ML20093N674
| ML20093N674 | |
| Person / Time | |
|---|---|
| Site: | Vermont Yankee  |
| Issue date: | 07/30/1984 |
| From: | YANKEE ATOMIC ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20093N651 | List: |
| References | |
| PGE-1021, NUDOCS 8408020097 | |
| Download: ML20093N674 (247) | |
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1 VERMONT YANKEE NUCLEAR POWER CORPORATION 3 1984 Refuel Outage Augmented In-Service Inspection Program - Final Report Prepared by Yankee Atomic Electric Company 8400020097 840730
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TABLE OF CONTENTS I.- INTRODUCTION AND OVERVIEW
'II. LIST OF ENCLOSURES Enclosure 1 - Details of 1984 Augmented Inservice Inspection Program Enclosure 2 - Recirculation and Residual Heat Removal (RHR) Piping Flow Indication Evaluations and Weld Overlay Repairs Enclosure'3 - Vermont Yankee Reactor Coolant Leak Detection Provisions Enclosure 4 - Augmented Inservice Inspection ALARA Information Enclosure 5 - Recirculation Loop Piping Tearing Stability Analysis
.III. LIST OF ATTACHMENTS Attachment A - Vermont Yankee I&E Bulletin 83-02 Examination Program
~ Attachment B - Projection Image Scanning Techniqu s Information Attachment C - Improvements in Flaw Sizing Capability Attachment D - Vermont Yankee Reactor Coolant Leakage Limits -
IV. LIST OF TABLES Table I - Verinont Yankee Weld Joint Inspection Matrix (1983 - 1984) Table II - Details of UT Indications and Weld Joint Stresses Table III - Sumnary of Predicted Growth During the Next Cycle cf Operation Table IV - Disposition of UT Indications Table V - Comparison of 1983 to 1984 Inspection Results (Large Diameter Piping) Table VI - Vermont Yankee Stress Information Table VII - Comparison of 1983 to 1984 UT Program Table VIII - 1983 Flaw Summary Table II 1984 Flaw Summary
- Table I - 1984 Examination Restriction Summary TABLE OF CONTENTS V. LIST OF FIGURES Figure 1 - Recirculation System Weld Map Figure 2 - Weld Joint Numbers - Recirculation Header and Risers ,
Figure 3 - Weld Joint Numbers - Recirculation Loop A Figure 4 - Weld Joint Numbers - Recirculation Loop B
- Figure 5 - Weld Joint Numbers - RHR-A Figure 6 - Weld Joint Numbers - RHR-B Figure 7 - Weld Joint Numbers - RHR-C l
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k LI. ImamssouriON AND PROGRAM OVERVIEW
>; In response to'the NRC's Generic Letter 84-11, dated April 19, 1984 y.- (Reference (b)], Vermont Yankee Nuclear Power Corporation performed an augmented in-service reinspection of Recirculation and Residual Heat Removal system piping during the 1984 refueling outags.
This report contains our assessment of indications found in piping as a result of that inspection, as well as the repair and/or evaluation techniques utilized _to ensure recirculation system integrity for the next operating cycle. Contained within, as part of this report, are numerous Enclosures. Attachments, Figures, and Tables which provide the details of our 1984 Augmented ISI Program. The' report also includes comparisons of our 1984 program with certain aspects of our 1983 program. II. 313lql_ o An extensive ultrasonic examination was conducted on welds in the recirculation and residual heat removal systems in accordance with the provisions of of Generic Letter 84-11, except as discussed in Item 7 of Enclosure 1. Results are contained in Section V and detailed in Enclosure 2 to this report. o Weld overlays applied during the 1983 refueling outage were reinspected in accordance with the criteria of Generic Letter 84-11. The inspection included weld overlay integrity and bond of overlay to base metal. No indications were found in any_ overlay. o Wold Joint 32 which had a mini-overlay applied at the 1983 refueling outage was further overlayed. The overlay at this joint is now structural. o Weld Joint RHR-32-4, which had a small axial indication, was overlayed in accordance with appropriate criteria. o In the 1984 inspection, no flaw indications were found in the 12" diameter welds. All 12" susceptible welds have been examined at least once during either the 1983 or 1984 inspection. o In the 1984 inspection, only one small axial flaw was found in the 20" diameter welds. All 20" susceptible welds have been examined at least once during either the 1983 or 1984 inspection, o In the 1984 inspection, no flaw indications were found on the 24" RHR piping. No flaw indications were found in the 1983 inspection. o For the large diameter 22" header and 28" suction and discharge i piping welds with indications of Intergranular Stress Corrosion i Cracking (IGSCC), linear elastic fracture mechanics analyses have been conducted which show that flaw growth during the next cycle of operation is sufficiently small so as to permit operation without s repair. .Acc.eptance criteria for the evaluation are established in
% Rnclosure 2.to this report. All susceptible 22" piping has been
, ' inspected at least once during the 1983 and 1984 inspections.
Twenty five out of'33 28" susceptible welds have been inspected at least once in the 1983 and 1984 inspection. o In the 1984 inspection, the new flaw find rate was 18% (10 out of 57) as compared to SPE (34 out of 58) in 1983. These results confirm the assessment that the most susceptible welds were selected
'for inspection in 1983 and that the selection criteria are sound.
o 'The twenty-two weld overlays applied during the 1983 refueling outage are now all structural overlays of low carbon (.025%) and high ferrite content. The structural integrity of these overlays was demonstrated in our letters dated March 13, 1984 [ Reference (c)] and May 15, 1984 [ Reference (d)]. o- Weld joints with indications of IGSCC were conservatively evaluated. These evaluations indicate that the flaws are relatively short and shallow. Predicted flaw growth is very steall in the next cycle of operation. o Our pipe replacement contractor studied the drywell arrangement, identified interferences, and established plans for the 1985 pipe replacement. Utilizing this extensive pre-planning, an efficient l' replacsment effort with a minimum of personnel radiation exposure
.will be conducted.
III. ADDITIOWAL EFFORTS TO ADDRESS IGSCC CONCERNS
- o. We are replacing all Recirculation System and stainless steel Residual Heat Removal (RNR) System piping with seamless Type 316 nuclear grade stainless steel during the 1985 refueling outage.
- o Reactor Water Cleanup (RWCU) piping was replaced during the 1980 and
-1981 refueling outages with low carbon stainless steel.
o Susceptible Core Spray piping was replaced in 1977 with low carbon stainless steel, o Recirculation Bypass piping was replaced in 1976 with cast stainless steel. o Sections of other nonsusceptible piping systems are also under consideration'for replacement in 1985. These include:
- Remaining Core Spray piping which operates at (2000F, and
- Vessel bottom head drain line, o- Plant procedures have been revlced to require enhanced Reactor Coolant System leak rate monitoring, curveillance frequencies, and corrective actions consistent with these described in Enclosure 3 to this report.
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o A local leak detection system will be installed to monitor eight (8) 28" uninspected joints. This system is discucsed in Enclosure 3 to this report. IY. JUSTIFICATION FOR CONTINUED OPERATION The evaluation of the overlayed weld joints and aff teted large bore weld joints indicate that flaw growth is acceptable for all design conditione. The justification for operation for a second cycle of operation with weld overlays was provided in our letter, dated March 13, 1984 [ Reference (c)]. The results of this inspection confirm the basis pecsonted for the integrity of the overlays. Acceptance criteria for the analyses of large and small bore piping are established in Enclosure 2 of this report. These analyses demonstrate that there is no loss of design safety margin over that provided by the l rules for Class I piping in the ASME Boiler and Pressure Vessel Code, Section III. For these reasons, we conclude that the operation of Vermont Yankee for another cycle of operation is juntified. i
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b ENCLOSURE 1 DETAILS OF THE VERMONT YANKEE AUGNENTED 15-SERVICE INSPECTION PROGRAM TO ADDRESS INTERGRANULAR STRESS CORROSION CRACKING
- 1. . 1984 Insoection Techniques The Ultrasonic Examination Program utilised in completion of the Vermont Yankee 1984 refuel outage was planned and executed with the following as its primary attributes:
- a. Utilize both equipment and personnel demonstrated as qualified in accordance with the EPRI NDE Center course, "U.T. Operator Training for the Detection of IGSCC".
- b. Utilise equipment capable of producing "hard copy" examination results,
- c. Utilize equipment capable of manipulating examination data "off-line" allowing for analysis of data in a non-radiation environment.
- d. Provide redundant levels of evaluation techniques to compliment the basic discrimination techniques,
- e. Size detected and discriminated flaws in accordance with a program demonstrated capable of providing accurate through wall dimensions.
The EPRI NDE Center, UT operator training for planar flaw sizing was utilized to provide assurances in this respect. To this end an examination program significantly different than that used in 1983 (see Attachment A) was devised and implemented. The primary detection phase of the program was relegated to the P-Scan System as deployed by Independent Testing Laboratory (ITL) of Searcy, Arkansas (see Attachment B). The P-Scan System, used in conjunction with the MWS-2 semi-automatic scanner, provided the primary means for acquisition of detection and discrimination data. This system was coupled to standard, contact type 2.25 megahertz shear wave transducers. The primary detection angle used was 450 nominal with 520 nominal used for additional investigation and to a very limited extent to compensate for coverage limitations of the 450 probe. Individuals qualified through the EPRI NDE Center analyzed all P-Scan data and provided disposition. P-Scan dispositions were made primarily on spatial parameters all of which were compared to construction documentation and actual as-tuilt measurements obtained during pre-examination investigation. Calibration of the system is established using a 10% ID notch in a basic code calibration standard. Once basic reference is established P-Scan records the presence of all ultrasonic reflectors to approximately -64 Db of this 10% notch reference reflector. It is the ability to look for flaws far below normal recording levels which permits P-Scan to detect small or off-axis flaws without swiveling the search unit. P-Scan presents a high confidence for detection of all indications having any circumferential component as is the case with most m-
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E'~ IGSCC flaws. In EPRI tests P-Scan has demonstrated an ability to detect pure " axial" flaws without benefit of additional compensatory scans. The information sul ,11ed by P-Scan can be further evaluated by several different methods. -Fxaminers demonstrated qualified through the EPRI Program supply signal characteristic and echo dynamic information from basic A-Scan analysis as well as supportive full or half scale plots of specific areas. h ALN 4060, programmed to discriminate actual ICSCC may also be applied. This manually-applied system, programmed by EPRI, digitises and analyses received RF signals and provides a detailed analysis of this information. This equipment has again been demonstrated by personnel utilised at Vermont Yankee as a reliable means of discriminating IG8CC flaws from other perturbations at the weld root. Evaluation scans, whether with the ALN or A-Scan units utilized probe motions intended to detect additional " axial" flaws in welds requiring further evaluation. The WsY 70 probe, utilizing ID " creeping" waves was used to confirm flaws in a number of welds. This tool was only used in confirmation of flaws since it was felt that significant potential for false-negative flaw interpretations exists. m examination with a P-Scan System is limited to some extent by the inspection fixture. N P-Scan System is capable of inspection of pipe-to pipe and pipe-to-elbow configurations on both sides of the weld. On pipe-to-tee, pipe-to-valve, and pipe-to-pump, only one-side exams were performed. scan limitations are noted on the P-Scan data sheets. The areas not scanned with P-Scan were manually examined with qualified examiners where possible. All pipe-to-pipe and pipe-to-elbow configurations were scanned on both sides, with minor areas not scanned due to interference of integral supports or branch connections. All pipe-to-pump, pipe-to-valve, and pipe-to-tee configurations were completed on the pipe side only. h heavy sections of the fitting and necessary weld taper preluded any examinations in these areas. Because ultrasonic examination of the component side of the weld joint is not
.possible, no relevant ultrasonic information is available on the component side of the weld. Tables VIII, IX and X summarize both 1983 and 1984 examination restrictions.
E151BE A number of different techniques were utilised in establishing through-wall flew dimensions. h oe techniques fall into four primary categories. High Angle Longitudinal Beam Techniques (HALT), were utilised to integrate the outer 4/10's of the pipe wall for crack faces or crack tips which may have propagated to that region. Flaws found to be located in that region can be confirmed with a full-vee examination. Pulse Arrival Time Techniques (PATT), are utt11:ed to interrogate the remaining volume to determine crack tips below the 0.D. region. As a complement to PATT, a siellar satellite Pulse Observation Technique (SPOT), can be used to both observe the crack tip and relate its position to the root of the flew through observations of both pulses simultaneously.
d complementing the aforementioned techniques is the Multi-pulse Observation sizing Technique (NOST), which insonifies the entire pipe well with several angles and modes of sound beam. Through observation of several constant and changing pulse relationships, determinations of through-wall depth can be made. It-is the combination of these techniques and their ability to complement-one another in establishing a given flaw size which serves as the basis for the 1984 flaw sizing program. All personnel utilized in sizing flaws at Vermont Yankee were trained in accordance with the spRI UT Operator Training for Planar Flaw Sizing. Three individuals, providing the basis for all sizing calls, have been designated as having passed a final examination at EPRI, thus establishing their overall ability. All flawed welds were evaluated on a weld-by-weld basis as to the need to
. grind for flaw sizing. Grinding, when necessary, was completed to enhance flaw sizing.
- 2. 1983 Inspection Techniques I
!~ The examination program in 1983 consisted of total manual scanning and evaluation of the weld joints with methods qualified per Is Bulletin 83-02. These methods generally consisted of 1/2 vee path 450 shear wave examinations performed at 1.5 MHz. Supplemental examinations were performed using 600 shear wave examination techniques. Sizing was performed with dual element search units using the amplitude drop technique modified to include beam path geometry. Details of the 1983 exams were included in the 1983 I&E Bulletin 83-02 Final Report (Reference (e)). Attachment A to this report is a sununary of the 1983 examination.
Scan limitations in the 1983 program were noted on the data sheets. Pipe-to-congonent configurations were scanned on the pipe side only. The configurations were pipe-to-valve, pipe-to-tee, and pipe to pump. Pipe-to-pipe and pipe-to-elbow configurations were scanned from both sides with minor areas not scanned due to interference with integral supports. Because ultrasonic examination of the component side of the weld joint is not possible, no relevant ultrasonic information is
.available on the component side of the weld. >
In 1983, welds were scanned for axial indications in full scope oxams. Based on Vermont Yankee /NRC meetings, some large bore piping was scanned only at locations 900 apart. This was referred to as a cardinal point exam. These cardinal point exams only scanned for circumferential indications. The extent of the weld exams, including those with only cardinal point examine, are included in Table VIII to this report. Cardinal point exams were performed by selecting four areas of the weld joint, 12" in length centered at 00, 900, 1800, and 2700 around ;
.the joint. This was an initial sample of 48" of inspection. The i inspections were on both sides of the weld joint where possibIe, as described above. When an indication was noted that extended beyond the original scan length, the examination was continued to determine the full extent of that indication.
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- 3. Wold Overley Examination Technique l-h examinations following the weld clad repair at Vermont Yankee ;
consisted of the following ;
- a. Clad Bond Examination i
- b. Clad Integrity Examination i f
The clad bond examination consisted of a straight beam examination from ' the clad surface. h principal area of concern is= the clad-to-base I i metal-Interface. A 3/8" diameter flat-bottomed hole at the clad-to-base metal interface of a clad calibration standard was used as the reference reflector. Scanning sensitivity were at least +6 dB gain. The acceptance criteria was 50% of the 3/8" diameter hole reference signal or i any indication with an area less than the reference reflector at ; reference sensitivity. This examination revealed no relevant indications. i The clad integrity examination consisted of an angle beam inspection of the clad and clad-to-base metal interface. The inspections were performed with a KB Aerotech gamma series, dual element, 3/8 x 3/4", 450, refracted longitudinal beam search unit, at r frequency of 1.5 Isis. N reference reflectors were 1/16" diameter side-dellied holes, i
< The holes were positioned such that an examination zone contained weld l metal, weld-to-base mots.1 interface, and base metal. The calibration was '
-performed on welded clad pipe of essentially the same material as the ,
piping components in the plant. These calibration standards were L manufactured in such a way as to duplicate the weld process and surface
'- conditions of the actual repairs. Overlay calibration standards were i fabricated at the minimum and maximum overlay thickness anticipated, thus ,
bracketing the overlays examined. Acceptance criteria were any indication less than 50% of the reference reflector. No cracks, lack of penetration, or lack of fusion were allowed. No elongated indications , greater than 1/4" were permitted. The results did not reveal any relevant indications in the overlay or overlay-to-base metal interface.
- 4. Flaw Evaluation Summary o UT Indications were found at welds in Vermont Yankee piping as shown in Table 2-1 of Enclosure 2. Indications in the recirculation system welds were evaluated and found to be acceptable for another 14-month !
fuel cycle without repair. h axial indication at weld joint RNR-32-4 was repaired by weld overlay as described in Enclosure 2.
'c UT indications were' evaluated for acceptability by fracture mechanics analyses for crack growth and ASME Section II, IW5-3640 flew size limits. End-of-cycle limits were used which included a 2/3 factor on Table IWB-3641-1 flaw sizes and included thetaal and prior repair l shrinkage stresses in the IWB-3641-1 evaluation, o Weld overlay thickness sizing is in accordance with ASME Section XI f c Table IWS-3641-1. The thicknesses recommended for circumferential'
> flaws include an additions 1 load factor margin of 1.5 for flaws less than 1800 in length. These factors are in addition to the safety factor of 2.773 incorporated in the above Section XI Table. This , ! approximately corresponds to the inclusion of thermal stresses in {
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sisias everlays for less than 1808 This methodology was used to l epply a full structural weld overlay to a previous repair at Wold i Joint 32 of the Boeirculatten system, i I e- The width of the wold overlay for circumferential flaws is computed as 1.5 (Et)1/2 The width for extal flows is centered on the antal flew length and extend 0.5 (24)1/2 past each end of he indication.
