analysis.

^;

,y. ,

3 m

• p.

-Q The results of the finite element analysis of the valve seat are presented >$

i 17 in Figures A.2-2 through A.2-h. The force-equivalent deflection curve $$%

Q n

of Figure A.2-2 was generated with the same procedure described in Section

• A.1. The maximum total strain as a function of total applied load for.

I .vo areas of interest, the face veld and the valve body parent material

[@

adjacent to the veld, is shown in Figure A.2-3. The growth of the plastic W I. 'h m

tone as a function of the applied load increment is shown schematically Q 6.

in Figure A.2-k.

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Figure A.2-4 Icad Increr.ent . Iced (Kips )', SEQUENCE OF GROWTH OF THE PIASTIC ZONE tl 1 AS A FUNCTION OF INCRFASING LOAD 4*

.I 3 2

573 688 802 yAtyg 3opy j k

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5 1.032 1.147 917 '

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Nuclear Services Corporation

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A.3 y^:5 E - Dise and Valve Body Interaction - Static 7;

r i.

v I In order to determine the total energy absorbed by the combination of the ,

(

dise and valve seat as a function of the impact force, it was ne'cessary to

\

combine the static models of the two components. This was accomplished by idealizing the respective force-equivalent deflection curves as I bilinear, elastic and strain hardening, and then combining the two

' 5, according to the lav of linear springs in series (i.e. k = h,g } to obtain a single force-deflection curve for the disc-valve seat system.- The resulting curve, Figure A.3-1, was composed of three linear segments corresponding to the following conditions: (1) both components elastic, I (2)discelksticandvalveseatyielded,and(3)bothcomponentsyielded. p i~

L The combined F-6 curve is useful for the purpose of determining strain

• energy, however the deflections do not directly relate physically to any r portion of the structure.

}

The area under the combined total force-equivalent deflection curve for E a particular impact force represents the energy absorbed by the system 7:"

I at that force. A plot of the impact force versus input energy is given in Figure A.3-2. g l a By utilizing this relationship between force and energy and recalling-the relationships between force and strain determined in Appendices A.1 and A.2, the =='H an= a strain as a function of input energy at regions of

- h

. p ,.

interest for each component was determined and is provided in Figure A.3-3.

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p -

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=

a In order to evaluate the isolation valve system for the effects of repeated spurious trips, it is necessary to consider the strain energy lost to strain hardening as a result of the spurious trip (s). ~ This .

energy loss is dependent of the input energy of each individual trip.

^

Since the input energy for each spurious trip is a random variable, the detennination of acceptable conditions at any given time was .[

accomplished by utilizing a limiting solution. For this analysis two I- bounds were determined, i

The first bound was based on the 130% full load spurious trip (kinetic energy = 0.15 x 10 in.-lbs, Table h.1) and represents the minimum number of allowable spurious trips. The allovable number of spurious trips was determined by recognizing that the capability of the valve to meet the design criteria at a given point in time is dependent on three energy values; that is, (1) the initial available strain energy, (2) the kinetic energy generated for. the design trip, and (3) the energy loss lj[%

during spurious trips. By applying the principle of isotropic hardening, the energy loss as a function of the number of spurious trips'vas determined and is shown in Figure A.3-4 It is required that:

UA t (K.E.)g + U (,) k where U

f = the initial available strain energy for a given d.

allovable strain as determined from Figure A.3-3 (K.E.), = the kinetic energy f r the design pipe break condition (1.35xlob in.-lbs)

-s

(

6 m . h3 ';

g!

.)

, . . , . . . . ~ . .

y 9 + ,

m um ,_ >.4;u

./

n

~

t a+

Y

'l Nuclear Services Corporation

" ,f I

Ug (n) = the strain energy lost to hardening after n; I spurious trips as determined from Figure A.3-h -

?> 1 L j a 'i; In particul' r, using the allowable strain of 9%, the appropriate quantities %

I for the disc are as follows -

h

[.1 ;

c- = 95 a l

.I 4~

6 U

A = 1.h0 x 10 in.-lbs I, I U3 (n) < (1.ho - 1.35) x 106 = 0.05 x 106 in.-lbs

? .}

'I i hence

, f n- =3 (Figure A.3-h).  !- d nw D.'

' ,; \

Therefore, the conclusion is that the isolation valve disc can be I

1 subjected to at least three spurious trips before the capability to i

,- J

,} d I sustain the design criteria for the design pipe break is jeopardized.

I The second bound was based on the occurrence of a number of spurious trips with undefined input energies and represents the allowabih upper I limit of permanent deformation at the center of the disc. As in the

}'")

3 previous case, isotropic hardening is assumed and the energy loss must '

r.

be less than or equal to 0.05 x 10. in.-lbs.

The allovable permanent Y set in the disc was obtained by first determining, from Figure' A.3-1, the 4

(

apparent impact force at the point where energy losses equal 0.05-x 10 I in .-lb s'. Utilizing the above' force value in conjunction with the curve tr' g

r in Figure A.1-6, the total centerline deflection (0.31 inches)..'of'the  ;,

[

[

Q.

. w -

- hh - ,

Y i*, ,1.

aa 3qau Jpa  ?

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y:.

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r..>

I disc van obtained for the allovable energy loss condition. When th'e

?.7 q

ti .r!

w elastic deflection (0.12 inches) is taken from the total deflection, I 0.19 inches is t'he resulting permanent (plastic) deflection. Therefore, 4

??[

it is concluded that the isolation valve vill sustain the design criteria for the design pipe breck if the permanent centerline deflection of the "

disc is less than 0.19 inches after a number of spurious trips at unspecified energy levels. 9~

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A nonlinear single degree of freedom analysis was perfonned on the disc 1,1 ;

to valve impact. The purpose was to determine the post impact' rebound velocities so the effect of the rebound centrifugal force upon the tail link could be analyzed. An additional purpose was to determine the potential for separation between the disc and valve following initial impact. A separation vould result in an additional energy I input which 'was not considered in the equivalent static analyses of I, Appendix A.3.

