ML041700124
| ML041700124 | |
| Person / Time | |
|---|---|
| Site: | Palo Verde  |
| Issue date: | 03/31/2004 |
| From: | Jirawongkraisorn S Westinghouse |
| To: | Office of Nuclear Reactor Regulation |
| References | |
| FOIA/PA-2004-0247 WCAP-15817-NP, Rev 1 | |
| Download: ML041700124 (104) | |
Text
Westinghouse Non-Proprietary Class 3 WCAP-15817-NP March 2004 Revision 1 (Revision 0 was never published)
Structural Integrity Evaluation of Reactor Vessel Upper Head Penetrations to Support Continued Operation:
Palo Verde Units 1 and 2 Westinghouse Oir Woa 1-e,11.,etcls
WESTINGHOUSENON-PROPRIETARY CLASS 3 WCAP-15817-NP Revision 1 (Revision 0 was never published)
Structural Integrity Evaluation of Reactor Vessel Upper Head Penetrations to Support Continued Operation:
Palo Verde Units 1 and 2 S. Jirawongkraisorn March 2004 Reviewer:
C.PngAN Piping Analysis rracture Mechanics Piping Analysis & Fracture Mechanics Westinghouse Electric Company LLC P.O. Box 355 Pittsburgh, PA 15230-0355 02004 Westinghouse Electric Company LLC All Rights Reserved
iii TABLE OF CONTENTS I INTRODUCTION ........................................ . 1-1 2 HISTORY OF CRACKING IN HEAD PENETRATIONS .................................. 2-1 3 OVERALL TECHNICAL APPROACH ........................................ 3-1 3.1 PENETRATION STRESS ANALYSIS ........................................ 3-1 3.2 FLAW TOLERANCE APPROACH ........................................ 3-1 4 MATERIAL PROPERTIES, FABRICATION HISTORY AND CRACK GROWTH PREDICTION .............................. 4-1 4.1 MATERIALS AND FABRICATION .............................. 4-1 4.2 CRACK GROWTH PREDICTION .............................. 4-1 5 STRESS ANALYSIS ............................... 5-1 5.1 OBJECTIVES OF THE ANALYSIS ........................................................ 5-1 5.2 MODEL ........................................................ 5-1 5.3 STRESS ANALYSIS RESULTS - OUTERMOST CEDM PENETRATION (51.5 DEGREES) ........................................................ 5-1 5.4 STRESS ANALYSIS RESULTS - INTERMEDIATE CEDM PENETRATIONS ........ 5-2 5.5 STRESS ANALYSIS RESULTS - CENTER CEDM PENETRATION ....................... 5-2 5.6 STRESS ANALYSIS RESULTS - HEAD VENT ...................................................... 5-2 6 FLAW EVALUATION CHARTS ...................... 6-1
6.1 INTRODUCTION
.................................. 6-1 6.2 OVERALLAPPROACH .................................. 6-1
6.3 RESULTS
AXIAL FLAWS .................................. 6-2 6.4 CIRCUMFERENTIAL CRACK PROPAGATION .................................. 6-3 6.5 FLAWACCEPTANCE CRITERIA .................................. 6-4 March 2004 Revision I
iv TABLE OF CONTENTS (cont.)
7
SUMMARY
AND EXAMPLE PROBLEMS ........................................... 7-1 7.1 SAFETYYASSESSMENT ........................................... 7-1 7.2 EXAMPLE PROBLEMS ........................................... 7-2 8 REFERENCES ........................................... 8-1 APPENDIX A . A-1 APPENDIX B.B-1 March 2004 Revision I
1-1 1 INTRODUCTION In September of 1991, a leak was discovered in the reactor vessel control rod drive (CRDM) head penetration region of an operating plant. This has led to the question of whether such a case could occur at Palo Verde Units 1 and 2. The geometry of interest is shown in Figure 1-1. Throughout this report, the penetration rows have been identified by their angle of intersection with the head. For each penetration of each unit, the angle is identified in Table 1-1 for Unit 1 and Table 1-2 for Unit 2. Note that the designations CRDM (Westinghouse and French designs) and CEDM (CE designs) are synonymous.
The issue resulted from cracking which occurred in the outermost penetrations of a number of operating plants, as discussed in Section 2. This outermost CEDM location, as well as a number of intermediate CEDM penetrations, the head vent, and the center penetration, was chosen for fracture mechanics analyses to support continued safe operation of Palo Verde Units 1 and 2 if such cracking were to be found. The dimensions of the CEDM penetrations for Units 1 and 2 are all identical, with 4.05 inch OD and wall thickness of 0.661 inches. For the head vent, the OD is 1.05 inches, and the wall thickness is 0.154 inches.
The basis of the analyses was a detailed three dimensional elastic -plastic finite element analysis of several penetration locations, as described in detail in Section 5. Results were obtained at a number of locations in each penetration, and used in the fracture analysis.
The fracture analyses were carried out using reference crack growth rates recommended by the EPRI Materials Reliability Project, which are consistent with service experience. The results are presented in the form of flaw evaluation charts for both surface and through wall flaws, to determine the allowable time of safe operation if indications are found. All the times calculated in this handbook are effective full power years (EFPY).
This report, WCAP- 15817-NP Revision 1, is the non-proprietary version of NVCAP- 15817-P Revision 1.
The non-proprietary version of WCAP-15817 Revision 0 was never published. WCAP-15817-P Revision 0 was revised and the record of revisions made is summarized in the table below.
Record of Revisions Revision Date Description 0 March . .
2002 Oriial Issue Hoop stress distribution plots shown in Appendix B are included in this report.
1 October The plots present the lengthwise hoop stress distribution on both the inside and 2003 the outside surfaces of the CEDM nozzles. In addition, typographical corrections were made to equations 4-4, 6-1, and 6-2.
Introduction March 2004 Revision I
1-2 Table 1-1 Palo Verde Unit I llead Penetration Nozzles, with Intersection Angles Identified Nozzle Angle Nozzle Angle Nozzle Angle No. Type (degrees) No. Type (degrees) No. Type (degrees)
I CEDM 0.0 28 CEDM 28.0 55 CEDM 40.7 2 CEDM 7.5 29 CEDM 28.0 56 CEDM 40.7 3 CEDM 7.5 30 CEDM 31.A 57 CEDM 40.7 4 CEDM 7.5 31 CEDM 31.4 58 CEDM 40.7 5 CEDM 7.5 32 CEDM 31.A 59 CEDM 40.7 6 CEDM 16.9 33 CEDM 31.4 60 CEDM 40.7 7 CEDM 16.9 34 CEDM 32.5 61 CEDM 40.7 8 CEDM 16.9 35 CEDM 32.5 62 CEDM 40.7 9 CEDM 16.9 36 CEDM 32.5 63 CEDM 40.7 10 CEDM 16.9 37 CEDM 32.5 64 CEDM 40.7 11 CEDM 16.9 38 CEDM 32.5 65 CEDM 40.7 12 CEDM 16.9 39 CEDM 32.5 66 CEDM 41.7 13 CEDM 16.9 40 CEDM 32.5 67 CEDM 41.7 14 CEDM 21.6 41 CEDM 32.5 68 CEDM 41.7 15 CEDM 21.6 42 CEDM 33.6 69 CEDM 41.7 16 CEDM 21.6 43 CEDM 33.6 70 CEDM 41.7 17 CEDM 21.6 44 CEDM 33.6 71 CEDM 41.7 18 CEDM 23.0 45 CEDM 33.6 72 CEDM 41.7 19 CEDM 23.0 46 CEDM 35.7 73 CEDM 41.7 20 CEDM 23.0 47 CEDM 35.7 74 CEDM 44.6 21 CEDM 23.0 48 CEDM 35.7 75 CEDM 44.6 22 CEDM 28.0 49 CEDM 35.7 76 CEDM 44.6 23 CEDM 28.0 50 CEDM 35.7 77 CEDM 44.6 24 CEDM 28.0 51 CEDM 35.7 78 CEDM 44.6 25 CEDM 28.0 52 CEDM 35.7 79 CEDM 44.6 26 CEDM 28.0 53 CEDM 35.7 80 CEDM 44.6 27 CEDM 28.0 54 CEDM 40.7 81 CEDM 44.6 Introduction March 2004 Revision I
1-3 Table 1-1 Palo Verde Unit I Head Penetration Nozzles, with Intersection Angles Identified (cont.)
Nozzle Angle Nozzle Angle Nozzle Angle No. Type (degrees) No. Type (degrees) No. Type (degrees) 82 CEDM 47.5 88 CEDM 51.5 94 CEDM 48.0 83 CEDM 47.5 89 CEDM 51.5 95 CEDM 48.0 84 CEDM 47.5 90 CEDM 48.0 96 CEDM 48.0 85 CEDM 47.5 91 CEDM 48.0 97 CEDM 48.0 86 CEDM 51.5 92 CEDM 48.0 87 CEDM 51.5 93 CEDM 48.0 March 2004 Introduction March 2004 Revision I
1-4 Table 1-2 Palo Verde Unit 2 llead Pcnetration Nozzles, with Intersection Angles Identified Nozzle Angle Nozzle Angle Nozzle Angle No. Type (degrees) No. Type (degrees) No. Type (degrees)
I CEDM 0.0 28 CEDM 28.0 55 CEDM 40.7 2 CEDM 7.5 29 CEDM 28.0 56 CEDM 40.7 3 CEDM 7.5 30 CEDM 31.4 57 CEDM 40.7 4 CEDM 7.5 31 CEDM 31.4 58 CEDM 40.7 5 CEDM 7.5 32 CEDM 31.4 59 CEDM 40.7 6 CEDM 16.9 33 CEDM 31.A 60 CEDM 40.7 7 CEDM 16.9 34 CEDM 32.5 61 CEDM 40.7 8 CEDM 16.9 35 CEDM 32.5 62 CEDM 40.7 9 CEDM 16.9 36 CEDM 32.5 63 CEDM 40.7 10 CEDM 16.9 37 CEDM 32.5 64 CEDM 40.7 11 CEDM 16.9 38 CEDM 32.5 65 CEDM 40.7 12 CEDM 16.9 39 CEDM 32.5 66 CEDM 41.7 13 CEDM 16.9 40 CEDM 32.5 67 CEDM 41.7 14 CEDM 21.6 41 CEDM 32.5 68 CEDM 41.7 15 CEDM 21.6 42 CEDM 33.6 69 CEDM 41.7 16 CEDM 21.6 43 CEDM 33.6 70 CEDM 41.7 17 CEDM 21.6 44 CEDM 33.6 71 CEDM 41.7 18 CEDM 23.0 45 CEDM 33.6 72 CEDM 41.7 19 CEDM 23.0 46 CEDM 35.7 73 CEDM 41.7 20 CEDM 23.0 47 CEDM 35.7 74 CEDM 44.6 21 CEDM 23.0 48 CEDM 35.7 75 CEDM 44.6 22 CEDM 28.0 49 CEDM 35.7 76 CEDM 44.6 23 CEDM 28.0 50 CEDM 35.7 77 CEDM 44.6 24 CEDM 28.0 51 CEDM 35.7 78 CEDM 44.6 25 CEDM 28.0 52 CEDM 35.7 79 CEDM 44.6 26 CEDM 28.0 53 CEDM 35.7 80 CEDM 44.6 27 CEDM 28.0 54 CEDM 40.7 81 CEDM 44.6 Introduction March 2004 Revision I
1-5 Table 1-2 Palo Verde Unit 2 Head Penetration Nozzles, with Intersection Angles Identified (cont.)