- l
' S. Camellance with 10CFRSO General Desian Criteria Appendia F to'the Vement Yankee Final safety Analysis Report (FsAs) I describes how Ve ment Yankee satisfied the Asc General Design Criteria ['
(appendix A to 10CFR50) when the plant was constructed. ( This discussion will demonstrate that IGsCC, weld overlays and/or the use [ of flawed pipe analysis have no effect on Vemont Yankee's compliance ; with the General Design criteria. f Of the General Design Criteria identified in Appendix A to 10CFR50, this ; discussion will address only those criteria that could be affacted by the ( esistense of 10 SCC in the reactor coolant pressure boundary. l Griteriga_M, "The reactor coolant pressure boundary shall be designed, fabricated, erected, and tested so as to have an entremely low probability of abnormal leakage or rapidly propagating ? failure, and of gross rupture." , Makhed of censliance - The potential for IosCC will increase the probability that flows may exist in reactor [ coolant piping. Vemont Yankee compensates for this
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probability by increasing the frequency of inspection. The existence of IosCC flows does not necessarily result $ in system leakage. Many studies, supported by actual : eperating emperience, have shown that IosCC flaws will i tend to arrest before penetrating the pipe well. l f Between 1943 and 1984, 90/113 weld jointe have been i inspected with very sensitive ultrasonic examination
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techniques. Indications in unrepaired joints are very shallow and have resulting very low probability of , propagating (see Enclosure 2). j structural weld overlays have been applied to weld joints ! whleh do not pass Asus Code flew evaluation criteria. ! These overlays are performed with a material which is ; s lemune to IosCC propagation. ! Flow evaluations on unrepaired joints were performed to i the criteria recosmonded in Generle Letter 84-11. Several I additional conservatisms were applied, as described in l
' Enclosure 2. Large margin between oised flows and I seceptable flaws exists for one additional operating eyele. !
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, p p' Me perfossed a Tearing stability Analysis of the j Boeirculatten System which demonstrated that meoumed !
through us11 flaws, having lengths which would result in ! o readily detectabl6 leake, were stable ueder ASNs Level D leads. Integrity is shown to estet with ample safety asesine. t t Fleued welde that are repaired by weld overlay or flowed f wolde that de not require repair because of compliance ! with 11 alt lead analysis techniques satisfy the desian margine required by the ASIE code. Thus, they are no more probable to emportance rapidly propesating failure or grees rupture then an unflowed weld. : Thus, we senclude that 00C 14 le satisfied. ! t d Critarian 30 "Ceapenents wh'.ch are part of the resetor coolant pressure boundary shall be designed, fabtleated, erseted, and ; tested to the highest quality standarde practical. IIeans j shall be provided for detecting, and te the entent j prestleal, identifying the loestion of the source of : reester seelant leakage." j 1 g hed af ' 18 3 - Testing for 108CC is perfossed C using ultrasenas testing methods that have been shown to [ have a hash degree reliability in detecting and sisins 1930C fleus. The detalle of the methode are described ) eleeuhere in thle report. In addition, se described in anslesure 3, we have Laplemented more restrictive leakage . deteetten previstene and will install a meisture sonettive l tape system en elskt (s) 2s" uninspeeted weld joints. [ t
, g Tip.s. we eenelude that 90C 30 le estisfied, critarian 31 "The reester ecolant pressure beuneary shall be deelsned i with suffletant margin to assure that when stressed under operating, maintenance, testing, and postulated aceldent j eendittenet (1) the boundary behaves in a nonbrittle {
manner, and (2) the probability of rapidly prepasaking , freature le minimised. The design shall refleet eeneideration of servlee temperatures and other eendittene L of the boundary material under operating, maintenance, ; teetlas, and postulated sealdent ennd!tlene and the uncertainties in detessinings (1) materlat properttee, f (2) the effects of irradiation en motorial properties (3) l reeldual, steady state and trenelent stresses, and (4) l slee of fleus." l Isothed af camellanne - staintese steel is very duettle ! material that le highly resistent to brittle behavior and o rapidly propesating freeture. The limit lead analyste i
;4 technique asseunts for the presence of flows ar.d the )
effect they may have on structural integrity. Ceay11ance ! with limit lead analyste requirements ensures that [ f
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unstable flaw propagation'will not occur. The tearing stability analysis discussed in Rnelosure 5 to this report demonstrates that even if a significant flaw should 1 propagate through unil, the plant leakage limits will initiate corrective action well before the potential for < unstable flaw propagation develops.
- Thus, we conclude that ODC 31 is satisfied.
Critarian 32 " Components whleh are part of the reactor coolant pressure boundary shall be designed to permit: (1) periodic inspection and testing of important areas and features to assess their structural and leaktight integrity, and (2) en appropriate material surveillance program for the reactor pressure vessel." Mathed of Cameliance - The application of weld overlays precludes the ability to inspect the pipe weld under the overlay. However, since the weld overlays are structural overlays only the integrity of the weld overlays needs to be inspectable. As described elsewhere in this report, the weld overlays are inspectable, and the requirements for inspection of overlays as defined by NRC Generic Letter 84-11 have been perfomed. Thus, we conclude that GDC 32 is satistled. In summary, the existenee of IOSCC in Vemont Yankee does not reduce Vermont Yankee's engliance with the General Design Criteria of Appendiu & to 10CFRSO.
- 5. Rania far Ianreved Inseestion Ensults The basis for better inspection results in 1984 is twofold. The validated eneminer and examination procedure certainly provide the most significant ressen for better performance. All personnel performing detection, diserimination, and sising, who are Level II or III, are qualified on an individual basis using the EPRI-NDR Center qualification programs. The 1943 exams used a team approach to the qualification preesse, rather than qualification on an individual basis.
The multifaceted examination procedure, using p-scan examinations, as us11 as manual evaluations, and the ability to compare results with 1983 eneminattens provide the second enjor reason for better 1984 examination results. The use of p-Sean equipment has allowed a greater examination work scope within the limite of available personnel and personnel exposure. Thus, more detail can be provided by the quellflod manual examiners doing indlestion evaluations. A more detailed comparison of key enemination vertables is included as Table VI to this repert. In contrast, espesure levels in 1983 were such that total exposure limited euen scope to the point that only cardinal point scans for aircumferential indlestions were performed on a large portion of large bore piping.
In summary, the 1984 examinations are performed with people who are better trained, with the training validated by performance exams. The equipment provides a greater amount of detail and a larger work scope, within the limits of total exposure. t Attachment C to this report provides Eraphic representation of improved sising capability based on training and qualification of personnel.
- 4. Maid Joint Samslina Criteria The sospling program was developed using four criteria for examinations o criterien 1 Inspect all unrepaired welds with IGSCC.
I o Criterien 2 r. Inspect all overlayed riser weld joints with previous cracks longer then 10E of pipe circumference. The inspection is for bond integrity with the base metal and a weld metal examination. o Criterien 3 Inspeet 205 of previously inspected joints without indications in each pipe sise (minimum of 2 weld joints). o criterien )
-Inspect 205 of the previously uninspected welds in each pipe size ,
(minimum of 4 weld joints), i The table below depicts the criteria and the first and second additional samples if defects were found in the oristnal sample. I k If defeats are found in the additional sample, then all remaining welds i of that oise in that line should be examined. ~ j The original sample has been expanded to include those we,1ds defined in ' Criteria 4 - 20", 22", and 20" lines. ; The table aise deplets the total number of welds in each criterion. I Criterien 1 - All unrepaired wolds with IGsCC indications in 1983. ! Original Sample - 28" - 64 28" - 58 ! 28" - 1A 24" - 59 28" - 2 22" - let i I t 28" - 9A 22" - 36B 28" - 45A 22" - 308 ! 28" - 15A 24" - RHR-31-1 i l Total - 12 Wales i
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Criterion 2 - Overlayed welds which had indications over 10% of circumference - overlay bond and weld integrity exams. 12" - 30* 12" - 16* 12" - 24 12" - 54 12" - 33* 12" - 23* 12" - 29 12" - 18 12" - 42* 12" - 36* 12" - 32 12" - 45* 12" - 50* 12" - 35 12" - 20* 12" - 53* 12" - 51
- Denotes sweepolet to riser welds.
Criterion 3 - 1983 inspection - no indications: 20% or minimum of 2 welds. { 13" 2q" 22" 28" Original sample 51A RHR-32-4 23A 98 54A 30A 17 41 44 Total population 18 1 3 4 Criterion 4 - Remaining welds - not inspected 20% or 4 welds minimum. 19" 12" 24" 28" ! criginal sample RHR-32-2 16A RHR-30-1 15 RHR-32-F1 47 RHR-30-3 15B RNA-32-5 48 RH2-30-9 27 RHR-32-1 36A RHR-30-10 26A 61 ! Total population 6 6 20 23 First Additional Sample RHR-32-6 238 17A RHR-32-7 49 15C 4 5A 178 second Additional sample 5 6 8 26 56
- 7. Justification for awaanded Sample of 28" Pipe Wolds !
During the initial inspections during the 1984 refueling outage, ICSCC indications were detected. The sample population was increased as required by NRC Ceneric Letter 84-11 and as described in our letter dated July 6, 1984 (Reference (j)). An additional 28-inch weld was found to have a flew in the second sampla population. Strict interpretation of Generic I,etter 84-11 would require that all remaining 28-inch welds f I (there are 13) be inspected. A third sample of five welds wee selected l for inspeetten. The five welds were selected to ensure that at least one j et each susceptible weld leestion in either loop was inspected. The a resulte of that sample showed that one weld use found with a small flew I (appreminately 3 inshee long). Vermont Yankee does not believe that { additional inopostlene are userented. Our justification le provided j below. j t During the inspections this year, the maximum cumulative flew length in ! any weld le less then 25 percent of the pipe etreumferences average flew lengths are in the range of 1 to 4 inches. The neutnum flaw depth detected in any flew to aces than 30 percent of well thlehneses average , flew depths are 15 to 20 percent of well thicknees. The weld oaeyle ! population was selostes to ensure that the welde meet probable to contain f 10s0C were inspected first. The sampling criteria addressed carbon j eentent, servlee stresses, and fabrication-related repaire. The 1egitioney of the selection eriteria le supported by the fact that even [ though additional flaws were detected in the expanded samples, the else [ of the flows le less then the first sample. Of the total length of all : weld joints inspected, less than two (2) percent of the total contained i flaws. Vessent Yankee believes there is suffielent evidence to suggest l that the remaining a welde de not contain a flew lorser than the first 67 i welde. I
'The safety significance of this situation can be shown to be neglislble, es follows ,
- 1. Using limit lead analysis techniques, the allowable end of eyele flaw f depth for a flow 25 percent of circumference is in escers of 100 j percent of well thlehnees ;
- 2. gesed en limit lead analysis, the allowable flew length for a 30 !
.persent deep flaw le in eseese of 100 percent of pipe etreumference8 l
- 3. The limit lead eve 1vattene aseeunt for potential flew growth during the nemt e,.reting eyeiei j
- d. The limit lead evolustions maintain full Asus Code design margine and ;
- 5. WRI studsee have demonstrated that a multiply-flowed pipe system hae [
et least the same margin of safety as a singly-flowed system.' (In estustity, it een be shown that the multiply-flowed system has ineressed esegin, but no eredit le taken for that.) ; thus, Veseent Yankee believes that further inspeettens will result in no f inesease in estety margin. Ineroesed inspeettens will have a significant ! redselegical Lopeek on the inspeetten pereenne1. speeifteally, we !
-estimate that an addittenal 22 man-ren would be empended to inspect the
- 1est eight'(g) 23" weld jelate, of which 16 men-ren would be to the UT personnel. There is appreuimately 12 man-res remaining among the ,
available Ut personnel, whteh is insuffielent to easytete the esame. ! further, it would take a week and a half to two wehe to obtain i addittenel qualified pereennel. gesed on the prinelpet of ALARA, we ! holieve ne further supesure to inspeet the remaining 23" welde le i jwetified. i lo- ! I f
( As a compensatory measure for the lack of inspections, Vermont Yankee ' proposea to continue in effect the more stringent " unidentified leaksge" limits adopted by management directive during the last operating cycle. 4 These limits are discussed in detail in Enclosure 3 to this report. i , purther, a local leak detection system will be installed to monitor eight (8) 28" uninspected weld joints. This system is also discussed in Enclosure 3. Finally, Verimont Yankee has conducted a tearing stability analysis on the recirculation system. This analysis includes consideration of the recently identified potential for low fracture toughness in austenitic submerged are weldsents. Even with these very conservative toughness considerations, it was demonstrated that structural stability was assured, even assuming a flaw of sufficient size to result in 10 gym leakage (five times the control limit). The results of this analysis are
~
provided in Enclosure 5 to this report. e
m V 9 ENCLOSURE 2 RECIRCULATION AND RESIDUAL HEAT REMOVAL (RHR) PIPING FLOW INDICATION EVALUATIONS AND WELD OVERLAY REPAIRS d
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ENCLOSURE 2 g. 'I Report No. SIR-84-018
.p : Revision 0 SI' Project No. YAEC-04 July 1984 f
Recirculation and RHR Piping Flaw Indication Evaluations and Weld Overlay Repairs at Vermont Yankee Prepared by Structural Integrity Associates San Jose, California 1
' Prepared for Yankee Atomic Electric Company
- Prepared by: Ed h M Date: 7/ 7/f '
J. F Cbpeland
^
Prepared by: # 7 h ' Date: 7/&7/8//
. S. S./ an / 0 Reviewed by: // If AA Date: ?!Z7 d P. C. Riccqpdell5 " '
Approved by: d d M At h h [ Date: 7/2-7[ry Projef M ager/
'J . Ff o land /
/'
f Date: 7!27 (h
' FTTncTpalysocfate
P. C. Riccardella , _ _ . STRUCTURAL INTEGRITY *wn f
r', SIR-84-018
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REVISION CONTROL SHEET SECTION PARAGRAPH (S) DATE REVISION REMARKS All All 7/27/84 0 Initial Issue l l l i i t ii j STRUCTURAL j INTEGRITY vx i.
.1
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' TABLE OF CONTENTS Page
.1.0 Introduction . . . . . . . . . . . . . . . . . . . . . . 1 2.0 Details of UT Indications and Weld Joint Stresses ... 3 3.0 Fracture Mechanics Evaluation ............. 4 3.1 Factors on Results ............... 4 3.2 Crack Growth Evaluation ............. 4 3.2.1 Applied Stresses ............. 5 3.2.2 Residual Stresses . . . . . . . . . . . . . 5 3.2.3 Crack Model and Crack Growth Analysis . . . 5 3.2.4- Crack Length Growth . . . . . . . . . . . . 6 3.3 Allowable Flaw Size Determination ........ 6 3.3.1 Circumferential Flaws . . . . . . . . . . . 6 3.3.2 Axial Flaws . . . . . . . . . . . . . . . . 8 4.0 Results and Disposition of. Indications . . . . . . . . . 10 4.1 Acceptable Indications . . . . . . . . . . . . . . 10 4.2 Re' pairs ..................... 10 5.0 Weld Overlay Repairs . . . . . . . . . . . . . . . . . . 11 5 .1 - Factors on Results . . . . . . . . . . . . . . . . 11 5.2 Repair Design Methodology and Results ...... 12 6.0 Summary ......................
14 7.0 References ...................... 15 1 Appendix A - Typical Crack Growth Result .......... A-1 Appendix B -- Allowable Flaw Depths ............. B-1 7
. STRUCTURAL IIi [ INTEGRITY usocm j
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t d List of Tables Table Title Page s
' 2 -l' Details of UT Indications and Weld Joint Stresses .. . . . . . . . . . . . . . 16 3-1. Summary of Predicted Crack Growth for a 13-Month Operating Period . . . . . . . . . 17 3 Circumferential Flaw Size Limits (ASME
-Section XI, Table IWB-3641-1) . . . . . . . . 18
.3-3 Axial' Flaw Size Limits (ASME Section XI, Table IWB-3641-3) . . . . . . . . . . . . . . . . . . . . 19
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, Figure Title Page I3 Residual Stress Curves Used In Analysis and Supporting Experimental Data . . . . . . . . . . . . 20
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^3-2= Crack-Growth Rate Curves Used in Analysis and' Supporting Data ........-........
21
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L3-3 Stress Distribution in a Cracked. Pipe - Basis
. for Net Section Collapse Equations . . . . . . . . . 22 3-4 'Circumferential Flaw Size Limits Versus Stress . . . 23 J4-17 Vermont Yankee Weld 1A Flaw Evaluation . . . . . . . 24 5-1 Weld Overlay Design for Circumferential Flaws ... 25
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STRUCTURAL lNTEGRITY x.ee.
1.0 Introduction During the current outage at Vermont Yankee, circumferential indications were observed by ultrasonic (UT) inspection at weld joints (listed in Table 2-1) of the recirculation system. An axial indication was observed at RHP weld joint 32-4. The indications are all located in weld heat affected zones and are judged to be intergranular stress corrosion cracking (IGSCC) in nature. Fracture mechanics evaluations of the observed indications were performed in
- accordance with References 1 and 2, in order to determine any need for repairs. This assessment is for one fuel cycle (14 months) of operation. The crack growth evaluation was performed for as-welded residual stresses plus operating stresses and shrinkage stresses from previous weld overlay repairs. The flaw was conservatively grown in depth as a 3600 circumferen-tial crack.
Results from the above evaluation are compared to the end-of-cycle (E0C) allowable flaw depth in ASME IWB-3640 (Ref. 1). A factor of 2/3 was placed on allowable E0C flaw depth to account for flaw sizing uncertainties, and thermal stresses and shrinkage stresses from previous overlays were con-sidered as primary stresses in the IWB-3640 evaluation, to acccunt for any possible low weld metal toughness. With these conservatisms included, a comfortable safety margin exists for - the indications observed in the recirculation piping welds at the end of one fuel cycle (14 months). Based on the above fracture mechanics evaluations, a weld overlay repair was
. performed on weld joint RHR 32-4.
~
Weld overlays have zbeen successfully implemented on Type 304 austenitic stainless steel pipe welds for the repair of intergranular stress corrosion cracks. (IGSCC) in boiling water reactors. These repairs consist of depositing a 3600 band of Type 308L weld metal (with controlled ferrite) on the pipe. outer' diameter and over the indication. 1 STRUCTURAL
% INTEGRITY usocca.
b-
,n; = g ..- -- gy. The weld overlay 1 repair serves a number of purposes toward restoring the
. piping integrity: (1) structural reinforcement, (2) compressive residual stresses on the pipe inner diameter due to weld shrinkage, and (3) an IGSCC-resistant weld metal pressure boundary. Consideration of welding residual stresses is notl necessary in cases of through-wall cracks. Structural reinforcement requirements are computed based upon the net sec+ ion collapse criterion '(NSCC), as justified by elastic-plastic fracture mechanics analy-sis-(tearing modulus) to show that the NSCC is the controlling mechanism for
-fracture.
Weld overlay repairs of IGSCC have been performed on a large number of welds, including Vermont Yankee. Weld overlays were designed for Vermont Yankee as' reported in this document, and bound the worst hypothetical cases (through-
-wall cracks in highly stressed weld joints).