The model is as shown in Figure A.h.1-1. The spring stiffness used ,'

is shown in Figure A.h.1-2 and is a bilinear representation of the disc I and valve body combined stiffness of Figure A.3-1. The combined stiffness was obtained from the static finite element analyses discussed in Appendix A. 'A constant forcing function of 358.538 lbs. vas utilized. 'This force correspondes to the pressure at impact of 965 psi , 7.f.

l per Reference 1 and the 21.75 inch inside diameter of the valve body.

The initial gap between the mass and spring was determined in correla-tion with the impact energy and the forcing function.

I The disc response for the spurious trip and closure following a, design pipe break is shown in Figures A.h.1-3 to A.k.1-6. It should be noted I .

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be'seen thak no separation occurs, thus maximum loads predicted by the

.I static energy analysis are not altered by rebound. The rebound angular h 1 velocities are so lov that additional plastic deformation of the tail link will not occur, thus the effectiveness of the seal vill. not be 2:

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INITIAL PEAK REBOUND ANGULAR z

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'fg symmetry of impact loads. w 5  ;

?. 1 The model used is shown in Figure A.h.2-1. IU:

The model consists oflthree hi-basic segments'. M The disc is represented by four identical crossed bars p which are uniformly spaced. The valve body is simulated by eight dis- '

crete springs located vertically beneath the terminsi points of'the r bars.

h gg A bar element representing the tail link connects the disc. to a pivot. ,.y

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Appropriate stiffnesses have been associated with the various segments. 4 3,9 The four crossed bars representing the disc have the effective disc  ;%

W stiffness as shown in Figure A.1-3.  %<

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concentrated at the center in order to simplify modeling of the effective .,

g..

I stiffness. . The eight discrete rprings in parallel have the effect'ive m.: 4m

j!3.y valve body stiffness as shown in Figure A.2-2. The stiffness of the .%

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G:t tg The results of the analysis are shown in Figures A.h.2-2 to A.h.2-6. [HP E A It can be seen that the maximum valve body loads are relatively uniform 4 j i.%. ,.

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E around the circumference. This confirms the validity of the axisymmetrical y

finite element models utilized in the disc and valve seat analyses.

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. U ig y __ B.1 Static Finite Element Analyses !E l A static analysis'of the isolation tail link was performed with the ' aid of the finite element computer code MARC-CDC. The model shown in Figure B.1-1 was constructed from 160 quadrilateral and triangular, isoparametric plane p stress elements which incorporated material strain-hardening effects. jf. w Analyses of the model were conducted for two valve closure loading 1 conditions; , g k, . that is, (1) centrifugal forces and (2) rotation restraint produced by the N axial deflection at the center of the disc following impact. 39 The analysis of the effects of centrifugal forces was accomplished by rigidly fixing the model boundary at the disc-link interface and applying . force at the rock shaft reaction point as shown in Figure B.1-1. By incre- M-, .u menting the applied load from zero to lh8 kips, the force-deflection curve shown in Figure B.1-2 was obtained for the link. In addition, total plastic g. strains as a function of applied load vere obtained and are shown in Figure , n 1-3 Utilizing the above results and noting the appropriate relationship ,4 between centrifugal force and angular velocity shown in Figure B.1 h, rela-tionships between maximum element strain and impact energy, and deflection and impact energy were detennined and shown in Figures B.1-5 and B.1-6 respectively. Figure B.1-7 represents, schematically, the growth of,the , plastic zone in'the link as a function of the increasing centrifugal force. s The effect of the restraint imposed upon the free are travel of the tail ~: , link mechanism by the disc after impact was determined by conservatively I. 7 m J-Tm ' ." . 1 7, . a . Gi [' k."@  ; u d ..i. .~ lQ l[ h,g6 M 4%- -< f;;7%C;%$ , W- . . s. y - ma y gp. .s , yo , o sy.,m7 , g - ~s A 34 e Nuclear Services Corporation restricting the travel of the link to the axial path of the center of the disc as shown in Figure B.1-8. Reaction for the restraint force was supplied at the rock shaft. A deflection, simulating the axial center deflection of the dise, was applied along the restricted path and was incremented to establish strain deflection curves. .. - -i Analyses were performed for two operating conditions. One was for the link- i initially unstressed which approximates the spurious trip. The resulting,- strain axial deflection relationship is shown in Figure B.1-9 Figure , B.1-10 represents, schematically, the growth of the plastic zone in the link as a function of the axial center dise displacement for this condition. f i The second condition was for the link initially stressed to correspond to j the design pipe break.. The resulting strain axial deflection relationship $ is shown in Figure B.1-ll. In both cases the state of strain due to the restriction of the link was related to the overall closure energy by d f. utilir ng the disc center deflection from Figure A.1-6. ([ . j J 4 As can be seen from the results, the sssumed conservative restraint imposed I by the axini displacement of the disc after impact adds to the maximum f 3 strains resulting from the centrifugal forces. j 5 As discussed in Section A.3, after a number of trips, the isolation valve $aq will meet the criteria- for a design pipe break if the total permanent cen- i 0 g terline deflection of the disc is less than 0.19 inches. During impact { l < . . .jj - 69.- 4 -? .]'.  : , Y .ll .f .' j ., hd ,'[ LJ , ' k A su  ;[& ' '