Nozzle Angle Nozzle Angle Nozzle Angle No. Type (degrees) No. Type (degrees) No. Type (degrees) 82 CEDM 47.5 88 CEDM 51.5 94 CEDM 48.0 83 CEDM 47.5 89 CEDM 51.5 95 CEDM 48.0 84 CEDM 47.5 90 CEDM 48.0 96 CEDM 48.0 85 CEDM 47.5 91 CEDM 48.0 97 CEDM 48.0 86 CEDM 51.5 92 CEDM 48.0 87 CEDM 51.5 93 CEDM 48.0 March 2004 Introduction March 2004 Revision I
1-6
!- Am 4 .
Figure 1-1 Reactor Vessel Control Element Drive Mechanism (CEDM) Penetration March 2004 Introduction Introduction March 2004 Rcvision I
1-7 Figure 1-2 Location of Head Penetrations for Palo Verde Units I and 2 Introduction March 2004 Revision I
2-1 2 HISTORY OF CRACKING IN HEAD PENETRATIONS In September of 1991, leakage was reported from the reactor vessel CRDM head penetration region of a French plant, Bugey Unit 3. Bugey 3 is a 920 megawatt three-loop PWR which had just completed its tenth fuel cycle. The leak occurred during a post ten year hydrotest conducted at a pressure of approximately 3000 psi (204 bar) and a temperature of 1940 F (90'C). The leak was detected by metal microphones located on the top and bottom heads, and the leak rate was estimated to be approximately 0.7 liter/hour. The location of the leak was subsequently established on a peripheral penetration with an active control rod (IH-14), as seen in Figure 2-1.
The control rod drive mechanism and thermal sleeve were removed from this location to allow further examination. Further study of the head penetration revealed the presence of longitudinal cracks near the head penetration attachment weld. Penetrant and ultrasonic testing confirmed the cracks. The cracked penetration was fabricated from Alloy 600 bar stock (SB-166), and has an outside diameter of 4 inches (10.16 cm) and an inside diameter of 2.75 inches (7.0 cm).
As a result of this finding, all of the control rod drive mechanisms and thermal sleeves at Bugey 3 were removed for inspection of the head penetrations. Only two penetrations were found to be cracked, as shown in Figure 2-1.
An inspection of a sample of penetrations at three additional plants were planned and conducted during the winter of 1991-92. These plants were Bugey 4 , Fessenheim 1, and Paluel 3. The three outermost rows of penetrations at each of these plants were examined, and further cracking was found in two of the three plants.
At Bugey 4, eight of the 64 penetrations examined were found to contain axial cracks, while only one of the 26 penetrations examined at Fessenheim I was cracked. The locations of all the cracked penetrations are shown in Figure 2-1. None of the 17 CRDM penetrations inspected at Paluel 3 showed indications of cracking, at the time, but further inspection of the French plants have confirmed at least one crack in each operating plant.
Thus far, the cracking in tubes not manufactured by Babcock and Wilcox Tubular Products has been consistent in both its location and extent. All cracks discovered by nondestructive examination have been oriented axially, and have been located in the bottom portion of the penetration in the vicinity of the partial penetration attachment weld to the vessel head as shown schematically in Figure 1-1.
]ac.e Cracking in of Cracking 1-lead Penetrations March 2004 History of History in Hcad Penetrations Marci2004 Revision I
2-2 Non-destructive examinations of the leaking CRDM nozzles showed that most of the cracks originated on the outside surface of the nozzles below the J-groove weld, were axially oriented, and propagated primarily in the nozzle base material to an elevation above the top of the J-groove weld where leakage could then pass through the annulus to the top of the head where it was detected by visual inspection. In some cases the cracks initiated in the weld metal or propagated into the weld metal, and in a few cases the cracks propagated through the nozzle wall thickness to the inside surface.
The cracking has now been confirmed to be primary water stress corrosion cracking. Relatively high residual stresses are produced in the outermost CEDM (or CRDM) penetrations due to the welding process. Other important factors which affect this process are temperature and time, with higher temperatures and longer times being more detrimental. It is interesting to note that no ICIs or head vents have been found to be cracked. The inspection findings for the plants examined thus far are summarized in Table 2-1.
Penetrations I-lead Penetrations March 2004 History of Cracking in Head Cracking in March 2004 Revision I
2-3 Table 2-1 Operational Information and Inspection Results for Units Examined (Results to December 30,2001)
Head Penetrations Plant Units Temp. Total Penetrations With Country Type Inspected K Hours (OF) Penetrations Inspected Indications France CPO 6 80-107 596-599 390 390 23 CPY 28 42-97 552 1820 1820 126 1300MW 20 32-51 558-597 1542 1542 95 Sweden 3 Loop 3 75-115 580-606 195 190 8 Switzerland 2 Loop 2 148-154 575 72 72 2 Japan 2 Loop 7 105-108 590-599 276 243 0 3 Loop 7 99 610 455 398 0 4 Loop 3 46 590 229 193 0 Belgium 2 Loop 2 115 588 98 98 0 3 Loop 5 60-120 554-603 337 337 6 Spain 3 Loop 5 65-70 610 325 102 0 Brazil 2 Loop 1 25 NA 40 40 0 South Africa 3 Loop I NA NA 65 65 6 Slovenia 2 Loop I NA NA 49 49 0 South Korea 2 Loop 3 NA NA 49 49 3 3 Loop 2 NA NA 130 130 2 US 2 Loop 2 170 590 98 98 0 3 Loop I NA NA 65 20 12 4 Loop 14 NA NA 899 287 35 TOTALS 113 _ _ 7134 6123 318 March 2004 History of Cracking Head Penetrations in Head Cracking in Penetrations March 2004 Revision I
2-4 j~FAIDm Figure 2-1 French R/V Closure Head CRDM Penetration Cracking EdF Plants- Penetrations with Cracking History of Cracking in Head Penetrations March 2004 Revision I
3-1 3 OVERALL TECHNICAL APPROACH The primary goal of this work is to provide technical justification for the continued safe operation of Palo Verde Units 1 and 2 in the event that cracking is discovered during inservice inspections of the Alloy 600 reactor vessel head penetrations.
3.1 PENETRATION STRESS ANALYSIS Three dimensional elastic -plastic finite element stress analyses have been performed to determine the stresses in the head penetration region [6A, 6B]. These analyses have considered the pressure and thermal transient loads associated with steady state operation, as well as the residual stresses which are produced by the fabrication process.
3.2 FLAW TOLERANCE APPROACH A flaw tolerance approach has been developed to allow continued safe operation until an appropriate time for repair, or the end of plant life. The approach is based on the prediction of future growth of detected flaws, to ensure that such flaws would remain stable.
If an indication is discovered during inservice inspection, its size can be compared with the flaw size which is considered allowable for continued service. This "allowable" flaw size is determined from the actual loadings (including mechanical, residual, and transient loads) on the head penetration for the plant of interest. Suitable margins to ensure the integrity of the reactor vessel as well as safety from unacceptable leakage rates, should also be considered. Acceptance criteria are discussed in Section 6.5.
The time for the observed crack to reach the allowable crack size determines the length of time the plant can remain online before repair, if required. For the crack growth calculation, a best estimate is needed; no additional margins are necessary.
The results of the evaluation are presented in terms of simple charts, which show graphically the time required to reach the allowable length or depth, which represents the additional service life before repair.
This result is a function of the loadings on the particular head penetration, as well as the circumferential location of the crack in the penetration tube.
Overall Technical Approach March 2004 Revision I
3-2 Schematic drawings of the head penetration flaw tolerance charts are presented as Figures 3-1 and 3-2.
These two types of charts can be used to provide estimates of the time which remains before a leak would develop from an observed crack. For example, if a part-through flaw was discovered, the user would first refer to Figure 3-1, to determine the time (tp) which would be remaining before the crack would penetrate the wall or reach the allowable depth (tA) (eg a/lt=.75). Once the crack penetrates the wall, the time (tB) required to reach an allowable crack length would be determined from Figure 3-2. The total time remaining would then be the simple sum:
Time remaining = tp + tB Another way to determine the allowable time of operation with a part-through flaw would be to use Figure 3-2 directly, in effect assuming the part-through flaw is a through-wall flaw. This approach would be more conservative than that above, and the time remaining would then be:
Time remaining =tB Overall Technical Approach March 2004 Revision I
3-3 Flaw Becomes Through - Wall Time ( Months )
Figurc3-1 Schematic of a Head Penetration Flaw Growth Chart for Part Through Flaws Overall Technical Approach March 2004 Revision I
3-4
- i- .t F _ tWxe3 Figure 3-2 Schematic of a Head Penetration Flaw Tolerance Chart for Through-Vall Flaws Overall Technical Approach March 2004 Revision I
4-1 4 MATERIAL PROPERTIES, FABRICATION HISTORY AND CRACK GROWTH PREDICTION 4.1 MATERIALS AND FABRICATION The reactor vessels for Pab Verde Units 1 and 2 were manufactured by Combustion Engineering with head penetration nozzles from material produced by Huntington Alloys and Standard Steel in the USA.