2 STRUCTURAL
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p. 9 2.0' Details of UT Indications and Weld Joint Stresses , The size,-location-and orientation of UT indications in the Vermont Yankee
. piping ; are ' presented in Table 2-1, along with the corresponding applied stresses'at the weld joint. UT indications are described in Enclosure A.
- Stresses for this analysis were taken from References 4 and 5, and are based onthe;pipingdesignstressreport(Ref.3).
.Thelvalues of weld shrinkage stresses _(Ref. 6) from previous weld overlay
-repairs'at Vermont Yankee have been determined (as shown in Table 2-1) and
.are added to residual plus operating stresses for flaw g'.owth calculations.
These" shrinkage stresses were .also included with primary and thermal
- stresses in the determination of allowable flaw sizes.
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\ s 3 STRUCTURAL INTEGRITY a,ocer. 6
( h 3.0 Fracture Mechanics Evaluation 3.1 Factors on Results Certain factors were employed to account for uncertainties in flaw sizing and weld metal toughness in the analysis. References 7 and 8 recommend using a factor 6f 2/3 on the end-of-cycle (EOC) flaw size limit from ASME Table IWB-3641-1 (Ref.1). Reference 8 recommends that thermal expansion stresses be considered as a primary stress in the use of IWB-3641-1 for end-of-cycle flaw size limits. Reference 8 also uses a conservative 3600 circumferential crack model to predict growth in the crack depth direction, whereas such cracks are
.usually less than 3600 All these recommendations were included in this analysis.
3.2 Crack Growth Evaluation
. Crack growth _was computed using the methodology of Reference 8. This methodology is based on growth by intergranular stress corrosion cracking (IGSCC) under sustained loading during operation, and has been found to be consistent with cracking experience (Ref. 8).
Contributions df fatigue loading to crack growth are considered negligible
-in this_-case. A major contributor to crack growth is the welding residual stress, which enters heavily into sustained loading calculations, but has only a mean stress effect in fatigue cycling. Furthermore, the available data suggests that the contribution of the conventional design operating transients to crack growth is negligibly small (because they comprise such a relatively small fraction of the life) 'and that most of the crack growth occurs under the. nominal steady-state operating conditions (Ref. 8). Small fluctuations in operating stresses are negligible from a fatigue standpoint
.(Ref. 8). Thus, in large diameter piping the f atigue crack growth associated with design loading histories is very small, and crack growth will be due primarily to IGSCC (Ref. 8).
4
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STRUCTURAL p lNTEGRITY anoccu
{. 3.2.1 Applied Stresses Pressure, dead-weight and thermal stresses, for the weld joint being studied, were employed with shrinkage stresses from previous repairs and with the following residual stress distribution, crack model and crack growth law to predict crack growth in the pipe thickness direction. These applied stresses are tabulated in Table 2-1, as discussed previously, and are all conservatively treated as through-wall membrane tensile stresses in the crack growth analysis. The residual stress distribution through the pipe wall is described below. 3.2.2 Residual Stresses The best estimate axial residual stress distribution as shown in Figure 3-1, was used with the above applied stresses for crack growth calculations. This residual stress curve is consistent with Reference 8. Due to the non-linear nature of the residual stress profiles, a third order polynomial equation was used to curve fit the test data and analytical data. The third order polynomial equation has the form Stress = Co + C1 X + C2 X2+CX3 3 (1) A least square curve-fit procedure was used to determine the coefficients in Equation 1, where X is location in the wall thickness direction. 3.2.3 Crack Model and Crack Growth tnalysis o A _ full 3600 circumferential crack on the pipe inside surface was con-servatively assumed for crack growth computations in accordance with the practice of Reference-8, even though the observed indications were finite length. Accordingly, the fracture mechanics crack model was a 3600 circumferential crack in a cylinder with a thickness to radius ratio (t/R) of 0.1. -The best estimate severely weld sensitized crack growth law (Figure 3-
- 2) was combined with the preceding stresses and crack model, and numerically integrated to predict flaw depth as a function of time. Results are shown in Appendix A.
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f 3.2.4 Crack Length Growth Crack growth was computed conservatively in the length direction by assuming a constant growth rate of 0.00025 in/hr (2.19 in/yr) at each crack end (Refs. [ 9 and 10). Table 3-1' presents a summary of predicted crack growth for the Vermont Yankee UT ' indications in a 14-month operating period.
'3.3 ' Allowable Flaw Size Determination Based on the concept of net section plastic collapse (Ref.11), ASME Section XI IWB-3640 contains end-of-evaluation period allowable flaw depths for circumferential flaws for normal and upset operation conditions for aus-
~tenitic piping material (Table 3-2). Results for Vermont Yankee are shown in Appendix B.
3.3.1 -Circumferential Flaws Briefly, the net section collapse theory for circumferential flaws considers a given crack of lengthl(corresponding to a crack angle 20), and depth a, with nominal primary membrane stress Pm and nominal primary bending stress Pb at force and moments' equilibrium in the longitudinal direction and with stress at the net section location equal to the flow stress of the material, (fr. This equilibrium is illustrated in Figure 3-3, along withf, the shift in neutral axis of the pipe due to loading the cracked pipe. The following equations are derived from the above concepts (Ref. 11): for 6 +g z F j, FT-da/t) - (Pm/#f) (2) Pb= [2 sing - (a/t) sin G] (3) 6 7 STRUCTURAL hI INTEGRITY usxms
.g; for e +/3 > 7 g/il-a/t-Pn/67) 7 2 - a/t P 2Gf b ".y (2-a/t) sin /3 (5) Using the above equations, the critical flaw size (k &G) can be determined through iteration.
~
The above basis leads to the formulation of the allowable end of evaluation-
' period-flaw depth for circumferential flaws for normal operating conditions in' ASME Section XI Table IWB-3641-1 (shown in Table 3-2). Several assumptions are used in obtaining Table IWB-3641-1. The primary membrane
.stressiis~ essentially due to operating pressure. It is assumed to be equal to half of the allowable' stress intensity (Sm). A safety factor of 2.773, from the consideration of the minimum margin on primary membrane stress as required by the ASME Code and the safety margin for pure bending in pipes, is used.
. An arbitrary cut-off at 75% for the allowable crack depth to thickness ratio is made for~ conservatism. Also, for crack lengths larger than 1800, a full
'circumferential crack. solution is. conservatively used, as illustrated in Figure 3-4.
-Itl can be seen _ that the allowable flaw dcpth in Table IWB-3641-1 (Table 3-2
-and Figure 3-4)- depends on the piping stress rat-:o (Pm + P )/Sm. b In accordance with the latest NRC guidelines (Ref. 8), service level A thermal expansion stres'ses are . included in.the stress ratio calculation to account for possible low weld metal toughness. Therefore, weld joint stresses due to
-q pressure, dead-weight, seismic (0BE), and thermal and prior repair shrinkage
-(shown in Table 2-1) were used to compute corresponding stress ratios with an Sm of 16.95 ksi. for austenitic stainless steel at 5500F.
7 STRUCTURAL INTEGRITY u.mn, L. .
f bly Stress ratios corresponding to the above stresses are shown in Table 2-1 and l -
- were used with Table IWB-3641-1 to determine the allowable end of cycle flaw depths. A factor of 2/3 is also included (Ref. 7 and 8), in the IWB-3641-1 results to establish the final allowable flaw size, in order to account for flaw sizing uncertainties. This results in an allowable flaw depth of 50% of the pipe wall thickness in all cases for the circumferential indications.
3.3.2' Axial Flaws
. Table 3-3 presents the. allowable end of evaluation period flaw depth to thickness ratio (a/t) for axial flaws for normal operation conditions. This
. table is formulated through emperical results for a through-wall flaw in pipe and extended to part-through-wall axial cracks with a curvature correction factor. Although an arbitrary flaw depth limit of a/t = 0.75 is shown in Table 3-3, Section XI IWB-3642 permits flaw acceptance based on applied
- stress and maintaining a factor of at least three against failure stress.
Thus, the source equations (Ref. 2) for Table IWB-3641-2 (Table 3-3) can be solved,_as shown below, to demonstrate a factor of at least three against
. plastic collapse for a through-wall. axial flaw, 0.5 in, long, in Vermont Yankee 20 in. RHR piping.
(Th
- 3 Sm (6)
M.=[1+4Rtm1.61f2]1/2 where: 6'h = hoop stress at failure 3 Sm = flow stress, with Sm = 16.95 ksi. at 5500F (from ASME Section III) M = curvature correction-factor E ]= through-wall axial flaw length R = pipe radius (10 in. for RHR) tm
- pipe min, wall thickness (1.095 in, for RHR)
The hoop stress, due to a design pressure of 1250 psi., in the 20 in. RHR pipe is given by: d'h = (8) tm 8 O STRUCTURAL I NTEG RITY ..w.u,
w m AR ' The above equation results : a design hoop stress of 11,416 psi. for the design pressure. Thus, the predicted failure stress should be at least three
- times the design hoop stress, or 34,248 psi., to give the required safety factor. Substituting a failure hoop stress of 34,248 psi. into Equation (6) o and solving for / in Equation (7) gives a through-wall axial flaw length of
.5.72 in. This flaw length of 5.72 in. is significantly above realistic axial flaw lengths at piping welds, which are generally limited to the weld heat affected zone width of less than 0.5 in.
-The above equations can also be solved for the more realistic through-wall axial flaw length of/= 0.5 in. to show a predicted hoop stress at f ailure of 50,617 psi., 'and a corresponding safety f actor of 50,617/11,416 = 4.43.
Another way to look at this is that the material flow stress could be reduced as low as'(3/4.43) (3 Sm) = 34,436 psi to still maintain a safety factor of three against plastic collapse for a 0.5 in. long through-wall axial flaw.
'~Thus, it can be seen that such an axial flaw is not a safety issue. The use of a thin weld overlay, simply to arrest further crack extension and to act as' a seal against potential leakage, is considered adequate in this case.
This conclusion is consistent with Reference 8, which states that analysis
~
can be use>l to justify long-term operation with weld overlays for relatively short axia.1 cracks. .This is true because errors on crack depth measurement or flaw growth predictions will lead at worst to relatively small leaks, which wilI be easily detectable long before the crack can grow long enough to cause failure (Ref. 8). 9 STRUCTURAL lNTEGRITY w.n.
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.4.0 ~Results and Disposition of Indications ,
- A typical change in flaw size for cne fuel cycle (14 months) is presented in Figure 4-1 and is compared to.the final allowable flaw size as described in
' the preceding sections. The disposition of the Vermont Yankee UT indications is. summarized:in the following paragraphs.
4.1 ' Acceptable Indications Even with'the preceeding conservatisms considered in the analysis, there is still a comfortable margin (from the allowable flaw depth of a/t =0.5) at the end of one fuel cycle of operation (14 months), for the circumferential indications observed in the recirculation system. Thus, these indications _ -are judged to be acceptable without repair for one 14-month period of operation. 4.2 Repairs The axial ' indication Jat weld joint RHR 32-4 of the RHR system was dispositioned to be repaired by weld overlay, as described in the preceding section. A. full structural weld overlay (through-wall, 3600 flaw assumed) was also applied to weld joint 32 of the recirculation system to add further margin to a previous repair. The weld overlay design methodology is described in the following section. 10 - STRUCTURAL INTEGRrr( mmu
5 5.0 Weld Overlay Repairs Weld cverlay repairs for Vermont Yankee were designed as described in the following paragraphs, using IWB-3640 (Ref. 1) as a basis, and including appropriate conservatisms and factors. 5.1 Factors on Results
. Weld . overlay repairs were designed based on measured indication length assuming a flaw completely through the original pipe wall thickness (through-wall crack). This is conservative, based on the measured finite depth of UT indications (as shown in Table 2-1), but is done to account for
. any uncertainties in depth sizing. It also avoids the need to consider further defect growth as influenced by overlay induced residual stresses in the pipe (an effect that gives further margin).
As in the evaluation of UT indications for acceptability, thermal stresses were considered in.the primary stress ratio for determining IWB-3640 table flaw limits. This is approximately equivalent to multiplying the primary stress' ratio (pressure, dead-weight and OBE) by a factor of 1.5 to account for potential low weld metal toughness. Overlay thicknesses corresponding to a load factor of 1.5 or the inclusion of thermal stresses in the primary stress ratio were used for piping repairs when flaws of less than 1800 length are assumed for overlay sizing. This is based on past experience with weld overlays, and results in reasonable thicknesses for the corresponding loads. This load factor is an extra level of conservatism to guard against any possible lower' toughness in existing butt welds and is in addition to the margin of 2.773 on loads in Table 2-1. The f actor of 1.5 becomes less
= important for larger flaws where more loading is supported in the controlled tougher. weld overlay material. For flaws greater than 1800 length, no credit
-is taken for the existing butt weld, and a load factor of 1.0 is considered adequate for the overlay deposited by the Tungsten Insert Gas (TIG) welding process.- For flaws less than 1800 length, the smaller overlay, based on
~s izing with actual flaw length and a load factor of 1.5 or by sizing with a 1800-3600 length with a load factor of 1.0, may be used.
-11 STRUCTURAL
=j[ lNTEGRITY xo,
k, f 5.2 Repair Design Methodology and Results
. The overlay designs are based on net section collapse theory, as described in
.the section of-this report on allowable flaw size determination, and includes the preceding conservatisms.
~
10verlays for circumferential flaws were designed in thickness to meet the
. flaw limits of' Table IWB-3641-1 (Table 3-2), and include additional factors,
! as discussed above. Two principal effects of the overlay are considered in using this table to size thicknesses of overlays: (1) reduction.in pipe stresses due to increased wall thickness from the overlay and (2) reduction. of the flaw depth / wall thickness, a/t, ratio as a result of the overlay. A maximum a/t of 0.75 is permitted. A weld overlay design miminum thickness of 0.2 in was computed for weld joint 32, based on an assumed through-wall 3600 flaw in the 12 in. diameter, 0.53 in. thick pipe, and an enveloping primary stress ratio of 0.522. The steps.followed for the weld overlay thickness sizing, for through-wall cracks.in.the unrepaired pipe, are:
- a. 0btairi allowable a/t using the given (Pm + Pb )/Sm ratio from Table 2-1.
- b. Reduce (Pm + Pb )/Sm proportional to the increase of wall thickness t due to the addition of assumed weld overlay thickness At.
- c. Recalculate the allowable a/t corresponding to the adjusted Pm +
b .Pb , due to the weld overlay. If the. calculated a/t from step c is larger than the allowable value given in Table 3-2 for the adjusted stress level, repeat steps b and c by increasing At until the solution converges to the allowable a/t at the adjusted stress level. If the calculated a/t is significantly smaller than the allowable value for the adjusted stress level, then the overlay thickness can be accordingly reduced.
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- 5 INTEGRITY amm 5.
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=
;The. minimum width of the weld overlay was computed as 1.5 (Rt)1/2 where R is
~the pipe radius and t is the pipe thickness. This is based on extending the overlay.a sufficient distance-from the crack that the effects of the local
' discontinuity _(crack) on the structural reinforcement are dissipated. This
, weld overlay design is shown in Figure.5-1.
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6.0 Summary
- 1. UT_ indications were found at welds in Vermont Yankee piping as shown in Table 2-1. Indications in the recirculation system welds were evaluated and ' found to be acceptable for another 14 month fuel cycle without repair. The axial indication at weld 32-4 was repaired by weld overlay as described in this report.
.2. UT indications were evaluated for acceptability by fracture mechanics analyses for crack growth and IWB-3640 flaw size limits. End-of-cycle limits were used which included a 2/3 factor on Table IWB-3641-1 flaw sizes, and included thermal and prior repair shrinkage stresses in the IWB-3641-1 evaluation.
- 3. - Weld overlay thickness sizing is in accordance with ASME Section XI Table IWB-3641-1. The thicknesses recommended for circumferential flaws include .an additional load factor margin of 1.5 for flaws less than 1800 in length. These factors are in addition to the safety factor of 2.773 incorporated .in the above Section XI table. This approximately corresponds to the inclusion of thermal stresses .in sizing overlays for less than 1800 This methodology was used to apply a full structural weld overlay to a previous repair at weld joint 32 of the recirculation
. system.
- 4. .The width of the weld overlay for circumferential flaws is computed as 1.5 (Rt)l/2 The width for axial flaws is centered on the axial flaw length, and extends 0.5 (Rt)1/2 past each end of the indication.
14 7 STRUCTURAL INTEGRITY ssoc.,s a
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'7.0 -References
. 1. ASME' Boiler and Pressure Vessel Code, Section XI, W83 Edition.
- 2. Ranganath, S., and Norris, D.M., " Evaluation Procedure and Acceptance Criteria for Flaws in Austenitic Steel Piping", Draf t No.10, sponsor-ed by subcommittee on Piping, Pumps, and Valves of PVRC of the Welding Research Council, July 1983.
- 3. ' Design Report, Recirculation System, 22A2615, June 1970, SIA File YAEC-02-200.
- 4. .NUTECH Computer Run AFT-ACFV, June 13, 1983.
- 5. Transmittal, Recirculation Piping Stresses, R. E. White to J. F.
Copeland, SIA File YAEC-04-102.
- 6. Telecon, R. E. White to J. F. Copeland, SIA File YAEC-04-102.
- 7. .Dircks, W. J., " Staff Requirements for Reinspection of BWR Piping and Repair of Cracked Piping", NRC Policy Issue SECY-83-267C, Nov. 7,1983.
, 8.' NUREG-1061, Report of Pipe Study Group, Draft Report for Comment, received May 17, 1984.
- 9. -" Guidelines for Flaw Evaluation and Remedial Action for Stainless Steel Piping Susceptible to Intergranular Stress Corrosion Cracking," Draft Report No. SIR-84-005, Rev. 3. Prepared by Structural Integrity Asso-ciates'for:EPRI BWR Owner's Group, April 13, 1984.
- 10. Bickford, R. l.. , et al, "Non-destructive Evaluation Instrument Sur-veillance Test on 26 in. Pipe," EPRI NP-3393, January 1984.
- 11. Ranganath, S., and Mehta, H. S., " Engineering Methods for the Assessment of Ductile Fracture Margin in Nuclear Power Plant Piping," Elastic-Plastic Fracture: Second Symposium, Vol. II, ASTM STP 803, C. F. Shih and J. P. Gudas, Eds., ASTM, 1983.