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"  ; - 1 { > 1 p; ~ ~ :y l Nuclear Services Corporation B.2 Dynamic Non-Linear Analyses f A single degree of, freedaan, non-linear analysis was performed on the tail t link with the disc mass during the angular acceleration prior to impact. H l j The purpose was.to determine the influence of the dynamic time dependent l-centrifugal force upon the tail link reaction load in comparison t$ a static ' l centrifugalforde. The model is as: shown in Figure B.2-1. The bilinear spring stiffness was s obtained from the finite element analyses of the tail link as shown in Figure B.1-2. .,, I I e The centrifugal forcing function which corresponds to an impact energy of j .a 1.07 x 10 in.-lbs. is shown on Figure B.2-2. p The variation of angular Y .I velocity with time was obtained from Reference 1. ~' y v., 'it results of the analysis shown in Figure B.2-3 indicate the maxisum cen- -{ trifugal force at $mpact is 93500 lbs. "his is 215 less than'the maximum force of 117800 lbs' as determined from the static analysis as shovu in , Figure B.1-h. Since the centrifugal force from static analysis has been shown to be conservative, it has been utilized for all tail link analyses. I H . . 1 it t' l ) n \ ~ 02 - .x l ..,; .a

h &

c .t YV ' 4 , .7 . , p. - . p ' ,, po Nuclear Services Co@A .. CARErnELL CArnotNIA ! ' b-.d

Figure 3.2-1

• I

l. MODEL OF DISC, TAIL LINK

,f' ' kM -p . . F k . - y. h " (E A. in ' i

• j K

a ut-w

?. a .w U ,4 YYl l A } } 6 Os i. ,.'4.- ., ?? w 4 ~ ? i $/ . Dise Weight + I;i -! + and Portion of & V" Tail Link 328 lbs. > -c. "j._ f I. F(t) e; ' "fl. /. , f -a N ' b* 5 Q . ,l a N' ' i e l i 4 ,. /.

y L

i, m <,e. 4 +a, ,,!^ .'s pF-lI 'p. .* voi 1.1 9 l' i ,,, J. 1,l' A@ e. ll by i , e

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s ,, ; r;,.,.  :..,. , ,,_..- ,g,,. fn g_ _,,3. ._] 'g '

, f ,

-j 9 _f. ? ,< x > c . ~y .< ;. ?g- ; : ,4q < . (, ~;g, 77,;, ,; ..., , r g . m .. + y,; ,g  ;}{n;fy.r-. T ~ Neelear & A C4jnerosion * $ji-- p ,. MM .g i-g d

• 2 l Figure B.2-2

~ MRCING FUNCTION , I Due to Angular Velocity g e. t. l , i:-y;-

-f; .

e MRC(, " Max. ENERGY. i -E f. E 227 8 2"- 88 6 "S** 1.07 x 10 in-lb l' j Impact of Disc g ' f[ 100 pf .d. ,% P 90 >. 80 .} 70 li s " 60 . F=M I e 50 't E w ho i g 2, 30 4r a. 20 * - E Approximation A 10 .005 .009 L- .013 .017 M , MODEL TIME i <..b - 0 , ,. .E 530 534 538 .5b2 .Sh6 REAL TIME  :?n I g .t g f i I! ~< v ., , ? I ' Y s ,n t y. __ ; 'N S ~ ' .h s, . D .': r' . jh 1 d )3 g)7 ,p. 8k ' ' "y'" r w fr?q MQP 'l l7 % ~- >' *i r r

m

t , q.7 .

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- k L- + w. t ., t o,_ +

4 k -

5 o . . r. '} , O. a ^ t ., N. i sg l - .g e. t , V. l l~ , IE! ! o- N pact Of DISC u ,i, , . . o a u? u. m= a n , J ~ I a;u Ls4 v. t c i me,. o . a .0 e e( v o W m ' i

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u:) + 1 u - l k.co o'.os o'. l a o'.15 c'.2 c 0'.25 0 30 3e ? >, TINE.(SEC) n10~ ik A+ k p- RESPONSE OF TAIL LINK REACTION LOAD , q# 4n N.,s k O: ' 34

' Figure B.2-3 7qv,.

g 'n k; i, k nr 'itl S

• y.

• f

! l.QY . f .:f &g'P' Y'Y s *f]h'

7.  ;

C}c- ',. if *. ,~  % l >&.- - b ' gf , .k ql 8 (;ll: . x., ,, .;$f,: ;py;l[;llQ;lll@ , , ,%.y : ( ,~ '.; j'lhd'L *" r 7,.y,7j .- u _ e. . . , . .yp msg .~ , .gw , .o; > .w.r.i-@,.m. crrn M3iYwnW;Q '0. 0MWA. s$,Wi"?iVg.g. gp ,. m - yg ;. .,. - r. 3.,n;p y -s y g-F: (GU' h i 'q q% 0.-6,. .y 4*

o. ,

a 4 0-'

..p; ,

4 f 's l.i,; 6-k 4FP :d[- ,L J .. ; ,+ m- .,;;: , y .p g.r. 't ) .i> . hyR ' y= ', . !t 4 fy a . ru . yg 3 v.:p 1. .i h (,# G ( + 4 7 , .et.- y 9. ' )- u w r n'( ,sp ~ t 3 -- e* . *: "v , p :/ ..,}=:. 4f , It h h'. 3. , t APPENDIX C 4 ". Analysis of the + .w . Dise & Tail Link Interaction .. rs .I . 1 I e . .I g l I I  !

• i

.,~. t I k[ .- yg 4-t .i >q ] .4 ' ..,,.t, o -i l / - ~ 4-b AtI. I s t j iM- CL< ? F. . S 5,.: ., .O j' .e,;,,' w ',.d' s' . g* @o' i ,e. 4 t-41/,. {gesff *, ^ # e4 A ' d

"f- ,%. r

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c. o : ~,-

j'.d ,7,-% w-

o- ,6"c,3> fy;; . rsf/p -

' i. i Syf , - jsl. ..i__, + s

• !fi,QQfQ( =

's/t 7 '^ * )  ; ~ f [4 a, i ..k' ^ .Euclear Services Corporation , .e, A. r$ ;ll C. An=1va4=' of the Dine and Tafi Link Interaction vn, Vl The reaction of the centrifugal force upon the dise through the disc. pin is 3

&p2 R examined in the following analysis. The model is shown in Figures C-1 and C-2.  !