The carbon content and mechanical properties of the Alloy 600 material used to fabricate the Palo Verde Units 1 and 2 vessels are provided in Table 4-1. The material CMTRs were used to obtain the chemistry and mechanical properties for the vessel head penetrations. The CMTRs for the material do not indicate the heat treatment of the material. However, Westinghouse records indicate that the Huntington materials were annealed for one hour at a temperature of 17000 F - I 8000 F, followed by a water quench. The Standard Steel materials were annealed for six hours at 16251F, and air cooled. Figures 4-1 illustrates the yield strengths and carbon content, based on percent of heats, of the head adapter penetrations in the Palo Verde Units 1 and 2 vessels relative to a sample of the French head penetrations which have experienced cracking. The general trend for the head adapter penetrations in Palo Verde Units 1 and 2 are a higher carbon content, higher mill annealing temperature and lower yield strength relative to those on the French vessels. These factors should all have a beneficial effect on the material resistance to PWSCC in the head penetrations.
4.2 CRACK GROWTH PREDICTION The cracks in the penetration region have been determined to result from primary water stress corrosion cracking in the Alloy 600 base metal, and in some cases the Alloy 182 weld metal. There are a number of available measurements of static load crack growth rates in primary water environment, and in this section the available results will be compared and a representative growth rate established.
Direct measurements of SCC growth rates in Alloy 600 are relatively rare, and care should be used in interpreting the results because the materials may be excessively cold worked, or the loadings applied may be near or exceeding the limit load of the tube, meaning there will be an interaction between tearing and crack growth. In these cases the crack growth rates may not be representative of service conditions.
The effort to develop a reliable crack growth rate model for Alloy 600 began in the spring of 1992; when the Westinghouse Owners Group was developing a safety case to support continued operation of plants.
At the time there was no available crack growth rate data for head penetration materials, and only a few publications existed on growth rates of Alloy 600 in any product form.
The best available publication was found to be that of Peter Scott of Framatome, who had developed a growth rate model for PWR steam generator materials [1]. His model was based on a study of results obtained by McIlree and Smialowska [2] who had tested short steam generator tubes which had been flattened into thin compact specimens. Upon study of his paper there were several ambiguities, and several phone conversations were held to clarify his conclusions. These discussions led to Scott's admission that reference 1 contains an error, in that no correction for cold work was applied to the McIllree/Smialowska data. The correct development is below.
Material Properties, Fabrication History and Crack Growth Prediction March 2004 Revision I
4.2 An equation was fitted to the data of reference [2] for the results obtained in water chemistries that fell within the standard specification for PWR primary water. Results for chemistries outside the specification were not used. The following equation was fitted to the data at 3301C (6261F):
da =2.8xlOl (K _9)1.16 m/sec (4-1) dt where K is in MPalm.
The next step described by Scott in his paper was to correct these results for the effects of cold work.
Based on work by Cassagne and Gelpi [3], he concluded that dividing the above equation by a factor of 10 would be appropriate to account for the effects of cold work. This step was inadvertently omitted from Scott's paper, even though it is discussed. The crack growth law for 330'C (6261F) then becomes:
d =2.8xlO12 (K 9) 1 16 m/sec (4-2) dt This equation was verified by Scott in a phone call in July 1992.
Scott further corrected this law for the effects of temperature, but his correction was not used in the model employed here. Instead, an independent temperature correction was developed based on service experience, as will be discussed below.
Iancbe There is a general agreement that crack growth in Alloy 600 materials in the primary water environment can be modeled using a stress intensity factor relationship with differences in temperature accounted for by an activation energy (Arrhenius) model for thermally controlled processes. Figure 4-3 shows the Material Properties, Fabrication History and Crack Growth Prediction March 2004 Revision I
4-3 recommended CGR curve along with the laboratory data from Huntington materials used to develop the curve.
]
Material Properties, Fabrication History and Crack Growth Prediction March 2004 Revision I
4-4
]ac,e The applicability of the MRP recommended model to the head penetrations at the Palo Verde Units 1 and 2 was recently confirmed by two independent approaches. The first was a collection of all available data from Huntington Alloys materials tested over the past ten years [4H]. The results are shown in Figure 4-3, along with the Scott model for the test temperature.
The MRP crack growth curve was structured to bound 75 percent of the 22 heats for which test results were available. Fits were done on the results for each heat, and the constant term was determined for each heat. The 75th percentile was then determined from these results. This was done to eliminate the concern that the curve might be biased from a large number of results from a single heat. The MRP expert panel on crack growth endorsed the resulting curve unanimously in a meeting on March 6'h and 7h 2002. This approach is consistent with the Section XI flaw evaluation philosophy, which is to make a best estimate prediction of future growth of a flaw. Margins are incorporated in the allowable flaw sizes.
The entire data set is shown in Figure 4-3, where the data have been adjusted to a single temperature of 3250C.
A second independent set of data were used to validate the model, and these data were obtained from the two inspections carried out on penetration 75 of D.C. Cook Unit 2, which was first found to be cracked in 1994 [4C]. The plant operated for one fuel cycle before the penetration was repaired in 1996 and the flaw was measured again before being repaired. These results were used to estimate the PWSCC growth rate, for both the length of the flaw and its depth. These two points are also shown in Figure 4-4, and are consistent with the laboratory data for Huntington and Standard Steel materials. In fact, Figure 44 Material Properties, Fabrication History and Crack Growth Prediction March 2004 Revision I
4-5 demonstrates that the MRP model is nearly an upper bound for these materials. The D.C. Cook Unit 2 penetrations were made from Huntington materials.
Since Palo Verde Units 1 and 2 operates at a temperature lower than 330'C in the head region, and the crack growth rate is strongly affected by temperature, a temperature adjustment is necessary. This temperature correction was obtained from study of both laboratory and field data for stress corrosion crack growth rates for Alloy 600 in primary water environments. The available data showing the effect of temperature are summarized in Figure 4-5. Most of the results shown here are from steam generator tube materials, with several sets of data from operating plants, and results from two heats of materials tested in a laboratory [4A].
Study of the data shown in Figure 4-4 results in an activation energy of 31-33 Kcallmole, which can then be used to adjust for the lower operating temperature. This value is slightly lower than the generally accepted activation energy of 44-50 Kcal/mole used to characterize the effect of temperature on crack initiation, but the trend of the actual data for many different sources is unmistakable.
IT Therefore the following growth rate model was used for the Palo Verde Units I and 2 head penetrations:
da =1.76x10-'2 (K_9)1-16 m/sec dt where K = applied stress intensity factor, in MPa F . This equation implies a threshold for cracking susceptibility, KIscc 9 MPal4.
Material Properties, Fabrication History and Crack Growth Prediction March 2004 Revis sion I
4-6 Table 4-1 Palo Verde Units I and 2 Head Penetration Material Information Palo Verde Unit I Heat CEDMI or Head Vent Yield Strength ksi Carbon wt% Supplier EO 4039 CEDM 51.0 0.087 Standard Steel EO 4163 CEDM 41.5 0.075 Standard Steel EO 4162 CEDM 49.0 0.079 Standard Steel NX 8445 Head Vent 37.0 0.03 Huntington Alloys Palo Verde Unit 2 EO 4358 CEDM 37.0 0.074 Standard Steel EO 4359 CEDM 49.0 , 0.078 Standard Steel EO 3045 CEDM 48.0 0.076 Standard Steel EO 4039 CEDM 44.0 0.089 Standard Steel EO 4707 CEDM 48.0 0.092 Standard Steel EO 4163 CEDM 40.0 0.073 Standard Steel EO 2845 CEDM 53.0 0.070 Standard Steel NX 8408 Head Vent 50.0 0.020 Huntington Alloys Material Properties, Fabrication History and Crack Growth Prediction March 2004 Revision I
4-7 50 5 EdF (11 Heats) 45 0 Palo Verde 1 (4 Heats)
_ 40 El Palo Verde 2 (8 Heats) 35 I~30 i 25 0
'~20 m 15 10 5
0 I Yield Strength (ksi)
Figure 4-1 Yield Strength of the Various Meats of Alloy 600 Used in Fabricating the Palo Verde Units I and 2 and French Head Adapter Penetrations Material Properties, Fabrication History and Crack Growth Prediction March 2004 Revision I
4-8 80 El EdF (11 Heats) 70 El Palo Verde 1 (4 Heats)
El Palo Verde 2 (8 Heats)
O 60 I
to ai 50 a) i.- 40 0
46-a 0 30 0
03 20 IL 10 1I n
Carbon Content (Weight %)
Figure 4-2 Carbon Content of the Various Heats of Alloy 600 Used In Fabricating the Palo Verde Units I and 2 and French Head Adapter Penetrations Material Properties, Fabrication History and Crack Growth Prediction March 2004 Revision I
4-9 I
I-.E-10 Cr i:.E-Il1 I.
U V
1.E-12 20 30 40 50 60 Stress Intensity Factor, K (MPavm)
Figure 4-3 Screened Laboratory Data for Alloy 600, with the MRP Recommended 75/50 Curve.
Note that the Modified Scott Model is also Shown Material Properties, Fabrication History and Crack Growth Prediction March 2004 Revision I
4-10 1.E-09
-1.E-10 to E
- 0 a:
L.
'5 0
0 J1~
() 1.E-1 1 I
0 10 20 30 40 50 60 70 80 Stress Intensity Factor, K (MPa*sqrt(m))
Figure 4-4 Model for PWSCC Growth Rates in Alloy 600 in Primary Water Environments (325 0 C), With Supporting Data from Standard Steel, Huntingdon, and Sanvik Materials Material Properties, Fabrication History and Crack Growth Prediction March 2004 Revision I
4-11 Note: All symbols are for steam generator materials, except the solid circles, which are head penetration laboratory data.