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- v. ..
j -} , i TABLE'2-1 Details of' Circumferential UT Indications and Weid . Joint 5 tresses Wall
~
Outer Stresses (psi.)' ( Shr.+Th.+ UT Indication Pipe Weld Dia. . . Thickness - Shr.+Th. P+DW Length Sire C n _ . _ . .t ISI 100. (in.) t(in.) Shrin6 age Therm. Press. +08E/16,950 1 att M DW 00E +P+0W f/ circ. 28 ELB0W IA 28.169 1.2 0 2122 5954- 1177 155 9253 0.555 0.22 5 0.057 28 ELBOW 2 28.169 1.2- 0 917 5954'- 635 371 7506 0.465 0.15 2 0.023 28 ELB0W 9A 28.337 1.29 600 393- 5534 259. 476 6786 0.428- 0.20- 5 0.057-28 TEE ISB 28.169 1.18 200 1887 6053 464 2164 8604 0.635 0.18- 3 0.034 28 ELBOW 26A 28.169 1.15 0 958 6210 637 636 7805 0.498 0.15 19 0.22~ 28 ELB0W 27 28.169 1.15 0 735 6210 475 182 7420 0.449 0.19 4.5 0.051 28 VALVE 61 28.337 1.25 150 325 5711 83 1158 6269 0.438 0.20 24 0.27 28 PUMP 59 28.337 1.34 200 389 5330 54 1221 5973 0.424 0.21 18 0.15 28 TEE 65A 28.337 1.29 700 537 5534 461 1149 7232 0.495 0.23 15 0.17 22 TEE 16A 21.879 1.05 1190 2303 5614 1417 758 10,524 0.666 0.20 7 0.10 22 TEE 16B 21.879 1.03 1190 2909 5718 1422 758 11,239 0.708 0.12 1 0.015 22 CAP 30B 21.879 1.04 0 0 5666 0 0 5666 0.334 0.20 20 0.30 28 ELB0W 17B 28.169 1.27 250 537 6023 196 227 7006 0.427 0.20 6 0.068 22 VALVE 49 21.879 1.09 2400 1136 5408 546 230 9490 0.574 0.22 1.5 0.022 28 PUMP 6 28.337 1.26 200 435 6068 173 1320 6876 0.484 0.17 3 0.034 22 CAP 23B 21.879 1.09 0 0 5408 0 0 5408 0.319 0.27 6 0.087 v,) F(n Tm $ oC aQ . 45 l? i
- _- _ _ - .-_ _ .-w.-
C u_[. .
- 7. .
TABLE 3-1 SUMARY OF PREDICTED CRACK GROWTl! FOR A 14-MONTH OPERATING PERIOD CIRCUMFERENTI AL FLAW SIZE PlPE~ START FINAc START FINAL FINAL SIZE WELO DEPTH DEPTH LENGTH LENGTH LENGTH (IN.) COMPONENT ISI NO. A/T A/T (IN.) (IN.) L/ CIRC. 28 ELBOW 1A 0.22 0.300 5 9.745 0.111 28 ELBOW 2 0.15 0.236 2 6.745 0.077 28 ELBOW 9A 0.20 0.261 5 9.745 0.111 28 TEE 15B 0.18 0.238 3 7.745 0.088 28 ELBOW 26A 0.15 0.246 9 23.745 0.270 28 ELBOW 27 0.19 0.268 4.5 9.745 0.111 28 VALVE 61 0.20 0.261 24 28.745 0.327 28 PUMP 59 0.21 0.245 18 22.745 'O.202 28 TEE 65A 0.23 0.285 15 19.745 0.224 22 TEE 16A 0.20 0.294 7 11.745 0.171 22 TEE 16B 0.12 0.254 1 5.745 0.084 22 CAP 308 0.20 0.244 20 24.745 0.35 28 ELBOW 178 0.20 0.254 6 10.745 0.122 22 VALVE 49 0.22 0.287 1.5 6.245 0.091 28 PUMP 6. 0.17 0.247 3 7.745 0.088 22 CAP 238 0.27 0.286 6 10.745 0.015 STRUCTURAL lNTEGRITY oems t
- TABLE 3-2 CIRCUMFERENTIAL FLAW SIZE LIMITS (SECTIONXI,IWB-3641-1)
TABLE IW8-M11 ALLOWABLE END-0F EVALUATION PERIOD FLAW DEPTH 1 TO THICKNESS RATIO FOR CIRCUMFERENTIAL FLAWS - NORMAL OPERATING GNCLUDING UPSET AND TEST) CONDITIONS p, + p, antio of F)se Lagth. /,, to Mye CA;.". .. e (Note 01
- s. ((n07 (t2.07 0.5 (ID-sy')
anste a)) a.o c.1 M 67 a2 c.3 asas o.4 er more 1.5 64) (4) - H) (4) (4) (4) 1.4 8.75 0.40 .177 0.21 0.15 (4) (4) 1.3 8.75 0.75 ,flo 839 062 7 1TJ 022 0.19 1.2 0.75 0.75 , MJ 8.S6 8.40 .373 0.32 SJ7 1.1 0.75 0.75 ,777 4.73 0.51 .$ 0.42 0.34 1.0 8.75 9.75 M G.75 SA3 .f"'O 031 0.41 e.9 e.75 e.75 .7f o.75 e.73 .WJ 0.59 0.47 8.4 47S 0.75 ,7f 8.75 8.75 .717 0 68 0.53 [ 9.7 0.75 0.75 ,7f E7S 6.75 75" 9.75 0.54 s 04 e.7S 0.75 ,7f 0.75 4.75 .7f 0.75 0.63 L 80TES: (1) Flan depth = a,for a surface flaw 2a.for a sutwrface flew t = nomweltNckness Linear intefpetation is permess41e. O) P = pnmary membrane stress p,= pnmary wnding stress 3,= allowab6e desiyi streu intsemity (in accordance wth Section Ill) 01 Caturnference bened on vervenal pipe dear 1eter. M) NFS 3514.3 shall be imod. 18 p i STRUCTURAL
> lNTEGRITY umwn
T'
. TABLE 3-3 AXIAL FLAW SIZE LIMITS (SECTIONXI,IWB-3641-3) )
s TABLE IWB-3641-3 ALLOWABLE ENtMW-EVA:.UATION PERIOD FLAW DEPTH 1 TO THICKNESS RATIO FtMt AX1AL FLAWS-IIORMAL OPERATING QNCLUDING UPSET AND TEST) CONDITIONS Ilandwnemanal Flus Langth Deste O)) (t,MT) Sams static 12.0 er 0.0 0.5,61.0 2.0 3.0 4.0 10 40 7.0 8.0 9.0 10.0 11 0 Greatee 3 (Note (2)) o.67 o 66 c.65 c.64 e 64 c 64 (4) s o.4 c.75 e.75.Wo.75 c.75 c.74 c.7o o.6s c.75 a.75,Wo.7s 0.72 c.65 c.61 c.59 c.se o.57 c.se o.ss (4) (4) (4) s.5 c.ss o.51 c.49 e.es e.47 (4) (4) (4) (4) (4) s.6 e.75 e.7s .70.75 e.64 0.75 7f o.73 0.53 c.44 c.4o 0.78 OJ7 W) (4) (4) (4) (4) (4) 0.7 S.75 8.75. Tag o.62 0.4e o.32 0.28 c.26 (48 (4) (4) (4) (4) (9) (4) 0.4 0.75 8.1s o.14 (4) H) (4) (4) (4) (4) (4) 6.9 8.75 0.7eM.42 S.23 0.17 W) .13 H) (4) (4) (4) M) H) M) (4) (4) (4) (4) (4) 1.0 14)
. NOTES:
(1) Fiem depth a, for a swface fleur aa, for a edewface Asw Unser intevyeleten is s .. "- (2) 5 tress retie = g, where P = masiamen preeswe fee normat operating constions O = menunal eutaier emmeter of the pipe f = nominalthsceness 3,= aBomatie design stress intensity (in accordance with Secten Ill) ' 01 /, = and-of.evolunten period for flee length a = nominalreden of the pipe f = nominal 10mctness (4) IWB.3514.3 shall be used. 19 - ;; S TRUCTURAL l
,', lNTEGRITY wce.
- traide won Outsid2 wsM E
40
, Curve B I t (Upper Bound) e E O y
10 - D j 00( d 8 7 go -- gb 1-10 -20
- eUC UA i
.o --
O e o t, Curve A
-30 - # (Best Estimate) l l I I I I I I I 40 O 01 02 03 04 05 06 07 08 09 10 Normal:ec Crack De;;m (a/t)
LEGEND: O GE 26 in MP944-1 D GE 26 in 1HS1 ref. pipe (4 azimutts) A ANL 26 in ND 944-2 (2 azimuths) A ANL 26 in KRB l e ANL 20 in T.-114 4 SWR 128 in (3 azimuths) Structural Integrity Curves ! Used in Analysis l l FIGURE 3-1 RESIDUAL STRESS CURVES USED IN ANALYSIS AND SUPPOR11NG EXPERI-f MENTAL DATA 20 t
- STRUCTURAL I INTEG RITY .w .. i.
.-s *
'
- Upper Bound Furnace Sensitiu d) y da/dt = 5.65x104(K)3.07
/
,,-A ,
- @ Best Estimate (Weld Sensitizet) m A da/dt = 2.27x10-8(K)?.26
/ T .Sa h S2 T e t t_0 I av S. t u. '.8,2, ,* C. 2 ee=
...Ia,
. 1 _
d St%S'TI21 d' 91M'8 2
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/ V h M%l064 .talt O St.t m*1332 2 atal 8 a* 35C i sc- 03
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& OfelatallO 33t*8. De a. evOtutLO'a*g paa 0 la aE 8 **378
. 3
=
8 8 8 i g.? I 8 M g 3 5 e M N g gtmess eartsees 1"v.s empKI FIGURE 3-2 CRACK GROWTH RATE CURVES USED IN ANALYSIS AND SUPPORTING DATA (FROM EPRI MP-2472) ) 21 __ STRUCTURAL
]
. ' lNTEGRITY . ...
9 NOMIN A L STRE $$ IN TMt UNCA ACa t D
. SECTiON OF PiPt (lrLACK LENGTH la 3#4 '****
* +-- PLour stress.o
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l'Q,_. R
?= k
*- i
- I s'
N- Q ______ A -
. j i + l s ; p ____._________i.
ZL l i Pm NEUTRAL , AMil STRE S8 DistmieUTION IN THE CRACKED SECTION AT TMt POINT OF COLLAPSE Pm* APPLit0 IMMeu ANE STRESS IN UNCR ACKE D MCTION lb = APPLIED NNDeNG ST AG N IN UNCRACKt O MCTION FIGURE 3-3. STRESS DISTRIBUTION IN A CRACKED PIPE -- BASIS FOR NET SECTION COLLAPSE EQUATIONS (REFERENCE 11) 22 STRUCTURAL lNTEGRITY .e .....
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.6 /,0 /./ 42. /,3 44 Ar i. 6
$7RCSS RATIO, L +f S,
FIGURE 3-4. CIRCUMFERENTIAL FLAW SIZE LIMITS VERSUS STRESS 23 -
} -
lNTEGRITY STRUCTURAL mceu
lIfi f A k
- 1. 0 ' - - .
(, r.
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RO.Q M TABLE IW3-369/-/ LIMIT
~
(Pg P, + %ERM/C/G,o
~F[- - 0. s%
u 0.6 _ -
.k.
2/3 0F IWB-369/-/
\ . UPPER BOUND fn., 0.4 _. ~
. PRED/CTED -
& SIZE RFTER g /f h/O.
(
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W 0.2 . CURRENT _ -
?! lND/CRT/ON -
.5IZE i
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O O.2 . 0.4 0.6 0.8 l.0 _ Fl.AW LENGTH lP/l'E CIMCUMFERENCE (f/Vb) FIGURE 4-1. Vermont Yankee Weld 1A Flaw Evaluation STRUCTURAL . INTEGRITY wruns - ~
, 45' M AXtMUM #
- j. . .
TYPE 308L A WELD OVERLAY I
.30",
A N ,r ,C eouu .o ,,,e I
/ l
/
... . O unN -
}
PtPE TO Etaow watos: I o e f.
. AS WELDEO SURFACE ACCEPTA8LE FOR
' OVERLAY TAPER 734 ANSITIONS
\, g RAosus !
EtsOw ! SCHEMATIC OF 12" ELBOW TO P!PE WELD OVERLAY j JOINT 32 - Figure 5.1 i
-Jt.- - .. - - - -
4s* uanuou A , i tv et wat. . A , artn ovra.4v - ?. .' l - 2.0" um .3 l % t *~ T l 10.125" l - um L e
) i
- Axial Inifica6 ion i
/ /
/ *
)Y J l 4.0" um h ,\ .
~
1 -
\ etre to e;. sow weu:,s-f I i i
t As wet.oso sunpace
' ACCEMASLE FOR i OVER4.AY TAPER y
. rnAmsmons N, '
We ld loi n t - HilR 4 i., , .. . 4 n i u t . u' Figure 5.2 7
,we n-., , - - - , - _ , , ,..-,-----,,.<y* , ,,,----,ww.-r-w y,-,,.ey,------,...-,,e---m.w,..evw,, --y--
f. 1 2 APPENDIX A CRACK GROWTH RESULT l lh f i
} STRUCTURAL )
A-1 > (NTEGRITY um. .
- - . . ~. .-
~
O l e d [ STRUCTURAL" INTEGRITY A N CIATEE STRESS CORROSIDN CRACK GROWTH ANALYSIS TITLE: VERMONT YANKEE 2e" SUCTION JOINT NO. 1A (NODE 8 l NIT AL CRACK DEPTH = 0.2640
+
WALL THICKNESS = 1.2000 MAX CRACK DEoTH DESIRED FOR SCCG= 0.9000 MATERIAL CONSTANT C.N OF DA/DT=C(DK)^N 2= 0.07000E-9 N= 2.260
") , HOUP MMAX A A/T DA/DT DA 97 +
. 1 ,1000.0 16.07 0.28 0.230 1.24163E-5 0.0124 2000.0 15.92 0.29 0.040 1.18187E-5 0.0118 3000.0 15.59 0.30- 0.250 1.1263BE-5 0.0112 2000.0 15.00 0.01 0.258 1.06910E-5 0.0107 5000.0 14.89 0.32 0.267 1.01604E-5 0.0102 5000.0
- 14.55 0.00 0.075 9.643S6E-6 0.0096 7000.0 14.22 0.34 0.280 9.15866E-6 0.0092 3000.0 10.91 0.35 0.290 8.70267E-6 0.0007 11 9000.0 10.59 0.36 0.297 8.26714E-6 0.0000
;0000.0 10.00 0.06 0.000 7.86492E-6 0.0079 11000.0 10.09 0.07 0.010 7.58542E-6 0.0076 10000.0 12.95 0.08 0.016 7.40696E-6 0.0074
'5000.0 *2.81 0.09 0.022 7.20498E-6 0.0072 14000.0 12.68 0.09 0.028 7.06976E-6 0.0071 25?00.0 12.56 0.40 0.000 6.91168E-6 0.0069
'6000.0 12.40 0.41 0.009 6.75906E-6 0.0068 17000.0 12.01 0.41 0.345 6.61165E-6 0.0066
;2000.0 12.20 0.42 0.050 6.47409E-6 0.0065 19000.0 12.09 0.43 0.355 6.34184E-6 , 0.0060 20000.0 11.98 0.40 0.360 6.21076E-6 0.0062 21000.0 11.87 0.44 0.366 6.09052E-6 0.0061 22000.0 11.78 0.44 0.071 5.97864E-6 0.0060 20000.0 11.68 0.45 0.375 5.86994E-6 0.005?
24000.0 11.59 0.46 0.080 5.76429E-6 0.0058 25000.0, 11.50 0.46 0.085 5.66215E-6 0.0057 STHUCTURAL 44 } _ ?- l N TEG RITY .sw.n.
k i STRUCTURAL INTEGRITY ASSOCIATES STRESS CORROSION CRACK GROWTH ANALYSIS TITLE: VERMONT YANKEE 28" SUCTrnN .in t h8T 2 /**m"" 9) INITIAL CRACK DEPTH = 0.1800 WALL THICKNESS = 1.2000 <- MAX CRACE DEPTH DESIRED FOR SCCG= 0.9000 MATERIAL CONSTANT C.N OF DA/DT=C(DK)^N Cr 2.27000E-9 N= 2.260 HOUR KMAX A A/T DA/DT DA 1000.0 16.50 0.19 0.161 1.28130E-5 2000.0 0.0128 16.09 0.21 0.171 1.24386E-5 0.0124 3000.0 15.99 0.22 0.181 1.19288E-5 0.0119 4000.0 15.69 0.23 0.191 1.14390E-5 5000.0 0.0114 15.35 0.24 0.200 1.08778E-5 0.0109
.. .6000.0 15.02 0.25 0.208 1.03588E-5 0.0104 7000.0 14.71 C.26 0.216 9.88692E-6 0.0099 8000.0 14.42 0.27 0.224 9.44840E-6 0.0094 9000.0 14.12 0.28 0.232 9.OOB80E-6 10000.0 0.0090 13.82 0.29 0.239 8.5769CE-6 0.0086 c 11000.0 13.53 0.29 0.246 8.17681E-6 10000.0 0.0082 13.23 0.00 0.252 7.77506E-6 0.0078 ~
10000.0 12.94 0.31 0.258 7.39859E-6 0.0074 14000.0. 12.67. O.32 0.264 7.04998E-6 0.0070 15000.0 12.40 0.32 0.270 6.71065E-6 0.0067 j e4 T STRUCTURAL A-3 [ INTEGRITY.wwt. p
.D
=
l tm STRUCTURAL INTEGRITY ASSOCIATES STRESS CORROSION CRACK GROWTH ANALYSIS TITLE: VERMONT YANKEE 28" DISCHARGE JOINT NO. 9A (NODE 52) INITIAL CRACK DEPTH = 0.2580
-- WALL THICKNESS = 1.2900 MAX-CRACK DEPTH DESIRED FOR SCCG= 1.0000 MATERIAL CONSTANT C.N OF DA/DT=C(DK)^N C= 2.27000E-8 N= 2.260 i
HOUR KMAX A A/T DA/DT DA 1000.O 14.80 0.27 0.208 1.00248E-5 O.0100 2000.0 14.50 0.28 0.215 9.56216E-6 0.0096 3000.0 14.21 0.29- 0.222 9.13222E-6 0.0091 4000.0 13.91 0.30 *0.229 8.71540E-6 0.0087 5000.0 13.61 0.30 0.235 8.29516E-6 0.0093 - 6000.0 13.33 0.31 0.242 7.90590E-6 0.0079 7000.0 13.05 0.32 0.247 7.53620E-6 0.0075 8000.0 12.76 0.33 0.253 7.17049E-6 0.0072
. 9000.0 12.49 0.33 0.258 6.83192E-6 0.0068 10000.0 12.24 0.34 0.263 6.51780E-6 0.0065 11000.0- 11.98 0.35 0.268 6.21355E-6 0.0062 10000.0 11.73 0.35 0.273 5.92514E-6 0.0059 13000.0 11.49 0.36 0.277 5.65721E-6 0.0057 14000.0 11.27 0.36 0.281 5.40784E-6 0.0054 15000.0 .
