ML20235M880

Residual Stress Analysis of Thermal Sleeve to Safe-End Welds

ML20235M880
Person / Time
Site:	Brunswick
Issue date:	04/18/1979
From:	Edwards N, Giannuzzi A, Riccardella P NUTECH, INC.
To:
Shared Package
ML20235M689	List: ML20235M686 ML20235M880
References
98.701.0006, XCP-01-003, XCP-1-3, NUDOCS 8902280474
Download: ML20235M880 (49)

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XCP-01-003" 98.701.0006

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f 4~ RESIDUAL STRESS ANALYSIS OF

, THERMAL SLEEVE TO SAFE-END WELDS BRUNSWICK STEAM ELECTRIC PLANT UNITS 1 AND 2

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Prepared for:

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- Carolina Power.and Light Co.

Prepared by:-

!" NUTECH.

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, Pre ed by Issued by:

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i. C.-Ri dardella, P.E. N.W. Edwards,P.Ef ,

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f' Reviewed by:-

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Date: April 18, 1979 j, f_ ~w@

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.A. 8 ' Giannuzzi,.'P.d.F e,-

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ABSTRACT An axisymmetric thermo-elastic-plastic finite element model was

- used to estimate through thickness residual stresses in the vicin-ity of the thermal sleeve to safe-end attachment weld for the

~~

Brunswick recirculation inlet nozzles. The resulting stresses were-then compared to previously calculated stresses for a similar )

thermal sleeve / safe-end design at the Duane Arnold Energy Center, in which severe cracking was observed. The results of this ana-

• -- lysis and comparison lead to the following observations:

1. - The peak axial, hoop and radial residual stresses in the nozzle safe-end,- near the tip of the crevice produced by the subject welds are similar to those reported for the Duane Arnold design.

2. The axial through-wall stress gradients for the Brunswick safe-end, in the vicinity of the crevice, attenuate more rapidly than for the Duane Arnold design.

3. On the basis of this more rapid attenuation, the growth .

rates, for a hypothetical crack originating at the tip of the crevice, and propagating through the safe-end, would probably be slower than was the case at Duane Arnold.

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" TABLE-OF CONTENTS l in i e1 PAGE~  !

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1.0 ............1 ~

INTRODUCTION . ...

2.0- COMPONENT DESCRIPTION . . . . .. . .. . . 2-1 l

. 3.0 ANALYTICAL PROCEDURE' . . . . . - . . . . . . 3-1

3.1' -THERMAL ANALYSIS PROCEDURE . . . . . .- 3-

~2 3.2 STRESS ANALYSIS-PROCEDURE . . - . . . .- 3-4 I i . .

'- . 4.0. THERMAL-ANALYSIS . . . ... . . . . . . . 4-1 r

4.1 INPUT . . . . . . . . . ~ . . . . . ' . .. . 4-1 q w '

4.2 RESULTS ....- . . . . . . . . . . 4 r l- .

5. 0 - . STRESS' ANALYSIS . . . . . ... . .... . . 5-1

! 5.1' INPUT. . . . . . . . . . . . . . . . . 5-1 5.2 RESUtTS .. . . . . . . . . . . . . .. 5-2 li 6'. 0 DISCUSSION OF RESULTS . . . . . . . . . . . . '6-1

7.0 CONCLUSION

S AND RECOMMENDATIONS . .. .. . . 7-1

8.0 REFERENCES

. . . . . .- . . . . . . . . . . 8-1

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XCP-01-003 LIST OF FIGURES PAGE 2-1 BRUNSWICK RECIRCULATION INLET THERMAL SLEEVE TO . 2-2 SAFE-END WELD CONFIGURATION . . . . . . . . . . . . .

3-1 MOVING HEAT SOURCE IN AN INFINITE SOLID 3-7 (REFERENCE 1) . . . . . . . . . . . . . . . . . . . .

4-1 SIMPLIFYING ASSUMPTIONS USED IN THERMAL ANALYSIS 4-5 MODEL . . . . . . . . . . . . . . . . . . . . . . . .

4-2 TEMPERATURE DISTRIBUTION FOR. WELD LAYER ONE AT . . . . . . 4-6

. TIME WHEN WELD METAL HAS COOLED TO 2100*F .

4-3 TEMPERATURE DISTRIBUTION FOR WELD LAYER ONE AT . . . . . . 4-7 TIME WHEN WELD METAL HAS COOLED TO 1100'F .

4-4 TEMPERATURE DISTRIBUTION FOR WELD LAYER FIVE.AT . . . . . 4-8 TIME WHEN WELD METAL HAS COOLED TO 2100*F .

4-5 TEMPERATURE DISTRIBUTION FOR WELD LAYER FIVE.AT . . . . . 4-9 TIME WHEN WELD METAL HAS COOLED TO 1100'F .

_ 5-1 FINITE ELEMENT MODEL USED IN RESIDUAL STRESS ANALYSIS OF BRUNSWICK RECIRCULATION INLET THERMAL . 5-3 SLEEVE TO SAFE-END WELD . . . . . . . . . . . . . . .

~

5-2 TEMPERATURE DEPENDENT PROPERTIES OF INCONEL 600

. . . . . . . . 5-4 BASE AND WELD MATERIALS (REFERENCE 1) 5-3 CONTOURS OF AXIAL RESIDUAL STRESS. DUE TO THERMAL

. . . . . . . . . . 5-5 SLEEVE TO SAFE-END WELD (KSI) .

a 5-4 CONTOURS OF HOOP RESIDUAL STRESS DUE TO THERMAL

. . . . . . . . . . . . 5-6 SLEEVE TO SAFE-END WELD (KSI) 5-5 CONTOURS OF RADIAL RESIDUAL STRESS DUE TO THERMAL

. . . . . . . . . . . 5-7 SLEEVE TO SAFE-END WELD (KSI) .

5-6 CONTOURS OF MAXIMUM IN-PLANE PRINCIPAL STRESS DUE

. . . . . . 5-8 TO THERMAL SLEEVE TO SAFE-END WELD (KSI).

