ML20024D150

Rev 1 to Analysis of Inaccessible & Potentially Rejectable Defects in Perry Plant.

ML20024D150
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Site:	Perry
Issue date:	07/31/1983
From:	Byron J, Egan G, Naughton P APTECH ENGINEERING SERVICES
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AES-8211352-1, NUDOCS 8308030212
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' AES 8211352-1 Final Report APTEG anginacring farvicci,inc esci~eeRise cO~SuL1A~1S

, 795 SAN ANTONIO ROAD . PALO ALTO . CALIFORNIA 94303 (415)858 2863 t

ANALYSIS OF INACCESSIBLE ~ f-AND POTENTIALLY REJECTABLE - - - -

DEFECTS IN PERRY NUCLEAR POWER PLANT Prepared by Warren P. McNaughton Geoffrey R. Egan Jeffrey D. Byron Aptech Engineering Services, Inc 795 San Antonio Road Palo Alto, California 94303 l

l

.f Prepared for Gilbert Associates, Inc.

Post Office Box 1498 Reading, Pennsylvania 19603 Attention: Paul B. Gudikunst I

eggeggg((2gg8 g July 1983

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Services in Mechanical a nd M etallurgical engineering, Welding, Corrosion, Fracture Mechanics, Stress Analysis

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QUALITY ASSURANCE VERIFICATION RECORD SHEET

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Title:

Analysis of Inaccessible and Potentially Rejectable Defects in Perry Nuclear Power Plant Originated by: / 8'86-83 Warren P. McNaughton N (Lw. %

Jeffrey D. Byron s-zc-ss

,0, Verified by: -

Md 63 feffre . O Approved by: F~20 O Geoffrey R. Egan

( Quality Assurance Approval: 4 // [ h 4!$3 bellC.Cipolla l

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1

1 A

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, TABLE OF C05 TENTS I

r SECTION TITLE PAGE

. SYNOPSIS 11 1 INTRODUCTION 1-1 References 1-4 2 ANALYSIS METHODS 2-1 2.1 Fracture Mechanics Backoround 2-1 2.1.1 Linear Elastic Fracture Mechanics 2-3 2.1.2 Elastic-Plastic Fracture Mechanics 2-5 2.1.3 Limit Load Analysis 2-6

, . 2.1.4 Summary of Fracture Mechanics Background 2-7 2.2 Fatigue Loading _2-8 2.2.1 Analysis Method _ f2-8

. 2.2.2 Crack Growth Rate Representation J-2-9 References 2-10 3 ANALYSIS OF STRESSES 3-1 3.1 Secondary Stresses 3-1 3.2 Primary Stresses 34 3.3 Combined Stresses 3-7

" References 3-14

! 4 FATIGUE CRACK GROWTH RATES 4-1 References 4-8 l

l 5 FRACTURE TOUGHNESS AND STRENGTH 5-1 i 5.1 Introduction 5-1 5.2 Fracture Toughness: Background 5-1 5.3 Toughness Values for Containment Welds 5-4 5.4 CrackOpeningDisplacement(C0D) Values 5-6 l

References 5-8 6 CHARACTERIZATION OF FLAWS 6-1 l 6.1 The Effect of Slag Inclusions on Structure Integrity 6-1 6.2 Defect Interaction and the Modelfng of Defects 6-8 l

l . 6.3 Digital Enhancement Methods Used in the Present Analysis 6-12 6.4 Results of Flaw Characterization 6-13 l ,

7 RESULTS OF ANALYSIS 7-1

.7.1 Results of the Linear Elastic Fracture Mechanics 7-1 (LEFM) Analysis i 7.1.1 Indications in Seams 1-1 and 2-1 7-1 l 7.1.2 Weld 1-4 7-3 l

. . .- ~ . _ . . . _ . . _. _ .m . _ ,_. m ...

.' ii

,, TABLE OF CONTENTS (Continued)

SECTION TITLE PAGE 7.1.3 Welds 1-7 and 1-9 7-3 7.2 Limit Load Analysis 3 7.3 Elastic-Plastic Fracture Mechanics (EPFM) Results 7-6 References 7-19 8 CONCLUSION AND

SUMMARY

8-1 APPENDIX A: Supplemental Toughness Data A-1 APPENDIX B: Background Information About Digital Imaging B-1 Techniques _

' 6-APPENDIX C: Details of Flaw Characterization Work I C-1 APPENDIX D: Controlled Documents D-1 I

i i

3 5 ' 55 l *

~' SYNOPSIS f

This report summarizes the results of fracture mechanics and f atigue

2 l evaluations which were performed f or the Perry Nuclear Power Plant Units 1 and 2. These evaluations were perf ormed f or several regions of the containment structures as follows

l e inaccessible locations in weld Joints 1-1 and 2-1 which had weld Indications

. e Three inaccessible weld locations in Joints 1-4,1-7, and 1-9 l

.- =

t' l In these last three locations, Incomplete radiographic Inf ormation exists to establish the existing def ect size (for example, whether or not f ull repairs were made); however, suf ficient data do exist to characterize the maximum extent of a defect that could remain in the structure and this potential defect has been analyzed.

The evaluations that were perf ormed required three types of input data.

These data were stresses, both applied and weld residual stresses, flaw geometries and material properties.

Recently, revised stress data were supplied by Gilbert Associates for the

,, contai nment. Bounding cyclical and steady state stresses were incorporated to provide an analysis that results in a conservative evaluation of the weld Indications. A conservative residual stress determination was also made by assessing the appropriate experimental data.

Flaw data were obtained from radiographic enhancement techniques performed on supplied radiographs. The radiographs provided contain defects which were deemed rejectable according to the criteria of ASME Boller and Pressure Vessel Code Section lil, Subsection NE-5320. These techniques were used to provide accurate sizing of the indications. Such

ii!

determinations remove some of the conservatism which has traditionally been

used to assess structures with defects in the absence of actual flaw size

. data.

Material properties such as strength, fracture toughness and crack growth rates were determined using available Certified Material Test Reports and by comparison to generic data obtained from the open literature.

The results confirm that crack growth by fatigue is small over the design plant lifetime, even assuming conservative stress levels, bounding initial flaws and worst case crack propagation rates. Furthermore, it is shown that the applied stress Intensity values reached during and af ter such growth are less than the critical value to cause structural failurdi. fThe conservative result of the linear elastic fracture mechanics methodol6sy used in this work is then confirmed using both elastic plastic and net section collapse (or Ilmit load) methods. It is shown that relatively long and deep flaws can be tolerated even with the conservative assumptions which have been made.

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I Section 1 INTRODUCTION

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A review of radiographs for the Perry Nuclear Power Plant Units 1 and 2, has found certain containment welds that contain potentially rejectable

Indications, when evaluated per the requirements of ASME BoIIer and Pressure Vessel Code Section lil, Subsection NE-5320 (1-1). This subsection provites accept / reject conditions based on workmanship

, standards. The radiographs can be grouped into two logical regions of Interest. .- ~

t-The first group of weld Indications are found in radiographs associated iY with weld joints 1-1 and 2-1. These weld locations are inaccessible. They

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j are located in the containment wall (see Figure 1-1) In double sided butt welds and are covered on the inside of the containment by a doubler plate.

Indications have been found that would be considered rejectable by ASME e

Boller and Pressure Vessel Code, Section ill criteria. These radiographs, determined by Level t il evaluation to contain rejectable Indications, include 21 radiographs of weld 1-1 and 43 radiographs from Inaccessile I

regions in weld joint 2-1 of Unit 2. The weld Imperfections on these radiographs were sized using enhancement techniques and conservative interaction criteria. The stresses in Units 1 and 2 are identical, so that bounding defects were developed considering both Unit I and 2 Indications.

The second group of Indications consists of three specific weld locations in joints 1-4, 1-7 and 1-9. Weld joint 1-4 has a defect at location (79-80) 11-i2, which has been sized at 2 3/4" long and 1/16" in height. It appears to be a plate delamination (CD-139, Attachment 4).

In weld joint 1-7, at location (110-111) 25-26, the film shows a i questionable Indication. If this location were accessible the disposition would be to grind and retake. In addition, the film is in question because I4

. - - , - - - . _ _ , . _ r,-

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1-2 9

, if __ Elev.

575'-10"

- Containment Shell Doubler Plate (InsideSurface)

(Inside Surface)

/ Containment Shell Doubler Plate (OutsideSurface)

(Outside Surface)

N Base Liner Concrete lev. 575'-1"

. Joint 1-1 /N Elev. [.FoundationMat 574'-10.187" +-

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l Figure 1-1 Typical Horizontal Weld Joint l (1-1shownhere)

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- - - - _-_._-- ~. -

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' 1: 1-3 l' It does not cover the f ull Indication. At maximum, it would be a 9/16"

long slag line (CD-139, Attachment 3).

in weld joint 1-9, at location (134-135) 24-25, the film has no

, Indications, however, the adjacent film did have repair which extended into this station. A slag inclusion may still be present, maximum length equal to 1 3/4" (CD-139, Attachment 3). These locations are also inaccessible and the existing radiographic Information can be used to provide worst case f

estimates of defect size remaining in the structure, or of the possible incomplete repairs.

All welds have been f abricated using E7018 weld metal. The defects are i

completely contained in the weld metal except in joint 1-4 where the f.

, Indication is in the base material-SA516 Grade 70. The concern in each case is that the indications if unrepaired, may lead to early structural failure. This report addresses that concern and does so by evaluating the

! potential for defect growth by a f atigue mechanism and concurrent or subsequent f ailure by fracture.

The remainder of this report consists of six sections. Section 2.0 outlines the analysis methods that have been used to evaluate the defects.

The next four sections introduce and discuss input to the analytical model.

They are: the evaluation of stresses-both applied and residual (Section

3.0), characterization of material properties (Sections 4.0 and 5.0), and results of the enhancement work perf ormed (Section 6.0). Section 7.0 provides the results of the analysis perf'ormed. Conclusions and summary are provided in Section 8.0. Throughout the report, reference is made to documents which have been used to provide input information to the analysis. These documents which are considered controlled under the

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i requirements of the Aptech quality assurance system are designated Controlled Document (CD) and are referenced in Appendix D of this report.

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1-4 r I.:t .

' lr Section 1 It- REFERENCES

.. 1-1 American Society of Mechanical Engineers, Boller and Pressure

. Vessel Code, Section !!!,1974 Edition, Winter 1975 Addenda.

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Section 2 ANALYSIS METHODS a

The following sections discuss those aspects of fracture mechanics and f atigue theory which were used in the analysis of the present problem. A

. presentation of general fracture mechanics background (2.1) is followed by

. a discussion of methods of analysis to assess f atigue growth (2.2).

2.1 Fracture Mechanics Background Thefailurebehaviorofstructuresundermonotonic(slowlyincreasingf.

loading can be classified into three regimes in which a specific type of f ail ure mode is appropriate. These three regimes cover brittle fracture.

ductile fracture and plastic collapse. The discip!!nes required to assess these regimes are:

e Linear Elastic Fracture Mechanics (LEFM) - The structure falls in a brittle manner and, on a macro scale, the load to failure occurs within nominally elastic loading.

e Elastic-Plastic Fracture Mechanics (EPFM) - The structure fa!!s in a ductile manner, and significant stable crack extension by tearing may precede ultimate failure.

e Fully Plastic Instability (Limit Load Theory) - The f ailure event is characterized by large deflections and plastic strains associated with ultimate strength collapse.

A diagram that shows the relationship between critical or failure stress and flaw size for the three f ailure modes is given in Figure 2-1. The shape and position of the f ailure locus will depend on the fracture toughness (K ) and strength properties (o ) of the material, as well as the structural geometry and type of loading.

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, limit load 2

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Fracture (LEFM)

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[ Limit Load N ~

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k. N'\ \ \

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Non-Dimensional Flaw Depth, 2a/t

,3 Figure 2-1 Schematic Showing the Relationship Between Failure Stress and Flaw Size For Two Limiting Failure Modes.

l 9

.2-3 2.1.1 Linear Elastic Fracture Mechanics (LEFM) t The principles of linear elastic fracture mechanics (LEFM) are applied to assess quantitatively the conditions for brittle fracture.

Brittle fracture consists of two separate events: (1) the initiation of a crack, and (2) the subsequent propagation of the crack to complete failure.

Each of these events, initiation and subsequent propagation, has dif ferent characteristics. For ferritic structural steels of the SA516 type and

, carbon manganese weld metal of the E7018 class, the resistance to a propagating f racture is usually lower than the resistance to fracture initiation under slowly applied loads. This is because steels of this type are sensitive to loading rate; the high loading rates associated with a running crack lead to higher yield strength and, hence, lower values pf fracture toughness. In constant load situations, therefore, continue'd crack propagation is expected once the fracture has initiated. For th is reason, no attempt is made to evaluate the characteristics of the propagating crack af ter it has initiated, and the criterion of fracture initiation is used as the definition of failure in the fracture analyses.

Fracture initiation occurs at a defect when the crack driving force exceeds the material's inherent resistance to crack initiation, o.- fracture toughness. The crack driving force is a function of the stresses acting on the defect and the geanetry of the defect. The stresses which act on the defect include both primary (applied) stresses and secondary (Internal) stres ses. Examples of secondary stresses are residual stresses and thermal stresses that are in equilibrium across the section. The manner in which a structure will fall will be determined by the Interaction of the defect geanetry, loading. and material toughness.

in linear elastic fracture mechanics, the most usef ul parameter for characterizing the behavior of cracks is the stress intensity factor K ,

which describes the magnitude of singular stresses ahead of a crack in a linear elastic body loaded in tension. For loading normal to the crack plane (Mode I), fracture Initiation occurs when the appiled stress intensity factor, K , equals or exceeds some critical value, which is m

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.. . . . . . .. . . . ~ . . . . . . . . - . . . . . . . . . . - . . . . , . . -

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^V 3 called the fracture toughness of the material. The applied stress intensity factor can be written in the form:

4 K = Cc6a- (2.1 )

, I where o is the acting stress, a !s the charactaristic flaw dimension, and C is a parameter which accounts for the flaw shape, structural geometry, and the type of loading.. In general, C is a_ function of a and in many cases must be evaluated numerically.. Fracture will occur under quast-static-loading when, K 2 K,- (2.2) x . .-

(i.e., when the applied stress Intensity factor equals or exceeds t,he[~ ,

static fracture toughness, K ). This reans x that.. the occurrence of Ic -

fracture is controlled by: (1) the stress level, (2) the flaw size, and (3) the fracture toughness. For small flaws, low stresses and high toughness, the applied K will not reach K , and frccture will not occur.

These relationships are relevant for material propertl6s determined under plane strain, ilnear elastic conditions.

To determine the significance of the inaccessible defects in question, it is necessary to know the material fracture toLghness, ecting stress level and actual distribution of defect sizes an"d shapet, Knowing any two of these parameters, one con solve for the third; For example, the critical flaw size to cause failuro is calculated from:

2 1 i (KIc \

a = (2.3) c -s u

---2(

IC oc /

if both the toughness (K ) and the stress level (o ) are known.

le c Conversely, the critical applied stress as a function of crack depth can be -

computed from, 1

- - - - - - , +

j 2-5

K m

Ic 0

C/ia (2.4) i.~

Although these conditions most appropriately describe t.he behavior of low l toughness, high strength materials where little ductility precedes fracture, the use of K as a toughness measure for either SA516 steel or E7018 weld metal ensures a conservative estimate of critical flaw size for brittle fracture, since r.o account is taken of the increased toughness which results from post-yield (transitional) behavior. Incorporating

, transitional behavior with more pre-fracture ductility gives increased toughness levels and decreases the susceptibility of the structure to fracture from a given sized flaw. The temperature dependence of toughness properties means that at emblent or higher temperatures, both SA516_ steel and E7018 weld metal are above their lower shelf values on a fracturef:

energy versus temperature curve. This in turn implies that the use of standard elastic fracture mechanics will be conservative. Elastic plastic

crack opening displacement (00D) concepts have been used as a check on structural integrity. Elastic plastic fracture mechanics concepts are discussed in the next section.