11.05 0.37 0.285 5.17210E-6 0.0052
' ' 16000.0 10.83 O.37 0.289 4.94559E-6 0.0049 17000.0 10.62 0.38 0.293 4.73427E-6 0.0047 10000.0 10.42 0.38 0.296 4.53679E-6 19000.0 0.0045 10.23 c.39 0.300 4.35193E-6 0.v044 20000.0 10.05 0.09 0.303 4.17863E-6 21000.0 0.0042 9.94 0,39 0.306 4.07065E-6 22000.0 0.0041 9.83 0.40 0.309 3.97201E-6 0.0040 20000.0 9.72 0.40 0.312 3.87705E-6 0.0039 24000.0 9.62 0.41 0.315 3.78540E-6 0.0038 25000.0 9.52 0.41 0.318 3.69748E-6 0. 00:.*.7
~ l STRUCTURAL A-4 - . _ }; inreoniry..,x .n w
ed
- STRUCTURAL INTEGRITY ASSOCIATES
. STRESS CORROSION CRACK GROWTH ANALYSIS TITLE: VERMONT YANKEE 28" SUCTION JOINT NO. 15B (NODE 13)
INITIAL CRACK DEPTH = 0.3186
- WALL THICKNESS = 1.1800 MAX CRACK DEPTH DESIRED FOR SCCG= 0.9200
.lATERIAL CONSTANT C.N OF DA/DT=C(DK)^N C= 2.27000E-8 -
N= 2.260 s i HOUR KMAX .A A/T DA/DT DA 1000.0 13.83 0.33 0.277 8.59195E-6 0.0096 - 13.52 0.04 0.284 8.16788E-6 0.0082 2000.0 3000,0 13.22 0.34 0.291 7.76269E-6 0.0078 - 4000.0' 12.93 0.35 0.297 7.38038E-6 0.0074 7.02685E-6 0.0070 " 0 T 5000.0 12.65 0.36 0.003 6000.0 12.45 0.36 0.309 6.77456E-6 0.0068 7000.0 12.31 0.37 0.314 6.60717E-6 0.0066 12.18 0.38 0.320 6.44616E-6 0.0064 8000.0 9000.0 12.05 O.38 0.325 6.29121E-6 0.0063 11.92 0.39 0.330 6.14000E-6 0.0061 10000.0 11000.0 11.7o 0.40 C.335 5.99332E-6 0.0060 11.67 0.40 0.340 5.85980E-6 0.0059 12000.0 0.0057 13000.0 11.56 'O.41 0.345 5.72648E-6 11.44 0.41 0.350 5.60082E-6 0.0056 14000.0 0.0055 i '15000.O ~ 11.33 0.42 0.355 5.4'7943E-6 STRUC fURAL A-5 [j INTEGRITY m.va.
- a. I
STRUCTURAL INTEGRITY ASSOCIATES STRESS CORROSION CRACK GROWTH ANALYSIE-
-e.
TITLE: VERMONT YANKEE 28" SUCTION JOINT NO. 26A (NODE 208) INITIAL CRACK DEPTH = 0.1725 WALL THICKNESS = 1.1500 MAX CRACK DEPTH DESIRED FOR SCCG= 0.0000 MATERI AL CONSTANT C.N OF DA/DT=C(DK) ^N C= 2.27000E-8 N= 2.260 A A/T DA/DT DA HOUR KMAX 1000.0 16.95 0.19 0.162 1.36096E-5 0.0136 2000.0 16.76 0.20 0.173 1.32768E-5 0.0133 3000.0 16.49 0.21 0.185 1.27967E-5 0.0128 4000.0 16.20 0.22 0.195 1.22850E-5 0.0123 5000.0 15.87 0.24 0.205 1.17245E-5 0.0117 6000.0 15.55 0.25 0.215 1.12094E-5 0.0112 7000.0 15.26 0.26 0.224 1.07343E-5 0.0107 - 8000.0 14.95 0.27 0.233 1.02542E-5 0.0103 14.64 0.28 0.242 9.77056E-6 0.0098 - 9000.0 10000.0 14.33 0.29 0.250 9.31023E-6 0.0093 14.01 0.30 0.258 8.84520E-6 0.0088 11000.0 12000.0 13.70 0.30 0.265 8.41566E-6 0.0004 13.39 0.31 0.272 7.99551E-6 0.0000 > 13000.0 0.0076 14000.0 13.09 0.32 0.279 7.59769E-6 15000.0 12.81 0.33 0.285 7.23011E 0.0072 16000.0 12.53 0.33 0.291 6.87423E-6 0.0069 12.26 0.34 0.296' 6.54080E-6 0.0065 17000.0 0.0062 18000.0- 12.00 0.35 0.302 6.23210E-6 19000.0 11.79 0.35 0.307 5.98838E-6 0.0060 20000.0 11.65 0.36 0.312 5.83311E-6 0.0058 21000.0 11.52 0.36 0.317 5.68405E-6 0.0057 11.39 0.37 0.322 5.54087E-6 0.0055 22000.0 23000.0 11.26 0.38 0.327 5.40253E-6 0.0054 ' 24000.0 11.14 0.38 0.331 5.26851E-6 0.0053
; 25000.0 11.02 0.39 0.336 5.13962E-6 0.0051 4-A-6 $
l l6""."wr.2. 3 NTEGRITY L_
y C. J. STRUCTURAL INTEGRITY ASSOCIATES STRESS CORROSION CRACK GROWTH ANALYSIS TITLE: VERMONT YANKEE 28" SUCTION JOINT NO. 27 (NODE 209) INITIAL CRACM DEPTH = 0.2185
- NALL~THICENESS= 1.1500 0.9000 MAX CRACL DEPTH DESIRED FOR SCCG=
MATERIAL CONSTANT C.N OF DA/DT=C(DK)"N C= 2.27000E-8 N= 2.260 A/T DA/DT DA KMAX A MOUR 0.0114 1000.0 15.64 0.23 0.200 1.13519E-5 0.0108 2000.0 15.32 0.04 0.009 1.08358E-5 0.0104 - 3000.0 15.02 0.25 0.218 1.03642E-5 0.0099 4000.0 14.74 0.26 0.027 9.92427E-6 9.45969E-6 0.0095 .v 5000.0 14.43 0.27 0.235 9.01632E-6 0.0090 6000.0 14.13 0.28 0.243 0.0086 - 7000.0 13.82 0.29 0.250 9.58703E-6 0.0082 8000.0 -13.52 0.30 0.258 8.16254E-6 0.0078 9000.0 13.23 0.30 0.264 7.77013E-6 0.0074 10000.0 12.93 0.31 0.271 7.~8788E-6 7.02465E-6 0 0070 l 1$000.0 12.65 0.32 0.277 6.68872E-6 0.0067 t
.12000.0 12.38 0.33 0.283 6.36916E-6 0.0064 13000.0 12.11 0.33 0.288 6.06401E-6 0.0061 11.85 0.34 0.293 5.78123E-6 14000.0 0.0058 15000.0 11.60 0.34 0.299 5.51852E-6 0.0055 16000.0 11.37 o0.35 0.303 5.34226E-6 0.0053 17000.0 11.21 0.35 0.008 0.0052 18000.0 11.08 0.36 0.312 5.00382E-6 0.0051 19000.0 10.95 O.36 0.017 5.07091E-6 0.0049 20000.0 10.83 0.37 0. 21 4.94324E-6 4.82008E-6 0.0048
-21000.0 10.71 0.37 0.025 4.70055E-6 0.0047 22000.0 10.59 0.38 0.329 4.58559E-6 0. 004 6 23000.0 10.47 0.38 0.333 4.47497E-6 0.0045 24000.0 lo. ; 0.39 0.337 0.0044 25000.0 10. d D.09 0.341 4.36846E-6
.W"g T STRUCTURAL A-7 INTEGRITY .u..
n STRUCTURAL INTEGRITY ASSOCIATES STRESS CORROSIDN CRACK GROWTH ANALYSIS t TITLE: VERMONT YANKEE 28" DISCHARGE JOINT NO. 61 (NODE 265) INITIAL CRACK DEPTH = 0.2500 WALL THICKNESS = 1.2500 .
' MAX CRACK DEPTH DESIRED FOR SCCG= 1.0000 MATERIAL CONSTANT C.N OF DA/DT=C(DK)^N C= 2.27000E-8 -
Nu 2.260 < HOUR KMAX A A/T DA/DT DA 1000.0 14.57 0.26 0.208 9.66873E-6 0.0097 2000.0 14.27 0.27 0.215 9.23156E-6 0.0092 3000.0 13.99 0.28 0.222 8.82465E-6 0.0088 4000.0 13.71 0.29 0.229 8.42912E-6 0.0084 5000.0 13.42 0.29 0.235 8.02725E-6 0.0080 6000.0 13.14 0.30 0.241 7.65468E-6 0.0077 7000.0 12.87 0.31 0.247 7.30048E-6 0.0073 8000.0 12.59 0.32 0.253 6.94753E-6 0.0069 9000.0 12.32 0.32 0.258 6.62067E-6 0.0066 10000.0 12.07 0.33 0.263 6.31735E-6 0.0063 11000.0 11.82 0.34 0.268 6.02266E-6 0.0060 12000.0 11.57 0.34 0.273 5.74232E-6 0.0057 - 13000.0 11.33 0.35 0.277 5.48196E-6 0.0055 14000.0 11.11 0.35 0.281 5.23966E-6 0.0052 - 15000.0 10.89 0.36 0.285 5.01049E-6 0.0050 16000.0 10.68 0.36 0.289 4.78913E-6 0.0048 17000.0 10.47 0.37 0.293 4.58275E-6 0.0046 10000.0 10.27 0.37 0.296 4.39000E-6 0.0044 19000.0 10.08 0.37 0. "l00 4.20968E-6 0.0042 20000.0 9.90 0.38 0.303 4.04071E-6 0.0040 21000.0 9.78 0.38 C.306 3.93090E-6 0.0039 22000.0 9.67 0.39 0.309 3.83245E-6 0.0038 23000.0 9.57 0.39 0.312 3.73781E-6 0.0037 24000.0 9.46 0.39 0.315 3.64677E-6 0.0036 25000.0 9.36 0.40 0.318 3.55916E-6 0.0036 S TRUCTURAL INTEGRITY == - .. L
h ;. 1 7 s STRUCTURAL INTEGRITY ASSOCIATES STRESS CORROSION CRACK GROWTH ANALYSIS f TITLE: VERMONT YANKEE 28" SUCTION JOINT NO. 59 (NODE 266) INITIAL CRACK DEPTH = 0.2814 WALL THICENESS= 1.3400 MAY CRACK' DEPTH DESIRED FOR SCCG= 1.0000
' MATERIAL CONSTANT C.N OF DA/DT=C(DK)^N
- C= 2.27000E-8 Nn 2.260 t
i HOUR KMAX A A/T DA/DT DA 1000.0 11.5G 0.29 0.214 5.75274E-6 0.0058 2000.0 11.35 0.29 0.019 5.50453E-6 0.0055 3000.0 11.14 0.30 0.222 ,5.27273E-6 0.0050-4000.0 10.93 0.30 0.226 5.04605E-6 0.0050 5000.0 10.71 0.31 0.230 4.82800E-6 0.0048
- 6000.0 10.51 0.31 0.033 4.62430E-6 0.0046 10.32 0.32 0.236 4.43393E-6 0.0044 7000.0 0.0043 8000.0 10.13 0.30 0.240 4.25551E-6 9000.0 9.95 0.30 0.243 4.0GG10E-6 0.0041 10000.0 9.78 0.33 0.246 3.92305E-6 0.0039 11000,0 9.60 C.33 0.248 3.76096E-6 0.0038 12000.0 9.44 0.34 0.251 3.62327E-6 0.0036 13000.0 9.28 0.34 0.254 3.48630E-6 0.0035 14000.0 ~9.12 0.34 0.256 3.35720E-6 0.0034 15000.0 8.98 0.05 0.259 3.23560E-6 0.0002 16000.0 0.93 0.35 0.261 3.12069E-6 0.0031 JL ...
0.0030 17000.0 B.69 0.35 0.263 3.01068E-6 10000.0 H.56 0.06 0.265 2.90485E-6 0.0029 j -19000.O G.43 C.36 0,267 2.00474E-6 0.0028 20000.0 H.30 0.*6 0.270 2.70994E-6 0.0027
! 21000.0 G.10 0.36 0.271 2.62006E-6 0.0026 0.0005 22000.0 0.06 0.37 0.273 2.53477E-6 0.0025 23000.0 7.94 0.37 0.275 2.45374E-6 24000.0 7.0; 0.07 0.277 2.37671E-6 0.0024
;' . 7 2 0.;7 0.279 2.30339E-6 0.0023 25000.0 l
A-9 ["i' STRUCTURAL l N T CGRITY .a e...,
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3 STRUCTURAL INTEGRITY ASSOCIATES . STRESS CORROSION CRACK GROWTH ANALYSLB TITLE: VERMONT YANKEE 20" DISCHARGE JOINT NO. 65A (NODE 250) INITIAL CRACK DEPTHt 0.2967 WALL THICKNESS = 1.2/00 - MAX CRACK DEPTH DESIRED FOR SCCG= 1.0000 MATERIAL CONSTANT C.N OF DA/DT=C(DK)^N - C= 2.27000E-B + N= 2.260 I HOUR KMAX A A/T DA/DT DA 1000.0 14.11 0.31 0.237 8.99564E-6 0.0090 2000.0 13.81 0.31 0.244 8.57103E-6 0.0086 3000.0 13.51 0.32 0.250 8.15694E-6 0.0082 . 4000.0 13.22 0.33 0.256 7.75786E-6 0.0078 1 5000.0 12.93 0.34 0.262 7.38863E-6 0.0074 6000.0 12.66 0.34 0.267 7.03964E-6 0.0070 7000.0 12.39 0.35 0.272 6.70110E-6 0.0067 8000.0 12.13 0.36 0.277 6.38746E-6 0.0064 9000.0 11.88 0.36 0.282 6.09627E-6 0.0061 - 10000.0 11.64 0.37 0.287 5.82019E-6 0.0058 11000.0 11.40 0.38 0.291 5.55668E-6 0.0056 12000.0 11.18 0.38 0.295 5.31145E-6 0.0053
- 13000.0 10.96 0.39 0.299 5.08278E-6 0.0051 14000.0 10.75 0.39 0.303 4.86921E-6 0.0049 15000.0 10.61 0.40 0.306 4.72397E-6 0.0047
. 16000.0 10.49 0.40 0.310 4.60725E-6 0.0046 17000.0 10.38 0.40 0.313 4.49498E-6 0.0045 10000.0 10.27 0.41 0.317 4.38693E-6 0.0044 a
19000.0 10.16 0.41 0.320 4.28290E-6 0.0043 20000.0 10.06 0.42 0.323 4.18268E-6 0.0042 21000.0 9.95 0.42 0.327 4.08641E-6 c. 0041 22000.0 9.85 0.43 0.330 3.99356E-6 0.0040
. 23000.0 9.75 0.43 0.333 3.90398E-6 0.0039 24000.0 9.66 0.43 0.336 3.81749E-6 0.0038
. 25000.0 9.56 0.44 0.338 3.73397E-6 0.0037 (18 "I STRUCTURAL A-10 j INTEGRITY .< ....
I
r i STRUCTURAL INTEGRITY ASSOCIATES STRESS CORROSION CRACK GROWTH ANALYSIS. TITLE: VERMONT YANKEE 22" JOINT NO. 16A (NODE 34) INITIAL CRACM DEPTH = 0.2100
- WALL THICKNESS = 1.0500
; MAX CRACK DEPTH DESIRED FOR SCCG= 0.G000 i MATERI AL CONSTANT C.N OF DA/DT=C (DK) ^N C= 2.27000E-8 N= 2.260 HOUR KMAX A A/T DA/DT DA
.' 1000.0- 16.34 0.22 0.212 1.23331E-5 0.0125
-2000.0 16.03 0.23 0.223 1. COO 73E-5 0.0100 3000.0 15.70 0.25 0.234 1.14912E-5 0.0115 4000.0 15.40 0.26 0.245 1.09555E-5 0.0110 5000.0 15.07 0.27 0.255 1.04036E-5 0.0104 6000.0 14.74 0.28 0.264 9.92260E-6 0.0099 7000.0 14.41 0.29 0.270 9.40527E-6 0.0094 .
8000.0 14.09 0.30 0.282 8.96723E-6 0.0090 9000.0 13.70 0.00 0.290 8.50105E-6 0.00G5 . 10000.0 13.48 0.31 0.298 0.11330E-6 0.0001 11000.0 10.19 0.32 0.305 7.72600E-6 0.0077 - 10000.0 13.02 0.00 0.012 7.50000E-6 0.0075 10000.0 12.91 0.03 0.319 7.05539E-6 0.0074 14000.0 12.00 0.04 0.026 7.21:21E-6 0.0072 15000.0 12.69 0.35 0.030 7.07958E-6 0.0071 16000.0 12.59 0.06 0.339 6.95172E-6 0.0070 17000.0 12.49 0.06 0.346 6.82744E-6 0.0068 18000.0 12.40 0.07 0.052 6.71671E-6 0.0067 19000.0 12.31 0.28 0.058 6.61003E-6 0.0066 20000.0 12.20 0.30 0.565 6.50597E-6 0.0065 21000.0 12.I"2 0.09 0.071 6. 41;551E-6 0.0064 22000.0 10.00 0.40 0.077 6.~3061E-6 0.0063
', 23000.0 12.01 0.40 0.380 6.24746E-6 0.0062 24000.0 11.95 0.41 0.~89 6.17401E-6 0.0062 25000.0 11.89 0.41 0.394 6.11116E-6 0.0061 I STRUCTURAL h INTEGRITY -..