5-7 CONTOURS OF MINIMUM IN-PLANE PRINCIPAL STRESS DUE

. . . . ,. . 5-9 TO THERMAL SLEEVE TO SAFE-END WELD (KSI)

5-8 DEFORMED SHAPE PLOT FOR THERMAL SLEEVE TO SAFE-END

. . . . . . . . 5-10 WELD (DISPLACEMENT SCALE FACTOR = 10)

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. LIST'0F. FIGURES:

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.. s 6-1; COMPARISON.OF THROUGH-THICKNESS IXIAL RESIDUAL STRESS' PROFILES"AT RECIRCULATION INLET THERMAL SLEEVE'TO SAFE-END WELD LOCATION-(BRUNSWICK o'

/ VERSUS DAEC) . . . . . .. . . . . . . . . . . . .. . .:-6-4 J 'l i

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XCP-01-dO3 LIST OF TABLES PAGE 1-TABLE 4 Welding Parameters Provided for i Thermal Sleeveito Safe-End Attachment' Weld (Reference 4) . . . 4-2 TABLE 4 Weld Thermal Parameters Assumed in Analysis . . . . . . . . . . . . 4-2 t e y U-e6 9

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1.0 INTRODUCTION

~The recent occurrence of intergranular stress corrosion cracking i

e in the recirculation inlet no::le thermal sleeve to safe-end 1

1 . attachment welds at the Duane Arnold Energy Center (DAEC) have f' raised concern regarding the possibility of similar cracking in-

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cidents in other Boiling Water Reactors (BWRs). The recircula-tion thermal sleeve attachment welds at the Carolina Power and 7- Light Brunswick Units'1 and 2 are fabricated from the same ma-1

'- terial (Inconel 600) and have a crevice geometry which is quite similar to that at DAEC. Although the recirculation safe-ends

{ are considerably thicker for the Brunswick Units compared to

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. DAEC, the similarity in materials of construction and in general

- crevice configuration at the thermal sleeve attachment weld have prompted the Nuclear Regulatory Commission to request that CPGL perform augmented inservice inspection to insure the in-tegrity of the safe-end.

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As part of the overall CPSL program designed to insure continued L safe operation at the Brunswick Units, NUTECH was contracted to perform a state-of-the-art finite element analysis of the state of weld-induced residual stresses existing at the crevice in the thetmal sleeve to safe-end attachment welds. The analysis was performed using a therme-elastic-plastic finite element model s-similar to that which was utilized by Battelle Columbus Labora-7" tories for the DAEC safe-ends (Reference 1), and using residual

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stress analysis methodology which has evolved over the last i

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^ XCP-01-003 several years (Reference 2) . The detailed analytical procedures used, and the results of the analysis are presented in this report.

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XCP-01-003 3.0 ANALYTICAL PROCEDURE As described in the prior residual stress analysis of the DAEC safe-ends, (Reference 1), the mathematical,model for predicting The first

~, residual stresses due to welding consists of two parts.

r part is a heat flow model that gives time dependent temperature

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distributions for the body being welded. The temperature model a incorporates the important aspects of the welding procedure such

- as arc velocity, power (heat input) and geometric location of the i

• arc. The output of this first portion of the model is the tem-perature distribution throughout the welded body as a function of time. The resulting temperature distributions are used as input to the second portion of the model, which consists of an

- elastic-plastic, finite element, thermal stress model. This

- stress model takes the previously computed temperature distribu-tions and determines the induced deformations, stresses, and strains which result throughout the body. By applying the tem-perature distributions to the stress model in chronological order, the nonlinear effects due to history dependent plasticity are i

- included in the analysis; the result being that deformations, stresses, and strains can be determined at any point during the fabrication procedure. It is the stresses which remain after

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all welding is completed and the body has reached a uniform am-bient temperature that are of prime interest, since these are

- the final welding induced residual stresses.

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The following paragraphs describe in some detail the heat flow and thermal stress procedures used in the analysis. Included in F 3-1 nu ech I

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7 XCP-01-003 this description, are the simplifications and assumptions which are used at various points in the analysis procedure as applied

-. to welding residual stress problems in general. Later sections F will describe the special modifications and assumptions which l~

were required for the analysis of the thermal sleeve to. safe-end I

t attachment weld,

. 3.1 THERMAL ANALYSIS PROCEDURE The thermal analysis model is based on the closed form so-lution for a point heat source moving at a constant velocity through an infinite body as developed by Rosenthal in Reference 5. The adaptation to thermal analysis of a weld-l.-

1; ing process is identical to that used in Reference 1, and the following description is taken directly from that report. The temperatures due to the moving heat source are given by:

T(r,C) = T o +4[Kr exp [- (E+r)vc/2 k] (1)

L .where: T = temperature (F)

T o = initial temperature (F) q = heat input (Btu /sec) k = thermal conductivity (Btu /in-sec-F)

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v = heat source velocity (in/sec)

C = distance from heat source to plane containing point of interest (this plane is normal to the

[_; path) (in) a r = distance from heat source to point of interest (in) 3

. c = heat capacity (Btu /in ,p)

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'The coordinates referred to in Equation.(1) and the assumedL velocity in.the'x coordinate direction are illustrated in Figure 3-1. The1 steady distribution', described by Equa-tion (1), moves with the velocity of.the heat source. 'For thisreason,rtimeappearsintheformulaonlyimplicith

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through the! equation E = C o - V t; where t is the time and

(, is the.value of C at t = 0.

The thermal solution represented by . Equation (1) does not

,  ; include spatial variations of the material thermal pro-perties or1 temperature dependence of these properties.

Therefore, it is assumed that temperature dependent changes in the thermal properties do not affect temperature dis-tribution at times :of interest for residual stress calcula-E

- tions. It is also assumed that any geometric discontinu- <

ities in the thermal properties due to different materials  ;

or the previous deposit of weld. metal does not signifi-cantly affect the final calculated residual stresses.

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Furthermore, due to the steady-state nature of Equation (1-),

the heat flow model cannot be used to simulate the begin-ning and ending points of a weld pass.