2.1.2 Elastic-Plastic Fracture Mechanics (EPFM)

The basic principles of EPFM have been developed over several years (2:L. 2=2. 2-4) and one national standard exists for crack opening

, displacement (00D) testing (2:1). This method uses critical 00D values (as measures of the material toughness) which are not available for the actual f iel d material . A review of the literature (notably 2:5) was made to check the appropriate material characteristics.

One of the best methodologies for EPFM evaluation, the British Standards Institution published document PD6493:1980 (2:1), utilizes crack opening displacement (000) concepts. In principle, the critical condition IL reached when the applied K or 00D (6) reaches the resistance level of toughness necessary to cause fracture (K or 6 ). The 00D method is Ic c completely compatible with the LEFM approach L2:8) and can be used in place

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2-6

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of the K Ic method. For applied stresses well below yleid, Ec a Y (" \

.I ' 6 = log (2.5) '

sec.l 9-o,;

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1 where 6 is the developed COD; e and o are the yield strain and yleid I stress of the region in which the defec is sited; o is the applied stress; and a is the hal f crack length of a center-cracked plate model.

It can be seen from Equation 2.5 that as o approaches o , the developed CX)D becomes Inf Inite. This only occurs for the elastic perfectly plastic material behavior that was assumed for the development of Equation 2.5.

, For materials that work harden, the relationship between COD and applied strain (for stresses above yield) has been determined by analytical, numerical, and experimental methods (2-9. 2-1.0). ..

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l. ..1 6-As in LEFM, once the stresses and material properties have been characterized, it is possible to determine the allowable flaw size to prevent f racture initiation. It is then possible to determine the expected margin of safety between the flaws that may be in the structure and those necessary to cause f ailure.

2.1.3 Limit Load Analysis .

<'a As the size of a critical flew increases, a regime is entered in which increasing material toughness no longer can prevent initiation of a crack under monotonic loading. The initiation criterion becomes independent of toughness and now becomes a f unction of the strength properties of the material and the remaining iIgament of material. In this regime, a 1imit I

load or plastic collapse analysis describes the governing f ailure mode, i

For limit load analysis, the critical stress to cause failure is calculated from an interaction relation common in the analysis of steel structures.

P This relationship between the appiled membrane load (P) and bending acment (M) at f ailure in a beam or plate with a rectangular cross-section is:

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pe ,- , - . - ..,r. ~. , . ~ - . - . .___.,.mm,,.,-y--...,,_.. %_- , - - _ , . , , , , , . . . , . - - . . , _ _ _ , , . , . . . . _ - . - , . - . _ - , - - , _ - , _ _ _ .

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2 fP\--

{M\

_ i l +t ,= I (Pg Mg (2.6)

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where P and M are the applled loads, and P and M are the 1Imiting values of P and M. The magnitude of P and M are functions of crack a

length ,a_ flaw geometry and material propert'es.

The limit load of P is determined from the gemetry of the section and

.. the material properties. Af ter the reduction in area due to the flaw is accounted, the limit load can be expressed In terms of a lirrit stress and the geometric variables. The limit stress is normally the material yleid strength when the material behavior is assumed to be elastic perfectly plastic. However, for materials which exhibit _

significant strain hardening, o couldbesomewherebetweenyieldond(( ,

ultimate strength, and the appropriate value to use should be determined by tests.

For this analysis, we use a flow stress which is the average of the yleid and ultimate strengths, i.e.:

o = (a +o )/2 ( 2.7) 4 y uts where o is the flow stress, o the specified minimum yield stress and a the specified minimum ultimate strength.

. uts Once the limit conditions have been calculated, Equation 2.6 and the expressions for applied membrane stress as a function of pressure and applied moment can be used to determine the f ailure condition.

A limit load evaluation has been made in conjunction with LEFM methods in the present case.

2.1.4 Summary of Fracture Mechanics Background The f ailure behavior of structures under monotonic loading can be I *

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. __ _ _ = _ _ _ . _ . . _ .

[ 2-8 i l classified into three regimes. Of these, linear elastic fracture mechanics has been determined to be most applicable to the current material and service conditions. Bounding studies based on elastic plastic and plastic I

limit load analyses have also been performed.

1 2.2 Fatigua Lnading 4

2.2.1 Analysis Method The preceding discussion eddressed the case of monotonic loading. In the

, present case, there are a small number of cyclic loads which may occur on the containment structure. This section discusses the way in which these loads can be evaluated in the light of the previous discussion. ' f-

.a Fatigue evaluation, based on fracture mechanics, assumes that initial flaws are present of size a and that the lifetime of a component is that required for a crack to grow from the initial size, ay , to the critical size, a C. Crack growth rate data may be correlated to the crack tip stress intensity factor range (AK) for the given load cycle in the following form:

da/dN = f(AK) (2.8)

where da/dN is the crack growth per load cycle. By integrating Equation 2.8 with the appropriate component stress field to calculate K, the number

.. of cycles, N, (residual life) for a crack to grow from a to a is I C computed from:

c da N =

(2.9) da/dN I

The final flaw size expected at the end of the des.ign li.fe, s , can be f

determined by integrating Equation 2.8, using the appropriate stress distribution to calculate K, and the number of total design cycles N from o

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i 2-9 E f da

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N - = 0 (2.10)

O s da/dN 1 a where Equation 2.10 is a transcendental expression involving a and must be solved by an Iterative process.

2.2.2 Crack Growth Rate Representation

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Many empirical relations to express da/dN behavior have been proposed; the earliest and most well known is the Paris rule L2-L1) which takes the form, da/dN = CAK" (2.11)

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where C and n are constants determined f rom the data, and AK is the cange of applied stress intensity f actor computed from the minimum and maximum stress in the cycle:

AK = AK - AK (2.12) max min The advantage of the Paris relation is that it is simple in form and it fits experimental data well in the middle range of AK. A disadvantage of the relationship is that it does not directly account for mean stress effects (R-ratio effect where R = K /K ) which can accelerate min max fatigue crack propagation. However, these ef fects are accounted for in the choice of experimental data used in the modeling procedure.

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, 2-10 r4 Section 2

!I REFERENCES

. 2-1 Wel ls, A. A. , " Notched Bar Tests, Fracture Mechanics and Brittle Strengths of Welded Structures," Houdremont Lecture 1964,

, British Welding Journal, No.1 (January 1965).

2-2 Sumpter, J.D.G. and C.E. Turner, " Fracture Analysis in Areas of High Nominal Strain," Penemedings Seennd international Cnnference 00 Pressure Vessel Technology, San Antonio, TX (October 1973).

- =

t' 2-3 Egan, G.R. , "The Application of Fracture Tougnness Data to ife Assessment of Pressure Vesel integrity," Prnceedinns Second International Cnnf erence on Pressore Vessel Technolony, San Antonio, TX (October 1973).

2-4 Burdekin, F.M. and M.G. Dawes, " Practical Use of Linear Elastic and General Yielding Fracture Mechanics With Particular Reference to Pressure Yessels," Conferenca on Practfcal Annl ication of Fracture Mechanics to Pressure Vessel Design, Institution of

Mechanical Engineers, London, UK (1971).

l 2-5 British Standards institution, " Methods for Crack Opening Dis-placement (000) Testing." (1972).

2-6 Stuber, A. , J.Wel lman and S. Rol fe, " Eighth Progress Report on Application of the 00D Test Method to the Fracture-Resistant Design of Pressure Yessels," int Suhenamittee on Ff factive UtlI17atton of Yield Strennth of .tha EyRC, (February 1980).

2-7 British Standards Institution, " Guidance on Some Methods for l

l l -

I

l 2-11

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the Derivation of Acceptance Levels for Defects in Fusion

(( Welded Joints," Published Document PD 6493:1980.

I; 2-8 Egan, G.R., "Compatability of Linear Elastic (K ) and

.-- General Yleiding (COD) Fracture Mechanics," Enginaar f ng

. Fracture Machanlcs, Vol. 15, (1973).

A 2-9 Merkle, J., " Analytical Applications of the J Integral," ASTM STP 536, Amartcan society fat Testing and Materials, (1972).

2-10 Hayes, D.J. and Turner, C.E., "An Application of Finite Element Techniques to Post-Yleid Analysis of Proposed Standard Three-Point Bend Fracture Test Pieces," International Journal gg(' f-Fracture, Vol. 10, (1974).

2-11 Paris, P.C. , M.P. Gomez and W.D. Anderson, "A Rational Analytic Theory of Fatigue," Ihn Trend in Fngf naarf ng, Vol .13, No.

1 (January 1961).

I e

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4 3-1 Section 3 1 ANALYSIS OF STRESSES

.; The analytical model discussed in Section 2 requires as input the characterization of the stress state present in the containment shell courses. This section discusses both the primary and secondary stresses.

The primary stresses are the applied stresses associated with dead load and cyclic service stresses. The secondary stresses, in this case, are the

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residual stresses due to welding.

3.1 Sacnndarv Straccas ~[-s

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The welds under consideration are double sided butt welds. A literature review was performed to characterize the resulting distribution of residual stresses in this type of weld. The weld has stress components transverse to the weld and longitudinal or parallel to the weld, each with through-thickness distribu* ions. A schematic of these applicable l distributions is shown Ir. Figure 3-1. For this analysis the flaw location j was assumed to be at the centerline of the plate which is the location for the maximum transverse stress. The flaw location was also assumed to be at the location of maximum longitudinal stress. The transverse distribution through-the-thickness wil l vary as shown in Figure 3-1.

l Figure 3-2 shows single sided butt we!d residual stress data transverse through the thickness. This figure is a composite of normalized l experimental data based primarily on work done by Nordell and Hall (.1-l).

In their work, the base plate was ASTM A212 Grade B (precursor to SA516 l Grade 70) with double-V butt welds of E7018 material. The applicable l r thicknesses tested were 1 Inch and 15/8 inches, requiring 12 and 30 weld passes, respectively. The two thicknesses comonstrated similar through l

thickness transverse stress distributions. Also shown in Figure 3-2 are residual stresses measured by others 0.-2, EI).

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Tension _

Compressi on

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Compression Tension __ .

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(c) l j Figure 3-1 Schematic Standard Assumed Residual Stress Distributions

! in Plates Without Fixed Ends for a Double-Sided Butt l Weld: (a) Longitudinal, (b) Transverse, and (c) Transverse Through Thickness.

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s Butt Weld (Single Side) Transverse Through Thickness 100 - - - 100 80 - -

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-80 ! -80 Percent Wall Thickness

-100 - -100

• Nordell and Hall (3-1),15/8" Double V, Piece 1 (half shown)

• Leggatt and Kamath (3-2),1" Double V, Un-notched (worst half shown)

A Leggatt and Kamath (3-2),1" Doub166V, Notched (worst half shown)

T Nordell and Hall (3-1),1" Double V (half shown)

Nordell and Hall (3-1),15/8" Double V, Piece 2 (half shown)

• Rosenthal and Norton (3-3), 1" Sinale V Figure 3-2 Residual Stresses of Single Sided Butt Welds, Transverse Through Thickness

3-4 t

From these data, a simpilfled through thickness transverse residual stress 4 distribution was developed. This is shown in Figure 3-3. The distribution i

assumes yield level residual tenslie stresses at the surf ace through 10% of the thickness. The tensile stresses then decrease to compressive residual

- stresses equal to one hal f yield at the mid-thickness. This distribution is then reflected about the centerline of the double sided butt weld to achieve a symmetric and complete distribution. Since any residual stress distribution must be sel f-equilibrating, the choice of values taken here will be conservative. The sum of the tensile portions is larger than the compressive portions and the maximum values have been assumed uniform except in the through thickness direction.

- =

t*

3.2 Primarv itgensas The primary stresses for analysis have been obtained from unit stress

~

calculations for joint 1 (CD-130) for wel ds 1-1 and 1-2; and from joint 5

, (CD-139) f or wel ds 1-4, 1-7, and 1-9. The stresses in welds 1-7 and 1-9 are substantially less than in weld 1-4 (see CD-139 attachment 6). A bounding case has been formulated for welds 1-7 and 1-9 using the stresses In joint 5 (elevation 592'-2"). These stresses are sunmarized by joint l

l number and Load Combination for each joint and the applicable load l combinations. Since the flaws found are oriented parallel to their welds, the stresses in the longitudinal direction apply. A schematic diagram of flaw orientation and applicable stress component is shown in Figure 3-4.

The approach taken was to determine the most highly stressed joint and load combination for the appropriate seam welds. These bounding cases could then be app!!ed to any weld def ect, regardless of location, to assure a conservative analysis, in f act, two primary stress distributions were obtained for each weld orientation; one for the f atigue analysis and the other for the fracture analysis.

In order to understand how these stresses were determined, it is necessary to review the tabular stress data as it was provided. Table 3-1 is the

,.,y - . , -ir. .. - _ - - - - 4

.--,er,.,%,w. .z ,pg,.,, yyy .y , _ _ -. mc:9<;==e. 2 _ ,

Assumed Double Sided Butt Weld Transverse Through Thickness 100

\ __ 100

\

l

• . /

yk

/

~~-

e

. g', ..- -

a a 0

\ i 0a5l

,- i Mr

> i

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0 w

E .

0.1 / / '

O.9 M  %' , e' /

\,',y,' ' '

'.;g"i -50

-50 y . s'

• M g Percent Wall Thickness

! Figure 3-3 Assumed Double Sided Butt Weld Tran@erse Through Thickness Residual Stress Distribution

B

?

. \L 3-6 6

a I- Vert cal Double Sided Dutt Peld longitudinal Stress

>r- i y .,

i

s-Horizontal I Double Sided Butt Weld Pt$ i l

Figure 3-4 Flaw Orientations and Applicable Service Stresses in Vertical and Horizontal Seam Welds i <

l l ..

l' 1e l

3-7 5

3 t

stress summary f or Joint No. 5. From this table one can see that the q longitud!nal and circumferential stresses are broken down into thirteen i load components and are then summarized at the bottom of the table into four load combinations, These load combinations represent maximum loading

. conditions. Some of these load components act at alI times, some act

, cyclically over the life of the plant, and some may act only once. The

load components f alling into these three categories are summarized in Table 3-2. For a fatigue analysis the cyclic loads are of primary significance and are superimposed with the continuous or steady state loads. For a fracture analysis the most significant loads are those producing the

, largest stress, which by observation of Table 3-2, are those load combinations that include the single event loads.

- =

s.

Table 3-3 is a summary of the cyclic stresses that govern the fatig'ue' evaluation. There are other cyclic load components. However, the controlling load components in a given load combination are those IIsted in Tabl e 3-3. Thus, in order to simpilfy the analysis and provide conservative results, all critical load combinations were evaluated f or 16,800 cycles, regardless of which cyclic load components the load combinations include. Table 3-4 shows the appropriate load combination components.

For weld seems 1-1 and 2-1, the stresses taken at joint 1-1 (elevation 575'-1") were used, as shown in Table 3-5. For wel ds 1-4,1-7, and I-9, weld seam 5 stresses were used to bound the applied stress condition.

These primary stress fields were combined with the assumed residual stress field for the analysis.

3.3 ovahI nad Stracca=

l The total stress considered for the evaluation of a defect consists of the sum of the primary and secondary components. Both sets of stresses have been chosen to bound the expected stress state conservatively. For 1

analytical purposes, they have been superimposed with elastic J

= tem %~ -t *. - e -vv,-.mm,,,,,wy,aw-.e*-r-we'-ev,* m ev+e=,*e+-e-v er= = em' v - * *-es4th + e - w w e w e e +w o - svNNaWr>'ePe -veme6 were- F * -N+w +-e<=as*4-euw-+'e--r--r-

t

. 3-8

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Table 3-1

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+

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[" lil 3-9 l 1 a

j ',$ Table 3-2 CLASSIFICATION OF LOAD COMPONENTS BY FREQUENCY OF OCCURRENCE Continuous or Steady State loads Dead Load External Hydrostatic Load Due to Annulus Concrete Pour

^

Hydrostatic Pressure 18'-6" s

Cyclic Loads --

.  ?'