, ~
7-. o (? g e J. crocce nn~~ - - - - __ r r ra c. uromnwr vnNKFF 77" JOINT NO. 16B (NODE 36)
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INITIAL CRACV. DFPru-WALL THICKNESS =- 1.0300
' MAX CRACK DEPTH DESIRED FOR SCCG= 0.8000 MATERIAL-CONSTANT C.N OF DA/DT=C(DK)^N C= 2.27000E-8 N= 2.260 t
HOUR KMAX A A/T DA/DT DA 1000.0 18.08 0.14 0.135 1.57581E-5 0.0158 2000.0 18.14 0.16 0.151 1.58618E-5 0.0159 3000,0- 18.07 0.17 0.166 1.57306E-5 O.0157 4000.0 17.90 0.19 0.181 1.53970E-5 0.0154 5000.0 17.65 0.20 0.195 1.49083E-5 0.0149 6000.0 17.31 0.22 0.209 1.42662E-5 0.0143 7000.0 17.00 0.23 0.223 1.36979E-5 O.013'7 9000.0 16.70 0.04 0.235 1.31589E-5 0.0132 9000.0 16.36 0.25 0.248 1.25724E-5 0.0126
. 10000.0 16.02 0.27 0.259 1.19821E-5 0.0120 11000.0 15.68 0.28 0.270 1.14090E-5 0.01.14 12000.0 15.30 0.29 0.281 1.08359E-5 0.0108 13000.0 14.98 0.30 0.291 1.00015E-5 0.0103 14000.0 14.64 0.31 0.300 9.78263E-6 0.0098
. 15000.0 14.33 0.32 0.309 9.31171E-6 0.0093 16000.0 14.22 0.33 0.318 9.14792E-6 0.0091 17000.0 14.11 0.34 0.327 8.98859E-6 0.0090 18000.0 14.00 0.35 0.335 8.83980E-6 0.0088
! 19000.0 13.90 0.35 0.344 G.69644E-6 0.0007 20000.0 13.81 0.36 0.352 8.56330E-6 0.0006 21000.0 13.72 0.37 0.360 8.44076E-6 0.0004 22000.0 13.63 0.38 0.369 8.32200E-6 0.0080 23000.0 13.56 0.39 0.377 8.22424E-6 0.0002 24000.0 13.49 0.40 0.384 8.12826E-6 0.0081 25000.0 13.43 0.40 0.392 8.04743E-6 0.0000 a ' - ~77 STRUCTURAL ! , I NTEG RITY .u r.-i.
f* a STRUCTURAL INTEGRITY ASSOCIATES STRESS CORROSION CRACK GROWTH ANALYSES - TITLE: VERMONT YANKEE 22" JOINT NO. 30B INITIAL CRACK DEPTH = 0.2080 WALL THICKNESS = 1.0400 -
- MAX CRACK DEPTH DESIRED FOR SCCG= 0.8000 MATERIAL CONSTANT C.N OF DA/DT=C(DK)^N C= 2.27000E-8 N= 2.260 HOUR KMAX A A/T DA/DT DA 1000.0 11.63 0.21 -0.206 5.31419E-6 0.0058 2000.0 11.40 0.22 0.011 5.55306E-6 0.0056
. 3000.0 11.18 0.22 0.216 5.31139E-6 , 0.0053 4000.0 10.96 0.23 0.221 5.08512E-6 0.0051 5000.0 10.76 0.23 0.226 4.86953E-6 0.0049 -
6000.0 10.54 0.24 0.230 4.65159E-6 0.0047 7000.0 10.33 0.24 0.234 4.44859E-6 0.0044 8000.0 10.14 0.25 0.238 4.25916E-6 0.0043
- 9000.0 9.95 0.25 0.242 4.08209E-6 0.0041 10000.0 9.76 0.26 0.246 3.90911E-6 0.0039 11000.0 9.57 0.26 0.250 3.74300E-6 0.0037
' 12000.0 9.40 0.26 0.253 3.56772E-6 0.0036 13000.0 9.23 0.27 0.256 3.44231E-6 0.0034 -
14000.0 9.06 0.27 0.260 3.30594E-6 0.0033 15000.0 8.90 0.27 0.263 3.17786E-6 0.0032 16000.0 8.75 0.28 0.266 3.05206E-6 0.0031 17000.0 8.59 0.28 0.268 2.93331E-6 0.0029 18000.0 0.45 0.28 0.271 2.82165E-6 0.0028 19000.0 8.31 0.28 0.274 2.71652E-6 0.0027 20000.0 8.17 v.29 0.276 2.61740E-6 0.0026 21000.0 8.04 0.29 0.279 2.52383E-6 0.0025 22000.0 7.92 0.20 0.281 2.43541E-6 0.0024 23000.0 7.79 0.29 0.283 2.35047E-6 0.0024 24000.0 7.67 0.30 0.286 2.26873E-6 0.0023 25000.0 7.55 0.30 0.288 2.19135E-6 0.0022 1 STRUCTURAL A-13 h INTecRiTy ...
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-5 .
. STRUCTURAL' INTEGRITY ASSOCIATES STRESS CORROSION CRACK GROWTH ANALYSIE TITLE: VERMONT YANKEE 28" SUCTION JOINT NO. 17B ( NODE 214)
INITIAL CRACK DEFTH= 0.2540 WALL THICKNESS = 1.2700 MAX CRACK DEPTH DESIRED FOR SCCG= 1.0000 MATERIAL CONSTANT C.N OF DA/DT=C(DK)^N C= 2.27000E-8 N= 2.260 b
+
HOUR KMAX A A/T DA/DT DA 1000.0 13.88 0.26 0.207 8.65933E-6 0.0007 2000.0 17.58 0.27 0.213 8.25239E-6 0.0083 3000.O 13.30 C.28- O.220 7.87470E-6 0.0079 4000.0 13.04 0.29 0.225 7.52347E-6 0.0075 5000.0 12.76 0.29 0.231 7.16481E-6 0.0072 6000.0 12.49 0.~0 0.236 6.82973E-6 0.0068 7000.0 12.24 0.31 0.242 6.51862E-6 0.0065 - 2000.0 11.c9 0.31 0.246 6.22280E-6 0.0062 9000.0 11.5*4 0.32 0.251 5.93430E-6 0.0059 10000.O i1.50 0.32 0.256 5.66627E-6 0.0057 11000.0 11.27 C.33 C.260 5.41678E-6 0.00"i4 10000.0 11.06 0.34 0.264 5.18411E-6 0.0052 13000.0 10.84 0.34 0.268 4.95672E-6 0.0050 14000.0 10.63 0.34 0.272 4.74437E-6 0.0047 15000.0 10.43 0.35 0.275 4.54596E-6 0.0045 16000.0 10.04 0.35 0.279 4.36026E-6 0.0044 17000.0 10.06 0.36 c.282 4.18620E-6 0.0042 18000.0 9.33 0.36 0.285 4.02OO5E-6 0.0040 19000.O 9.71 0.37 0.288 3.86216E-6 0.0039 20000.0 C.54 0.37 0.291 3.71379E-6 0.0037 21000.0 9.3G 0.37 0.294 3.57416E-6 0.0036 22000.0 c.23 0.38 0.297 3.44260E-6 0.0034 23000.O C.08 0.38 0.299 3.31848E-6 0.0033 24000.0 8.c3 0.38 0.302 3.20123E-6 0.0032 25000.0 8.30 C.39 0.304 3.11849E-6 0.0031 w g STRUCTURAL A-14 l y INTEGRITY.wwn u
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- STRUCTURAL INTEGRITY ASSOCIATES STRESS CORROSION CRACK GROWTH ANALYSIS
'ITLE: VERMONT YANKEE 22" JOINT NO. 49 (NODE 17)
INITIAL CRACK DEPTH = 0.2398 WALL THICKNESS = 1.0900 MAX CRACK DEPTH DESIRED FOR SCCG= 0.8000 MATERIAL CONSTANT C.N OF DA/DT=C(DK)^N C= 2.27000E-8 N= 2.260 HOUR KMAX A A/T DA/DT DA 1000.0 14.33 0.25 0.229 9.3174BE-6 0.0093 2000.0 14.02 0.26 0.237 8.86528E-6 0.0089 3000.0 13.72 0.27 0.244 8.44663E-6 0.0084 4000'.0 13.43 0.27 0.252 8.04056E-6 0.0080 5000.0 13.14 0.28 0.259 7.65209E-6 0.0077 6000.0 12.86 0.29 0.266 7.29232E-6 0.0073 - 7000.0 12.58 0.30 0.272 6.94473E-6 0.0069
'9000.0 12.32 0.30 0.278 6.61979E-6 0.0066 9000.0 12.07 0.31 0.284 6.31810E-6 0.0063 10000.0 11.02 0.32 0.289 6.03243E-6 0.0060 11000.0 11.59 0.32 0.295 5.76400E-6 0.0058 12000.0 11.36 0.33 0.300 5.51386E-6 0.0055 13000.0 11.15 0.33 0.304 5.28035E-6 0.0053 14000.0 11.04 0.34 0.309 5.16242E-6 0.0052 15000.0 10.94 0.34 0.314 5.05639E-6 0.0051 16000.0 10.04 0.35 0.318 4.95372E-6 0.0050
- 17000.0 10.74 0.35 0.323 4.85428E-6 0.0049 19000.0 10.65 0.36 0.327 4.76169E-6 0.0048 19000.0 10.56 0.36 0.332 4.67388E-6 0.0047 20000.0 10.48 0.37 0.336 4.58857E-6 0.0046 21000.0 10.39 0.37 0.340 4.50568E-6 0.0045 22000.0 10.01 0.37 0.344 4.42510E-6 0.0044 23000.0 10.24 0.38 0.348 4.3540SE-6 0.0044 24000.0 10.16 0.38 0.352 4.28502E-6 0.0043 25000.0 10.09 0.39 0.356 4.21770E-6 0.004 2 g-g l STRUCTURAL
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STRUCTURAL INTEGRITY ASSOCIATES
. STRESS CORROSION CRACK GROWTH ANALYSIS 6 (NODE 66)
.' T ITLE: VERMONT YANKEE 20" DISCHARGE JOINT NO.
INITIAL CRACK DEPTH = 0.2142 WALL. THICKNESS = 1.2600 - 1.0000 MAX CRACK DEPTH DESIRED FOR SCCG= , MATERIAL CONSTANT C.N OF DA/DT=C(DK)^N ~ C= 2.27000E-8 N= 2.260 A A/T DA/DT DA HOUR KMAX 0.23 0.180 1.22594E-5 0.0120 1000.0 16.18 1.17744E-5 0.0118 2000.0 15.90 0.04 0.189 0.25 0.198 1.12044E-5 0.0112 3000.0, 15.55 0.0107 4000.0; 15.22 0.26 0.006 1.06735E-5 0.015 1.01859E-5 0.0102 - 1= 5000.0 14.91 0.27 0.0097 6000.0 14.61 0.28 0.222 9.73390E-6 0.29 0.230 9.29518E-6 0.0093 7000.0 14.32 0.0088 8000.0 14.01 0.30 0.237 8.84729E-6 0.31 0.243 8.43240E-6 0.0004 9000.0 13.71 0.0080 10000.0 13.42 0.31 0.250 9.02810E-6 0.0074 11000.0 13.12 0.32 0.256 7.63561E-6 0.33 0.262 7.27245E-6 0.0073 10000.0 12.84 0.0069 - 13000.O- 12.57 0.34 C.267 6.92930E-6 0.0066 14000.0 12.30 0.34 0.272 6.59459E-6 0.35 0.277 6.28461E-6 0.0063 15000.0 12.04 0.0060 16000.0 11.79 0.36 0.282 5.99692E-6 0.36 0.287 5.72366E-6 0.0057 17000.0 11.55 0.0055 18000.0 11.32 0.37 0.291 5.46226E-6 0.00$2 19000.0 11.09 0.37 0.295 5.21913E-6 0.38 0.299 4.99257E-6 0.0050 20000.0 10.07 0.0048 21000.0 10.67 0.38 0.300 4.78109E-6 0.0046 02000.0 10.53 0.39 0.306 4.60807E-6 0.39 0,310 4.52031E-6 0.0043 23000.O 10.41 0.0044 24000.0 10.29 0.40 0.314 4.40716E-6 0.40 0.317 4.09839E-6 0.0043 25000.0 10.18
, fa
~;
STFt'" pet, A-16 I f INTEGRITY ..me.
.f' p
L 19 STRUCTURAL INTEGRITY ASSOCIATES STPESS CORROSION CRACK GROWTH ANALYSIS
-TITLE: VERMONT YANKEE 20" JOINT NO. 23B INITIAL CRACK DEPTH = 0.2943 MALL THICKNESS = 1.0900 MAX CRACf; DEPTH DESIRED FOR SCCGa 0.0000 MATERIAL CONSTANT C.N OF DA/DT=C(DK)'N C= 2.27000E-8 N= 2.260 f
u HOUR KNAX A A/T DA/DT DA 1000.0 7.27 0.30 0.272 2.01068E-6 0.0020 2000.0 7.17 0.30 0.274 1.946806-6 0.0019 3000.0 7.07 0.30 0.275 1.88619E-6 0.0019 L 4000.0 6.97 0.30 0.277 1.82842E-6 0.0018 5000.0 6.88 0.30 0.279 1.77337E-6 0.0018 6000.0 6.79 0.31 0.000 1.72087E-6 0.0017 7000.0 6.70 0.31 0.282 1.67061E-6 0.0017 [ 8000.0 6.61 0.31 0.283 1.62176E-6 0.0016 ' r 9000.0 6.53 0.31 0.285 1.57511E-6 0.0016 10000.0 6.44 0.31 0.286 1.53053E-6 0.0015 11000.0 6.36 0.31 0.287 1.48790E-6 0.0015 10000.0 6.29 0.31 0.289 1.44709E-6 0.0014 10000.0 6.21 0.32 0.290 1.40001E-6 0.0014 14000.0 6.14 0.32 0.291 1.37056E-6 0.0014 15000.0 6.07 0.32 0.293 1.33464E-6 0.0013 16000.0 6.00 0.32 0.294 1. COO 18E-6 0.0013 17000.0 5.93 0.32 0.295 1.26700E-6 0.00L3 10000.0 5.86 0.32 0.296 1.23529E-6 0.0012 19000.0 '5.00 0.32 0.297 1. 2047CE'-6 0.001: 20000.0 5.73 0.33 0.298 1.17533E-6 0.0012 21000.0 5.67 0.33 0.299 1.14704E-6 0.0011 22000.0 5.61 0.33 0.300 1.11981E-6 0.0011 L 23000.0 5.56 0.33 0.301 1.09625E-6 0.0011 24000.0 5.52 0.33 0.002 1.07890E-6 0.0011 25000.0 5.48 0.33 0.303 1.06198E-6 0.0011 T] STRUCTURAL A 17 E INTEGRITY wes [ . -
-7
? APPENDIX B ALLOWABLE FLAW DEPTHS f 6 S TFtuCTURAL B-1 (g[ l N7co nity .. ..,,
e L, n. t. STRUCTURAL INTEGRITY ASSOCIATEE CRITICAL FLAW St!E EVALUATION f t^N . CRITICAL FLAW S!lE FOR CIRCUMFERENTFA8 V TITLEsVERMONT VANKEE 20* SUCTION J0!NT NO. IA (NODE Sp teALL THICSNLGS* 1.2000 STPCSS RAflue 0.555 LOAD FACT 0pt= 1.09 L/ CIRCUM
.4 .2 .3-. 4 .5->1.0
.0 ALL0tKaDLE A/T 0.7500 0.7500 0.7500 0.7500 0.7500 0.e525 e TITLEsVERMONT YAND.EE 28' SUCTION JO!NT NO. 2 (NCDE 9)
WALL THICpNESS* t.2000 STRESS RAT 10* 0.465 LOAO F ACTOR = 1.00
.0 .* '
I- ALLOWASLE A/T 0.7500 0.7500 0.7500 0.7500 0.7500 0.6975 1 p- TITLEVERMONT VAfeEE 28" SUCT10N JOINT NO. 158 (N00E 135 I WALL THICKNESS = 1 1900 k' STRESS RAT 10= 0 635 . LDAD FACTOR
- 1.00 .
- L/ CIRCUM
.2 .5 4 .5->1.0
, .0 .I b' ALL m m a A/T 0,7S00 C.7500 0.7500 0.7500 0.7500 C.6425 TITLEsVERMONT YANKEE 29" SUCTION J0 TNT NO. 27 (N00C 2095 WALL THICkMSS* t.1800
. -STRESS RATID= 0.449 .
LOAO FACTOR = 1.00 L/ CIRCUM 4 .5-71.0 o .0 .1 .2 .5-ALLOWASLE A/7 0,7500 0.7500 0.7500 0.7*s00 0.7500 C.7055 (PODE 266i t1TLEaVERMONT YANKEE 29* SUCTION JDINT NO. 59 WALL THICDMSS= 1.3400 SIRESS RATI0a 0.424 . LOAO F ACTT= 1.00
- L/ CIRCUM
.I .2 .5 4 . 5- > 1. 0
.0 ALLOWASLE A/T 0.7500 0.7500 0.7500 C.7500 0.7500 0.71B0 IAA t en* "*
TITLEavfRMONT VAND EC 22" J0!NT NO. WALL THICDNESS= 1.0500 k SThESS RATIQa 0.644 LOAO FACTOR
- 1.00 L/CIFtCUM
.5 4 .5-)l.0
.0 .5 .2
' ALLOWASLE A/T 0.7500 0.7500 0.7500 0.7500 0.7500 0.5970
'6 f tTLCsVERMONT YAN>IEE 22" JOINT NO. 149 (NfW WALL THICDMSSa 1 0300 STRESS RAftO* 0.709 LDAD F ACTOR
- 1.00 L/ CIRCUM
.2 4 .5->l.0
.0 .1 .7 7
0.*500 0.7444 0.5 ALL0eenGLE A/T 0.7500 0.7500 0.7*00 l STRUCTURAL 8-2 - }; lNTEGRITY .wn,
g-9
.e-u 26A (N00E 2088
'TITLEIVERMONT VANKEE TS* SUCTION JOINT NO.
WALL THtCI'f4ESS= 1.1500 STRESS Raft 0= 0.490 LOAD FACTOR = 1.00 L/ CIRCUM
.1 .2 .0 .4 . 5- > t . 0
.0 ALLGWASLE A/T 0.7500 0.7500 0.7500 0.7500 0.7500 0.6910
+
(NODE 2658 TITLEtVERMONT VAre EE 29* DISCHARGE JO!NT NO. 41
.M4L TH1CkNESS* ! . ."500 STRCSS RATIO
- O.438 LOAD FACTOR
- 1.00 I'
L/C tRCtJM
.2 .3- 4 . 5- > t . 0
.0 .!