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5 Since the thickness of the welds being analyzed is not

,q large enough to be approximated by infinity, some modifi-cation to Equation (1) is required. This modification is l

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- " essentially no more than the superimposing of a numbe 7f the solutions represented by Equation (1) such that two l

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XCP-01-003 insulated-surfaces perpendicular to the y-coordinato' h '

j .,. direction (see Figure 3-1) are cree.ted. The positions of.

these two' surfaces 1are made to correspond to the actual 1 d

< inside and outside' surfaces of the pipe. ]

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One last' aspect of the therma 10model which must be consi- q 4 dered before it can be used-forepiping thermal analyses is f the fact that the solutian represented by the repeated a

superposi~cion of Equation (1) is that for a flat plate.

By applying this model to a cylindrical geometry (pipe),

oneLis essentially assuming that the heat flow in a curved-

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plate is not'different for that in a flat plate. This is a reasonabic approximation as long as the inner circumfer-

, ence of the pipe does not differ substantially from the 2 outer circumference. This condition is more closely sa-i~' tisfied as the diameter to thickness' ratio -(D/t) of .the pipe increases._ For less than a110, percent diffetance in T outer to-inner circumference, the D/t must exceed 20.

The' error in temperature due to this effect will again L_. depend'on the actual D/t of the pipe,_the welding velocity, e and.the power.

3 '. 2 STRESS ANALYSI

S. PROCEDURE

u The finite element-stress analysis program utilized for the y Reference 1 analysis was not available for the present

analysis. Therefore a similar elastic-plastic finite. ele-1

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ment computer program was set up and modified at NUTECH 3-4 nutech i-------__-_-______ _ - _ _ _ _ . _

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to accomplish the present residual stress analysis. The

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finite element program used is based upon the incremental tangent modulus approach to static, elastic-plastic analysis' as described in Reference 6.

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The program utilizes trian-gular. constant-strain elements and quadrilateral elements composed of a combination of constant strain triangles.

Special modifications to the program'were necessary to per-

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form a multipass weld residual stress analysis. The most significant of these modifications are highlighted below: ,

The program was modified to include fully temperature dependent material properties, including Elastic Modulus, Coefficient of Thermal Expansion, Poisson's Ratio, Yield Strength and Plastic Modulus (assuming a bi-linear stress-strain law).

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- The conventional isotropic strain hardening assumption utilized in the finite element program was augmented l

~- to incorporate an " isotropic softening" feature as well.

That is, the yield surface for any given element is assumed to expand uniformly'as the yield strength is exceeded, but then to contract in the same manner as the stress in that same element is reduced below yield.

- A capability to input two stress-free reference tem-peratures was added to the program. This permits the l

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solid bcse mstal-to expand or contract relative to f,

.one temperature, while the' molten weld metal contracts

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relative 1to a second, much higher. temperature.

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- Thermal ~: expansion and contraction was integrated in 1 .

y o crementally utilizing the instantaneous coefficient b- L..

of thermal' expansion-at,the appropriate temperature, mg.

Verification of the finite element computer. program used

'in the. Reference.1 analysis'has been documented through numerous. comparisons to experimental residual stress measurements on butt-welded pipe (Re f' e' rence . 2) . Since these same verification problems were not run with the t

program 1 described above, as. modified for the present ana--

. lysis', tifeiresults presented'cannot be' considered to have l

T the same level of design veri'fication as those presented--

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?in; Reference'I. - However, review of the - results presented 1 .. J

.in'Section 5kofLthisgreport indicates a highly consistent

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comparison;to.the? Reference l results, considering the; L ,

similarities and 'd'ifferences in the' geometries analyzed. .

b p On the basis4 of this comparison, Lit is 'strongly suspected - .

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thatifthepresenthprogramtwere:appliedto'theDAECL safe-end, the resultings residual stresses wouldfduplicate ,

'those reported in Reference 1. Thussthe experimental verification of the Reference 1 analytical approach is considered applicable to.the present analysis, s .

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FIGURE 3-1 MOVING HEAT SOURCE IN AN INFINITE SOLID (REFERENCE 1) .

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4.0 THERMAL ANALYSIS i.-

I 4.1- INPUT .i

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" -Welding' parameters were'obtained from.a welding procedure I b'  !

' specification for.the subject welds provided by CPGL (Ref-n- .

[ erence.4). The specific welding parameters provided are la t

summarized in Table:4-1, and a sketch of the attachment- i g

geometry . and weld ~ detail is shown in Figure 4-1 (a)~. As f.

a simplifying assumption in the analysis, the weld was j I

k assumed to be deposited in five. layers as illustrated in Figure'4-1 (c). Heat. input for each weld pass was.com-puted using the average values of the specified welding parameters, and, since travel speed was not provided,

,' travel speeds consistent with reasonable weld-deposition 1-rates, for the processes used were assumed. Table 4-2 I summarizes the actual weld thermal parameters used in the analysis..

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- Note that, in modelling each weld layer, which consists of

+- a number of weld passes, point heat sources were placed j at four locations within each weld layer,.and an upper bound temperature distribution enveloping all of the in-r 1

~dividual temperature distributions was used in the stress-3 analysis of the layer.

F Two additional thermal modelling assumptions used in the i analysis are the assumption of a constant thickness pipe j i;. l

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"N Welding Parameters Provided for Thermal Sleeve to Safe-End Attachment Weld (Reference 4)

I Process Pass Amps Volts ' Speed n

L GTA 1-5 140-170 12-15 ---

SMA. 6- 65-95 --- ---

GTA 140-170 12-15 ---

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'- Interpass Temperature = 300*F No Preheat TABLE ~4-2 Weld Thermal Parameters Assumed in Analysis '

Travel Speed Heat Input

• Heat Rate, Initial Layer Amps Volts' (in/ min) (KJ/in) Qo(Btu /sec) Temp (*F1 t -

1 155 13.5 4.5 27.9 1.49 70 2 80 25 3.5 '34.3 1.42 300

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34.3 1.42 300 3.5 3 80 25 e

4 80 25 3.5 34.3 1.42 300 5 155 13.5 4.5 27.9 1.49 300

• Includes'75% Efficiency Factor Conductivity = 3.086xiO'4 Btu /in-sec *F -

3 Heat Capacity = 0.0380 Btu /in *F k.

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• ~ XCP-01-003 model equal to the mtximum composite safo-end/ thermal sleeve / weld thickness - (Figure 4-1 (b)) and the use of an amplification factor on the basic heat rate, Qo, to ac-E count for the lack of material in the portion of the weld I

cavity which is not yet filled (Figure 4-1 (c)). Using the uniform thickness assumption, a simple coordinate map-ping routine was established. Each point in the actual body was associated with a corresponding point in an e

idealized uniform cylindrical pipe with thickness equal to the maximum thickness at the weld location. Axial coordinates were mapped directly without significant dis-tortion, while radial coordinates were normalized to the ideal geometry based upon the thickness of the actual nozzle connection at a particular axial location. Ampli-fication factors to account for lack of material were developed by -assuming no heat flow to an angular se3-ment of the region surrounding the point. heat source.