Figure 4-5 Summary of Temperature Effects on PWSCC Growth Rates for Alloy 600 in Primary Water Material Properties, Fabrication History and Crack Growth Prediction March 2004 Revision I
5-1 5 STRESS ANALYSIS 5.1 OBJECTIVES OF THE ANALYSIS The objective of this analysis was to obtain accurate stresses in each CEDM housing and its immediate vicinity. To do so requires a three dimensional analysis which considers all the pertinent loadings on the penetration [6A, 6B]. An investigation of deformations at the lower end of the housing was also performed using the same model. Five locations were considered: the outermost row (51.5 degrees),
rows at 35.7 degrees, 28.0 degrees, 7.5 degrees and the center location.
The analyses were used to provide information for the flaw tolerance evaluation which follows in Section 6. Also, the results of the stress analysis were compared to the findings from service experience, to help assess the causes of the cracking which has been observed.
5.2 MODEL A three dimensional finite element model comprised of isoparametric brick and wedge elements with midside nodes on each face was used to obtain the stresses and deflections. A view of the outermost CEDM model is shown in Figure 5-1. A similar model for the head vent penetration is shown in Figure 5-2. Taking advantage of symmetry through the vessel and penetration centerlines only half of the penetration geometry plus the surrounding vessel were modeled for the outermost and next outermost penetrations. The difference between the hillside penetrations and the center penetration was that there was no differential height across the weld for the center penetration.
In the models, the lower portion of the Control Element Drive Mechanism (CEDM) penetration tube or head vent, the adjacent section of the vessel closure head, and the joining weld were modeled. The vessel to penetration tube weld was simulated with two layers of elements. The penetration tube, weld metal and cladding were modeled as Alloy 600 and the vessel head shell as carbon steel. The models were consistent with, but slightly more refined, than previous models used to evaluate Combustion Engineering designed head penetration nozzles. The benchmarking of models is described in reference 6A.
The only loads used in the analysis are the steady state operating loads. External loads such as seismic loads have been studied, and have no impact, because the penetration tubes are captured by the full thickness of the reactor vessel head, over seven inches of steel into which the penetrations are shrink fit during construction. The area of interest is at the attachment weld, which is totally unaffected by these external loads.
5.3 STRESS ANALYSIS RESULTS - OUTERMOST CEDM PENETRATION (51.5 DEGREES)
Figure 5-3 presents the hoop and axial stresses for the steady state condition for the outermost penetration.
] awc.e Stress Analysis March 2004 Revision I
5-2 I
5.4 STRESS ANALYSIS RESULTS - INTERMEDIATE CEDM PENETRATIONS
[
Iace 5.5 STRESS ANALYSIS RESULTS - CENTER CEDM PENETRATION Iace 5.6 STRESS ANALYSIS RESULTS - HEAD VENT The head vent is a smaller penetration than the CEDM head penetrations, but is also constructed of Alloy 600 material, with a partial penetration weld at the inside of the reactor vessel head. The head vent is located 8.2 inches from the centerline of the head dome.
The head vent was evaluated using a three dimensional finite element model, as shown in Figure 5-2.
The critical stress location in the head vent is in the vicinity of the attachment weld, where residual and pressure stresses have the most impact. As with the CEDM head penetrations, the residual stresses dominate. Also similar to the CEDM head penetrations, the stresses in the pipe decrease quickly as a function of distance up the pipe away from the weld. The hoop and axial stresses are shown as contours in Figure 5-9.
Stress Analysis March 2004 Revision 1
5-3
- exisl L .raia (dul) ni1 DTnhifl Plaw No&s we O' f.ale LipWi Pla Nodes we SK)C(O's Sales TtheXo&Salier lS t Ii) ta1e00 ShebNodaSS S's a-nShe! ID (nmiied v.e OD) in well rclcm 6Su SMel ID above "Al rqrion 15's at e4v Msbell iwok Nde mn'an lisse by 1W up the latigIh athete and Act NodaNnbos arms by I Ieg the stilt smhel reuios Figure 5-1 Three-Dimensional Model of the Outermost CEDMI Penetration (51.5 Degrees)
Stress Analysis March 2004 Revision I
5-4 Figure 5-2 Three Dimensional Model of the Head Vent Penetration Stress Analysis March 2004 Revision I
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Stress Analysis March 2004 Revision I
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Figure 5-5 Stress Distribution at Steady State Condition for the 35.7 Degree Penetration (Hoop stress is the top figure; axial stress is the bottom figure)
Stress Analysis March 2004 Revision I
5-8
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Stress Analysis March 2004 Revision 1
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Stress Analysis March 2004 Revision I
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Stress Analysis March 2004 Revision I
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- 20400, DVA2-WS(.3di.CC 6S.i.S0/.742.0.AJ - tOF4YCIni Figure 5-9 Stress Contours in the Head Vent As A Result of Residual Stresses and Operating Pressure (Hoop stress is the top figure; axial stress is the bottom figure)
Stress Analysis March 2004 Revision I
6-1 6 FLAW EVALUATION CHARTS
6.1 INTRODUCTION
The flaw evaluation charts were developed from the stress analysis of each of the penetration locations, as discussed in Section 5. The crack growth law developed for Palo Verde Units 1 and 2 in Section 4.2 was used for each case, and several flaw tolerance charts were developed for each penetration location. The first series of charts characterizes the growth of a part through flaw, and the second series of charts characterizes the growth of a through-wall flaw in the length direction. The allowable safe operating life of the penetration may then be directly determined, using the combined results of the two charts. All times resulting from these calculations are effective full power years, since crack growth will only occur at operating temperatures.
6.2 OVERALL APPROACH The results of the three-dimensional stress analysis of the penetration locations were used directly in the flaw tolerance evaluation.
The crack growth evakiation for the part-through flaws was based on the stress distribution through the penetration wall at the location which corresponds to the highest stress along the inner surface of the penetration. The highest stressed location was found to be in the immediate vicinity of the weld for both the center and outermost penetrations.
The stress profile was represented by a cubic polynomial:
a(x) = AO + Ax+A2 x2 + A3 x3 (6-1) where x = the coordinate distance into the wall a = stress perpendicular to the plane of the crack Ai = coefficients of the cubic fit For the surface flaw with length six times its depth, the stress intensity factor expression of Raju and Newman [5A] was used. The stress intensity factor KI (4) can be calculated anywhere along the crack front. The point of maximum crack depth is represented by 0 = 0, and this location was found to also be the point of maximum K, for the cases considered here. The following expression is used for calculating K, (0), where 0 is the angular location around the crack. The units of K, (0) are ksiV4.
K Nt 053jGj(a/e a/t, t/R,D) Aj aj
[a)=_] (6-2)
The magnification factors Go (0), GI (4), G2 (0) and G3 (4) are obtained by the procedure outlined in reference [5A]. The dimension "a" is the crack depth, and "c" is the crack length, while t is the wall thickness.
March 2004 Flaw Evaluation Flaw Charts Evaluation Charts March 2004 Revision I
6-2 ja,cec
6.3 RESULTS
AXIAL FLAWS CEDM Surface Flaws The results of the calculated growth through the wall thickness of the CEDMs for surface flaws are shown in Figures 6-2 through 6-7 for inside surface flaws. For outside surface flaws the results are shown in Figures 6-8 and 6-9. These crack growth curves begin at a flaw depth that results in a stress intensity factor that exceeds the threshold value of 9 MPa 4m. This sometimes results in curves with different initial flaw sizes, as seen for example in Figure 6-3. Note that results are only provided for the uphill and downhill sides of each penetration nozzle; the stresses for the regions 90 degrees from these locations are compressive. If flaws are found in such a location, use the results for either the uphill or downhill location, whichever is closer.
Each of these figures allows the future allowable service time to be estimated graphically, as discussed in Section 3. Results are shown for each of the penetrations analyzed in each of these figures. The stresses are much higher near the attachment weld than below or above it, so separate figures have been provided for these three regions. Also, the stresses are different on the downhill side of the penetration as opposed to the uphill side, so these two cross sections have also been treated separately.
Flaw Evaluation Charts March 2004 Revision I
6-3 Examples have been provided in Section 7 for a range of possible flaw types, so the graphical approach can be completely understood.
CEDM Through-Wall Flaws Crack growth from surface flaws in the CEDMs is the primary concern in evaluation of the structural integrity of head penetrations, but in some cases the surface flaw may be sufficiently below the attachment weld that additional time may be required to grow the flaw up to the attachment weld. To provide a means to evaluate this time, a series of flaw evaluation charts for through-wall flaws were prepared.
Charts were prepared for each of the penetrations evaluated, for both the uphill and downhill locations, as shown in Figures 6-10 through 6-18. In each figure, the through-wall crack length is measured from the bottom of the nozzle itself, not the case. Note that in all the cases, the crack slows down significantly as it grows above the weld, due to the decreasing stress field. This provides further reassurance that axial flaws will not extend to a critical length, regardless of time, as the critical length exceeds 15 inches.
Head Vent The only flaw evaluation chart necessary for the head vent region is for flaws at and above the weld, since there is no portion of the head vent which projects below the weld. Figure 6-21 provides the projected growth of a part through flaw in the head vent just above the attachment weld. The growth through the wall is relatively rapid, because the thickness of the head vent is small.
6.4 CIRCUMFERENTIAL CRACK PROPAGATION Since circumferentially oriented flaws have been found at four plants (Bugey 3, Oconee 2, Crystal River 3, and Oconee 3), it is important to consider the possibility of crack extension in the circumferential direction. The first case was discovered as part of the destructive examination of the tube with the most extensive longitudinal cracking at Bugey 3, and the crack was found to have extended to a depth of 2.25 mm in a wall thickness of 16 mm. The flaw was found at the outside surface of the penetration (number 54) at the lower hillside location, just above the weld.