OBE .

SSE SRV Discharge - 19 Valves SRV Discharge - One Valve, 1st Pop SRV Discharge - One Valve, Subsequent Mean Condensation Oscillation Mean Chugging l .

Single Load Events LOCA Pool Swell DBA LOCA Thermal Stress 15 psig Static Internal Pressure Ib

,-- , ,e - + ,--+e - -- s-

l' 3-10 m

n Table 3-3

SUMMARY

OF CYCLIC STRESSES NUMBER OF NUMBER CYCLES TOTAL NUMBER SOURCES OCCURRENCES PER OCCURRENCE OF CYCLES

- SRV Actuation 1860 9 16,740 ,

i OBE 5 10 50 i

! SSE 1 10 10 f.

TOTAL 16,800

-_ =

I" 4

1 4

9 i

i I

l  : ..

i l ~

1 .

,(7 .

i

~ ' ' '

3-11

~

s Table 3-4 LOAD COMBINATIONS EVALUATED FOR ANALYSIS 4

Load Combination Number Load Components I

DL + OBE + CONC + SRV1 + HYDRO + PS

, II DL + SSE + CONC + SRV1 + HYDRO + PS III DL + OBE + CONC + SIP + SRVyg + CHUG + HYDRO IV DL + OBE + CONC + SIP + SRV2 + CHUG + HYDR 0 + LOCATHERM

. t-

..u-

. DL = Dead Load OBE = Operating Basis Earthquake SSE = Safe Shutdown Earthquake CONC = External Hydrostatic Load Due to Annulus Concrete Pour SIP = 15 psi 9 Static Internal Pressure SRV yg = SRV Discharge - 19 Valves SRV 3

= SRV Discharge - One Valve, First Pop SRV 2

= SRV Discharge - One Valve, Subsequent Pop CHUG = Mean Chugging HYDRO = Hydrostatic Pressure 18'-16"

, PS = LOCA Pool Swell l LOCATHERM = DBA LOCA Thermal Stresses l~

a 4

3-12

l..'f P t.

Table 3-5 BOUNDING SERVICE STRESSES FOR JOINT 1-1 lt

< Stress (psi) u Location Inside Outside Stress Component l,

'- Thermal 633 -5857 Hydrostatic 421 -818 i Design Pressure 2794 836 Dead Load -492 -659 PSRV 1085/-1847 -2078/3639 t-CO - 146 279 SSE 1731 12194 OBE 1555 11664 Load Combination Stress Range (psi)

' I; i I 1569/-2473 -1891/498 l

l II 1803/-2709 -1361/-32 III 4509/175 -776/1055 IV 5142/808 -6633/-4802 l

Ii I

(From C0-130) Gilbert Ref. Letter PY-STR-1555 e

0

?

Ii 3-13 perfectly-plastic material behavior.

. The cyclic (primary or service) stresses are added to the residual stress

'l distribution such that the maximum stress does not exceed the assumed yield stress of the material . The yield stress used for developing this distribution is 78.6 ksi (see Section 5).

.L

?

I

, -O t-1 t

b t

I 5

l t

e D

l l

I i

i

-.v . , - . -,s, ,--,- , - - - , , . . _ . - . -,., ,, , , - , , - , - , ,

e' 3-14 L

?

Section 3

.: REFERENCES e

. 3-1 Nordel l, W.J. and W.J. Hal l, 'rTwo Stage Fracturing in Welded Mild

, Steel Pl ates," Wel di nn Racaarch Sunnl amant, March 1965, Pp. 124-S to 134-S.

3-2 Leggatt, R.H. and M.S. Kamath, " Residual Stresses in 25 mm Thick Weld metal 000 Specinens in the As-Welded and Locally Compressed 1 .

States," The Welding institute, Report 145/1981, June 1981.