Nd It is assumed that the pressure on the lover half ring contact area between the disc and tail >1 ink is a cosine distribution. Thus to determine effective ,?ff, I reaction location-of P

+ s a. T I *l2 Y 2 / r,, 2 *. P = paa = 2pr sede rdr=pfr 2 2 ~ '# 1 JJ ' o M r. 1 5 ;p: q-l,'. l , ,y I 2 ,) Fy = 2p w/ 2 es0 2 2 r dr = l y w\ [7 2-r)' 3 3 3 1  ;' 6 r \ 2 l' N }- ' I 1 -A L ll thus I 3 3 + r - r . = Py = w 2 1 . Y 2 - P r 6(122'#1) / N, c.- f {r,- Ma4 - r 2- " 2'5 i"* 3 r 1 = .1.238 in. 58/ i;> i y y;(- g .. 5 The location is obtained ph , e, W, \' .I GW ) y =11.52h Nis , b Y and 2

I

_ n._ n l i P.' =,p n x h.7173 :s i fi W s M E, .i ~~ ;g y- ,) . " vi:

• ig y 3,.

?E S am +'87 - ' - 4, D 3m t-E ~ $$;& ll i*}. +s ?h;. J:%:?? I A '. ,~~ int % %.2 , , < ~ . . uin&p.?; ,:; -sn :.. . ., > L;Mi., &, < ' '~ l 1 p:( ,. ,; : mc;;g ve. m, . ,4 ,? ,.7 . , r. s g. w Sk. Nuclear Services Corporation n . L By examining the free body diagram in Figure C-3 the major forces are identified. V;

The ?=v4= nan centrifugal force which corresponds to the maximum e = 99 6h radf/sec. is

@

2 'F = 'mr e = x 15 75 x 99.6h .... p ' l I. F =: .122,536. Ib N , 7 The total contact force between the tail link and dise is: }j w y ,; 122536. x 1.913 = 153813. Ib I 1 524 x, The maximum contact pressure which is the maximum nomal stress of '9 I m-m becomes: j as ,. P ,153813 = 32.6 ksi g Pn ' 4.7173 h.7173 +, " A e..

Hence, the maximum stress produced due to the contact is less than the 1

ff yield stress 'of 70 ksi per Appendix E. I A.V:- 9. To determine'the stress in the disc pin the forces from equilibrium , w are: F 'I 4 1 T = F = 153813. Ib S = F = 122536. Ib

r. i.

-I, J ", i! < .h y bdI? 00 - iic" [ $ t { f R

• $ }Ml;Ly .q:

t .. u.z. r. z. ym ;x i, w , g :, .y ;py. ggy ,. . . a [kn , yb . [ ', f 2.hdP, , !J-fi! i-l' s , ; w.< /s his@!jd-Q % 1 4, - vW id 7 ~ ~ , 1 ,.4 y ' ' ' .. . .. . , . . g .. . , n./. m v . ym.er.~ n - , m.: gem r .. - w. r c mma;m,. - zsx, i , .xw .a . x . -E

a.g,w; - '

,g .. m LNuclear Services Corporation W.. ' ;E %c .. w.cfr;; ..y The maximum' stress which exists at point a in Figure C-3 ist a f., f ,x , 1A ,153813* = 31.87 kai ., . i. 4.827 EI l - 4,. a w kS = h_._ x 122536. a i o,. , 1, y xy = 3 A: 3 x 4.82T = 33.35 kai  % e , , :s w - f9 j p- ' 4: I! .i gp 'l. v The calculated principal stress at C are: y;. jw t .i p@ . . v:; C. o= 1 O+ 2 - (N 2- + 33.85 = 15.9h + 37.h1 ~ g a vh .T m. .gi . e I a' l = 2 53 35 kei tension ,-21.h7 kai compression MSG p.M. , - w X,y .I ' +- T max -= 37.h1 r, 6. ,.-r y r. e, 4 I By applying!the Mises yield criteria 4 wr x.e; , M8p ^ .$.* u A. y.g: e a 1 la -o\2 + 2 2 gg* n: g eff = (1 0, - c + 0 ~ 0 Y j- 2 2f - c 3 3 1 4:. g;r E

]-' with jkE e = 01, M m 3 M the effective stress is

{2 (; 9l$'.,. 7; qM-' ? 14g o aO, eff = 66.73 ksi W- ..I-a py sG; .es it?;k, ~ip 5 "hi*h i' h' " the yield 8trenEth f7 k81 Per Appendix E. gj9:l 3 u k "M kV l ay. h i* 4 ?: + ' , 4.tm.h.; n . ,.7 4l $*gts

• :. f If -

s i,  :;" - a?jj% y Q. .N ' 'g - b9

• e. Aa

,pJ-E .s ' ' r+ . 4 - iQ I.' n M. L t; y46 y9: %y .e. . g.., , . > w.c:,. .w -a ,. -- > ;e m- m s c, + <;s qs. g ..2r, tw t ;c-;. . m,,uc, ,e.glo A.m, c .r . . .s. . .w.y 4.s4..nt t -y ay y w /u.. - v.v.. .<, - tv. . s -. . llI?C .E. env. ... [ , FIGURE C-1 4 $. . , 5!l jE TAIL LINK & DISC ATTACHMENT AREA [, -j $< -l ' . s. 4Y D' 4,e $i m . n. : <. 4 e

• o b

e r 2 b J' s b. ep .J' - 4 e y : A [ -- 1 g m% C a - ,t u .. . y 6 ,. r l a 8 9tw 4, <- , 3 (.  !? < is . + r 34 1.tJ J. l -y < , y * ,- 'i 5 w ;e - i pc .y -

t m t

'a " .. A t  ;% '