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The heat which would have flowed to this segment was assumed to be added to the base heat rate in the bal-ance of the region. The size of the angular segment de-creases-as more of the weld is deposited, resulting in a gradual reduction.of the amp 1'ification factor from 1.125 for layer one to 1.0 for layer five (Figure'4-1

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(c)).

4.2 RESULTS The resulting temperature distributions are illustrated

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in Figures 4-2 through 4-5. Figures 4-2 and 4-3 illustrate 4-3 .

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, temperature distributions corresponding to weld layer one p

when the ' eld metal is at 2100*F and 1100*F, respectively.

Figures 4-4 and 4-5 show the same two 'seld temperature points for. weld layer five. '

. Note that the outer surface of the safe-end sees very little change in temperature during the welding process.

This factor, plus the circular nature of the temperature distribution indicates that the boundaries of-the model

s. . had very little effect on the solution, and thus justi-t fies the simplifying assumptions used in the analysis.

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, Distributions similar to these for each layer, with a re-turn to the isothermal 300*F interpass temperature follow-I t

., ing the completion of each weld layer were used as temper-ature input to the stress analysis finite element model.

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INNER FIVE i.

PASSES - GTA FINAL h

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LAYER - GTA L INTERMEDIATE PASSES - SMA

(a) ACTUAL GEOMETRY 6 WELD DETAIL s:

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(b) COMPOSITE THICKNESS REPRESENTATION I

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( WELD LAYERS AND HEAT AMPLIFICATION g

FACTORS ASSUMED IN ANALYSIS 2" Q = 1.125 Qo

.; j Q = 1.094 Qo

, -- Q = 1. 0 6 3 Qo f . g 0 = 1.031 0,o P i s s O = 1.0 On p

(c) CONSTANT THICKNESS PIPE MODEL FIGURE 4-1 SIMPLIFYING ASSUMPTIONS USED IN

, THERMAL ANALYSIS MODEL .

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[* 5.0 STRESS ANALYSIS a.. ,

p 5.1 INPUT j

%. The finite element model used for the . stress analysis is

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p- illustrated in. Figure 5-1. The model contains.a total' 1

. of 624 nodes and 559 elements. Boundary conditions im-posed on the finite element model included axial fixity c.

[ -at the interface between the reactor vessel and the safe-end. In addition, the thermal sleeve end cf the model

. {. .r b -was assumed to carry a net zero axial-force. All'other f points-in the model were allowed ho deflect as required  ;

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by the thermal loading.

i Temperature dependent mechanical properties for Inconel I 600 base and weld metal used in the stress analysis were 1 obtained from Reference 1, and are presented in Figure 5-2. l

,1. l Ambient temperature was assumed to be 70*F, with the inter- .

.. l pass temperature given as 300*F. '

{ _

7 h As stated earlier, each weld was considered to be placed 4 in five layers. For each layer, the temperature distribu-I tions used to represent the transient thermal history 3

.were approximated by three piece-wise linear segments.

's.

The first segment takes the model from uniform initial r

LL temperature to a time when the n'ewly deposited weld ma-

.7 terial reached a temperature of'2100 'F. The second.

l

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segment spans the times when"the newly deposited weld L-

[

5-1 nutech e ,

XCP-01-003

[ material cools from 2100*F to 1100*F, and the final seg-ment addresses cooling from 1100*F to the uniform inter-pass temperature (or ambient for the final layer). For each layer, the load segment taking the model from a uni-form ambient or interpass temperature to the temperature distribution at 2100*F was divided into nine smaller.in-crements. Four increments were used in going from 2100*F j to 1100*F, and five increments were used to return the

.. body to a uniform ambient or interpass temperature. All t-I told, ninety incremental solutions were used to describe the placement of the weld.

5.2 RESULTS The stresses in the model at the end of load increment

\

90 represent the calculated residual stresses at the com-

. pletion of the thermal sleeve to safe-end weld. Stress

I contour plots describing the axial, hoop, radial and j maximum and minimum principal residual stresses are pre-sented in Figures 5-3 through 5-7, respectively.

lp Figure 5-8 presents a deformed shape plot for the weld J.

l'- geometry following completion of the welding. The dashed'

[

t.

lines represent the original, undeformed shaue, while the solid grid represents the deformed geometry, after i

welding, only with the displacements amplified by a factor of ten for clarity. The shrinkage of the weld material is clearly shown in this figure.

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6.0 DISCUSSION OF RESULTS w

The general nature of the residual stress patterns presented in Figures 5-3 through 5-7 are similar to those reported for the l

DAEC safe-end in Reference 1, and many of the general observa-l tions noted in that report are repeated here for completeness.

Areas of differences in the two sets of results and the signifi-1 cance of those differences are then discussed.