3-3 Rosenthal, D. and J.T. Norton, "A Method of Measuring Triax'lai-4

~~~

Residual Stresses in Plates," Welding Research Supplement, May 1945, Pp. 295-S to 307-S.

l 1.

l l

l l

I?

l1 l

l l .

e

4-1 1

Section 4

~

FATIGUE CRACK GROWTH RATES In order to estimate the maximum extent of crack g. owth that could occur at an Indication over the design life of the plant, a fatigue evaluation

~

was perf ormed. This evaluation combined the cyclic stresses (Section 3) with the appropriate crack propagation rates (Section 4) to obteln the expected crack growth (Section 7).

The purpose of this section is to assess propagation rates for defect -

growth by a f atigue mechanism. With the except!on of the poss!ble-plate defect in weld joint 1-4, the defects are located in weldnents, thus.:

requiring an evaluation of carbon steel weld material crack growth data.

Referer.ces h5vc been drawn together to estimate a conservative (that is fastest possible) boured on potential crack growth. Although no data are available for the exect condition in ef fect, significant studies have been perf ormed to pennit b< unding values to be estimated.

The fcilowing engineering unit conventions are in ef fect unless otherwlse stated:

s AK (stress Intensity f actor range), ksi /In e T (temperature). 'F e da/dN (crack growth rate), Inches / cycle All weldments evaluated are composed of E7018 weld metal. Data available

~

In the literature were collected for all types of carbon steel weld metal with an emphasis on E7018. A study by Maddox (4-1), resulted in a substantial anount of crack growth data for four different weld metals including E7018. The four types of test specimens from Maddox are

t 4-2 summarized in Table 4-1 and the crack growth data for these specimens are

.: plotted in Figure 4-1. Crack growth data for the E7018 weld material (weld metal C) are shown separately in Figure 4-2. Also shown on Figure 4-1 is the bounding line from similar testing on plain steels perf ormed by Gurney

. (3:2).

. Other data from similar weld metals (4:1) with and without stress relief, fall within the upper bound shown in Figure 4-1 for Gurney (A:2). The literature also states that weld metals for joining steels such as A516 Grade 70 exhibit slower fatigue growth rates than the base metals (4-4).

Residual stresses may increase crack growth rate (da/dN), but if these stresses are included in estimating crack growth rates, the data l'n'dikate that the bounding line by Gurney (.d=2) will conservatively predict EFask grow th for E7018 weldments. Figure 4-3 shows Gurney's upper bound which is represented by:

-10 3.44 da/dN = 2,63 x 10 AK (4.1) i The f amination-like defect in weld 1-4 may propagate in either weld material or SA516 Gr. 70 base plate material depending on the exact defect

~

location and orientation relative to the weld. Both casos were analyzed to bound the possible effects. The growth rate used for SA516 Gr. 70 material was derived in previous work for Gilbert Associates (.azi), based on '

bounding curves developed from work by Bamford (4:fi), and ASE Code Section XI (4-7). This work provides a three point curve depending on the relevant cyclic stress range.

-10 3.76 da/dN = 3.8 x 10 AK AK < 4.9 I. -13 8.0 da/dN = 4.4 x 10 AK 11>AK)4.9 1

-6 1.4 da/dN =

3.16 x 10 AK AK>11 b

4

_--r w -w- ww--"'e'- " ^ ' "^

4-3 L

I Table 4-1 TEST SPECIMENS FROM MADD0X*

~

, YIELD ULTIMATE WELD AWS/A5TM STRESS STRESS METAL CLASSIFICATION (ksi) (ksi)

~

A None 74.4 88.0 A MIG deposit using CO2 gas shielding and 1 mm diameter wire Type A-17 to BS 2901, Part 2, 1960.

B E7013 68.3 73.9 A manual metal arc deposit of medium strength using a BS 1719 Class E317.rutile

. coated electrode l' C E7018.G 67.2 82.9 A manual arc deposit of medium strength using a BS 1719 Class E614 HJ low hydrogen electrode.

D E9018.G 89.6 105.3 A manual metal arc deposit of high strength made using a BS 1719 Class E614 HJ low ,

hydrogen electrode.

cger, (4_1)

L S

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9 e

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+ weldmetal D 15 /4-

. aa *a as l

aa (

8 a

10 5 i e i f*10 i i 15 20 30 40 50 60 70 80 90 100 Range of stress intensity factor AK MNniN2 Figure 4-1 Crack Growth data from Maddox (4-1) lJ

4-5 4K. ksiv7ii. pc 636-s 2

10 IS 20 30 40 50 sSO M 80 90 3 . . . .

i i, i, ,,

~ii 10-'

2 .-.-- . _ _ -. /  :

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4 I

. 15.7 ksi 3 )

I a 17.9 ksi -

,, f . 20.2 ksi . g-6

[ ( ~- ) x 31.4 ksi 2 h applied stress (ksi) -

p.

l lie.4 15.7 .

1 70-5 l i i , i

. 9 10 15 20 30 40 50 60 70 80 90 100 Range of stress in.*ensity factor AK,MNm' 2 Figure 4-2 Crack growth data for E7018 from M'ddox a (4-1)

- . - , , , . - - CL__.._

i

, 4-6

~

Stress intensity factor range, 8K (MN/m /2) r! .

p 4 6 10 20 40 60 80 100 10-3 ' I i i i i i i II l i i i l

. Bounding curve for SA516 _ 10 Grade 70 _-

~

T.

5

~

_ 10-4 -

~

b 3

R -

~4 m

- 10 5 -

i,

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-5 m

e

~

a  ? -

5 2 - 5 r -

10-5 y _

x o -

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=

, x U

i 2 k5-u Upper bound E7018 - "

30-6 __

Gurney (h2) _

l _

r

-.- 10-6 l,. _

I 10~7 i i i iiI i i i i I i i I-i 4 6 8 10 20 40 60 80 100 Stress intensity factor range, AK (ksi /i5)

Figure 4-3 Bounding Crack Growth Lines for E7018 --

weld metal and SA516 Grade 70 base material e

, , -.m,%--- .-*m- I

L 4-7 -

~

These values are shown in Figure 4-3. It should be noted that rhls curve

't is very conservative relative to tell experimental data reviewed and will I

,2 provide even more colservative results than the weld metal curve also shown

.. In Figure 4-3.

it l

.M l

,- - O

, , t-i*

't e

+

0 e

k

.--~---------~~~-~---------"~--.-~~vA----

i' . .t. g Section 4 REFERENCES 1

.. 4-1 Maddox, S.J., " Fatigue Crack Propagation in Weld Metal and Heat Af fected Zone Material," The Welding Institute, Report No. E/29/69, AbJngton, Great Britain, 1969.

4-2 Gurney, T.R. and S.J. Maddox. "A Reanalysis of Fatigue Data for Weided Joints in Steel," The Welding Institute, WatdInn Racaarch I nterna ti ona l , Vol . 3, No. 4, 1973.

o 4-3 ~ Y in Seeley, R.R. , L. Katz and J.R.M. Smith, " Fatigue Crack Growth Low Alloy Steel Submerged Arc Welds," FatIgua TantInn af, Watdmants, Pp. 261-284.

4-4 Gurney, T.R., "An Investigation of the Rate of Propagation of Fatigue Cracks in a range of Steels," The Welding Institute Members' Report No. E18/12/68.

4-5 Egan, G.R. , et. al ., "The Signif icance of Sensitized Stainless Steel Material in Drywell Vent and Containment Structures in the Perry Nuclear Power Plant - Fracture and Fatigue Evaluations,"

AES Report 81-11-88, November, 1981.

l g- 4-6 Bamford, W.H., " Application of Corrosion Fatigue Crack Growth I Rate Data to integrity Analysis of Nuclear Reactor Yessels,"

AmarIcan SncIctv af. Machanieni EnnInaars, Paper No.

79-PVP-116, 1979.

4-7 American Society of Mechanical Engineers, Boller and Pressure Yessel Code,Section XI.

] 5-1 1

Section 5

FRACTURE TOUGHNESS AND STRENGTH r

5.1 Introduc+1on Two additional model inputs to be discussed are the material properties; fracture toughness and strength. As discussed in Section 2, the applied stress intensity is compared to a critical value which is definod as the fracture toughness. Thus, to determine allowable flaw sizes, the fracture

, toughness must be characterized. Although no direct measurements of

,', fracture toughness were perf ormed in the course of this work, inference about the level of fracture resistance inherent in the material can by,made by reference to the Charpy impact values which are available. Th e . xl background is presented in Section 5.2. The data are discussed in Section 5.3 for the containment welds and base plates (for possible plate delanination of weld 1-4). Section 5.4 analyzes typical crack opening displacement values to be used in the elastic plastic fracture mechanics eval uation. Section 5.5 addresses the yield and ultimate strength values to be used in the Ilmit load assessment. Certified Material Test Reports (CMTR's) (CD-4, CD-7, and CD-127) were analyzed to determine Charpy (CVN),

yield strength and tenslie strength data. Controlled document 127 was provided specifically to confinn the CMTR's for E7016 used in Weld Joint 1-1 between seems 21-22 where the largest defects occurred. CD-4 and CD-7 .

were obtained in previous work for Gilbert and list data for many heats of E7018 used in containment welds. These data have also been included (see Table 5-1) to indicate the variation in material properties.

5.2 Frac +ure Tnunhneset Backnround To use the analyses described in Section 2.0, it is necessary to have the appropriate value of material fracture toughness in terms of the critical plane strain stress intensity factor (K ). Because of the excellent Ic toughness in this material, these data are net normally available for weld metals such as E7018 at temperatures around 70'F. Yalid K data for

'e 1

s; ,

('

5-2 Table 5-1

SUMMARY

OF WELD PROPERTIES BY HEAT (AS-WELDED)

WELD WIRE YIELD STRENGTH LIMIT STRESS AVERAGE TEMP.

QC# CMTR# (KSI) (KSI) CVN (FT/LBS) ( F) 77NNI518 456 66.3 72.3 77.3 -30 77NNI540 472 78.1 83.2 69.8 -30 77NNI563 493 63.4 70.3 45.0 -30 78NNI004 552 68.8 75.7 82.6 -20 78NNI013 557 65.8 72.0 95.2 -20 78NNIO14 557 68.4 74.3 109.6 -20 78NNIO15 557 65.3 69.9 24.0 -20

[ 78NNI016 557 65.5 71.4 85.0 -20

,, 78NNIO24 625 65.3 69.9 24.0 7, -20

• /8NNIl00 596 78.1 80.7 62.0 J-- -20 78NNI163 630 68.2 73.1 76.0 -20 78NNI164 630 63.8 69.0 115.3 -20 78NNI202 646 66.9 71.5 120.2 -20 78NNI221 653 68.1 73.9 86.8 -40

• 78NNI224 655 66.3 72.6 101.0 -20 78NN1255 663 70.2 73.6 118.4 -20 79NNI016 694 84.9 89.8 66.7 -20 79NNIO17 694 78.6 83.3 92.7 -20 j 79NNIO18 694 67.2 72.7 114.0 -20 79NNIO99 710 64.5 71.3 84.6 -20 79NNIl00 710 70.9 76.4 102.4 -20 79NNI131 716 72.7 79.2 80.3 -20 79NNI161 729 65.3 71.5 56.8 -20 79NNI172 737 65.3 71.5 56.8 -20 80NNI017 746 74.8 79.5 81.0 -20 80NNIO50 752 70.0 77.0 69.0 -20

• 76NNI182 224 68.9 74.9 42.3 -30

_ *76NNI218 256 68.5 74.3 85.7 -30

• 77NNIO58 398 70.0 73.8 138.0 -30 077NNI519 69.0 73.8 113.3 -30

• 77NNI589 520 69.7 73.7 109.7 -30 Taken from CD-4 and CD-7

• Included in CD-127

[Wem&-W @** *

• 7

\r 5-3

'?

e 1.5" thick material are generally only available at temperatures such as

, -100*F. However. It is possible to infer information about the relative i toughness of the present material from available CMTR's. There are several correlations that have been proposed to relate Charpy energy to K Ic values. These include two empirical relationships proposed and verified by Barsom and Rolfe (1-1). The relationship for the transition temperature regime is:

1..

• 2 3 K = 2 (CVN) /2 g (5.1)

E where _

, E-K =

Plane strain fracture toughness (psi /Tn)

E =

Young's modulus (psi)

CVN =

Charpy V-notch energy (ft-lbs)

The corresponding relationship for the upper shelf regime is:

[KIc) =

5 [ o y

)

CVN -

o 1 o i 20 1 (5.2) y/ y

\ /

, where o =

Material yield strength (ksi)

K =

Plane strain fracture toughness (ksi /T5) ic Barsom and Rol fe found that at 80'F. the upper shel f correlation was appropriate for al l material they t3sted. All their ivsts were with I material of yield strength greater than 100 ksi, although they claim that l

Equation 5.2 is valid for materials with yleid strength less tt.an 100 ksi If dynamic yield strength is used instead of static yleid str ength.

Another omnmon correlation, dLe to Sailors and Corten, which was developed i for A533B and A517F (1-2), is i-4

" '~

~ '

" - -~ -e". ,- 1, , --,.l _. _

~

5-4 1

• 0.5 K = 15.5 (CVN) (5.3)

,; ic

where

. K =

ksi /Tn CVN =

ft-lb Pisarski (.1-1) who reviewed and verified by experiment ten correlations 4

including those listed abcve, foana that good predictions can be obtained for high strength steels (o > 113 ksi). For lower strength steels, the y

correlations tend to be generally conservative with the degree of conservatism Increasing with decreasing yield strength. Thus, either Equation 5.1 or 5.3 should provide conservative estimates of critical fracture toughness. As a check, relations between critical crack opening displacement value and K Ic arealsoavailablefromRolfeandBarsom~(5[4) and Egan (.5-2), and take the form: "^'.

c . ,(K_ Iji\

(5.4)

'y \ *y )

where l 6 c

=

Critical crack opening displacement (In.)

I c

y

=

Yield strain (in/In) = c /E y

K =

Crit! cal fracture toughness (ksi /In)

IC

+

o =

Yield Strength (ksi)

Y A f urther evaluation of typical crack opening displacement (00D) values is found in Section 5.4.

l l

l e.

5.3 Tnunhness Valuan IDr_ Containment Welds Specific certified material test reports (CMTR's) were reviewed only for veld 1-1 between vertical joints 21-22 (CD-127). Furthermore, CNTR's for Perry containment stif fener welds fabricated using E7018 were evaluated in earlier work by APTECH (li::fi). These weld data are considered I

l<.

i . - .. , . _ _ . __ _ - .

5-5 s

representative of those that would be found elsewhere in the containment.

r 4 There is a large scatter of Charpy V-notch (CVN) daia as shown in Table 5-1. The range of test values represented there is 24.0 to 138.0 f t-Ibs.

The values given for CVN in the table are the " average." This is the average the 5 data points listed in the CMTR or 3 data points if 3 data points are given. In order to be conservative,1he lowest CVN value was a

used to determine fracture toughness (K ) of the weld material. Thus, IC the 24 f t-lbs. corresponds to a Barsom-Rolfe toughness value of 82.6 ksi

/Tn. With the exception of this one heat, all other heats have calculated K values greater than 132.3 ksiv77i.

These values represent tough welds, particularly since the CVN tests Yere perf ormed at a maximum temperature of -20*F. well below the operatinh ~

temperature. This fracture toughness value of 82.6 ksI/Tn will be conservative since:

e The Barsom-Rol fe correlation used to arrive at these values has been shown to be conservative for materials with these

, strength levels.

e Most calculated K values using this correlation are substantially abo 8 this level.

e The test temperature used to evaluate K is -20*F, whereas a higher temperature during opebEtion will result in correspondingly higher toughness.

A lower bound determination of SA516 Gr. 70 toughness expected in the containment structure was performed in previous work for Gilbert Associates c (ji::2) . This value was found to be 73.2 ksi /Tn.' Tne derivation of this result involves considerable conservatism.

l~

l 1

1 4

m

, 5-6 Other values determined from Table 5-1 to complete the analysis are yield

. strength and limit stress. The yield stress is used in determination of residual stresses, as they are a function of yield strength level. The higher the yield strength of the material, the higher the residual

, stres ses. Theref ore, the upper bound yield strength is used to determine the maximum possible residual stresses present. The !!mit strength is used In evaluating the limit load capacity of the structure. The limit strength (o ) is def ined as

, 4 o = (o +o ) (5.5) 4 y uts /2 where yo is the yield strength and o is the ultimate strength. For uts . ~ .

conservatism in the limit load analysis, lower bound values for yield (and ultimate strength are used in the determina tion of the limit strength 7 For yield stress, a value 78.6 ksi has been used ano for limit stress, 66.0 ksi.

(See Section 7.2) The conservatism is apparent in that the prescribed yleid stress is 12 ksi greater than the limit stress used in the analysis.

5.4 Crack Qpening Disniaramant fo0D) values Crack opening displacement testing is used es a direct measure of fracture j resistance. Literature data are available to provide typical COD values for E7018. These are presented in Appendix A. These data were used in two ways. First, as a check in the derivation of K and second, as direct Input to the EPFM analysis. Ic A check on derivation of the K value used can be provided by Equation 5.4. From the data in Appendix Ic A, the lowest COD value data at 32*F is I

.023". For th is val ue of COD, and f or c = 0.2%, o = 63.4 ksi (the Y Y lowest strength material given in Table 5-1), the resulting K value is i .ca lcu a el t d as:

Ic i

l l

1 .

l

~

6

.. ,~,-- ------ -, n .

L P

l t h, t 5-7

' E 6

_.c_ , (KIc)i

1. y (y/

K = 214.9 ksi /In

, Ic Thus, the value of 82.6 ksf/T5 taken in Section 5.3 corresponding to a CVN val ue = 24 f t-Ibs. , is very conservative.

e 4

i .e m **

i ,

g.

e I .

8 I'

l ,

9 e

\ .

9 s

e-4 l g l E m _

• - __4 _ . . _ . _ _ _ _ _ _ _ _ _ .

5-8 s

?.

Section 5 7 REFERENCES

5-1 Barsom, J.M. and S.T. Rol fe, " Correlations Between K and Charpy V-Notch Test Results in the Transition Temperature Range," ASTM STP 466, (1970), Pp. 281-302.

5-2 Sailors, R.H. and H.T. Corten, " Relationship Between Material Fracture Toughness Using Fracture Mechanics and Transition Temperature Tests," ASTM STP 514, ( 1973), Pp. 164-191.

5-3 Pisarski, H.G., "A Review of Correlations Relating Charpy f-Energy to K . . -

Ic, " Iha Waldt nn institute Reseacch

, Bul leti n. (December 1978), Pp. 362-367.

5-4 Rol fe, S.T. and J.M. Barson, Fracture and Fatigue Control in Str u ctur es . Annl lcations ai Frae+ure Mechanics, Prentice-Hal l (1977).

5-5 Egan, G.R. , "Compatiolilty of Linear Elastic (K ) and General Yielding (COD) Fracture Mechanics," Fnnineering Fracture Mechanics, (1973), Vol . 5, Pp. 167-185.

5-6 Egan, G.R. , W.P. McNaughton and J.D. Byron, "A Fracture Mechanics Analysis of Containment Stif fener Flange Welds in the Perry i Nuclear Plant," APTECH Report, AES-82-01-92 (April 1982).