. N I -

i b??: , ,. 3,i , f G FIGURE C-2 M 2 i I ,E SKEMATIC OF DISC & TAIL LINK AREA yy s. 3 .' . 4,:(0 -i .. M" ~ b- e Disc y ?E >, tg  :- Part of Y~ l .[ all Link- l '; l8 n / / / .e ~//, // $ / di ~, = [l' ,. 9 Critical Point Examined 4

U ,

( 5.00 in. 2hT9 in. C J- + f',;, ' ~2 .'-

~

jj s / ,; , $ . '/ / ,y5// / 'j li:i . (lE' 4 . /, i '(

i / Disc Pin  !?L y

l .' g i - L Total Contact Force a Eetween Tail Link and Disc .{ .. p 7 jf./g

4. ,

y +A'~ _ .F VT 5 y-F. Centrifugal Force hl t & b

• - ..'i, rh

,a . .m.. 4 E, I ? kW' .f. r'o a -- - , .,3 , s,. ?, * . s ..y ff', l ' ,e j i; Nuclear Services Corporation CAMrnau cAurowin } FIGURE C-3 FREEBODY DIAGRAM OF TAIL LINK s'; AND DISC CONTACT AREA * -4 o Y ,; A __ _ 1 913 in Y l il / - a / // ' l A Q g // .'/ / '/ ., Disc Center of $5

3 I l>

u T 2 d ' /*/ */*/ // ' a , Cravity / X l l P g j / / 4 v , f I 4 5 h F - \ a 'h i n .;i i s /y I- , ER [3 x .9 I s . Li'l ^ (t?) 64 kD a,

• $.i ' k . % ,; , , , _ , _: g. >

~ A f,' b , ' ' -. > ' . _' y & p em ey.3. e 4 ,.m.w .g c.

g. w u x, . .

.g , ,, s ._. ~,,.- u. ...w 1-' - . i - 31 w' ' ' Nuclear Services Corporation &.:. k @nc +4 A '< 9.*k <$) q

.= .

L, y% 5,' ~ , h,,.te m, , 1,: i .#..,te M y.. % k Ih.. h, ,s. a Js t. I!l a 'h .',/. ?. kh W. -G sgt 1:' ,sve# ( r3'/ p% ..; . ,, ,sr me . t,WR.. s 1 Fyti . , :+ , eP-- 49At APPENDIX D >Al9 , s,w V' JJ} h . 'r , +rh.Y ' d*k . Analysis of the Rock Shaft m hf I k.# vr!d 55  % .u iNyk g eg.d. e I. .h snJ > Qa: %,- y ;A a L, 'g Y N. 6 , l%9- %@f . '$k$ . ve, + .-l 5 h q'. '[f.

y. 2syj .-e a/

S wp39 l . e. y o ,9;f #k A , n-I. '.gQ 1

? f *)

j n A '9.~u n '(% i\$ s 50&c p.,,n.Nh j t eh -, . Ys Y' .\fu3%f.l sa. lv. '{ ' 8 p+ -/ rQ-)n"% rat'. c. > {J_I%f. ' 3.ty s %,,g l[gy; i .w p y ~ A' - 92 - $(&n%h$ L $g '; ;;- .i <> w- +y., .g. g '.f- s +2 A e' UN l . ,3 "f t.-- w;. * .. I,*- = P. 9 ' / %g .z.. . D'd' ? . ,. g .p J *,e 5- (,-

• f

, i- ' ._,.,) AA y , ' ' i yg ..q:g; s: <- 'ijyipr ;qr;7ny9.gyey qqmy.y;rys.< , pi:7 . x ycy s 9 9 y r Q: w , > * ;g Nuclear Services Corporation si. c . d4 . M, , D. , A_n'alysis of the R. oc.k_ S.ha_f.t, . 3

.?

The effect of the centrifugal force due to the effective weight of the tail ) link and disc is examined in the following analysis. The model is shown [)d in Figures D-1 and.D-2. 3 s; h, 1$ By referring to Figure D-1, it can be seen that the torque Tyis pmduced by pM.. 3 1 veight and pressure and T2 is the restraint torque produced by air pressure uyf in the cylinder acting on the piston. After rupture of, the piston in the t k l% cylinder, T = 0, there vould be no torsional stress in the rock shaft. Tj 2 t '~ 4 w 1' The mav4== torsional stress produced in rock shaft occurs when T 2is maximum. Qi When the piston ruptures, the pressure in the cylinder = 150 psi, per' Reference 5,

8. Also.5550 lbs.?on the piston stem produces 233. psi in the air cylinder. a G'

I_ Thus the force in the piston stem is: ,.c x 5550 = 3573 lbs, "? h 23 . N, and the maximum t rsional stress is: y 9 T 2 max = 3573 x 4.875 = 17hl8 lb.-ins. g; ! I Torsional stress in the rock shaft is W. ? 'I , max ,T2 # g. max - i; 1 , I* e m iu 62 y ,

;  ;.ft I =H h e-- o -r 2

= 1 57 I d '~, f T max 1= 1Tk18 x 1.0 = 11.1 ksi i l 1.57 rp,Q R [ 7 g ( $/ 8 [ , Ugj . >. w t , ~ 'L j% .t . . 'y?, 5%;;I ( G* 9 , Y N7/ v;. [ .[ g ,O.. -r s. ) '. ' 'e .. .. , 'd r _g 3y 3a - j Nuclear Services Corporation f i re 4.5 5 .f

S The maximum direct shear, V, in the rock shaft due to the centrifugal force-

'is F max l . l 2 vhere F,,x = m,ffrv ,,, = 2 x 15 75 x 99.6h = 1h9083 lb Ll V ,. = m = Th5h2 lb. I. , Maximum shear stress is S max

bV

b x Th5h2 = 31.6 ksi,

I F 3A 3xlkk I where ,

A = cross section area of the rock shaft. i By conservatively assuming that the rupture of the piston does not occur, the total maximum shearing stress = 31.6 + 11.1

= h2 7 ksi '

I which is also the principal shearing stress. This is below the yield stress of 85 ksi for the rock shaft, jf ,

./?'

a r3

/

I F  ;

_ gh _

!l 3

p. -

1

, , . 3 :. ; , , , s. M ^ *

. ~h ,  ;.,

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i.

y Sp :4(t, - ,,py;, , , , ,,

.'. - t_

m, e 9 3cy.; 9,,, .: a.g 7 , ,

4 . w -- y>

w

. Figure D-1 r rg_;

' SCHDfATIC OF TORQUE PRODUCING STEMS ON ROCK SHAPI' M

aq, I-w 4 ..

r ?7.o.. .