The circumferential flaws in Oconee Unit 3 were discovered during the process of repairing a number of axial flaws, while the circumferential flaw in Oconee Unit 2 and Crystal River Unit 3 were discovered by UT. Experience gained from these findings has enabled the development of UT procedures capable of detecting circumferential flaws reliably.
It is important to realize that a flaw would have to propagate through the penetration or the attachment weld, and result in a leak, before the outer surface of the penetration would be exposed to the water.
Cracking could then begin for an outside surface flaw. (This is believed to have been the case at all four plants in which circumferential flaws were found). This time period was conservatively ignored in the calculations to be discussed.
To investigate this issue completely, a series of crack growth calculations were carried out for a postulated surface circumferential flaw located just above the head penetration weld, in a plane parallel to the weld Flaw Evaluation Charts March 2004 Revision 1
6-4 itself. This is the only flaw plane which could result in a complete separation of the penetration, since all others would result in propagation below the weld, and therefore no chance of complete separation because the remaining weld would hold the penetration in place.
[]
The results of this calculation are shown in Figure 6-20, where it may be seen that the time required for propagation of a circumferential flaw to a point where the integrity of the penetration would be affected (330-350 degrees) would be at least 46 years. Because of the conservatisms in the calculations, as discussed above, it is likely to be even longer.
6.5 FLAW ACCEPTANCE CRITERIA Now that projected crack growth curves have been developed, the question which remains to be addressed is what size flaw would be acceptable for further service.
Acceptance criteria have been developed for indications found during inspection of reactor vessel upper head penetrations. These criteria were developed as part of an industry program coordinated by NUMARC (now NEI). Such criteria are normally found in Section XI of the ASME Code, but Section XI does not require inservice inspection of these regions and therefore acceptance criteria are not available.
In developing the enclosed acceptance criteria, the approach used was very similar to that used by Section Flaw Evaluation Charts March 2004 Revision I
6-5 XI, in that an industry consensus was reached using input from both operating utility technical staff and each of the three PWR vendors. The criteria developed are applicable to all PWR plant designs.
Since the discovery of the leaks at Oconee and ANO-1, the acceptance criteria have been revised slightly, to cover flaws on the outside diameter of the penetration below the attachment wveld, and flaws in the attachment weld. These revised criteria are now in draft form, but they are expected to be acceptable to the NRC, and will be used in these evaluations. The draft portions of the acceptance criteria will be noted below.
The criteria which are presented herein are limits on flaw sizes which are acceptable. The criteria are to be applied to inspection results. It should be noted that determination of the future service during which the criteria are satisfied is plant-specific and dependent on flaw geometry and loading conditions.
It has been previously demonstrated by each of the owners groups that the penetrations are very tolerant of flaws and there is only a small likelihood of flaw extension to large sizes. Therefore, it was concluded that complete fracture of the penetration is highly unlikely and, therefore, protection against leakage during service is the priority.
The approach used here is more conservative than that used in Section XI applications where the acceptable flaw size is calculated by putting a margin on the critical flaw size. In this case, the critical flaw size is far too large to allow a practical application of this approach so protection against leakage is the key element.
The acceptance criteria apply to all flaw types regardless of orientation and shape. The same approach is used by Section XI, where flaws are characterized according to established rules and then compared with acceptance criteria.
Flaw Characterization Flaws detected must be characterized by length and preferably depth. The proximity rules of Section XI for considering flaws as separate, may be used directly (Section XI, Figure IWA 3400-1). This figure is reproduced here as Figure 6-22.
When a flaw is found, its projections in both the axia I and circumferential directions must be determined.
Note that the axial direction is always the same for each penetration, but the circumferential direction will be different depending on the angle of intersection of the penetration with the head. The "circumferential" direction of interest here is along the top of the attachment weld, as illustrated in Figure 6-23. It is this angle which will change for each penetration and which is also the plane which could cause separation of the penetration tube from the head. The location of the flaw relative to both the top and bottom of the partial penetration attachment weld must be determined since a potential leak path exists when a flaw progresses through the wall and up the penetration past this weld. A schematic of a typical weld geometry is shown in Figure 6-24.
Flaw Evaluation Charts March 2004 Revision I
6-6 Flaw Acceptance Criteria The maximum allowable depth (a) for flaws on the inside surface of the penetration, at or above the weld is 75 percent of the penetration wall thickness regardless of the flaw orientation. The term a, is defined as the maximum size to which the detected flaw is calculated to grow in a specified time period. This 75 percent limitation was selected to be consistent with the maximum acceptable flaw depth in Section XI and to provide an additional margin against through wall penetration. There is no concern about separation of the head penetration from the head, unless the flaw is above the attachment weld and oriented circumferentially. Calculations have been completed to show that all penetration geometries can support a continuous circumferential flaw with a depth of 75 percent of the wall.
Axial inside surface flaws found below the weld are acceptable regardless of depth as long as their upper extremity does not reach the bottom of the weld during the period of service until the next inspection.
Axial flaws which extend above the weld are limited to 75 percent of the wall.
Axial flaws on the OD of the penetration below the attachment weld are acceptable regardless of depth, as long as they do not extend into the attachment weld during the period of service until next inspection. OD flaws above the attachment weld must be evaluated on a case by case basis, and must be discussed with the regulatory authority.
Circumferential flaws located below the weld are acceptable regardless of their depth, provided the length is less than 75 percent of the circumference for the period of service until the next inspection. Flaws in this area have no structural significance but loose parts must be avoided. To this end, intersecting axial and circumferential flaws shall be removed or repaired. Circumferential flaws at and above the weld must be discussed with the regulatory authority on a case by case basis.
Surface flaws located in the attachment welds themselves are not acceptable regardless of their depth.
This is because the crack propagation rate is several times faster than that of the Alloy 600 tube material, and also because depth sizing capability does not yet exist for indicatbins in the weld.
These criteria are summarized in Table 6-1. Flaws which exceed these criteria must be repaired unless analytically justified for further service. These criteria have been reviewed and approved by the NRC, as documented in references 7 and 8, with the exception of the draft criteria discussed above, for OD flaws and flaws in the attachment weld. These criteria are identical with the draft acceptance criteria now being considered for Section XI, for head penetrations.
It is expected that the use of these criteria and crack growth curves will provide conservative predictions of the allowable time of service.
Flaw Evaluation Charts March 2004 Revision I
6-7 Table 6-1 Summary of RA'. Head Penetration Acceptance Criteria (limits for future growth)
Axial Circ Location ag I of Below Weld (ID) t no limit t .75 circ.
At and Above Weld (ID) 0.75 t no limit Below Weld (OD) t no limit t .75 circ.
Note: Surface flaws of any size in the attachment weld are not acceptable.
- Requires case-by-case evaluation and discussion with regulatory authority.
af = Flaw Depth as defined in IWB 3600
= Flaw Length t = Wall Thickness Flaw Evaluation Charts March 2004 Revision I
6-8 70 , I ' I '
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0 50 100 150 200 Crack Half Angle Figure 6-1 Stress Intensity Factor for a Through-Wall Circumferential Flaw In a Head Penetration Flaw Evaluation Charts March 2004 Revision I
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Figure 6-2 Crack Growth Predictions for Axial Inside Surface Flaws Below the Attachment Weld by More Than 0.5 Inches- Nozzle Uphill Side Flaw Evaluation Charts March 2004 Revision I
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Figure 6-5 Crack Growth Predictions for Axial Inside Surface Flaws Near the Attachment Weld - Nozzle Downhill Side Flaw Evaluation Charts March 2004 Revision I
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Figure 6-6 Crack Growth Predictions for Axial Inside Surface Flaws Above the Attachment Weld - Nozzle Uphill Side March 2004 Flawv Evaluation Charts Flaw Charts March 2004 Revision I
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Figurc 6-7 Crack Growth Predictions for Axial Inside Surface Flaws Above the Attachmert Weld - Nozzle Downhill Side Flaw Evaluation Charts March 2004 Revision I
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0 1 2 3 4 5 6 7 8 9 10 11 12 perlod(yrs)
Figure 6-11 Crack Gronlh Predictions for Through-W'aD Axial Flaws Located in the Outermost CEDI Pcnetrations (51.5 degrees),
Downhill Side Flaw Evaluation Charts March 2004 Revision I
6-19 6.84 _
70' E 6.4 E 6.0___-_wl 5.2-5.0-0 52 3 4 5 6 7 8 perlod(yrs)
Figure 6-12 Crack Growth Predictions for Through-_all Axial Flaws Located in the 35.7 Degree Row of Penetrations - Uphill Side Flaw Evaluation Charts March 2004 Revision I
6-20 3.4-3.2 7
E I o 3.0 0
7
- a 2.8 0
z E
iB 2.6 S
U C
.m 2.4 -
a 2.0 1.8 _ _
0 I 2 3 4 5 6 7 8 9 period(yrs)
Figure 6-13 Crack Growth Predictions for Through-Wall Axial Flaws Located in the 35.7 Degree Row of Penetrations - Downhill Side Flaw Evaluation Charts March 2004 Revision I
6-21
- 6. 3 - l l l l l l l l l l t l l l l l l l [ t l 1
- 6. 1_I 1 1 V _
9 - I1 I171IF M 1 5.
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0 z _ I HII1 H E
.25.
.; 5.
V U 97 _ l l l l l l _ Wl I~5. 4.'
0 75_ l1 lLl 1 l § Ll +9=+0 4.
4.
4.