i 5-7 Egan, G.R. , et. al . , "The Signi ficance of Sensitized Stainless Steel Material in Drywell Vent and Containment Structures in the Perry Nuclear Power Plant - Fracture and Fatigue Evaluations,"

AES Report No. 81-11-88, November, 1981.

a

, f.f 6-1

?

7 Section 6 c- CHARACTERIZATION OF FLAWS I

a ,

The final input required for the fracture mechanics evaluation is flaw size. The applied stress Intensity factor calculated by linear elastic fracture mechanics methods and the net section stress of the Ilmit load method will both require an accurate description of flaw dimensions. This l wilI include both depth and length Information. Length information is generally easier to obtain as the projection of length onto film is

.' obtained by standard radiographic methods. Depth data have been less easily obtained without resort to volumetric examination by ultrasonly, techniques or destructive testing techniques. For structural integrity evaluations an assumption has been generally imposed that confines the flaw depth to one weld pass in multipass welds for certain defect types. Th is assumption wilI be conservative for porosity and slag inclusion defect types. However, in many Instances it may be overly conservative. Such an s.

assumption confining the expected defect depth to one weld pass will not however guarantee conservatism for " linear" defects like cracks, lack of fusion and lack of penetration. To more fully characterize both types of defects in the weld joints of interest, a radiographic enhancement technique has been used. This is discussed in Section 6.3 below. The enhancement procedure also allows eccurate length sizing of defects. When combined with equations of interact!on (discussed below), this allows the analyst to determine if two adjacent defects or a series of defects should be most accurately represented as single Imperfections or treated as continuous. The details of defect interaction are discussed in Section 6.2. The following section discusses the ef fect on structural Integrity of the rounded defect types, particularly slag inclusions.

6.1 Iha Ef f ac+ nf. Slag i nct union = an s+ruc+ura in+aneI+v Work by Harrison (6_1) has Indicated that slag inclusions have little ef fect on the tensile strength of butt welds up to considerable percentages

_ _ o

J 6-2

> 1 I  ?

of cross-sectional area. In support, he shows results of work by Ishil

~

(.6-2) and by Kihara (f=1). These results are shown in Figure 6-1.

Harrison further points out that by their nature slag inclusions are unlikely to occupy a large proportion of the cross-sectional area of a given weld and the weld metal will usually overmatch the base metal in strength. The conclusion to be drawn from these f actors is that the ef feui

, of slag inclusions on static tensile strength in materials like E7018 is

, negligible. Harrison conf irms that size-for-size, slag inclusions will be less detrimental than cracks because of their roundness and limitations on their through-thickness size.

A similiar conclusion is reached considering low cycle fatigue. Work by F

ishil and lica L6:d) is shown in Figure 6-2 and Indicates that slag _{;

inclusions have little4 ef fect on load-controlled low-cycle f atigue and up io lives of about 10 cycles. The design can thus be based on the static tensile behavior. For the analysis of these inaccessible defects, the structure may be subjected to as many as 18,600 cycles. This is still considered low-cycle fatigue for the purposes of our analysis, and the l'.

l ef fect of slag inclusions will be well characterized by the steile loading case, particularly in light of the relatively low magnitude of the cyclic stresses (relative to the fully reversed limit level stresses used to generate the S-N curves of Figure 6-2). Additional results given in figures 6-3 though 6-5 Indicate the ef fects on fatigue strength for high cycle fatigue. The number of cycles required to enter a regime characterized by substantial ef fects on life is shown to be at least an l- order of magnitude greater than the design life in the present case.

(

in su;amary, Harrison L6-1) states that there seems to be suf ficient evidence to indicate that under load-controlled conditions, low-cycle fatigue is not a problem which will be influenced by the presence of slag inclusions.

The tensile strength, o , of a defective butt weld will be either u

. . _ _ m ._ _ . , . --

6-3 s.

ifA

>K

\*

y yy , _ . m -. . .. . _ . . , - . .

, _e-Starter is't res* te r4 7h s.~: y - -- - ,,,,,

i

'{'

3x! .

4,

~ ~ ~ . , _ _

a i 3::- -

% a , . - -

o 5-r .; .

.s  ; - g. ,; . .. -

c.wr. a ,, iw ,,e ruo.

Figure 6-1 Effect (taken fromof Slag 6-1 ) Inclusions on Tensile Strength i

85 ( s e s

7) s s:c-n;-

,- . 1 -

t ., J. -

k 3 JJJ'- 7 Em-

. w m,-

,i m-e

} ,

,,_ - _ \ '

i

, 'N L -

• L=---_K - n ' .,5 as e v e' e- c> ca e-Figure 6-2 Results of Load-Controlled Repeated Stress Fatigue Tests on Butt Welds Containing Slag Inclusions

. (taken from 6-1) .

4

c 6-4 L

tt P

rr

' a* .

e

~

.d

.o:- P.'. ' ** :.::i ,',

.m 253r\ ' """

j '.

~ 2m[ . . . . , ,,1

j l25N' g g;

.: .., )

1ll  !,7 i

)'*[N N

, l 5a- N- ..

• • c'r.,.,.,,,,,, 5'

. t-

. w-Figure 6-3 Results of Tests on Low-Hydrogen Welds Containing Slag Inclusions up to Smm Long (taken from 6-1) m o

.ex =

.scT 73 su ..

N..

1 h

} :=w .

N..< -

• v

' </.4 ..< g/ ,'-

i7 < 4., ' :;y

, , ni s eu -

. .s .s

c -

.e;

. -,n e<

j '$

~

,, t.eI 1e- .  !'

a' a=

imewes,s. tyvers 4 , Figure 6-4 Results of Tests on Low-Hydrogen Welds Centaining Containing Slag Inclusions up to 25m Long

. (taken from 6-1)

7 t.

t s 6-5 x

I

r '

f 2

I-

'7

,i l

lI w

w. n

}40 aw. ,a 4

ng

~ we .

g .

• 'A ,sg.

s<

a. e 4

'"= }

f h ,w.

. ;.'..:!.:r;.-. ' ' f., . , , ?> l

~ $.

g *y ,

.o.-

.4 .. ,

saf 4 .

g 12 e

1

. )9[i N,s 4

,',. 3 ue c. ex .

, -+

Figure 6-5 Results of Tests on Low-Hydrogen Welds Containing Slag Inclusions up to Continuous Slag Lines (taken from 6-1) g.

e jP 9 e

0 e w -- __w__ __

L

6-6 IL (1 - AA/A) 0 u,w it or o u,p 6

whichever is least. Where:

o = the tensile strength of the weld metal u.w o = the tensile strength of the parent material

u,p AA/A = ratio of the loss of area due to porosity or slag inclusion to the total area The ef fect of porosity on structural integrity is similiar to that fof-slag inclusions. Figure 6-6 shows the ef fect on tensIIe strength of a weId as a f unction of vol ume of pores. This figure is from work by Harrison (4:1).

. Figure 6-7 shows the ef fect on f atigue life for porosity defects for the case of low cycle fatigue. The behavior is similiar to that for slag inclusions. Harrison concludes for porosity (5:1) that, "There seems to be suf ficient evidence to Indicate that, under load controlled conditions, Icw cycle fatigue is not a problem which will be influenced by practical porosity levels." Furthermore:

I "In view of the probable necessity to limit porosity to some percentage probably well below 10% because higher levels would obscure other defects, there is no need to give further consideration to the ef fect of porosity on static ductile strength. This is because weld metals normally overmatch parent material strength and even where this is not the case the percentage reduction in strength due to porosity is equal to the percentage by volume of porosity and at a maximum of 10% this would not in any normal circumstances be significant."

l 6

s k -- _ ~ g

6-7

' . . . . 50 Ret 6-6 Ret 6-2 T 700 _, _ _ _,_,

mM" ,,,,*9)4%l%%Ms#A.;W8,,

k%2 r 600 - ..'$. _ Y5.Y- .Nhjfc'd!- Ref. 6 '#0

• g50u kYd%W^

u ww w,www#

1. w *4nn w Cp n n a>>:v e , .

_ ggg

%g.

S 400 -

"N s I

U @

t Ret 6-7 t

% 300 - -

20 %

g e Ref. B-E g E E

& 200- 2 10 10 0-

$ t_

I I I 0 2 4 f

O-6 8 10 Volume of pores %

Figure 6-6 Effect of Porosity on the Tensile Strength of Butt Welds (taken from 6-5) 700, . . 45 i

600'- l

~

W 500 -

35 e i Percentati 30 *e G 400 -

Porosity - 25 S

& I by volume

• g 20 h300 l 1%

. E E n l 3% -

15 m 5 200 w

10 9

6 10 0 7

05 10 W2 WJ 10 ' W3 M' N Endurance. cycles J

Figure 6-7 Mean S-N Curves for Load Controlled Tests on Porous Welds from Ref. 6 '9 (taken from'6-5)

. 6-8

~

~

6.2 Defect intaraction and .tha Morial Inn ni Def ects in Iight of the discussion In the previous section it is clear that the rounded defects such as porosity and slag inclusions wilI have less detrimental ef fects on structural Integrity than linear type defects.

~

However, for the purposes of this analysis we wilI continue to model the rounded defects as a sharp crack-like defect of the same size. This assumption will lead to a very conservative assessment of the potential for 4

failure as caused by these imperfection types. Thus both linear and rounded defects will be treated as linear imperfections.

Fracture mechanics analyses Indicate that the ef fect of a surf ace linperfection on structural integrity will be much more severe than_that of a buried Imperfection. Since we presently have no information about khe relative location of the observed Imperfections within the weld (in one or two cases some location data are available in th3 form of ultrasonic inspection records, but generally speaking this additional Information is not available), the assumption has been made that the defect is surf ace-connected. This will give the most conservative result in the calculation of applled stress intensity factors and is consistent with the

!, treatment outlined in the flaw characterization methods of Section XI, i

, Subsection IWA-3370 of the ASE Boller and Pressure Yessel Code (fi .LO).

l The model that has been used incorporating these considerations is the t

l ,

surf ace connected elliptical flaw as shown in Figure 6-8.

The available information which is contained in the radiographs of the j welds of Interest has been assessed using a computer enhancement system.

That procedure is discussed in the following section. However, as a prelude to that discussion, sme preliminary remarks should be made regarding the interaction of adjacent defects. Two International standards are commonly used to evluate such interactions. These are the ASME Boller and Presure Yessel Code,Section XI (.6-LQ) and the British Published i . Document PD6493:1980, " Guidance on Some Methods For the Derivation of l

Acceptance Levels For Defects in Fusion Welded Joints" (.6-1.1). Figure 6-9 demonstrates the ASE method of evaluation, and Figure 6-10 illustrates the

6-9 4

9 r

~

Ifl = 305 . ILLIPilCAL SURTACC CRACK 4

M W L Cf0PiffR)

,Y o..(m..)

N "

i t 11 ls ,

j l I

+, - - ll *2 l l

l lI - - -

_ y I '3 I  !'

i g ..-

I a

• • 1 -

g l /.

y, I

i y vg .__

+

a

'C 70CCL DESCRIPTICN

!10_ DEL T[ATURES

_ PARNIETER OPTION F[ATURED l'.adel Inden Number IFI 335 tbr.ber of Degrees of Freedom IDUF 3 Crack Front Shape --

, Semi. Elliptical Crack Opening Itode --

Mode I Finite Width [ffects w No Variable Thickness Effects NTH No Figure 6-8 Schematic of Surface Elliptical Flaw Model JL i.

6-10 GENERAL REQUIREMENTS Fig. IWA 33301 y >*

i

1. , *- 2a -** S > 0.4a >

,) .

" '# 1 *~~

~-

-> dg s h h

.1 '

SCL k

Surf ace flaw .1 g

f e 4

W Subsurface flaw .2

~

C 5 < 2a \ \

. t )*

y U Uncied surf ace ----* * # 2 d-2d 3

- m- d, + # # I#3 *' I#2

( (whichever.5, ~

-,2d; g- s greater)

-> n l . _.

1 1 r

y Subsurf ace flaw .3 *-Clad surface S > 0.40 g ~n 4-S C 2dj or d2 4--- e- Pressure retaining (whichever is greater) 1 I surf ace o' unciad

• 1. cornponent or clad -

- #3 M >0.4d 3 -> base rnetal interf ace

• ---28 5 of clad component S > 0.4d g -> e. 2d g .* 2d y $ y n,a 4 2

4->w z JL

= .

A 3 c . ,,

Y Subsurf ace flaw .4 Y --> e-

- S < 2d3 or 2d2 (whichever is g' eater) 7

, NOTE: __1 L d,dg,d2# 3'2# 1=2# 2-and 2d3 are depths of 4----- 2a -

individual flaws. Surface flawe5 S < 2dj or 2d2 -

I 2d C (whtchever is greaterl

+ -

e. dh d

FIG. IWA-3330-1 MULTIPLE PLANAR FLAWS ORIENTED IN PLANE NORMAL TO

, ~/

PRESSURE RETAINING SURFACE lilustrative Flaw Configerations and Determination of Dimensions o, 2a, and l' r Figure 6-9 ASME Section XI Defect Interaction Criteria (taken from 6-10)

Schemitic defects Criterion for interactio3 Effective dimensions I; .

Lf ter interection .

1 Cnpf;we surf acs defoct:

t = (2 I +12

- _ _ , g ,7 s

u h r

' ~ ~ ;x 0 3 y I = 1,+12 +s fy - qY! g -f

!,1. s! la  !

2 Coplanar embedded defects

') 7 -

t=tg+t2+3

.s ., . I g, + t i

)

{ 2

~

a-[ f=f

! 1- ,E ,

)

. ._._I t, _ _j 3 Coplanar embeaded defects

. _  !. . _ . . _ - h I ", I2_

~

. - - I .. o yg1*5 5 2

{.

q,_, 'l

.'_~ I, f s i I"I*$*# I 2 y ,

la l 1

(

4 Coplanar surf ace and embedded defects

) t=ti+t2+8

l. h

') I j

[ s<U+f 2

!, V' ' ~~l,

) I l ' l~s 1

\ '

r& f , >

L <t 5 Coptanar embed 6ed defects I s, 4 51 52 t=t1 + tg + s2

, I q and

'll L.'......

1_.s,

_(2 q

,+t2

, z t g , g, , y, , ,,

. 6 Coplanar surface end embedded defects si < 1 +A12 1

t=ti + r2 + 82 I

I r -------

l and 1 !t f'

l lM l

e I, l s, is I

l l

Sa I,

s2 < ti+5 2 l=l+ i 1 2+ si Figure 6-10 Planar Defect Interactions (takenfrom6-11)

.% pe om . 4 e . -  %

i

'[

6-12 PD6493 method. In eliher case, a separation distance, s, is calculated as

, either a f unction of the through thickness dimension (ASE) or length

dimension (PD6493) of the Imperfection, if the actual distance between h

adjacent imperfections is greater than the required separation, s, then the

defects are treated as separate for the structural integrity evaluation.

The criteria of PD6493 were found to be more conservative for the long,

relatively shallow Imperfections which provided the bounding conditions.

l 6.3 ninital Enhan< ament hthodt USA.d 1D .the Pracon' Anal vs t s l

Digital enhancement methods have been used in this analysis to provide accurate length and depth values to be used in the fracture mechanics '

~

calculations. The image of a radiograph is digitized and then computkr manipulated to provide accurate meesures of film density. Indicatio'ns on the radiograph can then be Interrogated to determine their extent in the through-thickness direction. By comparing nisnerical measurements of density of Indications with known density changes from image quality

, Indicator wires (penetrameters) or the plate thickness, the depth of Indications can be determined. In this case, the density at the defect locations has been compared with the general density of the surrounding weld area to detennine the through thickness extent of the defects in terms of a percentage of the weld thickness. A further use of the technique, l .

! particularly important in this case, was the use of digital information to determine the extent of weld defects. This was performed by determining the " ends" of any given defect, that is the point at which The dansity was found to be Indistinguishable from that of the weld remote from the defect.

Once this location was found then the interaction criteria given in the previous section could be applied to determine the defect length to be used

. In the fracture mechanics analysis.

The digital enhancement techniques used hcVe been applled to several nuclear applications including both the enhancement of radiographic records and the real time signal enhancement of Inspections by remote video camera.

Appendix B provides further information about the previous uses of thess h

e (N

R 6-13

, [ techniques and includes a recent paper presented at the International Conference on Fracture Toughness Testing - Methods Interpretation and Appl ication in London, June 9-10, 1982, (4-12), which gives further background Information about the methods used to establish the depth of Indications in radicgraphs.

6.4 Results nf. .the f_ Lax CharacterI2ation A total of 21 radiographs from weld 1-1, an 43 radiographs from weld 2-1, as well as film for weld 1-4 (79-80) 11-12, were examined using the enhancement techniques discussed above. The results are given in Appendix C, which lists all the welds examined, the maximum def ect length (af ter applying the .nost conservative interaction criteria availabfeij. and depth found on that radiograph also listed. In many cases, the maximum depth and length were on separate flaws. For example the maximum dapth

. defect was usually a rounded Indication, but the longest defect was linear.

It should be noted that the resulting bounding or worst case flaws will be much more accurately sized than simply assuming a depth equal to a full weld pass. Furthermore, the linear Indications tended to be shallow, although in some cases, relatively long.

In most of the radiographs, the defects were singular or not separable by analysis using the Interaction equations. Several regions marked as containing rejectable Indications were found to have separable defects using the Information provided by digital enhancement techniques. The defects deemed separable by use of the wore conservative PD6493 criterion were checked using the ASME Code 53ction XI criteria and confirmsd to be separable under that Code as well.

The interaction criteria were not used in the cases of the bounding defects given below, as they are continuous. The enhancement work allowed the following worst caso defect types, lengths and depths to be entered into the fracture mechanics analysis:

- -. -i.----,-- . - - - - , - _ , , ,_m-- w . , .-e. - - -- , - - - . --

6-M i

Weld Location Maximum Length Depth Weld ID

't

,; Weld 1-1, 2-1 Inaccessible Defects 4.0" 12.0% 1-1(21-22), 5-6 1.125" 14.3% 2-1(35-19), 14-15 R1 (slag defect)

Additional inaccessible Welds Wsld 1-4, 2.75" 10.0% Estimate of worst condition frcun CD-139, attachments 3-5