Nl

. j I

15 T Rock Shaft N N, g;.

~

I O- - - -

\= x nh:

s +n

/

/ 1 e h.875 1 % .

N t v .,

jL N! -

Disc "3

')i T

/ / (m 1 't .

L

/

l a.,

t~

( T.

2 mru 150 psi b[

~

NT:c Figure D-2 '

7 SCHEMATIC OF ROCK SHAFT, TAIL LINK & VALVE BODY ASSEMBLY :li k\\\\\\\\

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APPENDIX E e'i

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- Material Properties a. e.

w .s.

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f- *p.e  %.

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I E. MATERIAL PROPERfIES Since highl strain rate test data at the operating temperature was not available,-it was conservative not to utilize enhanced yield strengths.

Temperature dependent material properties were extrapolated from room temperatureflot test data and high temperature properties of identical I or similar' materials. Conservative bilinear approximations for the I" stress-i. train curves as shown in Figures E-1 to E-5 vere utilized for "

all analyses.

d.r The details of how material properties at the operating temperature were determined are shown in the following sections.

s I 4

,;l.9 I  :

l  :

I

I 4?

.,1

'I .

l pI

- 97 -

. k,c

? .a ?' 5 da, a ' 'J u ; , y - s. A s <~ $ 64$ss W .A >t' I "%

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, .. . m.n 3~

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g , ,

,, ,r g g .

.. . a ..

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-l

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, c Nuclear Services Corporation  :

E '

1

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E.1 Disc h10 CB Martensitic Stainless Steel

] Per Reference' 2 the following room temperature lot test data vas:obtained: f j' e = 79.'kai

. Y 2 '

au = 103. kai eu= 22. - 2k % -

Per Reference 3 the following lot test data at 600 F was obtained:

o7 = 65 9 - 67. kai o = 81.2 - 82.2 kai

• cu= 185: .'l 1

Utilizing the 'information from Reference 3 and Figure F3.0312 in Reference 10, .

.I which reflects the effect of temperature on bl0 CB, the following properties

l. at Sh6*F.are obtained:

)

' i o# = 65 9 + 600. - Sh6. { 110. - 95. }

600. 't 0y = 67 2 kai .

a = 81.2 + 600. - Sh6. {125. - 105.} .,

600.

o,= 83.0 kai ll

" cu= 18 + 600.200.

' - Sh6. { 22. - c- 0.}

r

].! cu= 18.5% F' n .

~

y ,, 98 ;Q

{ ;p., -

. ;..g y;, r ,.t - -

N'! ,

4 -

g e-q_ );?l ll$l- q ,

Nl ,

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. , c ..

Nuclear Services Corporation

?.;

fi}

. ,s

? :, h lI The elastic modulus obtained from Reference 9 Figure 17 at Sh6*F'- .,:.

ti eq

,.l  ;

i is: (J .1, I F ,

E = 2h.5 x 106 i .

1 Actual properties utilized in the analysis were: '

A o = 70. kei

• s

~.I '

y '

r.,

o = 80. k'st ' '

s c " 18. I

. u -

m E .' 2k.5 x 10 poi g

.5 pe For constructing the stress strain curve the ultirate stress was assumed to occur at c .15 c which is typical for stainless steels. .-

u

. 3 ..

The stress strain curve reflecting these properties is in Figure E-1. yk E # .

I .

A-

' \ '.

,p

, r y

99 ,, rg

, f?g.

^

• + f  :

,p15 '

5fb,! v il :' .';L; 5 L

, . . .

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,. ,e c n ,y ,

1. n , , cr p .

el-hf'~'

i.,,',:d ,'7 g

y =

m . i,;. ,;

s g

...c p

~ . . . .

7.;1 Nuclear Services Corporation 0 .

E.2 Dise Seat '

lI ,

Weld E-309-Cb-15 Austentic Stainless Steel I

{.'

h 1

I Per Reference 12 the veld material can be treated as annealed. : Annealed properties ac* room temperature per Reference 11 are:

.h

'. o = h5; - 55. kai ~,

Y di ou = 85. - 95. kai N

c".= 30. - h5.5 From Figure 17, Reference 9 the elastic modulus at temperature ic obtained: '

E = 24.5 x 106 psi As 309 is a modification of 30h with similar properties, the properties at elevated temperatures for'30k in Reference 10 are applicable. From Figure 3 03132 in Reference 10 the following were ottained: "I oY = 40. kai a = 72; kai c = 20.%

l The actual properties utilized in the analysis vere: ,..

oY = h0. kai '

=

ou = 72. kai cu= 20.5 [

E = 2k.5 x 106 psi '

l

'l '

i

a. ,D

- 100 -

,4, ,,

.r,.; ..e ;3e x y , u m .~. ,

.x ;g, 7v-p

.. +

c' Nuclear Services Corporation ,

't The ultimate stress was assumed to occur at c .15 c which is typical

y for stainlese steels. The strese strain curve reflecting these properties I

is in Figure E-2.

u,

+

w J

J )

I i

I .

1 P

I l I '

I l l I ,

l -

t 4

ll

_ ici .

y ~. <

..%,p p , .