0 t 2 3 4 5 6 7 perlod(yrs)
Figure 6-14 Crack Growth Predictions for Through-Yall Axial Flaws Located in the 28.0 Degree Row of Penetrations - Uphill Side Flaw Evaluation Charts March 2004 Rcvision I
6-22 3.7 -T- -J -L- -
3.5 3.3- _ _
E S
- . . - . . . I - _ _ _
0 0 3.1I I II I I T I I I I I I I -II 0
CD II I 1 I I [- = 7 I 1 Z-[. I I I - 1I-F-I IT 0
z I2I91L I _ CEMWe d E
0 2.7 9------ __ __ _ _ ____/ ll__- [lg(1 01 C
c 2.7 - -
0 2.3_
l = .l___ 1 1= §[= lIr4^l 4.1 0 I 2 3 4 5 6 7 8 9 perlod(yrs)
Figure 6-15 Crack Growth Predictions for Through-Vall Axial Flaws Located In the 28.0 Degree Row of Penetrations - Downhill Side Flaw Evaluation Charts March 2004 Revision I
6-23 4.2 4.0 E 3.8 0
to
.R 3.6 0
z E 3.4 3
00 3.2 - l a
l l
~.W 0 I 2 3 4 5 6 7 8 period(yrs)
Figure 6-16 Crack Growvth Predictions for Through-Wall Axial Flaws Located In the 7.5 Degree CEDIN Uphill Side Flaw Evaluation Charts March 2004 Revision I
6-24 3.8 3.6 3.4 0_
E
_ 3.2 CV 1 3.0 E
0 4h 2.8 Z 2.6 2.4-2.2--
2.0 0 1 2
3 4 5 6 7 8
_9 perlod(yrs)
Figure 6-17 Crack Growth Predictions for Through-Wall Axial Flaws Located In the 7.5 Degree CEDI1 Downhill Side Flaw Evaluation Charts March 2004 Revision I
6-25 3.8 I 3.7 3.6, 3.5-3.4 3.3
- 0 3.2 0
In 3.1
~13.0 04d Z 2.9 2.8 2.7 2.6 2.5 2.4 2.3 2.2 2.1 2.0 0 1 2 3 4 5 6 7 8 period(yrs)
Figur3 6-18 Crack Growth Predictions for Through-Wall Axial Flaws Located in the Center Penetration Flaw Evaluation Charts March 2004 Revision I
6-26 1.0 0.9 0.8 0.69- - - - - .I /
0.7 IIlT 1llTTS V.8 1 A. I I I I I I I I II I _
CL
- ~0.4 0.4 S: XX 0.5 0 =_ 6 10 1 20 25 30 35! -
Period (years)
Figure 6-19 Crack Growth Predictions for Circumferential Surface Flaws Near the Top of the Attachment W~eld Flaw Evaluation Charts March 2004 Revision I
6-27 180, 150
' 120-C, -
0 I I I I I -
30- t t --- X X
_II ll 111IIIlL0 0 5 10 15 20 25 30 35 40 45 50 period (year)
Figure 6-20 Crack Growvth Predictions for Circumferential Through-Wall Cracks Near the Top of the Attachment Weld Flaw Evaluation Charts March 2004 Revision I
628 1.0 0.9 0.6 0.7-0,6 CL 0.6 0.5 0,
0.1 0.0 0 1 2 3 Period (years)
Figure 6-21 Crack Growth Predictions for Axial Inside SurfacenFaws- head Vent Flaw Evaluation Charts March 2004 Revision I
6-29 r~t-.2
- AX --
- Figure 6-22 Section XI Flaw Proximity Rules for Surface Flaws (Figure IWA -3400-1)
Flaw Evaluation Charts March 2004 Revision I
6-30 Figure 6-23 Definition of "Circumferential" Flaw Evaluation Charts March 2004 Revisio'n I
6-31 Figure 6-24 Schematic of Head Penetration Geometry Flaw Evaluation Charts March 2004 Revision I
7-1 7
SUMMARY
AND EXAMPLE PROBLEMS An extensive evaluation has been carried out to characterize the loadings and stresses which exist in the head penetrations at Palo Verde Units 1 and 2. Three-dimensional finite element models were constructed, and all pertinent loadings on the penetrations were analyzed [6A, 6B]. These loadings included internal pressure and thermal expansion effects typical of steady state operation. In addition, residual stresses due to the welding of the penetrations to the vessel head were considered.
Results of the analyses reported here are consistent with the axial orientation and location of flaws which have been found in service in a number of plants, in that the largest stress component is the hoop stress, and the maximum stresses were found to exist in the circumferential locations nearest and farthest away from the center of the vessel. The most important loading conditions were found to be those which exist on the penetration for the majority of the time, which are the steady state loading and the residual stresses.
These stresses are important because the cracking which has been observed to date in operating plants has been determined to result from primary water stress corrosion cracking (PWSCC). These stresses were used in fracture calculations to predict the future growth of flaws postulated to exist in the head penetrations. A crack growth law was developed specifically for the operating temperature of the head at Palo Verde Units 1 and 2, based on the EPRI recommendation, which is consistent with laboratory data as well as crack growth results for operating plants.
The crack growth predictions contained in Section 6 show that the future growth of cracks which might be found in the penetrations will be very slow, and that a number of effective full power years will be required for any significant extensions.
7.1 SAFETY ASSESSMENT It is appropriate to examine the safety consequences of an indication which might be found. The indication, even if it were to propagate through the penetration wall, would have only minor consequences, since the pressure boundary would not be broken, unless it were to propagate above the weld.
Further propagation of the indication would not change its orientation, since the hoop stresses in the penetration are much larger than the axial stresses. Therefore, it is extremely unlikely that the head penetration would be severed as a result of any indications.
If the indication were to propagate to a position above the weld, a leak could result, but the magnitude of such a leak would be very small, because the crack could not open significantly due to the tight fit between the penetration and the vessel head. Such a leak would have no immediate impact on the structural integrity of the system, but could lead to minor wastage in the ferritic steel of the vessel head, as the borated primary water concentrates due to evaporation.
Any indication is unlikely to propagate very far up the penetration above the weld, because the hoop stresses decrease in this direction, and this will cause it to slow down, and to stop before it reaches the outside surface of the head.
Summary and Example Problems March 2004 Revision I
7-2 The high likelihood that the indication will not propagate up the tube beyond the vessel head ensures that no catastrophic failure of the head penetration will occur, since the indication will be enveloped in the head itself, which precludes the opening of the crack and limits leakage.
7.2 EXAMPLE PROBLEMS The crack growth prediction curves in Figures 6-2 through 6-21 can be used with the acceptance criteria of Section 6.5 to determine the available service time. In this section, a few examples will be presented to illustrate the use of these figures. The example cases are listed in Table 7-1.
Example 1. For an axially oriented inside surface flaw, located below the weld, on the uphill side of penetration 53, first find the angle of the penetration in Table 1-1 or 1-2. The angle is 35.7 degrees. The crack growth curves of Figure 6-2 are appropriate and Figure 6-2 has been reproduced as Figure 7-1. In this case the flaw initial depth is 24 percent of the wall thickness, so project a line horizontally at a/t =
0.24, intersecting the crack growth curve. The service life is then determined as the time for this flaw to grow to the limit of 100 percent of the wall thickness, or approximately 2.5 years (labeled Service Life in Figure 7-1).
Example 2. In this case the flaw is identical in size to example 1, but located at the outside surface, in the penetration row at 28.0 degrees, and at the uphill side. The curve to use is in Figure 6-8. The determination of service life is illustrated in Figure 7-2, where we see the result is approximately 2.3 years.
Example 3. The axial flaw is at the weld, in penetration 89, whose angle can be determined from Table 1-1 or 1-2. The table shows that this penetration is in the row at 51.5 degrees, and not as deep as the flaw considered in example 2. It is oriented on the uphill side. The curve from Figure 6-4 is used to determine the service life. The flaw depth is 18 percent of the wall thickness, so project horizontally at this value to intersect the crack growth curve. The allowable service life is then determined as the time for the flaw to reach a depth of 75 percent of the wall. As shown in Figure 7-3, this time is approximately 2.2 years.
Example 4. This case is for a circumferential flaw which has been discovered above the weld, in penetration 22, which is in the 28 degree row, as seen in Tables 1-1 or 1-2. The appropriate figure for this type flaw is Figure 6-19, which has been reproduced as Figure 74, where the flaw size has been plotted.
The additional service life is obtained by plotting the flaw depth (a/t = 0.18) on the vertical axis and projecting horizontally to the crack growth curve. The service life is the time for the flaw to reach 75 percent of the vessel wall, which is approximately 12.2 years, as seen in Figure 74.
Example 5. Here we have postulated an axial flaw that will require two charts for its evaluation. The flaw has a depth of 3 mm, and is located on the inside of CEDM number 26, which has an angle with the head of 28.0 degrees. The flaw is 10 mm long, at 190 degrees, and its upper extent is 1.0 inch below the weld. The first step is to estimate the time required to grow to within 0.5 inch of the weld, and this is done in Figure 6-2, reproduced here as Figure 7-5A. The flaw will grow to within 0.5 inch of the weld when its depth reaches 45 percent of the wall thickness, and the time to reach that size is estimated as 3.0 years from Figure 7-5A. Then, use Figure 6-4, for flaws within 0.5 inch of the weld, and start with the Summary and Example Problems March 2004 Revision I
7-3 flaw depth at 45 percent. Figure 7-5B shows an additional 1.0 year of service, for a total of 4.0 years service.
Example 6. This case is an axial through-wall flaw whose upper-most end is 0.40 inches below the weld region in penetration 49, which is in the 35.7 degree row of penetrations, as seen in Tables 1-1 or 1-2.
From Figure 6-12 we obtain the appropriate curve for the crack growth prediction, and this is reproduced as Figure 7-6. This figure gives a service life estimate of approximately 2.6 years to grow to the bottom of the weld. This is illustrated in Figure 7-6.
It is clear from these examples that the most important figures for use in evaluating flaws in head penetrations are the surface flaw Figures 6-2 and 6-9 for axial flaws and 6-19 for circumferential flaws.
The figures which project the growth of through-wall flaws are valuable, but may be of limited practical use with the acceptance criteria. There is an important safety aspect to the through-wall flaw charts, however, in that they demonstrate that flaw propagation above the weld will be very limited.
Several guidelines are important to understand When using these charts.
- 1. If a flaw is found in a penetration nozzle for which no specific analysis was done, interpolation between penetrations is the best approach, when there is a uniform trend.
- 2. If a flaw is found in a penetration nozzle not analyzed, and there is no apparent trend as a function of nozzle angle, use the result for the penetration with the closest angle.