- r s-Welds 1-7, 1-9 1.75" 10.0% Estimate of worst.:

condition from CD-139, attachments 3-5 g.

r s i

• O ie a

d #

{

___.__,.,E. , - - _ . - _ . - . . - . . - . _ _ , . . . . . . , . _

- . . . . . . - . . . ~ . . . . . . -

re .

i 6-15 T

a-Section 6 REFERENCES

. 6-1 Harrison, J.D., MBasis for a Proposej Acceptance Standard for Weld Defects. Part 2: Slag ine,lusions," The Welding Institute,1972, Abington, Cambr!dge.

~

6-2 ishil, Y., H. Kihara and Y. Tada, "On the Relation Between the Non-Destructive Test!ng information of Steel Welds and Their Mechanical Strength." Journal of Non-Destructive Testing (Japan), Vol .16, No. 8,1967, Pp. 319-345. IlW Document Xill-466-67.

- f, 6-3 Kihara, H. , Y. Tada, M. Watanabe, and Y. Ishil, "NctrDestructive Testing of Welds and Their Strength," 60th Anniversary Series.

The Society of Naval Architects of Japan, Vol. 7.

6-4 ishil, Y. and K. lida, " Low and Intermediate Cycle Fatigue Strength of Butt Welds Containing Defects," Journal of the Society of Non-Destructive Testing (Japan), Vol.18, No.10, 1969, llW Document Xil-560-69.

6-5 Harrison, J.D., "The Basis for a Proposed Acceptance Standard f or Wel d Def ects. Part 1 : Porosity," The We! ding Institute Report 26/3/71, March 1971, Abington, Cambridge.

6-6 Norrish, J. and D.C. Moore, " Porosity in Arc Welds and its Ef fect on Mechanical Properties," Second Conference on the

~

Significance of Defects in Welds, London, May 1968 (Published by the Welding Institute).

O 4

I. g

T 6-16 i

f4 I 6-7 Green, W.L. , M.F. Harmad, and R.B. McCauley, "The Ef fects of

, Porosity on Mild Steel Welds," Weld J. Res. Supp. Vol. 23, No. 5, Pp. 206s-209s.

6-8 Masi, O. and A. Erra, "L'Esame Radiograf ico del le Saldatura, Una Completa Yalutaziono del Diffeti in Termine di Resistenza Statica ed a Fatica," La Metalurgica Italiana, Vol. 45, No. 8, 1953, P. 273-283.

6-9 lshi, Y. and K. lida, " Low and Intermediate Cycle Fatigue Strength of Butt Welds Containing Defects," Journal of the Society of Non-Destructive Testle.g (Japan), Vol .18, No.10, 1969. IlW Document Xl il-560-69. - r

, l'

.u.

6-10 American Society of Mechanical Engineers, Boiler and Pressure Vessel Code,Section XI,1977 Edition, Winter 1977 Addenda.

i .

6-11 British Standards Institution, Guidance on Some Meihods for the Derivation of Acceptance Levels for Defects in Fusion l

Welded Joints," PD6493:1980, British Standard institution.

, 6-12 Egan, G.R. , M.F. Elgart and A. A. Smith, " improved Radiographic Flaw Sizing by Digitai Image Processing," International Conference on Fracture Toughness Testing - Methods, inter-pretation and Application, June 9-10, 1982.

0 l

l l

h s

• m

7

7-1 3

~

Section 7 RESULTS OF THE ANALYSIS l

c The following sections present the results of the analysis which were perf ormed using the methodologies discussed in Section 2.0 and the inputs of Sections 3.0-6.0. The results of the linear elastic fracture mechanics analysis are discussed in Section 7.1, followed by the results of the elastic plastic and limit load analysis in Sections 7.2 and 7.3 res pectivel y.

7.1 Rennits ni fha Linaar Elastic Frneture Mechantec (LFRM) Ann t y [s The LEFM analysis was performed using the BIGIF (Boundary integral -a-Generated influence Function) computer program. This program performs the numerical Integration of equation 2.10. The program can be used to evaluate tho ef fect of both cyclic and steady state loading on a structure.

7.1.1 Indications in Seams 1-1 and 2-1 Table 7.1 sunmarizes the input conditions and results for the four bounding cases which were established. For seems 1-1 and 2-1, the maximum length defect could be modeled as 4.0" long and 12% of the weld thickness In depth (see Section 6.4 for details). The worst cycIIe stress condition was found (see Section 3.3) to be load combination ill (CD-130). These conditions and the bounding fatigue growth rate developed in Section 4.0 were combined in case number 1, which is BIGIF run 1H1. The results are shown In Table 7-1. These Indicated that very small growth over the full design life of the plant for these conditions can be expected. The possibility for fracture, given a maximum one time stress loading of this final flaw size was then assessed. The results are listed in Table 7-1 under case 1, run 1H2. The appropriate bounding stress case was load

~

canbination IV (CD-130). The results Indicate an applied stress Intensity factor (K ) for this bounding case to be 60.0 ksi In. This compares to the critical value 82.6 ksi in (see Section 5.3). This in turn Intplies that this size defect will not propagate by a fracture mechanism.

IT

m

~

~ -

.. . , ._ _ . . . . , < . , - . .. a

. Table 7-1

SUMMARY

OF RESULTS Run Weld Data Stresses Bounding Flaw K Evaluation a 7 Case No. Orientation Seam No. Joint Load Combination Lenath Depth Forl ksi/In K

-ICIK 1 1H1 Horizontal 1-1, 2-1 1 III-+3 4.0 12%(.180) Fatigue 60.0 1.37 I

1H2 Horizontal 1-1, 2-1 1 IV- 4 3 4.0 60.0 1.37 12%(.180) Fracture

, 1H5 Horizontal 1-1, 2-1 1 III-$

3 1.1 14%(.214) Fatigue 57.5 1.44 1H6 Horizontal 1-1, 2-1 1 IV- 43 1.1 14%(,pla) Fracture 57.4 1.44 2 6H1 Horizontal 1-4 5 III-4"'S 2.75 10%(.150)' Fatigue 54.68 1.345 6H2 Horizontal 1-4 5 IV- 4"'9 2.75 10%(.150) Fracture 54.68 1.345 6

l 3 6H3 Horizontal 1-4 5 III-&9 2.75 10%(.150) Fatigue 54.68 1.51 6H4 Horizontal 1-4 5 IV- 4"'9 2.75 10%(.150) Fracture 54.68 1.51

!, 4 8H1 Horizontal 1-7, 1-9 5 III-4" 1.75 10%( 150) Fatigue 53.48 1.54 7

} 8H2 Horizontal 1-7, 1-9 5 IV- 4" 1.75 l')%( .150) Fracture 53.48 1.54 "

l l NOTES:

l 1. Fracture and fatigue susceptibility checked for worst load combination j 2. Maximum at end of design life j 3. Taken from CD-130 1

4. Taken from CD-139 l 5. Check using base plate growth rates and toughness ,

t ,

l 6. Check using weld metal growth rate and toughness properties S

7. K Ic taken to be 82.6 ksi/T5 for weld metal, 73.2 ksi/tw for base metal i 8. Governing case is load combination III l 9. Note true applicable stresses will be in "z" direction and thus substantially lower than longitudinal i stresses used as bound

j 7-3 o

4 A deeper def ect was found during the enhancement work. This led to the

!l development of additional computer runs 1H5 and 1H6 using a bounding defect 1.125" long and .214" deep. The resulting value for K was less than a

that for the longer flaw (57.5 ksi in versus 60.0 ksi in).

7.1.2 Weld 1-4

, The plate imperfection of weld 1-4 was modeled using stresses from weld joint 5 (CD-139). Since the location of the imperfection could be af fected by either weldment or base plate properties, both possibilites were run.

l Case 2, runs 6H1 and 6H2, use base plate properties. It should be noted that the actual orientation of the defect is such that stresses to_af fect

!- it would be the through-thickness stresses which are very small (CD-139).

l '

For purposes of this bounding analysis the flaw has been modeled as if oriented normal to the maximum stresses in the longitudinal defect. The results will thus be very conservative when obtained. As with the Imperf ections in seams 1-1 and 2-1, the combination of stresses and flaw size modeled predicts structural f ailure will not occur for either set of

• material properties.

7.1.3 Welds 1-7 and 1-9 a It was noted that weld 1-7 is at worst a 9/16" long slag line. In weld 1-9, an excavation for repair was made and that area is not covered by RT to assure repair was completed. Attachment 3 in CD-139 reasons that a bounding case of a flaw 13/4" long will bound both 1-7 and I-9 defects.

Case 4 runs 8H1 and 8H2 have been established to evaluate this bounding flaw taken to have a depth equal to .150". The stresses of joint 5 have been used in this evaluation although the true stresses in the upper parts

~

of the containment near joints 1-7 and 1-9 are substantially lower (compare CD-139 attachments 2 and 6). Even with the conservative bounds, this model predicts structural failure will not occur.

7.2 Limit LQad A_nalvsls As discussed in Section 2.0, limit load provides a bounding method of e

_a**- -*P--,--**--**7 *

• M**f e t A te p4-* s > p . g ,_w_ g_+s..sg yg,_. g '*7'*_-

.T 7-4 L

i -

k- . analysis for structural failure. It is based on two theorems. The.first

., gives rise to lower bound solutions and states that a structure will not

, fall if the applied forces can be balanced by a redistribution of stress

, such that the Induced stresses do not exceed the yield or flow stress. The second theorem, which is an upper bound theorem, states that the structure will. collapse when the rate of external work done by the applied forces exceeds the rate of Internal plastic work for any collapse mechanism. As the first theorem provides lower bound results, it will be used in this analysis.

Many solutions have been developed to calculate the critical stress for various geometries and loading conditions using the lower bound theorem.

Sevaral of these are reported in (7-1). Foracenter-crackedplateuider uniform tension (Figure 7-la): #

o = a-(1-2a/t) (7.1) c 4 and for a single edge-cracked plate in tension (Figura 7.1b):

1 .

l (7.2a) c" Ot (1-a/t)[ 1+1n and _1-a/2t}]

1-a/t ; for a/t 5 0.884 "c = 2.571og (1-a/t) for a/t > 0.884 t

i (7.2b)

These solutions are for infinitely long flaws, which provide overly conservative solutions. The ef fect of flaw aspoct ratio may be considered by replacing a with:

l'

) 2 2 a(1-(1+1/2t 2

)-I )-

2 1-a(1+t /2t )-ijt (7.3) i<

' For the present case, o has been taken to be:

l

-f

'I

~

.. .- r .

, . . n . .. . . .. ...- - , .. a . . . . . ,

4

g

~ 2a, _,

Y, u

i

^ t

a. Center-cracked Plate .-b. Single Edge-cracked Plate

i 5., p1 Figure 7-1 Flaw Geometries for Limit Load Analysis

,i 7-6

. a

'7 i

~~

o = 1/2(o + o-) (7.4)

J y u Where o and o are the minimum values of yield strength and tensile strengtk. The minimum limit stress In E7018 can be taken to be 66 ksi

,; which corresponds to the specified minimum tensile and yield strengths of 72 ksi and 60 ksi respectively (taken from AWS A5.1-78 " Specification for Carbon Steel Covered Arc Welding Electrodes" (2:2).

Simplification to the governing equations occurs in this case because any bending stresses applied to the structure will appear as a un! form stress across the section. Further, the size of the detall is suf ficiently small so that a crack will not af fect the integrity of the structure by net _ loss _

of sectional area. Limit load will thus indicate f ailure when the aphlied stress in the net ligament is equivalent to the maximum stress which the i ,

! section will support, or the limit stress. The critical area loss to cause l

l Ilmit load failure was determined from the maximum appiled tensile stress.

The maximum design tensile stress is found to be in Joint 3, load combination IV, longitudinal stress component o , inside surf ace and is equal to 33,136 psi (CD-139).

This worst case stress will only result with a section ligament loss of l 49.8% . Since the maximum depth defect is 14.35 the ratio of worst defect depth to critical depth is 3.5. Thus, there is a large margin against failure by a limit load mechanism.

7.3 Elastic-Pinstic Fracture Mechanics (FPFM ) Rasults To assure that the linear elastic fracture mechanics analysis was conservative, an EPFM analysis was completed utli tzing the concepts

_ outlined in Section 2.1.2. The basis for this~ analysis is the "Draf t British Standard of the Rules for the Derivation of Acceptance Levels for Defects in Fusion Welded Joints" (7:1). Since a rigorous elastic plastic analysis is very complex compared to LEFM, the Draf t Standard has a

s

--,-,,,-e,.--,n - -,-, - . - , . - - - - , s -

7-7 9

[ simplified the procedure by using a semi-empirical design curve set in its current form by Burdekin and Dawes (2=d). Experimental work was performed

[, by Burdekin and Stone LI:5). For a defect in a uniform stress field, the

, relationships between 000 and applied strain are:

$ = (e/ey )2 for 0 < e/ey < 0.5, (7.5) y <2

$ = e/e - 0.25, fo- 0.5 < e/e y

where 4 is the non-dimensional 00D, i.e.,

$= 6 - --

(7.6) O ,

2ne,.a The 00D design curve, shown in Figure 7-2, relates the non-dimensional 00D, 4, to the ratio of applied strain to yield strain. The applied strain is taken as the local strain which would exist in the vicinity of the crack if the crack itsel f were not present. It has been shown (2:6) that this design curve is conservative, and thus the allowable flew size, a, will be a smaller than the critical flaw size, a' .

crit The design curve is based on a through-thickness defect (a =

through-th ickness crack hal f-length). Although not rigorously justified, It has been suggested that the ef fect of crack shape is the same for contained yielding problems as for linear elastic problems (2-1). Thus, a pseudo-elastic plastic solution to part-through cracks can be developed.

For a through flaw with the gaometry shown in Figure 7-3a, the LEFM expression is:

K = ovia" g

For a surf ace crack with the goemetry shown in Figure 7-3b, the LEFM

• ~v. __ , u ., ,  ? ,, - r --,--w

L .,

w.w 6

_ a,

. [ 4

)

. 6 7

f s

I 3 e R

m

_ o r

f ,

re

" (

T

_. e

.

• v r

u C

n g

i s

e

. w D r D u 0 c C s

Y n

g t

2 e i / '

s e e 2 D -

7 0

0 e C r g

u

_ i F

i t 1 3 , I

_ :6

' , e

, Ii;

t! .

l...

. ?w

_ ~

i

~.

l M = s

- a w

~ .

1 h l F

< d t  : e N)B

, i a

r r 2 u h >

- c 4

= - < 2 c

_. y r

t

= e 5

}

e w Eo

. a l * ,. G F

t e c c e t a f

= f e h r D u

S

= 3

) 7 c a b e

=

r u

c F g

2 i F

~

~

1. h

= w a

l a F m2 = -

g u

o vTh r

v)a i b

7-10 s

'P P

expression is:

1 K

g

=

YS cha (7.8)

$2 4

where M T is the magnification factor due to finite th!'ckness ef fects, M is the free surf ace magnification factor, and & is the elliptic

' 2 integral of 1he second kind. It can be seen that K = "3 i surface flaw $2 K lI (through-thickness flaw ] (7,g)

, .i ,

From thIs relattonshIp, solving for an equiyalent through-thIekness erack size gives:

. [ y1 3 ',

a a l I"E ( '2 ) (7.10) t The values of s= f ~ ~ ~ ~

ere taken from a survey by Maddox (2-8), and j are shown in figure 7-4.

It can be similarly shown that for a buried elliptical crack with the gecrnetry shown in Figure 7-3c:

/

IMg M)2 i *i (7.11) l where M is the magnification factor at a point due to the nearest free surf ace and M,Is the magnification factor at that point due to the more i

remote free sLrface. These values were derived from the finite element 4

~

. ~ ~ - - n -- n= ,. ,

=

7-11 1i so , ,

• s s ' '

~.

• / c

• 0 o y 1t =i

]; I

,/

, 40 ~

. i.

3

~

3C# .

.i NN g g

'CeeJt

-C 5-

..A 20 ~

\ '

l i

%ce01 2C-

. we.e u

&ec 9

(

I I (-

i i o , e i i , , ,

o os 02 0-3 04 os os o1 as t

i a Figure 7-4 Variation of (M3g/4) with Crack Depth for Various Shapes (from Ref. (7-7))

e. .

i N , er 4 e=h ee + e .=

=u_ ._e

+ * + - + - + * *

=_ ===-^e * --- * = = -

7-12 y

i work by Shah and Kobayashi-(2-2), and M 'and M,are shown in Figures 7-5 or o and 7-6 respectively. The elliptic integral, 4 , is plotted in Figure 7-7.

2 For a/c = - 0; Mou M was derived from Fedderson's relationship (2-1.0)

n

". f i 1'/

~ ~7 M = 1 sec wa l

, ( tj 4

Thus, combining the C00 design curve with equations 7.6, 7.10, and 7.11, a relationship can be developed between allowable flaw size and allowable strain levels f or each crack conf iguration. The Draf t Standard contains criteria which require that when a crack tip approaches a free surface, the j

flaw should be reassessed as a surf ace-connected flaw (a buried flaw is '

l recategorized as a surf ace flaw and a surf ace flaw is recategorized bk a through-thickness flaw).- The buried flaws present can be conservatliely modeled as surf ace flaws and in the present case, on!y such surf ace flaws

, and through flaws have been considered. -

The critical value for crack opening displacement was taken to be 0.023" es outlined in Section 5.4. The yield strain is assumed to be 0.2%. The calculated allowable flaw sizes for surf ace defects are shown in Figure 7-8 for 1.5 inch thick weldments.

The theoretical development used is valid to strain ratios (e/o ) up to Y '

)- 2.5 but are plotted beyond that point to demonstrate that at flew depths of i

approximately 10% of the wall, a very large margin to fracture exists. The results are shown for a variety of aspect ratios. For e/2c = 0.0, an infinitely long. flaw is modeled. The theory provides a recategorization

,, process at a depth of 50% of the wall thickness. The flaw should then be considered as a through-flaw.

[ Figure 7-9 shows the results of the elastic-plastic through-flaw analysis.

The length of a critical flaw as a function of strain is given. For through- flaws, residual- strains are not included.

9

-,.-g # -m-- v, . -,-- . , , , ----y m-,, ,-,,.y,.m. .----.,.-.---......m - v.- -,,-,-,.------m,---m.----_ -

\ - --

.- c . v. c. .. ~,

-MAGNIFICATION FACTOR FOR NEAREST FREE SURFACE 2.0 , , , , , , , ,

I,,,, , , , , , , , ,

a/c -

1.8 0.10 ' -

1.6 M -

0.20 _ y N

M

_ L w

1.4 0.40 -

0.60 _

0.80 -

1.2 1.00

~

~

' ' ' ' l l 1.0 ' ' ' ' ' ' ' '

O 0.2 0.4 [B 0.8 1 e/h G

Figure 7-5 e c

. s . _ ... . . .. .

. . . . . -...a.

MAGNIFICATION FACTOR FOR FARTHEST FREE SURFACE 1.20 .,,, , , , ,

, , , , ,, , . i , , .

j a/c-1

1.15 0.10 O.20 _

y 1.10

.y N -

M 0.40 - -

1.05 O.60 _

0.80 1.00 -

z 1.00 s --_ ... ,,,,I, , , ,

~

, 0 0.2 0.4 dI$ 0.8 1 l c/h1 i

O Figure 7- 6 0 e

i

-.,,,,.s..

, ,. ... .c,.... v-,,

ELLIPTIC INTEGRAL OF THE SECOND KIND 1.6 _

1.5 _

1.4 -

N 1.3 9

_ u m

l 1.2 -

,1 _

1.1 2

' ' ' l ' ' ' ' ' ' '

1.0 ' ' ' ' ' ' ' '