.g, g agg ., _.

., , y, ,,,,g .

4 ,

p r .

p. x

\

- + 4.1

+

y .m ;

t t

Nuclear Servica Corporation y a s a

Y ll '

, s E.3 Valve Body & Tail Link

\ _ ,

S,i n e i

5-l, A216 WCB Cast Carbon' Steel y

h Per Reference 2 the following room temperature lot test data was obtained: pl "y = h1 5 - 46.:ksi

+j .

l

, ou= 73 5 - 75.5kai *:

., 4 i Ni cu= 26 5 - 31.5 .

t 8

From Reference 9 the; elastic modulus obtained at Sh6'F from Figure t

16 is. f{

0 E = 27.5 x 10 psi

,w Properties at the operatir4 temperature from Reference 13 are:

a f

I =37.-f5h6.Th00 400. {37.-25 II t.

md oY= 32.6 kai t

<h o

= 70.- f Sh6. h0a [70. - 65. - j l boo. >

y u = 68. kai M

43 c = 30.+ Sh6 ho 38.-30.f cu= 33.5 &

f4 0F Actual properties utilized in the analysis were: "f}

m

?/.

oy = 36. kai s:

d' ou = 65. ksi T

4 -

I

s. .:

N' 5 95

l .q

, . ,.s - m i; y _Q,

_' . ,' I , ' i ,

, , , q),h.; ,' f,fjj ]l , , ,

, g'

- ~y. - gn . p. ..y,g , m,.,

. .. s p ,. E o f t- ' ' ' .;;g, i 4

9-74

, ~:'{f L :1- ,

,e*

an Nuclear Services Corporation y a

E .;QQ c = 2 14 . 5 m u

6 E = 27 5 x 10 psi .

r. ,

which is a reasonable The ultimate stress was assumed to occur at .5c u 1 assumption for a medium carbon steel. The stress strain curve reflecting these properties is in Figure E-3 x E ,l l

ft , a i

( k Y'

t _!

'4'-

, J 1

• t; k

B ..;-

1

.g I

~

- }

1 l

1 i

i J

.T I

i O h,W a

i I

i i, . *

• j 4 '; l

. )

A i, 5

'l[ l

$t I

i

'l Q ', i 7 >

,osr 3'Yc;i- ,

13. .

a%

+

-k ,t .

- 103 - a-;

,Eff v.

g,, q@ f '3,' ,

y 4.

if M 4 s

• * ;) ., , ;<ry s j t- Qg ~ ,

%g n;. nmy g . .

y ,,,,,,g. ,

9 t

' ^ .,.m

x t

Nisclear Services Corporation

} d, 2

$o-

,' E.h Rock Shaft h)

- VJ h16 Martensitic Stainless Steel gL, .

7 Room temperature properties per Ref 3rence 2 from lot test data vere:

M e 1

j;

.~ J 9

oy = 108'5 - 124. kai

. al y f.- +M 0 = 127.- 1h8. kai g u

D.. ,j 1! cu = 15 0 - 16.7%  %

h t.

w pp I

Figure 3 03122 in Reference 10 presents data for h10 which is similar to, .,,

hf I h16. From this figure at Sh6 F: m;,+

.N.M 1, p _,'

,a

.wr, W.

I' '

ay = 85. ksi .: MQ 5,16 ptr

..w I vhich was utilized-in the analysis. This information is reflected in.

I'$

$?;

wp np

?O n.,

i Fir.,*e E h.

.a

• -,?.x-g s,,  ;.

w N;&

'$c

/..

hy,I4 e y vm A

ff.fllld v,

,y v_ ,y

.,.l Y

. 1d.,

~ :v _ ,

'l.

s 9,,,

w' gg .y ,

,..} j

. [f .

'?.sM:

t

%kkn I

n I '5)f;p<I j:,

y. .

s;mn n, w Ih s

NNh D

6)

':. - lok - . y ir

.g <

+

a;

.v, a yy

$sm(:  :. a .

4. , W-?a w.-

.m

>,..,. ,A 6 -

.; n;kf g q y; ;9_.

, I, ,, i

'a a .' s

.) ? E,,8j < b':(l;gi ;_

j pw - -.

} } .t y

.w..

Nuclear Services Corporation E.5 -Valve Body seat I

+

g ,

[E Weld 316 Austenitic Stainless Steel Per Reference 12 the velded material can be treated an annealed. Typical room teunperature properties for 316 were obtained from Reference 9 and noted below:

\

'E oy = 35. kai ,' I

ou =. 80. kai i

4

!c , cu = 65.5

't i  :

Also from Reference 9 typical tensile properties at operating temperature '

are obtained for.316. ~

I

' g c y = - 25 0 - Sh6.- SOG

25. - 23 SP i 200 1 0

7

=

2h.kkai ou Sh6.- 50 r

= 73.0 200 {T3.-725 e = 72.8 kai u

h9. - 5h6.- 500 - '

c =

h9. - k7.

u 200 cu= h8.5% s-and A'

E =. 2h.5 x 106per Figure 17 4

y s

l - 105 -

c' '

.. e ~

-){

,,,,  ? :~ 4, ,,,

m<_

Qs '-p;. & ,r .h& L

. y l2 , ,,  ;< f ':,.Q , ;g,, . y% ~ ,;;.g n . . . . . . . .

. . . . . . . . , . , c s, - ,

y yccy y ~ , nny mg m 7.mxmucw ~.w rP g;

,  ? F;'y 3 n

-e .g' r-j

.- Nuclear Services Corporation fm k.

n 3<

. ./

J The actual properties utilized in the analyses were: .

j A

1*- o Y

= 19. kai .;

b.