- 3. If a flaw is found which has a depth smaller than any depth shown for the penetration angle of interest, assume the smallest depth which was analyzed for that particular penetration, and make the time calculation with that flaw size instead of the actual flaw size.
Summary and Example Problems March 2004 Revision I
7-4 Table 7-1 Example Problem Inputs Example Vertical Radial Penetration Source No. Orientation Location Location Row Length Depth Figure I Axial-ID 1.2" Below Uphill 35.70 10 mm. 4 mm. 6-2 lWeld 2 Axial-OD 1.2" Below Uphill 28.10 10 mm. 4 mm. 6-8 Weld 3 Axial-ID At Weld Uphill 51.50 14 mm. 3 mm. 6-4 4 Circumferential Above Weld Downhill 28.10 8 mm. 3 mm. 6-19 5 Axial-ID 1.0" Below Uphill 28.10 10 mm. 3 mm. 6-2, 6-4 Weld 6 Axial 0.40" Below Uphill 35.70 _ 16.79 mm. 6-12 Through-Wall Weld Summary and Example Problems March 2004 Revision I
7-5 1
1 I I 1)Xt9 40L c11 rzlIXL TlIe Nc1l, 49~18l1 N.l.1;_AIg 0.9 X l 0l_ i 5.rdi 9 1-t - I I rl ceA I11.1e IY Dcog 0.8 i~~~~0 3E 3EEInX
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- -7 m 11 H e11 1 I 33 e 0.0 2.0 4.0 6.0 8.0 10.0 12.0 time (year)
Figure 7-1 Example Problem I Summary and Example Problems March 2004 Revis ion I
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Figure 7-2 Example Problem 2 Summary and Example Problems March 2004 Revision I
7-7 1
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U o 0.6 C) 0.
2 0.4-
- 0.3 u.
0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 time (year)
Figure 7-3 Example Problem 3 Summary and Example Problems March 2004 Revision I
7-8 1.0 I I I 0.9 0.8 0.7 111 1__I1_I I__II II ITI 1 I I1 1II I .I I 1 0.6 0.5 0.3 0.2 0.1 0.0 0 5 10 15 20 25 30 35 Period (years)
Figure 7-4 Example Problem 4 Summary and Example Problems March 2004 Revision I
7-9 1
0.9 0.8 It I~ III Xi -I! f I/I le I8I I I1 1 r I 1I 1l 111 l 11 H A 1 /1IJ I I I Igl: l11111 4 1
, 0.7 0.Y IIAII:7_
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0.0 2.0 4.0 6.0 8.0 10.0 12.0 time (year)
Figure 7-5A Example Problem 5 (See also Figure 7-5B)
Summary and Example Problems March 2004 Revision I
7-10 IN
- I. I II I A{ Irrr _ I 0.8 0.3 lllllllilIf IAI tI I I I I L+
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.20.2' 0.1' 0.
0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 time (year)
Figure 7-SB3 Example Problem 5 (See also Figure 7-SA)
Summary and Example Problems March 2004 Revision I
7-1l 7.0 1111 1 I 6.8 6.6 6.4 0
5.8 5.6 5.4 5.2-5.0 0 1 2 3 4 5 6 7 8 perlod~yrs)
Figure 7-6 Example Problem 6 Summary and Example Problems March 2004 Revision I
8-l 8 REFERENCES
- 1. Scott, P. M., "An Analysis of Primary Water Stress Corrosion Cracking in PWR Steam Generators," in Proceedings, Specialists Meeting on Operating Experience With Steam Generators, Brussels Belgium, Sept. 1991, pages 5, 6.
- 2. McIlree, A. R., Rebak, R. B., Smialowska, S., "Relationship of Stress Intensity to Crack Growth Rate of Alloy 600 in Primary Water," Proceedings International Symposium Fontevraud II, Vol, 1,
- p. 258-267, September 10-14, 1990.
- 3. Cassagne, T., Gelpi, A., "Measurements of Crack Propagation Rates on Alloy 600 Tubes in PWR Primary Water," in Proceedings of the 5th International Symposium on Environmental Degradation of Materials in Nuclear Power Systems-Water Reactors," August 25-29, 1991, Monterey, California.
4A. Crack Growth and MicrostructuralCharacterizationofAlloy 600 PWR Vessel HeadPenetration Materials, EPRI, Palo Alto, CA. 1997. TR-109136.
4B. Vaillant, F. and C. Amzallag. "Crack Growth Rates of Alloy 600 in Primary Water," Presentation to the EPRI-MRP Crack Growth Rate (CGR) Review Team, Lake Tahoe, NV, August 10, 2001.
4C. Vaillant, F. and S. Le Hong. Crack Growth Rate Measurements in Primary Water ofPressure Vessel Penetrationsin Alloy 600 and Wfeld Metal 182, EDF, April 1997. HT-44/96/024/A.
4C. Framatome laboratory data provided by C. Amzallag (EDF) to MRP Crack Growth Rate Review Team, October 4, 2001 (Proprietary to EDF).
4E. Cassagne, T., D. Caron, J. Daret, and Y. Lefevre. "Stress Corrosion Crack Growth Rate Measurements in Alloys 600 and 182 in Primary Water Loops Under Constant Load," Ninth InternationalSymposium on EnvironmentalDegradationofMaterials in NuclearPowerSystems-Water Reactors (Newport Beach, CA, August 1-5, 1999), Edited by F. P. Ford, S. M. Bruemmer, and G. S. Was, The Minerals, Metals & Materials Society (TMS), Warrendale, PA, 1999.
4F. Studsvik laboratory data provided by Anders Jenssen (Studsvik) to MRP Crack Growth Rate Review Team, October 3, 2001 (Proprietary to Studsvik).
4G. Bamford, W. H., "D. C. Cook Unit 2 Upper Head Penetration Crack Growth Determined from Inspection Data," Westinghouse Electric Report LTR-SMT-01-72, November 2001.
4H. Materials Reliability Program Crack Growth Rates for Evaluating Primary Water Stress Corrosion Cracking of Thick Wall Alloy 600 Material," EPRI MRP Report 55, to be published April, 2002.
5A. Newman, J. C. and Raju, I. S., "Stress Intensity Factor Influence Coefficients for Internal and External Surface Cracks in Cylindrical Vessels," in Aspects of Fracture Mechanics in Pressure Vessels and Piping, PVP Vol. 58, ASME, 1982, pp. 37-48.
References March 2004 Revision I
8-2 5B. Hiser, Allen, "Detemninistic and Probabilistic Assessments," presentation at NRC/Industry/ACRS meeting, November 8, 2001.
6A. Broussand, J.E., CEDM and ICI Stress Analysis for CE Reactor Vessels," Dominion Engineering Calculation No. C7736-00-0i, January 2002.
6B. Fleming, M.A., Palo Verde Unit 1 and 2 CEDM and Head Vent Stress Analysis, Dominion Engineering Calculation No. C7736-00-6, February 2002.
- 7. USNRC Letter, W. T. Russell to W. Raisin, NUMARC, "Safety Evaluation for Potential Reactor Vessel Head Adapter Tube Cracking," November 19, 1993.
- 8. USNRC Letter, A. G. Hansen to R. E. Link, "Acceptance Criteria for Control Rod Drive Mechanism Penetrations at Point Beach Nuclear Plant, Unit 1," March 9, 1994.
- 9. Fernandez, E., "Reactor Vessel Head Temperature Required for SCC Nozzle Flaw Evaluation,"
Arizona Public Service Report xxx, February 2002.
References March 2004 Revision I
A-l APPENDIX A ALLOWABLE AREAS OF LACK OF FUSION: WELD FUSION ZONES There are two fusion zones of interest for the head penetration nozzle attachment welds, the penetration itself (Alloy 600) and the reactor vessel head material (A533B ferritic steel). The operating temperature of the upper head region of the Palo Verde Units I and 2 is 314'C (597F), so the materials will be very ductile. The toughness of both materials is quite high, so any flaw propagation along either of the fusion zones will be totally ductile.
Two calculations were completed for the fusion zones, one for the critical flaw size, and the second for the allowable flaw size, which includes the margins required in the ASME code. The simpler case is the Alloy 600 fusion zone, where the potential failure will be a pure shearing of the penetration as the pressurized penetration tube is forced outward from the vessel head, as shown in Figure A-I.
The failure criterion will be that the average shear stress along the fusion line exceeds the limit shear stress. For the critical flaw size, the limiting shear stress is the shear flow stress, which is equal to half the tensile flow stress, according to the Tresca criterion. The tensile flow stress is the average of the yield stress and ultimate tensile stress of the material. The criterion for Alloy 600 at 318sC (604°F) is:
Average shear stress < shear flow stress = 26.85 ksi This value was taken from the ASME Code,Section III, Appendix I, at 6000 F.
For each penetration, the axial force which produces this shear stress results from the internal pressure.
Since each penetration has the same outer diameter, the axial force is the same. The average shear stress increases as the load carrying area decreases (the area of lack of fusion increases). When this increasing lack of fusion area increases the stress to the point at which it equals the flow stress, failure occurs. This point may be termed the critical flaw size. This criterion is actually somewhat conservative.
Alternatively, use of the Von Mises failure criterion would have set the shear flow stress equal to 60 percent of the axial flow stress, and would therefore have resulted in larger critical flaw sizes.
The allowable flaw size, as opposed to the critical flaw size discussed above, was calculated using the allowable limit of Section III of the ASME Code, paragraph NB 3227.2. The criterion for allowable shear stress then becomes:
Average shear stress < 0.6 Sm 13.98 ksi where Sm,= the ASME Code limiting design stress from Section III, Appendix I.
The above approach was used to calculate the allowable flaw size and critical flaw size for the outermost and center penetrations. The results show that a very large area of lack of fusion can be tolerated by the head penetrations, regardless of their orientation. These results can be illustrated for the outermost CEDM penetration.