O 0.2 0.4 ';.6'B .. 0.8 1

(.

blC i

j Figure 7- 7 e t

i

. . . . ... 2 - . .

STRAIN vs. CRITICAL FLAW SIZE (SURFACE FLAW) 5 . . . , , , , , . . , , , i , , , . . i , .

_ l l l l _

_ c/2c-0.4 6c -0.023 in. _

t =1.500 in. _

4 i i - _

a/2c-0.3 _

3

[ _ _

y

) -

a/2c-0.2 -

E 2

l _ _

a/2c-0.1 1

[

l a/2c-0.0 t

j i t I 1 1 l I I I I I t i i t I t i t i i I I I O 0.25 0.5 0.75 1 1.25 1.5 j FLAW DEPTH, INCHES Figure 7-8 Strain versus critical flaw size for 1.5" thick weldments D

- --- =-

.;. m..  ::

.. . . , .. a v..

STRAIN .vs THROUGH-THICKNESS CRITICAL FL AW: SIZE 5 , y . . . . i. . , i , , , , , , , . . .

_ so -o.023 in. . ._

4 3

x Z -

- y e

_ c 2

1 O ' ' ' ' ' ' ' ' ' ' ' '

O 2 4 ~6' 8 _O FLAW LENGTH, INCHES Figure 7-9 Strain versus through-thickness flaw (EPFM)  ;

e

IA ! - 7-18 r

For the present case, the ratio of maximum applied strains to yleid strain

.; Is approximately 0.5 so that long through-flaws are predicted prior to t

.' fracture by EPFM methods. Most relevant cases have lower strains and correspondingly larger predicted allowable flaw lengths.

I i .

, Two notes-are of interest here. First, the toughness used was taken from

[ ,

generic data and not plant specific Information. Thus ths' flaw values

( should only be used to gauge the reasonableness of the LEFM approach as was their Intent. Second, this EPFM approach has an inherent minimum factor of j safety of 2 on flaw size. G-11) i w,

l l

1 ,

I l  ; 7-19

't 5

Section 7 I REFERENCES 1 .7-1 Chel l, G.G. , " Elastic-Pl astic Fracture Mechanics," Central Electricity Rasaarch I mhnra+nr f as, Report No. RD/L/R 200F, (August 1979).

7-2 American Weiding Society, AWS A5.1-78, " Specification for Carbon Steel Covered Arc Welding Electrodes", 1978 i

7-3 British Standards institution, "Draf t Standard Rules for the Deriv-ation of Acceptable Levels for Defects in Fusion Welded Jointh".

"~

WEE /37, February, 1976.

. ~7-4 Burdekin, F.M. and Dawes, M.G. " Practical Use of Linear Elastic and Yielding Fracture Mechanics with Particular Reference to Pressure Yessel s". Ins +f +u+1on cd Mechanical Fngineers C5/71.

7-5 Burdekin, F.M. and Stone, D.E.W. , "The Crack Opening Displ,acement Approach to Fracture Mechanics in Yleiding Materlafs". Journal of Strain Analysis, Vol .1, No. 2, (1966), p.194.

7-6 Harrison, J.D. et al, "The 00D Approach and its Application to l'

Welded Structures". lha Waldinn In=+1+u+e Rasmarch bptqri 55/1978/E, January (1978).

t 7-7 Dawes, M.G., " Fracture Control in High Yield Strength Weldments",

Welding Journal Research Supplement, Vol. 53, (1974), p. 369s.

7-8 Maddox, S.J., "An Analysis of Fatigue Cracks in Fillet Welded

! Joints", I n+arnn +I ona l Janenal sd. Frac +ura Machanics, Vol. 11, No. 2, April (1975), pp. 221-243.

l r

g----. ~-

~v

,.m..

,--n, ,, n,

-m

--,.,,-v,-

- - + . - - , - --- - - ,,, ,-,,,,-, ,,- n~~*

n,,-- -

w,-- w e.e- , , , , ,,

4 7

1 7-20 r?

9 4 .

7-9 Shah, R.C. and Kobayashi, A.S., " Stress Intensity Factors for an j Elliptical Crack Approaching the Surf ace in a Semi-Inf inite Solid",

i n+arna ti ona l Journal ni Frac +ura Machanics, Vol . 9, No. 2, (1973) p. 133.

i i

+

7-10 Fedderson. C.E., " Discussion to Plane Strain Fractur6 Toughness Testing". ASTM STP 410, (1967), p. 77.

i e

7-11 Kamath, M.S., "The C0D Design Curve: An Assessment of Validity

_ Using Wide Plate Tests". The Welding institute Report 71/1978/E.

September, 1978.

h t

i 1

T I .

y s

a V-

'?

g 8-1

?

Section 8 1 CONCLUSION AND

SUMMARY

i 9

This report sunmarizas the results of an evaluation of the ef fect on

, structual integrity of several weld imperfections. These imperfections were assessed from Indications in radlographs and have been analyzed in separate regions as follows:

e inaccessible defects in weld seam 1-1 and 1-2 e Potentially rejectable inaccessible regions in weld 1-4 l'

e Potentially rejectable inaccessible regions in welds 1-7

~

and 1-9 It has been found that assumed bounding defects of each category will not cause structural failure over the plant design lifetime. This analysis has included the ef fects of:

e Flaw growth by a fatigue mechanism o Residual stresses e Worst case applied stresses (steady state plus cyclic) e Maximum defect sizes Throughout the analyses, conservative estimates of input data have been used. These are sunmarized as follows:

e Methodologies. The LEFM method used has been shown to be O

t

, ~ - . - - - , - , - - - ~

--,,--r-, , ~ -m ~ -nc--- ,-v--

8-2 7

conserystive in this analysis as compared with both j elastic-plastic f racture mechanism and limit load methods.

The LEFM calculated final flaw conditions do not account for the increased ductility that will be available at operating

temperatures.

e Residual stresses. Bounding values from distributions published in the open literature for like weld details have been used.

e Flaw location in the through-thickness direction e Flaw geometries. In all cases semi-elliptic surf ace flaws have been used to bound buried and near surf ace flaws. Manyoffhof-flaws thus bounded may be buried, in which case, substantially ~

, larger margins against f ailure occur than those listed.

l e Crack growth rates. Upper bound crack growth rates for E7018 weld metal were developed and used to predict the snount of of crack growth.

l e Choice of applied stresses. Bounding stress cases (including

[ numbers of cycles) have been used.

e Material toughness. Material toughness values derived from 00D data were shown to be conservative. In addition, data was collected at -20'F. which is much lower than the expected service temperature.

l e Limit strength of the material. Values were estimated that result in conservative estimates of remaining Ilgament.

l l

l .

4 s

, n

.np- - y.-gm--n- 7-----,,w, - < - .-- ,n--w--, ,-. -

, A-1

')

Appendix A SUPPLEMENTAL TOUGHNESS DATA t

?-

1 1

e i

I l

j. . . , ... .-. - . . . . _.-.

..~, - - , .-

., . . . . . .- - 7 . , ,

m  ;, , , _ , ,

i APPENDIX A CRACK OPENING DISPLACEMENT (C0D) VALUES - E7018 WELD METAL TEST TEMPERATURE C00 DIAMETER (oF) (in.) (inches) C0FNENTS REFERENCE I

l -20 .0182 5/32 Lincoln LH-70 I Ductile Ductile

-50 Ductile Ductile

-100 .0107

-20 .0133 5/32 Lincoln LH-72 I ,

.0192 4 Ductile

-50 .0018

.0113

-100 .0011

+14 .0224 1/8, 5/32 Lincoln LH-72 I

.0118

.0209

-76 Tearing 10nn2 C00 specimens for E7018 weld metal II

-76 >.030 1 in2 C0D specimen. E7018 weld' metal

\ mj-l -

\

b

_ ~

m .

N ,. i ._. . e . .s APPENDIX A (Continued)

TEST TEMPERATURE C0D (OF) (in.) DIAMETER COMMENTS REFERENCE

-94 10 values ranging British equivalent to E7018 III' from .008 .034 '

R = 20.00 s = 10.01

-40 22 values ranging British equivalent to E7018 III from .016 .048 R = 34.45

$ s = 9.246

?

32 22 values ranging British equivalent to E7018 III

'd

. from .023 .052 1

i = 39.68

s = 7.305 s

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i 'l APPENDIX A

., REFERENCE LIST FOR FRACTURE TOUGHNESS. DATA - E7018 i

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,' I. Personal correspondence between W. McNaughton (Aptech Engineering Services) and R.C. Shutt (The Lincoln Electric Company), Dated i May 22, 1979.

II. Dawes, M.G. , " Designing to Avoid Brittle Fracture in Weld Metal,"

Metal Construction and British Welding Journal (February 1970),

Pp. 55-59.

f.

III. Tait, P. and D.M. Haddrill, " Fracture Toughness of Some Mild' Steel Manual Metal-Arc Weld Deposits," Welding and Metal Fabrication (September 1970), Pp. 370-375.

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. Appendix B BACKGROUND INFORMATION ABOUT DIGITAL IMAGING TECHNIQUES 1

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{ g[ [ DIGITAL lMA21N3 SYSTEMS 795 SAN ANTONIO ROAD

• PALO ALTO

• CALIFORNIA 94303 (415) 858

• 2866 4

I INSPECTION SIGNAL ENHANCEMENT SERVICES FOR THE NUCLEAR POWER INDUSTRY Aptech Imaging, Inc., has developed unique digital signal enhancement schemes which are applicable to a wide range of inspection signals. Three particular applications of this technology have been developed as follows:

Enhancement of existing inspection information, such as x-rays, radiographs, or videotapes.

Real-time x-ray imaging systems or filmless radiography.

_ .. ~

i Records management to store and retrieve in digital format, radia* graphs,

} engineering drawings, etc.

• These enhancement techniques have already had wide application in the nuclear l power industry to aid resolution and interpretation of inspection signals. We l- have outlined below some relevant experience in the nuclear power industry.

i l' CLIENT SERVICES PROVIDED Southern California Edison Enhancement of indications in radio-San Onofre Generating Station Unit I graphs of welds in main steamlines Boston Edison Company Enhancement of images from underwater l

Pilgrim I television cameras of core spray sparger.

i This work was performed by enhancing existing video tape recorded in 1980. We

{

performed on-site real-time enhancement j in conjunction with the inspection at the 1981 outage.

Gilbert Commonwcalth Asspciates/ Enhancement and interpretation of Cleveland Illuminating Company radiographs of containment welds.

Perry Nuclear Rolls Royce Nuclear Ltd Enhancement and interpretation of old radiographs.

,, General Electric Company Enhancement and restoration of out-of-specification radiographs for cast nuclear valve bodies.

l-I Digital truaging Systems for inspection, Records Management and Signal Enhancement f

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Nuclear Regulatory Commission Enhancement of radiographs of stress-corrosion cracks adjacent to welds in Type 3D4 stainless steel.

l Bechtel Power Corporation Demonstration of digital records management system to store and retrieve radiographs and engineering drawings.

J Battelle, North West Demonst. ration of enhancement techniques for underwater viewing of reactor internals.

Consolidated Edison Company Radiographic enhancement for defects and restoration of old radiographs.

General Electric Company Enhancement and interpretation of defects in radiographs of nuclear plant valve bodies.

GPU Nuclear Enhancement of video tape inspection records of core spray sparger.} -

Enhancement of video tape inshection records of core spray riser piping and reducer.

On-site real time enhancement of video signals from core spray sparger and annulus inspection.

Northern States Power Enhancement of radiographs of stress corrosion cracks in primary piping.

Power Authority, State of New York Enhancement of original fabrication radiographs from steam generator closure weld.

[ Bechtel Power Corporation Determination of defect dimensions

( South ~ Texas Project (length and depth) in emergency cooling water lines.

Duke Power Company Enhancement of radiographs from pipe to valve welds.

North East Utilities Enhancement of piping radiographs from Millstone III.

Further details about these services and individual utility contacts can be provided to you on request.

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CONFERENCE "hactura Toughness Tcsting -

0 M AUTHOR (S)

• '" '" ' '"' * """* '*'* " ' ^ "" " " As receive d G .R.Esran , M .F.Elgart & A. A. Smith iate type script

i IMPROVED RADIOGRAPHIC FLAW SIZING BY DIGITAL IMAGE PROCESSING G. R. Egan,* H. F. Elgart,** and A. A. Smith *"

~

I 1.0 SYNOPSIS This paper describes the calibration and application of a digital image processing method to determine the depth of indications seen on radiographs. The image of the radiograph is digitized and then computer manipulated to provide accurate measures of film density. Indications on the radiograph can then be interrogated I

to determine their extent in the through thickness direction. By comparing numerical measurements of density of indications with known density changes from image quality indicator wires or the plate thickness, the depth of indic~ations can be determined.

A calibration procedure is also' described wherein known defects were introduced

(

into a steel sample which was then radiographed. A comparison of the digital

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processing size estimates and the actual dimensions of the defects shows excellent correspondence. An example of the application of this technique to lack of root fusion defects is also described.

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eco Consulting Engineer

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2.0 INTRODUCTION

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Fracture mechanics principles have been used to characterize a wide range of f ailure mechanisms including brittle fracture, ductile fracture, fatigue, stress corrosion : cracking, Md also weld cracking phenomena. The analysis required to characterize any of these crack extension mechanisms is based on determining a material property (e.g., toughness ia the case of fracture problems), the acting 1 >

, stresses that can cause crack extensioncand calculating the critical flaw size for crack initiation or continued crack propagation. In any of the analysis methods that have been developed there are uncertainties in the input data that are used. For example, in the very simple case of assessing.the critical flaw 4

size to cause. brittle fracture, scatter in data of K tests, uncer_tainty in the IC definitionofstressesincludingresidualstressesandtheerrorsassbciatedwith flaw sizing techniques mean that when critical flaw sizes are calculated, margins of safety must be established before the results can be applied.

i

, , lSince, in most cases, there are insufficient data to perform a full probabilistic

' analysis of the problem, it is necessary to choose bounding values of the input j

data to fracture mechanics equations so that the resulting calculated value of critical stress or critical flaw size is regarded as a conservative estimate.

Pursuing this procedure-to its logical conclusion often results in such J pessimistic predictions that currently operating equipment is deemed to be in i

jeopardy from some predicted failure mechanism Q). It may well be the case that f' operating limits are so restricted that major economic penalties are incurred by i performing so-called " conservative analyses." One answer to this dilema is to perform the complete probabilistic analysis to establish the probability of l , failure by the mechanism that is described by the deterministic fracture mechanics equations (_2_). Having derived this number, however, there is usually l .

uncertainty in establishing the significance of such numbers @) and comparisons of risk between different events can be made to judge the comparative significance of an outcome.

By looking at uncertainty in the input parameters, it is possible to establish a l ranking of importance of input variables to any fracture mechanics analysis. For example, by the very nature of the equations that describe the interrelated

_- . _ -. o I

!!' 3 7 variables, we are able to establish a ranking in order of the importance of l particular input parameters. With this ranking, we can then concentrate on refining methods for decreasing the uncertainty in any set of input data.

t The importance of establishing the initial flaw size can be determined by studying the equations which describe some of the fundamental failure mechanisms.

Table 1 outlines the relationship between calculated critical flaw size and the measured flaw size for three different failure mechanisms--fracture, high cycle fatigue, and stress corrosion cracking in sensitized Type 304 steel (4). It can be seen that the importance of flaw size dimensioning increases as the exponent of K or AK in the equation that describes the failure phenomenon. This has also i

led to the establishment of the half life concept for both fatigue and stress corrosioncracking(5).

In this paper, we concentrdte on developing methods to improve sizirig of defects discovered by either x-ray or radiographic techniques. We anticipate that this work will lead to better and more reliable methods of flaw sizing.

I Although much research is currently being undertaken to improve the reliability of inspection systems, it is clear that a key element that contributes to the 1-l uncertainty is the presence of a human interpreting and recording information.

The work that we describe following is aimed, in the long term, at providing j automatic pattern recognition systems for flaw detection and sizing.

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( IMPORTANCE OF FLAW SIZE IN FRACTURE MECHANICS ANALYSIS FLAW SIZE FAILURE HECHANISM GOVERNING EQUATION DEPENDENCE K

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Brittle Fracture o = I 5 a

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Fatigue N = I 1-a}2 g - ?y

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m-2 Stress Corrosion Cracking -

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, of Sensitizer! 304 T = o ,m ,x a j a 1

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,, 3.0 DIGITAL IMAGE PROCESSING TECHNIQUES l

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, 3.1 Background

j. The digital image processing techniques that we have developed are described in

_ detail elsewhere, however, a short sumary of the basic background technology will be provided here (6). A diagram of the equipment that is used to capture images in digital format is shown in Fig.1. The image collection system (i.e.,

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a television camera) can deal with signals from x-rays, radiographs, pictures, and printed pages. Once the information is digitized and stored in the computer, b manipulation by mathematical methods of analysis can be performed to provide