As in the DAEC safe-end, the axial residual stress pattern (Figure 5-3) illustrates the existence of a significant residual flexure (and bending moment) in both the safe-end and the sleeve.

i The sense of this flexure is the same in both portions of the l

assembly, leading to a discontinuity in stress at their junc-ture. It can be seen that this flexure decays repidly as one F moves away from the weld region. It is also interesting to J

note that the majority of the weld' region itself remains in a This is not surprising since I state of residual axial tension.

a basic feature of welding is the contraction of the weld metal L upon cooling. The weld region is also subjected to high tensile

' residual hoop stresses (Figure 5-4). Again, this is as might a

be expected based on the contraction of the weld metal. A cor-relation between these hoop stress contours and the temperature contours of Figure 4-5 can be seen. For instance, the zero hoop stress contour is very similar to the 320*F contour. This cor-relation implies that the regions of the safe-end and sleeve I-which exceeded 320*F, are trying to. contract radially in a man-I ner similar to the weld metal and are being kept from contracting t.

6-1 Ilutech

^- XCP 0 03

.to some extent'b'y the surrounding material which is, as a result, in it state of! hoop compression. In Figure 5-5, the radial stress contours are' plotted. It'is interesting to note in this figure, the region-of radial tension extending axially from the' crevice

.tip along the base of the fused zone.. This tensile zone indicates I that the safe-'end is resisting a tendency for the sleeve to pull

-away'and deflect radially inward. The contours of maximum in-7; plane- (r-z) tensile. and compressive stresses (principal stresses)

are plotted'in Figures 5-6 and 5-7, respectively. Combining

'" these figures.with that for the hoop stress (the third principal r stress)-it can be seen that a region of triaxial tension exists p-j ust ' above ~ the crevice ' tip. Similarly, a' region'of triaxial com-3.

j pressive... stresses exists in the area immediately below the. crevice tip.

1

'Despite these.' similarities, a potentia 11y'significant difference exists between'the present Brunswick residual stress results and b the Reference l'DAECL results in theLrate of attenuation of the g

residual stresses through the wall thickness. If, for example, h one plots the axial: residual stress through a plane originating

[ at the crevice tip and extending radially through the safe-end as shown in Figure 5-3, and. compares the through thickness result fy to that for DAEC, the'results are as presented in Figure 6-1.

g The peak axial stress level at the tip of the crevice is ex-b sentia11y the same for the two designs, but the Brunswick curve drops off more rapidly through'the thickness. The hoop resi-dual stress compares quite favorable between the two designs ja' along the same radi'ai through-thickness plane.

l.

l

.; l' w-.. , /XCP-01-003

'n l

- e.. o, This result _ suggests thatl: veld residual stresses play approximately an' equal-role in contributing to potential crack initiation in

~

both the DAEC and-the Brunswick designs, but that the role played -

by residual stress in crack propagation may be somewhat reduced

~

in the Brunswick design as compared to DAEC.

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RESIDUAL STRESS (KSI)

FIGURE 6-1

- COMPARISON OF THROUGH-THICKNESS AXIAL RESIDUAL STRES

'i AT RECIRCULATION INLET THERMAL SLEEVE TO

- SAFE-END WELD LOCATION (ERUNSWICK VERSUS DAEC)

'~ 6-4 nuteCh

XCP-01-003

7.0 CONCLUSION

S AND RECOMMENDATIONS q

The results of this finite element residual stress evaluation indicate that the axial, hoop and radial residual stresses at the tip.of the crevice in the Brunswick design are similar to those levels which were present in the DAEC safe-ends. The residual stress through-thickness distribution, however, is

- somewhat lower in the Brunswick design versus the DAEC design thus providing a reduced driving force for an initiated crack.

. 1

-~

In order to estimate the crack propagation rate for a hypothetical crack, one must first determine the total through-thickness stress distribution (including contributions from pressure,- temperature, external loads, etc.) and apply suitable crack propagation laws to the resulting stress intensity factors. Such an analysis was not performed as part of this study, but is highly recommended to obtain a more c entitative assessment of the significance of the residual strsss results reported herein.

h er d

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e W

___m.__ - _ _ __ _ _ . - _ _ _ _ _ _ _ __ _ _ _ _ _ . _ __

l, 1. ; ,

, .XCP-01-003 y

L* -8.0' REFERENCES

/f f

j g.

1.'R. B. Stonesifer~and J. M. McConaghy, Residual Stress Analysis t

9 of Sleeve to Safe-End ' Welds , Battelle Columb'us Laboratories i Report to NUTECH,_ December 126, 1978~.

2. E.-F. Rybicki, et al, Residual Stresses at Girth-Butt Welds u

-in Pipes and Pressure Vessels, Final. Report to U.S. Nuclear

+:

Regulatory' Commission Division of Reactor Safety Research under Contract No. AT(49-24)-0293, NUREG-0376, published

. November,'1977..  :

i

3. Reactor Vessel' Stress Report, 218" &' Boiling Water! Reactor, Chicago Bridge and-Iron Co., Contract No. 68-2471, August, V , 1968.

.c P

4. Recirculation Nozzle Thermal Sleeve to Nozzle-Butt Weld with'

-Integral Backing Strip,-Brown and Root, Inc., Weld Procedure

~

No. SP88-WTS No. 1, Revision 2, August, 1974. ,

l E< 5. D. Rosenthal, Mathematical Theory of Heat Distribution During

~ "

Welding and Cutting, Welding Journal'Research Supplement, May, 1941, pp. 220-234.

g

6. J. L. Swedlow, A Procedure for Solving Problems of Elasto__

Plastic Slow, Carnegie-Mellon University, Report SM-73, Septemost,- 1973.

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ATTACdMENT.4' 4 r:

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SSAFE"END CRACK-GROWTH = ANALYSIS '"

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AT THERMAL SLEEVE' ATTACHMENT WELD LOCATION . . ,

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• tm pc-CRACK (C) COPYRIGHT 1984, 1987 STRUCTURAL INTEGRITY ASSOCIATES, INC.