V

au = 61. kai '- J9 w

c = 40 % C u #

3 = 2h.5 x 10 psi 6

E O J_ i

[' '

4e 1 The ultimate stress was assumed to occur at c-u .15 c which is typical" 3,'

u -

,$i2 for stainless steels.., The stress strain curve reflecting these properties is in Figure E-5 > @>

w

,h

,, ey M

ww w

kf. h ,-

tra 4 )

1, W

p g&

r n

kv'i hkj I

n t

Mi I + sc_'

<r

-> i ;

1!<y l

W, p:

b 1j (M

.-I -

Vk

$k 2?

' yy e 9:-

p . .,

v.y s;a 1 'J. p l

E ,

.h i

I g $$h k.ig

~

[

g

~

- 106 - b[d.

(

q l,f?

,_'i'. '(j f f'*

-a yl

, c; , Mi,Ir!

n kg

- i, ,e 0J:1.: -

. :49l ., c . .< 7, : , ,S

1. ]i IN ' " %! i+Ey 93y%$ip? fcw mew;m g ,w ou.. .. .%

,.. an, .

m y ,y s., - <

NAr Seradees Corpensden

R~

• i cdarsar4, cAuroaNiA

' j gr.

= Ts , :M 1

i r At Room Temp ,

(Ref. 2) ,

\

5 e = 79 kai b*

r ou= 103 kai -

g t 80 - .v

, e.

o c = 22-2k% -

l Q " "

h e 60 < -

Linear Approximation ,<

! $ LO h 6

+> bE = 2h.5 x 10 psi

- us ,

,f 20 , .g

, 4.'

A a

n.

A.  ;

15 0 20.0 N' f 5.0 10.0 6 '

Strain 5 ,

~

e Stress Stre.ir. Curve for L10-Cb (Dise) at Sh6'F from References 3, 9, & 10 *

\

Figure E-1 l -

• V, l

I Ji .=

_O {

l h

';C

~

At Room Temp (Ref. 11 and 12I.

5 o- = h5-55 kei

80 .. Y

=

" c = 89-95 ksi;-

u .

G 60 .-

f

' c"= 30 h5%

i.

! ko '

h ,

E = 2h.5 x 10 psi '

20 b:I -

20.0 u

5.0 10.0 15.0

' Strain - 5 .

Stress St' rain Curve for E-309->15 (Dise Seat) at SL6'F, frem References 9 &s 10h t'.

p ,

Figure E-2 "?1 s 4ji

%0 E -

h-t g

a,..

,[

Ufy] .x

,a j ,, < . l , Kjl: jfk

un <, . . >-g: e,g . . . y p ; <s

, , g7,ypp y .,,. - ? . .,

.4

~ . . , rp .

i s

. A.s yy . ,

,, , f.,.__

• y y n -f:g-' . y t .O.

(

9 Nanclear Services Corpbration g y,

, i CAastsath CAuronNIA '.* Q '

.f,.-e ~ ;; F <

, t l, Ji:e At Room Temp

,. 4 (Ref. 2), [d .

s, m.

! 80 - -

k oy' = kl.5 h6.0;n;,ai ,:

60

-- ou= 73.5-77 5lk'si

, y-c m

to ..

-Linear Approximation u =26.5-30.0{.

3 w ,

1- >

l 20 '

N E = 27.5 x 10 psi m

g 'I O

e #

1 e a a a 3

g.

u) 20.0 5.0 10.0 15 0 v

?5 Strain - % A '

Stress Strain Curve for A216-WCB (Valve Body & Tail Link) at Sh6'F from References 9 & 13 .s', m

. Figure E-3 s?

ao

_t.

1 yl y'

I At Room Temp

/- (Ref. 2) .

- /

g, 80 --

/ o y

= 108.5-12L:;ksi

.I W

60 "

o"= 127-lh8 ksi.. "

=

m c"= 15.0-16.75 :

I. y vs LO "" E = 2h.5 x 106 psi .

20

<c

:  : +

O.5 1.0 15 2.0 +

Strain - % NY d1 Stress Strain Curve for h16 (Rock Shaft) at Sh6'F from References 9 & 10 i .,J.~ e, Figure E h y u a, ; . 7

-t

',3 . , . ,;,

~.

<<<.f.c- . :. . . , j. pm ,

. . .f .

.; 3  ;- ~

?e y * ,,_3 7

.0 gb.n.

i(

Nuclear Services Corporation ,

CARE 9Nt2. calif 0tNIA

'",  ;[ ize:/% a:

l.  ;;; Q[a 1 p, a M

,y t,' e

?

sa "

, n-g -s

. 1f

• ; 7 n

T.

6 a y 4.,

.-y 4v p ..

1.

I 4- .

At Room Tmp - R"

. O.

(Ref. 9 & 12), #

il

[.

v I ey= 3h h2 kai f

9 '

b

.c eu= 81 ksi " '

9 a

e cu= 50-55% - '

! s..

a j $'gr

$ .; $'I I -

M

!!, 80 A y!_

3,l.<

.- 60 '

m

= ,T e

a 2 ho -- .<!

f l I ca 20 <

NE = 2h.5'x 10 pai 1

3. i r't e

-/ E ,I i6 I 5 0-10.0 15 0 20.0 25.0

~30.0

> c V

? 10 a g:

Z I Strain - % . fE gs

^

Stress Strain Curve for 316 (valve Body seat) at Sh6'F from Reference 9 k

b.,

l i

g ?f ., f Firure E-5 . r.

h,T 7  %

• + y i $

I

.,s e:.,

( m-

• M3 f' f$

3 .

m I- 'j Y

?

s >.,

}i ?h.i, L

u  :::,:.

, f l Q$t '

{ r$

k. I

?

{T j #-

",. $Y:

-" 29 9.L k

4 h$

i

., c

- 109 -  : c^ " ,.

w f&

. w,'

f .

y- -

C;f

,+ , , . -[

p'ht .,

%N .

lt- _i$; s . . . ,

, ,y , , pp[ , , ;f'

_.