Appendix A March 2004 Revision I
A-2 The total surface contact area for the fusion zone on the outermost head penetration is 17.4 in2 . The calculations above result in a required area to avoid failure of only 1.45 in2 , and using the ASME Code criteria, the area required is 2.79 in2 . These calculations show that as much as 83.9 percent of the weld may be unfused, and the code acceptance criteria can still be met.
To envision the extent of lack of fusion which is allowable, Figure A-2 was prepared. In this figure, the weld fusion region for the outermost penetration has been shown in an unwrapped, or developed view.
The figure shows the extent of lack of fusion which is allowed, in terms of limiting lengths for a range of circumferential lack of fusion. This figure shows that the allowable vertical length of lack of fusion for a full circumferential unfused region is 84 percent of the weld length. Conversely, for a region of lack of fusion which extends the full vertical length of the weld, the circumferential extent is limited to 302 degrees. The extent of lack of fusion which would cause failure is labelled "critical" on this figure, and is even larger. The dimensions shown on this figure are based on an assumed rectangular area of lack of fusion.
The full extent of this allowable lack of fusion is shown in Figure A-3, where the axes have been expanded to show the full extent of the tube-weld fusion line. This figure shows that a very large area of lack of fusion is allowable for the outer most penetration. Similar results were found for the center penetration, where the weld fusion area is somewhat smaller at 16.1 in2 .
A similar calculation was also carried out for the fusion zone between the weld and the head, and the result is shown in Figure A-4. The allowable area of unfused weld for this location is 84.8 percent of the total area. This approach to the fusion zone with the carbon steel head is only approximate, but may provide a realistic estimate of the allowable. Note that even a complete lack of fusion in this region would not result in rod ejection, because the weld to the tube would prevent the tube from moving up through the vessel head.
The allowable lack of fusion for the weld fusion zone to the head may be somewhat in doubt, because of the different geometry, where one cannot ensure that the failure would be due to pure shear. To investigate this concern, additional finite element models were constructed with various degrees of lack of fusion discretely modeled, ranging from 30 to 65 percent. The stress intensities around the circumference of the penetration were calculated, to provide for the effects of all stresses, as opposed to the shear stress only, as used above. When the average stress intensity reaches the flow stress (53.7 ksi), failure is expected to occur. The code allowable stress intensity is 1.5 Sm, or 35 ksi, using the lower of the Alloy 600 and ferritic allowables at 316'C (600 0F).
The results of this series of analyses are shown in Figure A-5, where it is clear that large areas of lack of fusion are allowable. As the area of lack of fusion increases, the stresses redistribute themselves, and the stress intensity does not increase in proportion to the area lost. These results seem to confirm that the shear stress is the only important stress governing the critical flaw size for the head fusion zone as well.
Appendix A March 2004 Revision I
A-3 Figure A-I Typical Head Penetration Appendix A March 2004 Revision I
A-4 Figure A-2 Allowable Regions of Lack of Fusion for the Outermost Penetration Tube to Weld Fusion Zone: Detailed View Appendix A March 2004 Revision I
A-5
-r- -'
nrl -
Figure A-3 Allowable Regions of Lack of Fusion for the Outermost Penetration Tube to Wcld Fusion Zone Appendix A March 2004 Revision I
A-6 Figure A4 Allowable Regions of Lack of Fusion for all Penetrations: Wlbld to Vessel Fusion Zone Appendix A March 2004 Revision I
A-7 Figure A-5 Allowable Regions of Lack of Fusion for the Weld to Vessel Fusion Zone March 2004 Appendix A March 2004 Revision I
B-i APPENDIX B: HOOP STRESS DISTRIBUTIONS IN THE CEDM NOZZLES This appendix presents the lengthwise hoop stress distributions on both the inside and the outside surfaces of the CEDM nozzle, as shown in Figures B-1 through B-9. The data in the plots begins at the bottom of the weld and extends towards the bottom of the nozzle. As can be seen, the stress peaks at or near the bottom of the weld location and decreases quickly as it approaches the bottom of the CEDM nozzle.
Based on experiences from other operating plants, the entire length of the CEDM nozzle may not be accessible for inspection. For Palo Verde Units 1 & 2, the accessible space, L, below the weld on the downhill side of the CEDM nozzle is assumed to be 1.3 inches. On the uphill side, the accessible space, L, below the weld is greater and is determined as below:
L downhill side = 1.3? = constant for all CEDM Nozzles L uphill side = 1.3? + ? H where ?H is the bottom-of-weld elevation differential between the uphill side and downhill side of a CEDM nozzle. It should be noted that the ?H values are different for CEDM nozzles with different intersecting angles as identified in Tables 1-1 and 1-2. The plots shown in this appendix also identify the assumed "Inspection Zone" for each CEDM nozzle.
Appendix B March 2004 Revision I
B-2 Hoop Stress vs Distance from Bottom of Weld, 0.0 Degrees (Uphill and Downhill) 80,000 70,000 4- InspectionZone lI____ I rflnllnnn I -MK ------- i -i I,. ....... _ I 1 40,000 I N Cn oD 30,000 2 0,000 x10,000 0
-10,000 X. . :-1-J-J-T-I ... I I. .I .... I ... I ..
-20,000 0.00 0.50 1.00 1.50 2.00 2.50 3.00 Distance from Bottom of Weld (in)
I--- Inside n Outside]
Figure B - Length)wise Hoop Stress Distribution in CEDM Nozzle, 0-deg Appendix B Marc 2004 Revision I
B-3 Hoop Stress vs Distance from Bottom of Weld, 7.5 Degrees (Downhill) 80,000 4-- Inspection Zone l 70,000 60,000
_ 50,000 a,
C' U,
40,000 I! 30,000 en L I I =V Oa
- 0. 20,000 10,000 0
-10,000
-20,000 . . . I. . . . . . . . . . . . . . . I . . .
O).00 0.50 1.00 1.50 2.00 2.50 3.00 Distance from Bottom of Weld (in)
I--Inside - Outsidel Figure B -2 Lengthwiise Hoop Stress Distribution in CEDAM Nozzle, Downhill Side, 75-deg Appendix B March 2004 Revision I
B-4 Hoop Stress vs Distance from Bottom of Weld, 7.5 Degrees (Uphill) 70,000 60,000 50,000 40,000 U,
30,000 0.
0 20,000 0 3 I1 I 10,000 0
-10,000
-20,000 l I . . . . . . j . . . . . . . . . . . . . . .
Cp.00 0.50 1.00 1.50 2.00 2.50 3.00 3.50 Distance from Bottom of Weld (in)
-- Inside - Outside Figure B-3 Lengthwise Hoop Stress Distribution in CEDM Nozzle, Uphill Side, 7.5-deg Appendix B March 2004 Revision I
B-.5 Hoop Stress vs Distance from Bottom of Weld, 28.0 Degrees (Downhill) 100,000I 10n0ection Zone 80,000 60,000-0.
(n
,~40,000 C,,
0.
0 20,000 0
-20,000 -'~-- ~-a 0.00 0.50 1.00 1.50 2.00 2.50 3.00 Distance from Bottom of Weld (in)
I--*-Inside v Outside I Figure B 4 Lengthiwise Hoop Stress Distribution in CEDMI Nozzle, Downhill Side, 28.0-deg Appendix B March 2004 Revision I
B-6 Hoop Stress vs Distance from Bottom of Weld, 28.0 Degrees (Uphill) 70,000 60,000 50,000 1 t' Inpcto Zon1-- 1-G 40,000 en Cn u.
30,000 l2 20,000 I k' o 10,000 0
I 00
-10,000
-20,000
-30,000 .. . . .... ..-. . ... . .I . .. II. .. .I aw.00 1.00 2.00 3.00 4.00 5.00 6.00 Distance from Bottom of Weld (in)
I s inside -nOutside Figure B-5 Lengthwise Hoop Stress Distribution in CEDM Nozzle, Uphill Side, 28.0-deg Appendix B March 2004 Revision I
B-7 Hoop Stress vs Distance from Bottom of Weld, 35.7 Degrees (Downhill) 100,000 Inspection Zone 80,000 60,000 -
U, 420,000 - . .. , .
0 0
-20.000-
-40,000L...L L.____________
0.00 0.50 1.00 1.50 2.00 2.50 3.00 Distance from Bottom of Weld (in)
I- Inside + Outside Figure B-6 Lengthwise Hoop Stress Distribution in CEDM No~zle, Downhill Side, 35.7-deg Appendix B March 2004 Revision I
B-8 Hoop Stress vs Distance from Bottom of Weld, 35.7 Degrees (Uphill) 60,000 -
50,000 Inspection Zone 40,000
°h 0L 30,000 u- 20,000 co 10,000 0.
o 0
-10,000
-20,000 -
-30,000 ...... .
0.00 1.00 2.00 3.00 4.00 5.00 6.00 Distance from Bottom of Weld (in)
I- Inside - Outside Figure B -7 Lengthwise Hoop Stress Distribution in CEDM Nozzle, Uphill Side, 35.7-deg Appendix B March 2004 Revision I
B-9 Hoop Stress vs Distance from Bottom of Weld, 51.5 Degrees (Downhill) 100,000 - I 14 Inspection Zone 80,000 -
60,000 -
a.
0 40,000 -
C-
, 20,000 0
0 0
-20,000
-40,000 .....................
0.00 0.50 1.00 1.50 2.00 2.50 3.00 Distance from Bottom of Weld (in)
I- Inside -uOutside I Figure B -8 Lengthwise Hoop Stress Distribution in CEDM Nozzle, Downhill Side, 51.5-deg Appendix B March 2004 Revision I
B-10 Hoop Stress vs Distance from Bottom of Weld, 51.5 Degrees (Uphill) 60,000 -
50,000 :
40,000
= 30,000 _
0.
' 20,000 -
_ 10,000 -
0 0
M -10,000 -
-20,000 _
-30,000 _
-40,000 l 0.00 1.00 2.00 3.00 4.00 5.00 6.00 7.00 8.00 9.00 Distance from Bottom of Weld (in)
I I inside w Outside I Figure B-9 Lengthwise Hoop Stress Distribution in CEDM Nozzle, Uphill Side, 51.5-deg Appendix B March 2004 Revision I