SAN JOSE, CA (408)978-8200 VERSION 1.2 STRESS CORROSION CRACK GROWTH ANALYSIS BSEP-1 TS - N0ZZLES A,B & H INITIAL CRACK SIZE = 0.6250 WALL THICKNESS: 1.1250 MAX CRACK SIZE FOR SCCG= 0.9000 STRESS CORROSION CRACK GROWTH LAW (S)

LAW ID C N Kthres K1C INCONEL 1.0780E-08 2.2600 0.0000 200.0000 STRESS COEFFICIENTS CASE ID C0 C1 C2 C3 PRES-AX 3.4500 0.0000 0.0000 0.0000 BEND 1.0000 0.0000 0.0000 0.0000 RESID-D 48.8303 87.0490 -513.3800 325.2888 RESID-B 57.7614 -217.5516 166.8299 -14.0500 RESID-W -30.7892 -16.4059 137.3668 -55.7917 Kmax CASE ID SCALE FACTOR PRES-AX 1.00 BEND 4.32 RESID-B 1.00 TIME PRINT TIME INCREMENT INCREMENT 87600.0 730.0 730.0-CRACK MODEL: ELLIPTICAL LONGITUDINAL CRACK IN CYLINDER (T/R=0.1,A/L=0.2)

CRACK ---------------STRESS INTENSITY FACTOR----------------

DEPTH CASE CASE CASE CASE CASE PRES-AX BEND RESID-D RESID-B RESID-W l 0.0180 0.713 0.207 10.287 11.411 -6.398 l 0.0360 1.013 0.294 14.867 15.515 -9.143 l 0.0540 1.251 0.362 18.542 18.255 -11.298 0.0720 1.453 0.421 21.733 20.236 -13.144 0.0900 1.635 0.474 24.589 21.703 -14.785 0.1080 1.802 0.522 27.181 22.788 -16.274 0.1260 1.959 0.568 29.544 23.569 -17.637 0.1440 2.106 0.611 31.701 24.104 -18.893 0.1620 2.248 0.651 33.667 24.435 -20.057 N - - - - - - - - * - - _ _ _ _ _ . _ _ _ _ _ _ _ . _ _ , _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ , _ _ _ __ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

(

Q^,

p. pc CRACKL VERSION;1.2 PAGE 2 y

L 0.1800' 2.383 0.691 35.452 24.594 -21.135-0.-1980 2.514- 0.729 37.061 24.606 -22.133-LO.2160. 2.642 0.766 38.498 24.490 -23.056 0.2340 2.767 0.802' 39.795 24.271 -23,923.

0.2520 2.892 0.'838- 40.954. 23.959 -24.737 0.2700 3.015 0.874 41.952 23.559 -25.484 0;2880- 3.136 0.909 42.789 23.079 -26.166 O.3060 .3.255 0.943 43.467 22:530 -26.782 0.3240J 3.372 0.977. 43.988 21.920 -27.334-0.3420 3'.489 1.011 44.352. 21.260 -27.822

0.3600 3.6041 1.045 44.560 :20.567 -28.249 0.3780 3.719 1.078 44.614 19.835 -28.612 0.3960 3.833 1.111' 44.517 19.071 -28.911 0.4140 '3.946 1.144 44.270 18.279 L-29.147 0.4320 4.058 1.176 43.877 '17.466 -29.319 0.4500 4.170- 1.209. 43.340- 16.636 -29.428 0.4680: 4.280 1.'241 42.670 15.769 -29.472' O.4860 4.390 1.273 41.865 14.893 -29.453 0.5040 4.500 1.304 40.930 -14.010 -29.372 0.5220 4.609 1.336 39.869 13.125 -29.229

'0.5400 .4.718 1.367 38.687 12.241 -29.023 0.5580 4.826 1.399 37.389 11.363 -28.755 0.5760 4.940 1.432 35.982 10.469 -28.441

'0.5940: 5.055: 1.465 34.460 9.574 -28.067

'0.6120 5.171 1.499 32.826 8.690 -27.626; 10.6300: 5.286 1.532 31.086 7.820 -27.118 0.6480 5.401 1.566 29.246 6.967 -26.544 0.6660 5.517. 1.599 27.312 6.135 -25.904

'0.6840- 5.632 1.632 25.296 5.307 -25.197 0.7020 5.747 1.666 23.209 4.484 -24.424 0.7200 5.862 ~ 1.699 21.051 3.689 -23.586 0.7380 5.977 1.732 18.832 2.924' -22.683

-0.7560 6'.092 1.766 16.560 2.193 -21.715 0.7740- 6.208 1.799. 14.244 1.498 -20.683 r

'0.7920 6.324 1.833 11.886 0.855 -19.586 0.8100 6.440 1.867 9.476 0.292 -18.423 0.8280- 6.557 1.901 7.048 -0.222 -17.195 10.8460 6.674 1.935 4.611 -0.684 -15.905 0.~8640 6.792 1.969 2.176 -1.092 -14.553

.0.8820 6.909 2.003 -0.245 -1.442 -13.139 0.9000. 7.027 2.037 -2.641 -1.733 -11.664 TIME KMAX DA/DT DA A A/THK 730.0 19.89 9.2849E-06 0.0068 0.6318 0.562 1460,.0- 19.67 9.0461E-06 0.0066 0.6384 0.567 2190.0 19.45. 8.8214E-06 0.0064 0.6448 0.573 2920.0 19.24 8.6052E-06 0.0063 0.6511 0.579 )

3650.0 19.03 8.4010E-06 0.0061 0.6572 0.584 I

4380.0 18.84 8.2078E-06 0.0060 0.6632 0.590

pc CRACK -VERSION.~1.2' PAGE 3 5110,0 18.65' 8.0215E-06 0.0059 0.6691 0.595 5840.0 18.46 7.8423E-06 0.0057 0.6748 0.600 P =6570.0L 18,28 7.6697E-06 0.0056 0.6804 0.605 L L7300.'O. 118.10 7.5031E-06 0.0055 0.6859 0.610 8030.0 ~17.93 7.3424E-06 0.0054 0.6912 0.614 l .8760:.0- ~1 7.76 7.1880E-06 :0.0052 .0.6965 0.619 9490.0- 17.60 7.0386E-06 0.0051 0.7016 0.624 10220.0. 17.44 6.8939E-06 0.0050 0.7067 0.628 10950.0 17.29 6.7602E-06 0.0049 0.7116 0.633 11680.0' ~ 17.14 6.6309E-06 0.0048 0.7164 0.637 12410.0 .17.00 6.5055E-06 0.0047 .0.7212 0.641 13140.0- 16.86 6.3855E 0.0047 0.7258 0.645-13870.0 16.73 6.2740E-06 0.0046- 0.7304 0.649