As.L0esabLE A/T Ov7500 0.7500 0.7500 0.7500 0.7500 0.7110 TITLEsVERMONT Y h EE 29" OISCHARGE JO!NT NO. 654 (NODE 2508 teALL THICKNESta 1.2900 STRESS Raft 0* 0.495 LOAD FACTOR = 1.00 L/ CIRCUM
.2 .3 4 .5 >t.0
.0 .1 ALL0tsABLE A/T- 0,7500 0.7500 0.7500 0.7500 0.7500 0.6825 Juls TITLEsVLHMUNT YANKEE 22' JQtNT NU.
i , blALt. TH1CKNESGe 1.0400 STRESS RAT 10e 0.334
-s LOAD FACTOR = 1.00
- L/CIRCurt
. 3 .. 4 . 5 + > 1. 0 0 .8 .2 e
ALL m E A/T OsF500 C.7500 C.7500 0.7500 0.7500 0.7431
#***** TD TITLE 6 VERMONT VAN >:EE 29* D!SCHARGE.1f1IMt M 9%
alALL THICKNESS = 1.2900 .. STRESS RATIOe 0.42t3 LOAD FACTOR = 1.00 L/C!RCurt
.3 .4 . 5- > t . 0 0 .1 .2 ALLommeLE A/T O,7500 0.7500 0.7500 0.7500 0.7500 0.7860
+. T STRUCTUHAL B-3 i 3 lNTEG RITY .wwn s
F A-S,
,g:
(~ . 1 f t fLEsVERMONT VANVEE 28' SUCTION JO!NT NO. 179 (NODC 2148 68ALL THICDNESS= 1.2700 3fMESS RAT 10e 0.427 . LOAD FACTOR
- 8.00 n./C!RCUM
.0 .1 .2 .3- 4 .5->l.0 ALLutshSLE A/f Os7500 c.7S00 0.7S00 0.7500 0.7S00 0.7165
?!TLEeVERMONT VANFEE 72* JOINT NO, 49 (NnfW 9Tl i
asALL fMICDNESS= 8.0900 SimESS RAT 10= 0.573 LOAO FACf0Re 1.00 L/ CIRCUM
.0 .8 .2 .3a 4 .5->l.0 ALL0temSLE A/T 0,7500 0.7900 0.7500 0.7900 0.7500 0.6433 f!TLEevfRNONT YANVEE 29" DISCHamtst JOINT NO. A (NMM ***
tsALL THICD. NESS
- 8.2600 .
STRESS Ref!Os 0.494 *
- j. .
40A0 FACitste 1.00 - 7 L/ CIRCUM
~'
i 4 . S *)l . O
.0 .8 .2 . 3. =
- ALL0 tem 8LE A/f 0,7500 0.7500 0.7900 0.7500 0.7S00 0.64S3 I
L TITLEsVEMMONT YANKEE 22* JOINT NO. 7tm
'heALL THICD_ NESS 8,0900 STEESS NATIDa 0.389 .
. L OAO F AC f 0Re 1.00
- L/ CIRCUM
.o .8 .2 .2 4 .t,->t.o ALLoesaste A/T o.7too o. 72,00 0.7500 0.7S00 c.rtoo o.7462 un a
J 5 .. i E
-w g4 STRUCTURAL
{ INTLGRITY w v.
Y ENCLOSURE 3 VEP.MONT YANKEE REACTOR COOLANT LEAK DETECTION PROVISIONS i I I
y f
,i,
.y EscLosURE 3
,t. VE MDET YAMEEE REACTOR COOLANT 1.EAEAGE DETECTION PROVISIONS m
h
- 1. Raasige Coelant Leakana Limits
. /o 9;By Inther dated June 27, 1983 (Reference (1)], the NRC issued a Confir: set.ory Order which included provisions for reactor coolant
- ' ,/ Meekage; These provisions were incorporated into plant procedures prior to restart from our 1983 refueling outage and will continue to be in
- ,' effeet'during the'1984-1985 cycle of operation. These provisions are
/c , :provided in Attaebeant D to this report.
a'& ,
'-(( .
. 97n t.ase9 wherv these limits, frequencies, or corrective actions conflict
' 'r , < ; with the current Technical speelfication 3.6.c/4.6.c. It is our intent i to fotlow the provisions of Attachment D in lieu of our present Technical specifications.
2.'l Eg}Ejarg1_3gMjdte,Tase System
,O fi As/previously, 'liscussed between representatives of Virmont Yankee and c 'imtere of your staff, a moisture sensitive tape leak-detection system will be installed during the curre.it refueling outage, six (6) detector
.. ~1ecations have been selected such that the remaining eight (8) uninsp9cted 2P weld joints;1n the Reelreulation System will be y < . monitor)d, < The exact detector locations have been discussed in detail j '.with'the supplier of.the system and assurance has been provided that the
' eight (d) weld joints will be adequately monitored.
t
'Due t'o operational problems associated with the moisture sensitive tape L - ' detectors-dur).ng the last operating cycle, a modified detector will be
, used. Tida sipification consists of relocating the electronics portion of the dessetoes to a separe'.o junction box uJeh that the electronic
~
~~ >
~
- b. (devices'arill W remete fres high temperatures at the detector 3
' j f ? \Ltr9neter-locatia,as; free the ' High temperatuns Recirculation. System'at piping the detectors is caused to the stainless steel by heat fi .;'; detector housins. As a result of this neodification, we believe that the
'syrtem nl. cued be more reliable.
-j' pastedon'theabave,we'willvnrballynotifytheVermontYankeeNRC
#[-
Project 4anager of kny significant chanes in the ' status of the moisture sensitive tape systein during the 1984a1985 operating cycle. Further, we
)'M: . will sake evety.reasonstle effort to maintain the cystem fully operable. tin the event of partial or intermittent system operability l$ (eleller to the condition that owleted durlag our last operating cycle),
i;Verment Yankee' would be responsibjo for determining the frequency during
% ! '. whish.the o system will be used yo chek for weld joint leakage.
,0 .
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y-ENCLO::URE 4 AUCMENTED INSERVICE INPSECTION ALARA INFORMATION
ENCLOSURE 4 AUCMENTED ISI ALARM INFORMATION A total of 190 man rem have been expended for work associated with the 1984 Augmented Itt'-Service Inspection Program. The man-rem expsoure levels are as follows: , Man-Rem UT Inspections: 48 Repair: , 9 Insulation: 62 Shielding: 26 Weld Preps: 33 Moisture . sensitive Tape System Installation 12 Total: 190 We estimate that an additional 22 man-rem would be expended to inspect the last eight (8) 29" fields, of which 16 man-rem would be to the UT personnel. There'is approximately-12 ran-rem remaining among the available UT personnel, which is insufficient to complete' the exams. Further, it would take a week and a half to two_ weeks to obtain additional qualified personnel. Based on the principal of ALARA, no further exposure to inspect the remaining 28" welds is justified. , 6 l
e ENCLOSURE 5 RECIRCULATION LOOP PIPING TEARING STABILITY ANALYSIS
- u -.
ENCLOSURE 5 YAEC Contract 104300 Final Report 84-3 45, Rev. 1 July 18,1984 VERMONT YANKEE NUCLEAR POWER STATION RECIRCULATION LOOP PIPING TEARING STABILITY ANALYSIS Prepared by FRACTURE PROOF DESIGN CORPORATION 77 Maryland Plaza St. Louis, MO 63108 Principal Investigators Keyren H. Cotter Steven W. Slocum Prepared for-Yankee Atomic Electric Company 1671 Worcester Road Framingham, MA 01671 YAEC Proj ect Manager Robert White
CONTENTS Eagg Section 1-1
'1 INTRODUCIION 1-1 1-1 Background 1-1 1-2 USNRC Safety Assessment Criteria ,
1-2 1-2.1 USNRC SEP Criteria for Break Postulation 1-4 1-2.2 USNRC PRC Criteria for Break Postulation 1-6 1-3 Vermont Yankee Criteria 2-1 2 CRACE STABILITY CRITERIA 2-1 2-1 Crack Driving Force 2-2 2-2 Fracture Toughness Considerations 2-3 2-3 Tearing Stability Considerations 2-3 2-3.1 Theory 2-6 2-3 . 2 The J-T Diagram Curve 2-6 2-3.3 Extrapolations of the Jant-Tmat 2-8 4 Structursi Ductility 3-1 3 SSY BASED ANALYSIS 3-1 3-1 J-Integral Estimation 3-1 3-1.1 Circumferential Cracks 3-3 3-1.2 Longitudinal Cracks 3 -3 3-1.3 Tearing Stability for SSY Conditions 3-4 3-1.4 Plastic Zone Instability Failure 3-4 3-2 Leak Rate Analysis [L
4 LSY BASED ANALYSIS 4-1 4-1 Structural Response and T,pp 4-2 4-1.1 Compliance 4-2 4-1.2 Plastic Hinge Behavior 4-2 4-2 Cracked Section Parameters 4-3 4-2.1 Plastic Limit Moment 4-3 4-2.2 J-Integral 4-3 4-3 Stability Analysis 4-4 4-4 The J-T Diagram 4-5 5 RECIRCULATION SYSTEM, RESULTS AND DISCUSSION 5-1 5-1 Recirenlation Loop Piping System 5-1 5-1.1 System Description 1 '5-2 5-1.2 Piping Code Stress Analysis h-3 5-2 Leak Detectability 5-4 5-2,1 Circumferential Flaws 5-4 5-2,2 Longitudinal Flaws 5-5 5-3 Crack Stability; Pressure + Thermal +SSE 5-5 5-3.1 Circumferential Flaws 5-5 5-3 . 2 Longitudinal Flaws 5-6
'5-4 Crack Stability, Upper-Bound Loads 5-6 5-4.1 Applied Loads , 5-7 5-4.2 Crack Stability 5-7 6'
SUMMARY
AND CONCLUSIONS 6-1 7 REFERENCES 7-1
' APPENDIX A PROGRAM: JTPIPE A-1 APPENDII B MATERIAL PROPERTY DATA B-1 APPENDIX C USNRC SEP + PRC ALTERNATIVE BREAK POSIULATION C-1 APPENDIX D CALCULATIONS D-1
C. [ ; . c.
~ ~
TABLES Page h T: 5-10 5-1 ;Section Properties,and Maximum Stresses 5-10 5-2 Summary of GE. Stress Analysis Results 5-11
;5-3. :J,pp for Leak Rates of 1 and 10~gpm g --- g yy.(- gy. --y-m.-, . . y. -m7-,- - .s .,-y- +-w g ,.we g_rpye,w,,gnp ,9y,, ,,, y..,q_, .p,,,-5..--yy,yym,%"
ILLUSTRATIONS Finure Pane 2-1 Schematic of J-T Stability Diagram 2-9 4-1 Geometry of Cracked Cross-section of a Pipe 4-6 4-2 Fully Plastic Bending of a Pipe 4-7 4-3 Stability of Part-through Crack Under Fully Plastic Bending 4-8 5-1 Vermont Yankee Recirculation Loop Piping Isometric 5-12 5-2 Location of Anchors, Supports and Pipe-whip Restraints 5-13 5-3 Leak Rates for Circumferential Cracks 5-14
.5-4 Leak Rates for Longitudinal Cracks 5-15 5-5 Node Locations in Idealization of Recirculation Loop Piping 5-16 5-6 Crack Locations in Idealization of Recirculation Loop Piping 5-17 5-7 J-T Stability Diagram, Load Case 4, 20 = 60 s 5-18 5-8 J-T' Stability Diagram, Load Case 4, 20 = 120 e 5-19
^
Section 1 INTRODUCTION 1-1 BACKGROUND The Vermont Yankee Nuclear Power Station is a 540 Mw GE BWR design. Because of the implications of having flaws or undetected flaws in the recirculation loop piping at Vermont Yankee, it was decided that a safety analysis should be performed. And, the methods of evaluating the stability of any flaws in that piping, should be conservative. Recent advances in crack stability criteria, based on structural ductility concepts, are appropriate for the analysis of ductile piping such as that found in the Vermont Yankee recirculation loop. Over the past four years, several criteria, based on elastic plastic fracture mechanics, tearing stability methods and structural ductility, have been proposed for evalrating defects in piping. These are reviewed in the following paragraphs. 1-2 USNRC SAFETY ASSESSMEKr CRITERIA A review of recent criteria proposed for the analysis of nuclear piping is presented below.
Page 1-2 1-2.1 USNRC SEP Criteric In 1981, the USNRC developed criteria which permits plant operators to use alternative methods to obviate the need to consider pipe rupture -events and the pipe-whip protection requirements (1) imposed on " older" plants 'being reviewed under the USNRC Systematic Evaluation Program (SEP). The complete USNRC Alternative Criteria (2) is presented in Appendix C and in abbreviated form below. In order to be exempt from the requirement to protect against the ef fect of pipe whip and jet impingement resulting from postulated pipe breaks under SEP Topics III-5A+B, it must be demonstrated that the particular piping in question exhibits: A) ,Detectability Reanirements. Provide a system to detect leaks, resulting from both longitudinal and circumferential through cracks having lengths equal to Ai t (Ay times the wall thickness) under normal operating loads, where A3 )2 and A1 is to be determined and depends on the method of leak detection; plus, B.1) Intenrity Requirements, Level D. Stability of both longitudinal and circumferential cracks that have a length equal to "Ay t + 2t" under Level D loads must be demonstrated; integrity of anchors is presumed; plus,
-B.2) Intearity Reanirements. Extreme Conditions. Stability of a
-circumferential crack that is the greater of "Ai t + 2t" or 90 degrees circumferential length under fully plastic bending loads; hangers are to be assumed inef fective; snubbers are to be assumed ineffective unless specially justified; integrity of anchors is presumed; plus,
Page 1-3
.B.3 ) Na t e ri a l P_g. ope rt ie s . Lower bound material properties are to be c used.
C) Sub-critical Crack Growth. Consideration shall be given to the types of sub-critical cracks that might exist in the piping system. D) Anamented ISI. An optional approach involving special inspection procedures may be used if other corrective measures are not practical. (Not considered applicable to primary coolant system cracking problems.) The satisfaction of the USNRC Criteria A) requirement involves an analysis which utilizes linear-clastic fracture mechanics methodology and the computation of leak rates for- longitudinal and circumferential cracks. The normal (or Level A) operating stresses (3) are used to J compute the crack length that would result in e detectable leak rate. These calculations are reasonably straight forward and require the gathering of the stress analysis results and review of the plant Technical Specifications as to detectable leakage rates. The USNRC Criteria B.1) calculations require the postulation of a crack having a length of "A gt + 2t" (t = the wall thickness of the pipe). The
"+ 2t" amount is included to permit a margin for ,ub-critical crack growth due to fatigue or stress-corrosion. The crack is assumed to be oriented longitudinally or circumferentially and extends through the wall of the pipe. Z~U rr7 N e cow N1 USA /PC CnYenbn 2.2. 1) oz & v a nt g w er k an<0 m c d w & A
~
Page 1-4 properties, the piping will not exhibit instability under upper-bound loading. The upper-bound loading assumes that the section containing the
. crack is fully plastic and the crack is oriented circumferentially. The upper-bound loads are assumed (non-deterministically) to be equivalent to the bending moment required to induce a fully plastic section at the crack location. This upper-bound is one means of accounting for extremely low probability events such as water hammer and snubber failure under seismic loading. Note that such high load levels are not expected to occur, but are felt more realistic than the load inferred by the currently accepted method of postulating breaks for typical pipe-rupture analysis; namely, a load that occurs instantaneously and is equal to the ultimate strength of the uncracked section of the pipe.
1-2.2 USNRC PRC Proposed Criteria for Break Postulation The SEP criteria described in Section 1-2.1 was developed with the specific intent of providing relief from compliance with current pipe rupture criteria for older design plants. The older plants were not
' designed to meet the curren,t criteria and the imposition of such criteria on old designs is nearly impossible or impractical in certain cases.
Thus, the criteria had a specific rather than general purpose. The Piping Review Committee (PRC), formed by the USNRC in 1983, assumed the responsibility for developing a break postulation criteria that could be applied to any class of plant. The following "draf t" criteria have been prepared by the PRC's Task Group on Break Postulation for use with primary coolant system analyses. The
)
(
Page 1-5 current draft of the criteria is limited in application to sections of primary. coolant systems that are not prone to IGSCC, thermal fatigue or water-hammer. It consists of a step-wise approach as follows:
- 1) Postulated Defect Sizes. Select the highest stress - poorest materials properties location in the pipe under consideration.
Then, postulate a crack that may be missed during fabrication and pre-service inspections or would be permitted by Code, whichever is larger. And, demonstrate by analysis that the crack would not grow significantly during service either by fatigue, corrosion or impact forces (water-hammer).
- 2) Detectable Leak. age Ri te. Demonstrate that even if the crack propagated through tle wall that: a) the leakage through the crack is significantly greater than the minimum leak detection capability under normal operating loads so that detection of the crack is assured; and, b) even if undetected prior to an earthquake, the crack is stable under normal plus SSE loads, and, (grewth, if any, is minimal for long periods of time).
- 3) Safety Narain. Crack Sizes. Show that adequate sa fety margin exists based on crack sizes by: a) comparing the leakage crack size computed in 2a) with the critical crack size under normal plus SSE loads; and, b) demonstrating that there is adequate margin to account for uncertainties inherent in the analyses and leak i
detection.
F ', Pege 1-6
- 4) Safety Marsin, Loads. Demonstrate that the leakage size cracks of 2a) will not result in unstable crack growth even if larger loads are applied and that the final crack size is limited (that is, a double-esied pipe break will not occur).
1-3 VERMONT YANKEE CRITERIA W2 This The following criteria, mar selected for this analysis. selection was made af ter reviewing the USNRC approaches described in Section 1-2. It was designed to be as conservative or more than any that would ultimately be approved as the USNRC Guidelines.
- 1) Detectability Reanirements. Provide a method of detecting leaks, resulting from both longitudinal and circumferential through-cracks having lengths equal to Ay t (Ay times the wall thickness) under normal operating loads, where A y>2 and A 1is tc be determined and represents the minimum given the method of leak detection in operation at the plant in the area in question; plus,
- 2) Intearity Reaufrements, Necessary Conditions. Stability of both longitudinal and circumferential cracks that have a length equal to "Ag t + 2t" under loads equal to SSE plus Thermal plus Pressure must be demonstrated; integrity of anchors, supports, hanger and snubbers is presumed; plus,
- 3) Intearity Reauirements. Sufficient Conditions. Stability of a circumferectial crack that is the greater of "Ag t + 2t" or 90 degrees ci rcum fe renti al length under conditions that insure structural ductility; the loading for insu ing structural ductility IC Cb includes thermal, dead weight, pressure, thet is.wltlag free support failures and inertial; snubbers and other supports are to be assumed
n - Page 1 7 ineffective unless specially justified; integrity of anchors is presumed, but verified; plus,
- 4) Material Pronerties. Representative material properties are to be used; and,
- 5) Sub-critical Crack Growth. Consideration shall be given to any sub-critical crack. growth that might occur in the piping system.