! images that can be interpreted by the non-expert. The APTECH system shown in Fig.1outputsinformationtoatelevisionmonitorfromwhichhardi:op)canbe obtained by normal photographic means. In addition, copy can be provided on either magnetic tape, video tape, or video disc.

3.2 Treatment of Radiographs The primary advantage of radiography as a nondestructive testing technique is associated with the fact that an image is involved. The use of an image enables both expert and non-expert to interpret effectively the meaning of the test results. However, because of his experience, the expert radiographer will usually extract more information from an image than a non-expert will. The amount of discernible information is unfortunately biologically limited.

Although film contains sufficient information to detect density differences of 0.05% to 0.1%, the human eye can only resolve grey levels which differ by at least 1.5%. Therefore, the film has captured much more information than the eye can extract.

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Additionally, the eye discerns a boundary or edge condition only when two

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adjoining areas of an image differ by more than 15% in density. The full range of film density information can be made available to the observer by use of digital techniques of analysis. Small density differences not discernible to the unaided eye can be made visible on a television monitor by expanding a small density range on film to the full white to black grey level information. Then

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I! 7 since the image is in discrete digital form, mathematical methods can be directly applied to provide image enhancement. Images are currently captured using a 512 x 512 matrix, and each pixel in this matrix contains grey level information.

{ Grey level infor: nation or radiographic density is provided on a scale from 0 to approximately 4000.

m 3.3 Procedures for Depth Measurement of Indications Since the computer can discern density differences from white to black, on a grey level scale of 0 to 4000, the sensitivity on depth measurement in a full grey

. scale is 1/4000 of the depth of the piece being examined. Furthermore, for defects that lie normal'to the plane of the radiograph, it is possible to

+

, establish a density or defect size profile using information derived from image quality indicators or the plate thickness. For example, since the topputer provides such a sensitive measure of density, it is possible to calibrate a radiograph by using density changes associated with image quality indicators.

Details of this procedure are outlined later.

In addition, for dimensional verification, for establishing the dimensions of radiographic indications in the plane of the radiograph, it is possible to s

perform automatic integration schemes that will yield defect length and width.

The sensitivity of these measurements is 1/512 of the extent of the image being processed. Since mathematical magnification models are available, it is possible

e to develop the required sensitivity in the length and width directions. An example of dimensional verification by computer is shown in Fig. 2.

The foregoing description leads to the natural conclusion that we are able to i

plot out in three dimensions, defect profiles. From a single radiograph, it is possible to do this; however, one important piece of information is missing and that is the location of the indication within the thickness of the part being examined. The procedure has certain limitations related to the eccentricity of buried defects, however, for surface connected defects, the procedure can be used in a straight forward manner. Some connents are, made later on how the procedure may be changed to establish positional information rel' a tive to the two free s'urfaces of the parts being examined.

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2b. Dimensioning by Computer.

n, Figure 2 - Example of Dimensioning by Computer. ,

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] 3.4 Calibration To determine that the procedure does provide accurate measures of defect depth t

into a plate from a single radiograph, we have performed a calibration procedure using blocks containing defects of known dimensions. Details of the calibration i blocks are shown in Fig. 3. For convenience, we shot the calibration blocks in real time and captured the information on videotape which was subsequently digitized and processed. The image could also have been collected on film.

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. Once the image had been collected on videotape, the information was presented to the computer in digital format. We were then able to establish measures of defect height through the thickness for the range of defects contained in the blocks (fig. 4). The defacts consisted of machined notches and drilled holes, and these l were also measured using gauge blocks. A comparison of the two methods of j, measurement is shown in Table 2. Itcanbeseenthatasmalluncertajntystill exists in the dimensioning of defects by this method, but it is significant i improvement on existing methods of determining size information by nondestructive means.

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RESULTS OF CALIBRATION - BLOCK 4 l

Digital Information Estimated Depth Measured Depth Defect Full Plate Defect (inch) (inch) 451 2550 3390 0.107 0.110 2775 452 2775 3380 0.800 0.075 2900

-t 4S3 2900 3370 0.068 0.060 2925

=

6-4H1 2200 3260 0.162 ..J 0.113

i. 2200 i 2200 2200 4H2 2350 3050 0.125 0.098 2350 2250 2200 2335 l.

{ 4H3 2400 2870 0.086 0.067 2400 l 2350 I

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5 4.0 APPLICATIONS t

, ,  ; In this Section, we outline a specific application of this technique to

,, indications derived from radiographs of offshore platform welds. In each case, we

have determined a calibration coefficient from the image quality indicator and the

! part thickness and used this to establish the through thickness extent of

)'[' indications in the radiograph. These dimensions can then be compared with fracture mechanics calculations to support continued operation of these structures.

4 (f 4.1 Offshore Pipeline Welds f.. In this particular example, we were supplied with radiographs of weldf from the

j. main legs of fixed platforms located in relatively shallow water (174T, less than

! 100 feet). The radiographs were generally of poor quality and out of focus p because of the large object to film distance. We processed the radiographs by j' applying an artificial focus procedure which essentially consists of squeezing the i digital information down until the picture becomes focused. The results of this 1 .

procedure are shown in Fig. 5. After this procedure was completed, we then 1 interrogated the image quality indicator wires to establish a unique calibration f for each radiograph. Having done this, we then established the depth of j indications that were apparent in the root of the welds. These were surface lH connected defects so that the calibration procedure was straight forward, and the

, missing information for through thickness position was not necessary. An example

'l of the procedure used is shown in Fig. 6.

j ,

From Fig. 6, it can be seen that not only can depth information be established, buttheprocedurealsoprovidesinformationaboutthetypeofindication(i.e.,

rounded or linear). The techniqu' e has also been used for defect identification.

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j . 5.0 CONCLUDING REMARKS

t It would be unrealistic tc conclude without a few cautionary coments. First, the technique that we have developed can only be used on images that have not been enhanced using non-linear transforms. Obviously, in such enhancements, where particular parts of the grey scale are compressed, a calibration on one part of

,, the radiograph would not apply at other locations on the raciograph. Second, where the indications are smaller than one pixel (1/512 of the image), unique depth information cannot be determined. This is not really a practical limitation '

because it is possible to capture a smaller part of the radiograph.

I I~

In addition, from a single radiograph we cannot determine the location, of the e

.t indication in the through thickness direction. It is necessary to ta b one otFer

' I shot at an angle to ~oe able to reconstruct all of the positional iniormation.

4 Even 1.n spite of these limitations, we have found that for surface connected

, defects there is a need for an accurate method of estimating depth of indications.

The work is continuing with the objective of developing an automated process for providing defect dimensions and positions.

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REFERENCES q

< . 1. Snaider, R. P., J. M. Hodge, H. A. Levin, and J. J. Zudans, " Potential for Low Fracture Toughness and Lamellar Tearing on PWR Steam Generator and Reactor Coolant Pump Supports. Resolution of Technical Activity A-12," NUREG-0577 for Comment, Office of Nuclear Reactor Regulation, U.S. Nuclear Regulatory Commission (October 1979).

2 Thomas, J. M. and D'. C. Peters, "Probabilistic Decision Model for Structcres Subjected to Crack Growth and Fracture," ASME Winter Annual Meeting, San

Francisco (1978).

3. C. Starr, " Benefit-Cost Studies in Socio-Technical Systems," Colloquium on Benefit-Risk Relationships for Decision Making, Washington, D.C. (April l

1971).

i

4. - Egan, G. R. and R. C. Cipolla, " Stress Corrosion Crack Growth and. Fracture Predictions for BWR Piping," ASME, New York (1978). ~ f.

. J..

5. Hayes, D. J., " Fracture Mechanics Based Fatigue Assessment of Tubular Joints.

Review of Potential Applications," AES 61-01-45, Shell Oil Company (January

1981).

6. Elgart, M. F. and R. H. Richman "Real Time Digital Techniques for Radiography," Proceedings of Workshop on Nondestructive Evaluation of Turbines and Generators, EPRI WS80-133 (July 1981).

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4 Appendix C DETAILS OF THE FLAW CHARACTERIZATION WORK Y

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SUMMARY

OF RADIOGRAPHS REVIEWED INACCESSIBLE WELDS l

JOINT STATION DEFECT LENGTH (Max) DEPTH (Max)

Unit 1

, 1-1 (21-22) 4-5 Elongsted Slag 0.75" 5.8%

1-1 (21-22) 5-6 LOF 4.0" 11.0%

1-1 (21-22) 11-12 LOF 1.0" 11.4%

'. 1-1 (21-22) 12-13 LCF 0.3" 10.7%

l .s 1-1 (24-25) 9-10 Elongated Slag 0.5" 6.0%

, 1-1 (24-25) 16-17 LOF 0.63" 13.3%

1-1 (25-26) 10-11 LOF 0.4" 11.1%

1-1 (26-27) 9-10 LOF 0.75" - ;; 12. %

1-1 (27-28) 4-5 LOF 0.2" .' 8.8%

1-1 (28-29) 3-4 LOF 0.6" 3.9%

1-1 (28-29) 4-5 Slag 4.5" 8.8%

0.32" 8.7%

1-1 (28-29) 5-6 Elongated Slag 1-1 (28-29) 18-19 LOF 0.5" 13.6%

1-1 (29-30) X-17 LOF 0.5" 12.5%

1-1 (30-31) 8-9 Elongated Slag 0.75" 5.4%

11-12 Elongated Slag 1.25" 10.0%

1-1 (31-32) 1-1 (32-33) 2-X Elongated Slag 1.0" 11.3%

1-1 (33-34) 24-25 No Defects Observed 1-1 (34-35) 17-18 Crack / Pore 0.5" ' 8.8%

1-1 (35-19) 1-2 LOF 0.5" 1.7%

, 1-1 (35-19) 6-7 Isolated Slag 0.19" 5.3%

~-

i , 1-4 (79-80) 11-12 Slag . 0.7" 4.5%

l l, Unit 2 2-1 (19-20) 1-2 Slag /LOF' l.0" shallow 2-1 (19-20) 3-4 Pore /LOF 0.6?" 9.0%

l 2-1 (19-20) 4-5 Slag / Pore /LOF 0.25" 5.0%

2-1 (19-20) 12-13 Slag 0.75" shallow l 2-1 (19-20) 15-16'R1 Slag / Pore 0.4" shallow l 2-1 (19-20)' 17-18 R1 Slag /LOF 0. .' 4"' , ' shallow

"-~

2-1(19-20) 18-19 Slag /LOF 0.75" 4.0%

p, 1. Distance between successive stations reoresents approximatel 12" of weld p 2. After flaw characterization, including interaction criteria.y -

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. JOINT STATION DEFECT LENGTH (Max) DEPTH (Max) n' 2-1 (20-21) 10-11 Slag 0.5" shallow

,, 2-1 (20-21) 11-12 S1ag/LOF 0.62" 6.2%

2-1 (20-21) 12-13 LOF 0.5" shallow

{

2-1 (20-21) 14-15 Slag /LOF 0.4" very shallow 2-1(21-22) 4-5 Slag 0.5" shallow

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2-1 (21-22) 5-6 Slag /LCF 0.5" 7.4%

2-1(22-23) 11-12 LOF 0.4" shallow 2-1 (24-25) 7-8 Slag /LOF 0.5" 5.3%

2-1 (27-28) 0-1 R3 LOF 0.31" shallow i

2-1 (27-28) 1-2 Slag /LOF 0.62" shallow

-1 (27-28) 11-12 Slag /LOF 0.4" rounded 2-1 (27-28) 13-14 LOF 0.5" 3.0%

2.-1 (27-28) 16-17 LOF 0.6"  ; shallow

.. 2-1 (27-28) 19-20 Slag / Pore 0.4" / ~

rounded 2-1(28-29) 12-13 Slag 0.75" shallow 2-1(28-29) 13-14 Slag 0.5" shallow 2-1(28-29) 24-25 Slag 0.62" 6.0%

2-1(28-29) 25-0 Slag / Pore 0.62" 7.6%

2-1 (30-31) 0-x Slag 0.5" shallow

,. 2-1(30-31) 6-7 Slag /LOF 0.4" shallow-2-1(30-31) 11-12 Slag 0.19" very shallow 2-1(30-31) 12-13 Slag /LOF 0.5" 7.3%

2-1(30-31) 13-14 Slag /LOF 0.5" very shallow 2-1 (32-33) . 8-9 Slag 0.25" very shallow

[ 2-1(32-33) 10-11 R2 Slag 0.4" very shallow 2-1(32-33) 11-12 Slag /LOF 0.4" 8.0%

2-1(32-33) 16-17 Slag 0.4" 10.0%

2-1(32-33) 19-20 Slag /LOF 0.5" 7.7%

2-1(32-33) 23 Pore /LOF 0.5", 0.06" 5.7%, 9.0%

2-1 (33-34) 7-8 R1 Slag /LOF 0.5" 12.5%

2-1(35-19) 0-1 Slag 0.19" shallow 2-1(35-19) 1-2 Slag /LOF 0.19" 7.1%

2-1(35-19) 10-11 Slag /LOF 0.25" 3.8%

2-1(35-39) 14-15 R1 Undercut / Slag 1.125" 14.3%

2-1(35-39) 15-16 Slag 0.5" 10.7%

2-1(35-39) 20-0 LOF 0.5" 7.7%

1. Distance between successive stations represents approximatel

2. After flaw characterization, including interaction criteria.y12"- of weld

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'I Appendix D CONTROLLED DOCUMENTS 6-

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l D-2 APTECH ENGINEERING SERVICES DOCUMENT LOG 1

CONTROLLED DOCUMENT NUMBER ITEM

] '

4 (AES-8110276) Volume 1 of 1 Containment

' Horizontal Stiffener Ring Flanges for rings 1-4 including CMTR's.

7 (AES-8110276) Volume 1 of 1 Containment

+

Horizontal Stiffener Ring Flanges for rings 5 and 6, including CMTR's. .

125 Letter PY-CEI/GAI-5519 dated 11-15-82 from W.T. Melia to R. Alley.

126 NNIC letter dated October 26, 1982 to R.W. Alley from B.R. Cofer on subject

of Perry Nuclear Power plant contain-ment vessel analysis of weld 1-1. ,,

, i.

. 127 Certified material test reports of weld . J:

1-1 between vertical welds 21-22, 128 R. Dail preliminary report to R.W. Alley dated November 16, 1982 on limited re-view of containment vessel radiographs. -

129 Longitudinal stresses weld 1-1 showing Rev. O thennal, hydrostatic, design pressure, dead load, SRV, OBE and SSE stresies.

130 Letter PY-STR-1555 information, dated November, 29, 1982, from R.W. Alley to W. McNaughton, including Rev.1 of stresses included in CD-129.

131 Letter dated 11-5-81. from B.R. Cofer, NNIC to i

R.W. Alley, GAI on subject of containment Vessel i Analysis of Weld 1-1 Indication .

' ~

l 132 NNIC letter dated 9-24-82 on prelimi-l nary summary containment vessel embednent analysis - Figures 1 and 3.

133 UT reports - inaccessible shell joints.

134 Letter to J. Keppler from D. Davidson dated 9-30-82.

135 Letter to R. Dail from W. McNaughton results of enhancement of radiographs 1-50 A/B (1-2), 1-1 (17-18), 1-1 (24-25).

F 136 Letter to W. McNaughton from R.W. Alley.

transmitting radiographs 1-1(24-25),9-10 ll-

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and 1-2 (45-46), 14-15.

l 9

==-es - -e ema sw+ = , m = . - * - " * * " " ~

i D-3

CONTROLLED DOCUMENT NUMBER ITEM 137 Lette report dated December 7,1983, 3 W. McNaughton to P. Gudikunst preliminary

results of inaccessible defect evaluation, weld 1-1.

138 BIGIF computer runs 139 Letter from R.W. Alley to W. ficNaughton PY-STR-1587, February 16, 1983 Attachment 1: Memo R. Dail to R. Alley Unit I containment radio-graphs Attachment 2: Stresses in Joints 2-6.

Attachment 3: Memorandum E.M. Horeth to B.R. Cofer, December 7, 1982 i Attachment 4: Memorandum B.R. Cdfer to M. Lastovka, December 17, 1982 Attachment 5: Memorandum R.L. Dail to R. Alley January 4, 1983

" Review of NNIC Evaluations -

of Indications in Contain-ment vessel circumferential l

welds 1-4.

I Attachment 6: Latter B.R. Cofer to R.W.

Alley, February 3,1983 c' " Data for Analysis of Welds 1-7 and 1-9.

l 140 Receipt of Design Materials letter from W. McNaughton to P. GudikiinstTDRemb'er 20, 1982.

,, 141 Letter from W. McNaughton to K. Nebb (site) returning radiographs dated 3-11-83.

142 Acknowledgement of receipt of radiographs

' listed in CD-141 (from K. Webb) 143 Letter from K. Webb (site) sending (29) radiographs with acknowledgement 3-23-83.

! 144 Letter from K. Webb (site) sending (15) radiographs with acknowledgement 3-24-83.

145 Letter from R. Alley to W. McNaughton summarizing radiographs sent and stresses which correspond. PY-STR-1607 3-31-83.

l i i. '

- . . ; y- ,.. . ., ; -,. .~. ..f7 ; _------

i D-4 n CONTROLLED

, DOCUMENT NUMBER ITEM

., 146 Letter from K. Webb (site) to W. McNaughton

sending 6 radiographs 4-4-83.

147 Letter from W. McNaughton to R. Alley

'l '

April 19,1983 "Aptech Evaluation of

, 4 shell courses defects - background information".

. 148 Summary of radiographs reviewed cons-

. isting of notes and measurements.

149 Documentation of locations on each radio-

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graph of CD-148 as to depth measurement location.

. 150 V!deotape with balance of Gilbert i

enhancement.

151

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l Stress calculations - load combinations as taken from CD-139 and checked.

152 Developed integrated stresses - service and residual stresses.

153 Limit load analysis. .

>- 154 BIGIF runs used in final report - 1H1, 1H2, 1HS, 1H6, 6H1, 6H2, 6H3, 6H4, 8H1, 8H2 I :.

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