~14600.0 16.60 '6.1656E-06 0.0045' O.7349 0.653 15330 0 . 16.47 6.0600E-06 0.0044 0.7394 0.657 16060.0 16.35 5.9593E-06 0.0044 0.7437 0.661 16790.0 16.24 5.8658E-06 0.0043 0.7480'0.665 17520.0 16,12 5.7745E-06 0.0042 0.7522.0.669

.18250.0- 16.01 5.6855E-06 'O.0042 0.7564 0.672 18980.0 15.91 5.5992E-06 0.0041 0.7604 0.676 19710.0 15.81- 5.5210E-06 0.0040- 0.7645 0.680 20440.0 15.71 5.4444E-06 0.0040 '0.7684 0.683 21170.0 15.61 5.3696E 0.0039 0.7724 0.687 21900.0 '15.52 5.2963E-06. 0.0039 0.7762-0.690 .:

'22630.0 0.0038 15;43 5.2295E-06 '0.7800 0.693 23360.0 15.35 5.1677E-06 0.0038 0.7838 0.697

'24090.0 15.27 .5.1069E-06 0.0037 0.7875 0.700-24820.0l 15.19 5.0473E-06 0.0037 0.7912 0.703 25550.0 15.11' 4.9887E-06 0.0036 '0.7949 0.707 26280.0 15.05 4.9410E-06 0.0036 0.7985 0.710

-27010.0- 14.99 4.8965E-06 0.0036 0.8021 0.713

'27740.'O 14.93' 4.8526E-06 0.0035 0.8056 0.716 28470.0 14.87- 4.8094E-06 0.0035 0.8091 0.719 29200.0 14.81 4.7667E-06 0.0035 0.8126 0.722 29930.0 114.76- 4.7298E-06 0.0035 0.8160 0.725 30660.0 14.71 4.6951E-06 0.0034 0.8195 0.'728 31390.0 14.67 4.6607E-06 0.0034 0.8229 0.731 32120.0 14.62 4.6268E-06 0.0034 0.8262 0.734

-32850'.0, 14.57 4.5932E-06 0.0034 0.8296 0.737 9-33580.0 14.53 4.5632E-06 0.0033 0.8329 0.740 N<d 34310.0 14.49 4.5372E-06 0.0033 0.8362 0.743 35040.0' 14.46' 0.0033 0.8395 0.746 49 4.5113E-06 g% -

135770.0 14.42 4.4857E-06 0.0033 0.8428 0.749 36500.0 14.38 4.4603E-06 0.0033 0.8461 0.752 t.

37230.0~ 14.35 4.4353E-06 0.0032 0.8493 0.755

^ .37960.0 14.32 4.4172E-06 0.0032 0.8525 0.758 38690.0 14.30 4.3993E-06 0.0032 0.8557 0.761 (39420.0 14.27 4.3815E-06 0.0032 0.8589 0.764 40150.0 14.24 4.3638E-06 0.0032- 0.8621 0.766 J40880.0 14.22 4~.3462E-06 0.0032 0.8653 0.769 141610.0 14.20 4.3316E-06 0.0032 0.8685 0.772 42340.0 14.18 4.3212E-06 0.0032 0.8716 0.775 43070.0 14.17 4.3109E-06 0.0031 0 8748 0.778 43800.0 .14.15 4.3006E-06 0.0031 0.8779 0.780 i j

pc-CRACK VERSION 1.2 PAGE 4 0

44530.0 14.14 4.2903E-06 0.0031 0.8810 0.783 45260.0 14.12 4.2801E-06 0.0031 0.8842 0.786 45990.0 14.12 4.2748E-06 0.0031 0.8873 0.789 46720.0 14.11 4.2718E-06 0.0031 0.8904 0.791 47450.0 14.11 4.2689E-06 0.0031 0.8935 0.794  :

48180.0 14.10 4. 2659E- 06 0.0031 0.8966 0.797 48910.0 14.10 4.2629E-06 0.0031 0.8997 0.800 49640.0 14.09 4.2599E-06 0.0031 0.9028 0.803 CRACK DEPTH EXCEEDED 0.9000 AT TIME 4.9640E+04 END OF pc-CRACK l

1 s , . Pc-CRACK

'(C) - COPYRIGHT 1984, 1987 STRUCTURAL'-INTEGRITY ASSOCIATES, INC.

SAN JOSE, CA.(408)978-8200 VERSION-1.2 STRESS' CORROSION CRACK GROWTH ANALYSIS

'BSEP-1 TSL-#N0ZZLES C & E-

! INITIAL" CRACK' SIZE = 0.4050.

WALL THICKNESS - 1.1250-(MA%_ CRACK SIZE FOR SCCG= 0.9000

. STRESS CORROSION CRACK GROWTH LAW (S) 4

' LAW ID ' . C N Kthres K10 h .INCONEL 1.0780E '2.2600- 0.0000 200.0000 STRESS COEFFICIENTS CASE ID. CO' 101. 102 C3

' PRES-AX' , :3.4500: -0.0000- 0.0000 0.0000 BEND 1.0000 0.0000' O.0000. 0.0000

.RESID-D: '48.8303. 87.04901 -513.3800 325.2888 RESID-B '57.7614. -217.5516 166.8299- -14.0500

.RESID-W- -30.7892' -16.4059 137.3668- -55.7917' 1

'Kmax' CASE-ID SCALE FACTOR L

, LPRES-AX: 1.00 BEND" 3.34 RESID-B- E1.00 TIME PRINT' TIME- INCREMENT INCREMENT 87600.0- 730.0 '730.0

' ~

CRACK-MODEL: ELLIPTICAL LONGITUDINAL CRACK IN CYLINDER (T/R=0.1,A/L=0.1)

. CRACK ---------------STRESS INTENSITY FACTOR----------------

DEPTH- -CASE CASE' CASE ~ CASE CASE

. PRES-AX -BEND- RESID-D RESID-B RESID-W H '0.0180 - 0.775 0'.225 11.176 12.413 -6.953 L ?O.0360 1.108 0.321 16.217 16.970 -9.979

!. 0.05td -1,371: 0.397 20.308- 20.082 -12.383

.0.07ts . 600 0.464 23.899: 22.393 ' -14.467 p

0.0900 .1.808 0.524 27.149 24.164 -16.342-O.1080 2.001 0.580 30.134 25.534 -18.062 L- 0.-1260 -2.183 0.633 32.895 26.582 -19.660  !

1 ,

0.'1440: 2.357 0.683 35.455 27.369 -21.154 0.'1620 2.526 0.732 37.827 27_.937 -22.556

[

1 L

k u -

s :pc-CR'ACKL ' VERSION '1. 2. PAGE- 2' 9:

O 1800 2.689 0.779 40.020 28.319 -23.874-0.1980 -2.848 0.825 42.040. 28.540- -25.114 I0'.2160- 3.003 0.870 43.890 28.621 -26.281

.0 2340. 3.165- 0.917 45.693 28.668 -27.451 E

'0.2520- .3,333' 0.966- 47.459 28.696 -28.631

'0.2700' 3.500 1.014. 49.072- 28.625 -29.752 0.2880 3.667. 1.063 50.530 28.467 -30.813

,0.3060 3.833 .1.111 51.832 28.227 -31.815" 0.3240 4.000~ 1.159 52.977 27.916 -32.756 0.3420' 4.166 1.207 53.964 27.540 -33.636 0.3600 4.332- 1.256 54.788 27.109 -34.454 0.3780 4.499 1.304 '55.452 26.627 -35.210-0.3960. 4.666 1.352 ~55.957 26.097 -35.902 0.4140. 4.833 1.401 56.301 25.526 -36.531' l'

O.4320 5.000 1.449 56.486 24.918 -37.094 0.4500 5.169' 1.498 56.513 24.279 -37.591 0.4680 5.337 1.547 56.386- 23.600 -38.023

l. 10.4860 5.506 1.596 56.103 22.897: -38.387 0.5040 5.676 1.645 '55.666 ~22.174 -38.682 0.5220 5.847- 1.695 55.077 21.435 .-38.909

( ~0.5400 6.018 :1.744 54.'338 20.684~ .:-39.065 0-5580- 53.451 19.924

. 6.190 1.794 -30.151 0.5760 6.386 1.851. 52.678 19.238 -39.329

). 0.5940 6.591 1.910 51.846 18.566 .-39.486.

.0.6120 6.798 1.970 50.871 17.884 -39.568 L 0.6300 7.007 2.031 49.757 17.194 -39.574 0.6480- 7.217 2.092 48.508 16.500- .-39.501' L 0.6660' 7.429- '2.'153' 47.128 15.806 -39.349

~0.6840 7.643- 2.215 .45.641' 15.126 -39.124' O.7020' -7. 860 2.278- 44.057 14.463 -38.824 0.7200 8.078 2.341 42.364 13.811 -38.444

.0.7380 8.297 2.405 40.569 13.175 -37.981 0.7560 8.518 2.469 38.681 12.557 -37.436 0.7740 8.741 2.534 36.707 11.962- -36.807 0.7920 8.966- 2.599 34.642. 11.390 --36.090 0.8100 9.191 2.664 32.462 10.839 -35.274 0.8280 9.418 ~2.730 30,219 10.323 -34.371 0.8460 9.647 2.796 27.923 9.845- -33.381

'O.8640 9.877' .2.863 25.583 9.409. ' -32.303 0.8820 10.109 2.930 23.212 9.019 -31.139 0.9000 10.343 2.998 20.822 8.678 -29.886

. TIME- .KMAX DA/DT DA A A/THK 730.0 35.16 3.3625E-05 0.0245 0.4295 0.382 1460.0 34.80 3.2849E-05 0.0240 0.4535 0.403 2190.0' 34.38 3.1971E-05 0.0233 0.4769 0.424 2920.0. 33.92 3.1011E-05 0.0226 0.4995 0.444 3650.0- 33.44 3.0030E-05 0.0219 0.5214 0.463 4380.0 32.96 2.9049E-05 0.0212 0.5426 0.482 u' ,

( .

• ~ '

PA D. - pc-CRACK VERSION 1.2 0.5631 0.501 2.8083E-05 0.0205 5110.0' 32.47 0.0199 0.5830 0.518 5840.0 132.02 2.7220E-05 0.6024 0.535 2.6612E-05 0.0194

-6570.0- 31.70 0.0190 0.6215 0.552 7300.0. 31.41 2.6060E 0.6401 0.569 2.5513E-05 0.0186 8030.0' 31'.12 0.0182 0'.6583 0.585 8760.0 30.83 2.4981E-05 0.6762 0.601 2.4469E-05 0.0179 30.55 0.6937. 0.617 9490.0 2.3993E-05 0.0175 f 10220.0 '30.28 0.0172 0.7109-0.632 10950.0 30 04 2.3565E-05 0.7278 0.647 2.3177E-05 0.0169 11680.0 29.82 0.0167 0.7445 0.a62-C 12410.0-- 29.62 2.2825E-05 0.7609 0'.076 2.2511E-05 0.0164~ 9 13140.0 29.44 0.0162 0.7771 0.691 T 13870.0 29.28 2.2236E-05 0.7932 0.705 2.2003E-05 ,

0.0161 14600.0 29.14 0.0159 0.8091L0.719 15330,0- 29.03 2.1807E-05 0.8249.0.733 eg'$

2.1648E-05 0.0158 16060.0 28,93 0.0157 0.8406 0.747 16790,0 28,87 2.1542E-05 0.8563 0.761 '

2.1489E-05 0.0157 17520.0 28.84 0.0157 0.8720 0.775 18250.0 28.84 2.1491E-05 0.8878 0.789 2.1553E-05 0.0157 18980.0 28.88 0.0158 0.9036 0.803:

19710.0' 28.95 2.1680E-05 0.9000 AT TIME 1.9710E+04 CRACK-DEPTH EXCEEDED END OF pc-CRACK I

J

_ _ _ _ _ _ _ _ _ ___ J