\
/
Section 2 i CRACK STABILITY CRITERIA 1 l In order to analyze the stability of cracks in nuclear piping using l fracture mechanics methodology, it is necessary that the material properties, distribution of crack sizes and shapes, applied loads (or stresses) and. the crack stability criteria be specified. The crack stability criteria are discussed in this Section as the other factors or parameters follow directly from it. 2-1 CRACK DRIVING FORCE Crack stability is usually evaluated by comparing the value of a crack driving force parameter with the resistance of the material to crack
' extension. The crack driving force paraueters can be grouped into those applicable to cases involving limited, (sm:11) amounts of crack-tip plasticity and those with large amounts of crack-tip plasticity, including net-section yielding. The former is otten referred to as small-scale yielding (ssy) and the latter as large-scale yleiding (1sy).
- For ssy cases, the crack driving force is usually described in terms of the associated vaines of the crack-tip stress-intensity factor, K. For lay _ cases, K is not applicable and the parameter is described in terms of thevaineoftheJ-lategral,J,jp. J,pp is also valid for the say regime, e
-y - - - , - -,.m-- -r . , - - - ,n , , , , , , - _ , - ,,- , ..n, ,_
Page 2-2 The analysis of crack problems is the ssy regime involves the use of the methods of linear-clastic fracture mechanics (LEFM). For the analysis of 1sy problems, elastic plastic fracture mechanics will be relied upon. 2-2 FRACTURE TOUGHNESS CONSIDERATIONS Recall that for ssy problems the crack driving force based on LEFM i s' defined in terms of the stress-intensity factor, K. And, for stability, the K computed for the applied stress and crack size of interest must be less than the fracture toughness, Kye, of the piping material. Stability can also be defined in terms of the J-integral, which, for LEFM, can be computed from J=(K*/E') as further discussed in Section 3-1. The parameter Jy , can be considered a toughness that is equivalent to the fracture toughness, K ye, or crack initiation toughness, and thus for J<J ye, stability is insured. Because J is used throughout this report, consideration of LEFM methods is presented in terms of J. A Jy, approach to stability (that is, not including stable growth above Jye) is not acceptable for lay problems because it is far too conservative. Estimates of J for ssy conditions are developed for the crack geometries of interest using the accepted practice of basing J estimates on plastic zone corrected stress-intensity factor solutions (i.e., K(a+r )). (Note y that estimates of J based upon K solutions, that is, LEFM, result in unconservative estimates of J as the limit moment of the cracked section is spproached.)
Page 2-3 2-3 TEARING STABILITY CONSIDERAT]ONS , I 2-3.1 Theory Before considering the application tearing stability methods to a typical piping problem, it is worthwhile to review a bit of theory. In the application of LEFN to brittle materials, crack instability is assumed to be . incipient when K)K re. Physically, this is interpreted as an instability that accompanies the onset of crack extension. But, for tough materials, it is known that crack instability does not generally
. accompany the onset of crack extension. Rather, the K (or J) at Einstability can be well above the K y , (or Jye) point. It is important,
'from design and safety considerations, to be able to take advantage of the higher J values (or loads) that co-exist with the stable crack
-extension but, until the recent development of the tearing modulus concept, it ws: not possible. Analysts had been faced with the problem of using a Jyg value for instability predictions unless representative R-curves could be developed which were typical of the significant material dimensions actually used in the structures of interest.
Solutions _to problems that rely on the tearing stability approach involve _. expressing the intensity of the crack-tip deformation field by an appropriate elastic plastic crack driving force parameter. Based on the I fracture parameter, the behavior or growth of cracks can be expressed functionally. It follows that the use of a parameter like J infers that crack- growth is controlled or determined by the value of the parameter. This logic ~1eads to the te rm "J-cont rol l ed growth". Typically, the l ~ quantifying of the fracture parameter is accomplished by computing the l 'value of the path independent J-integral, developed by Rice (i), either by use of direct integration arcund the crack-tip or by use of any one of a number of acceptable estimation schemes. Relative to any J computation,
'it is interesting to note that the phrase " clastic plastic f racture
Page 2-4 mechanics" infers that problems involving plasticity can be analyzed for ) cny type of loading. But Rice (i) proved the path independence of J only for the idealized case of no crack growth and a material which exh ibit s "non-linear elastic" behavior. Unfortunately, real materials do not behave exactly as non-linear elastic materials and the problems of interest involve crack growth. However, the violation of this idealized behavior is not sufficient to invalidate the path independence of the J-integral if certain restrictions are met. Based on a need for these restrictions, Hutchinson and Paris (1) set forth strict theoretically based guidelines for J-controlled crack growth. Extensions beyond those limits are possible under the conditions discussed in Reference (6). Although the value of J is indicative of the intensity of the crack-tip deformation field, it is not sufficient by itself for resointion of the question of stability. To resolve this, Paris, et al.(1) defined a non-dimensional parameter, called the tearing modulus, which assumed the validity of J-controlled growth. It is applicable to material property data and applied loads alike. For the applied case, it is expressed as m.. dJ (2-1) Tapp = - app _E_ t da e
.where E is the clastic modulus, a is the crack length, a is a flow stress, and I is the J-integral. J controlled growth requires that the crack extension, da, occurs under the equilibrium condition
- - J,pp = J mat, (2-2)
I which applies whether or not stability of the crack extension is present. t l
Page 2-5 In this . expression, J aat is the value of J on the material J-resistance curve, and the J,pp is the computed value of the J-integral for a given load and crack length. For a crack under the preceeding equilibrium conditions stability is determined from T,pp < Tmat (stable) (2-3) T,pp > Tmat (unstable) (2-4) where T,,g is . determined from the material J-R curve and T,pp is dependent upon the crack geometry and loading existing in the actual structure. The very power of this approach stems, in part, from the fact that the use of tearing stability methods is applicable (6.8-11) to both the ssy and 1sy regime. This stability criteria has been experimentally verified for several specimen types. Paris, et al.(ll), were the first to demonstrate applicability through experiments using A471 steel 3 point bend bars in a test systen of variable compliance. The variable compliance feature was used as a means of controlling the T,pp. Similarly, Zahoor and Kanninen(12) tested circumferential1y cracked 4-inch diameter TP304
. stainless steel pipes in 4 point bending, rnd Gudas and Joyce (13)
, evaluated several materials of varying degrees of toughness in 4 point l bending. l l _ . . . - . . _ . - . _ _ . , . _ . _ . - ~ _
- Page 2-6 2-3.2 zThe J-T Diagram For safety assessments of nuclear piping systems that are based on the tatring modulus ' stability concept, it is convenient to present the results using the J-T diagram due to Paris (f) . The J-T diagram compares
.the applied (or calculated) values of J and T with the material (invariant characteristic of a material) values. That is, the J,pp vs.
T,pp response is compared with the J aat vs. Tmat curve to determine whether.the T,pp value is less (or greater) than the Tmat value for the J,pp value specified. If the T,pp(Tmat, then stability is assured and, conversely for instability. A sample J-T diagram is shown in Figure 2-1. The schematic material curve shown on Figure 2-1 was derived from a typical J-resistance curve. Note in Figure 2-1, that the T,pp values are dependent upon-the J,pp. For cases where the T,pp is less than the Tmat values, . stable crack behavior is assured. On the other hand, a lower Tmat value corresponding to higher J,pp values can cause unstable
. behavior.
2-3.3 Extrapolation of the J aat-Taat Curve For applications that require Jaat values greater than those available from the J-resistance curve, the assessment of whether a system is stable or unstable based on a J-T diagram may require extrapolation of the material curve. One way of extrapolating the resistance curve is to assume that the material continues to tear with the same slope. This will mean that in the extrapolated regime, the T remains constant. mat The extrapolation of the material curve on the J-T diagram would then be
Page 2-7 the vertical line extending from the maximum Jmat value point on the material curve. This is shown as 1-C in Figure 2-1. An alternative to this extrapolation is to assume that there is no further- lucrease in the J-resistance with crack growth. Such a behavior would imply that the Tmat reduces to zero in the extrapolated regime.
~That is,~ on a J-T diagram, the extrapolated material curve wonid take the form of the horizontal line noted as 1-0 in Figure 2-1.
These two extrapolations represent the upper and lower bounds of resistance curve behavior for continued growth. In reality, the Tmat
.value is expected to decrease gradually with increase in the Jmat value, leading to the possibility of a zero value of the Tmat at some higher J
ust value. One accepted approach is to follow Paris (6) and construct a tangent to the material curve. This approach is noted as line 1-T in Figure 2-1. The validity of J-controlled growth is dependent upon the satisfaction of several requirements (1). One of these is that w be "large". For typical
, Type 304 stainless steel, valid J-resistance curves may have over 1 inch of crack growth, Aa, and u values that range from 10 to 20. Considering T
mat to decrease abruptly to zero simply implies that the e value, which is proportional to T, also decreases to zero. This would invalidate the assumptions of J-controlled growth, and any assessment of the stability of piping would be subject to serious error. F-resistance curves need to be developed to include extended amounts of crack growth while satisfying the J-controlled growth requirement. Because of these limitations, the
Page 2-8 assumption of the tangent extrapolation of the Tmat curve is felt to be the most acceptable method. 2-4 STRUCIURAL DUCTILITY Recent discussions of the criteria to be applied to the analysis of cracks in nuclear piping have resulted in Paris (li) proposing that such criteria must insure that " structural ductility" is maintained. Paris e argued (11) that one of the fundamental tenants inherent in the ASME Code (1) that insures the safety of nuclear plants is the concept of structural ' ductility. In the simplest sense, this concept insures that the stored elastic energy in a structural system can be absorbed by plastic work. The plastic work takes the form of local or gross plastic deformation of sections of the structure. Its purpose is to provide y assurances that no brittle type failure can occur should the loads portion of the analysis be in error. Paris repeatedly cited (11) examples that this is the basis used by Nathan Newmark in his work on the safe response of structures to seismic excitation (inertial loading). For piping in typical nuclear plants, this criterion is always met, because of Code (1) requirements, if there are no cracks present. To insure that it is met when cracks are present, then Paris (11) and i*e r iY and Cotter (la) showed that ; T,pp ( q Tmat (2-5) for I *3 1ocal + Iglobal (2-6) mpp where q is a constant between 1 and 2, J3oc,3 describes the response of the structure to the " worst-case" loads and J global satisfies the requirement of absorbing the stored elastic energy.
'r t
) L E A
. L I B R A E T T S A N M U
( m f a r g D a E i I D L y . P t P ) i l A E i L b B a A t S T - S T ( T - J f o c i t C a l l 1 m e h c S
\ 1 2
e r . u g i r v'4 0 - c J 1 J
Section 3 , SSY BASED ANALYSIS 4 3-1 J-INTEGRAL ESTIMATION
-For the ssy regime, the J-integral, J app, can be estimated using the relation J,pp = K*/E' (3-1) where E'=E for plane stress, E'=E/(1 p*) for plane strain, K y is the opening' mode- plastic zone corrected stress-intensity factor, E is the clastic modulus and p is Poisson's ratio.
3-1.1 Circumferential Cracks For circumferential cracks, the Ky consists of contributions from three types of loads: axial load, bending moment and membrane stress due to Pressure. 'The Ky due to pressure loading, K,, was obtained by utilizing the sointions from Reference (15), giving K,= ag/nRO F, (3-2) where a,is the membrane stress (axial) and F,is a non-dimensional shell correction factor that depends upon the length of the crack and the geometrical dimensions of the sheII.
, ,. - -- . - , , - . - - - , . , . . , , - . - - - - - - - - - - ~ , , . , , , , , , , ..-----,n. ,,., - . . . , - , - , , - . , , , - --,w-.
Page 3-2 The K y due to the applied axial tension load is Kg = af/nRO Ft (3-3 ) where Ft depends upon the same parameters as F,. The function tF can be derived from the recent work of Erdogan and Delale(16). FPDC has developed its own approximate, but conservative, expression for F t which was used in this study. at is the stress (tension) due to the axial load F, 3
't = F,g/(2nRt) (3-4)
Similar to the tension loading case, FPDC had previously developed an estimate- of K for the externally applied bending load; and the K due to this loading is Eb " 'b,/nRO Fb (3-5) where bF is a correction factor for a circumferential crack in a shell -subjected to a bending load, ob is the maximum bending stress due to the external moment, M, ab = M/Z (3-6) -where Z is the elastic section modulus. The total Kydue to these three types of loading is
. _ . . __ _ . . . _ _ , _ ._ ~_ .. ._ _ _ . , _ _ . _ . , _ _ . -- . _ . _ _ . _ _ _ _ _ - ,
Page 3-3 Ky = K, + Kt+Eb (3-7) Equations (3-7) and (3-1), when combined together, give the functional form for J,pp. 3-1.2 Longitudinal Cracks The computatation of crack stability for longitudinal flaws is based on plastic zone corrected stress-intensity factor solutions. For a longitudinal through crack in a pipe '
~
K = age F(A) (3-8) where oh is the hoop stress, c is half the crack length, A=c//Rt and the shell correction tern F(A)=(1.+1.3A*)** for A<1 and F(A)=.5+.9A for 1(A(4.45. J,pp can be found as before from Equation (3-1). 3-1,3 Tearing Stability for SSY Conditions The form for T,pp can be found by differentiating the equation for J,pp, following Equations (3-1) and (3-7) or (3-8), with respect to crack length, giving dJ app E_ (3-9) T,pp = da e, Then, using Equations (2-2) through (2-4), the stability of the crack can be determined.
Pogs 3-4 s. 3-1.4' Plastic Zone Instability Failure Vasquez and Paris (11) have shown that situations exist in which the gradient with respect to the crack size of the clastic stress field at the tip of the. crack becomes sufficiently large that the plastic zone cannot maintain stable static equilibrium and plastic zone instability occurs, followed by:the prcpagation (or unstable extension) of the crack. This mode of unstable extension is called a " plastic zone instability failurc" or PZIF). 'The functional form of the PZIF criterion is given by
., K*ggg=2na$c,fg/P, (3-10) i s
4 where P g =112AF'/F, and c,gg, A sad F( A) are the plastic zone corrected terms described in Equation (3-8) .
. 3-2 LEAK RATE ANALYSIS The estimate of the leak rate for various cracks was based upon the LEFM
/ " based methods given in Reference (LR). In general, the leak rate depends upon the applied stress and crack length. Thus, the calculation of leak rate necessitates the development of a fluid flow model for fluid leaking j;,;; , through a crack. It also requires consideration of the thermodynamics of
, the flow and the surface roughness of the crack.
,(!>
- p;
-1 is s
r
Section 4 LSY BASED ANALYSIS Tada,- et al.(11) were the first to apply the tearing modulus stability criteria to actual structural problems using a rigid plastic idealization for the cracked section. They applied it to a piping system for the purpose of evaluating the stability of a circumferential crack in a BWR recirculation loop. The methods were subsequently refined using an elastic plastic approach (11). In this Section, the method of analysis and the crack stability criteria are discussed. The analysis method is consistent with both References 19 and 32, but takes into account the sekavior of structures having more complicated boundary conditions. Additionally, the dependence of T mat 0" I mat is included through the use of a J-T stability diagram. Details of the cracked section are shown in Figures 4-1 and 4-2. Only one through-the-thickness crack (TC) is assumed to esist and that crack is oriented circumferentially. Use of a TC assumption per Criteria 3 was justified. by Zahoor(1Q) for fully plastic bending. He considered circumferential vs. radial instability for part-through cracks (PIC) and concluded that the FTC becomes a TC. See Figure 4-3. Based on the approach that he used, this conclusion can be eatended to cases of fully plastic bending having "small" axial loads. Under the postulated loading, the following conditions are assumed: a) The cross section containing the circumferential, through-the-thickness crack (Figure 4-2), is fully yielded. b) The material local to the cracked section (or hinge) exhibits elastic perfectly plastic behavior.
Page 4-2 4-1 STRULTURAL RESPONSE AND T,pp. The behavior of the pipe is idealized as sections which behave elastica 11y, separated by a plastic hinge. To compute T,pp, there are two system parameters which must be evaluated. Ile first is the compliance of the elastic section and the second is the rotation of the plastic hinge at the assumed crack section under the prescribed loading. 4-1.1 Compliance. By using finite element methods along with the assumption that a plastic hinge is developed at the cracked section of the pipe , the rotational compliance of the elastic section about the hinge location is determined using the JTPIPE programs 21) described in Appendix A. Note that the elastic compilance does not depend on the crack size because the crack section has been idealized to behave as rigid perfectly plastic; thus, only the uncracked section of the pipe behaves clastically. 4-1.2 Plastic Hinge Behavior The rotationni response at the plastic hinge simulating the cracked section requires, computing' -the finite discontinuity in rotation taking place at the cracked section, C,7 (See Figure 4-2). The solution is dcveloped by satisfying compatibility at the hinge. This discontinuous rotational angle is due to the localized deformation at the fully plastic cracked sect ion. ,
, , s A
r 4
Page 4-3 4-2 CRACKED SECTION PARAMETERS 4-2.1 Plastic Limit Moment l l The plastic limit moment of the cracked section, (M,)p, can be defined in terms of simple parameters. For a thin pipe, (t/R)<<1, Tada, et al (19) have shown that the (M,), can be expressed as (M,)p = 4ao R*t(cosp - 1/2 sin (0)) (4-1) where E = (a + n S )/2 , S = at/ "y t t and ao is the flow stress and R, t, and 0 are, respectively, the mean radius .and thickness of the pipe and the cngle defined by the through wall crack. (See Figure 4-1). 4-2.2 J-Integral For the fully yielded cracked secticn nd the rigid perfectly plastic material behavior assumed above, the J-integral can be expressed as follows J epp =c o RF j O ct, (4-3) where O c , is the rotationti angle caused by the plastic hinge at the cracked section, and Fj = sin D + cos 0 (4-4)
Pate 4-4 The above estimate of J,pp based on a rigid plastic idealization was shown to be valid (32) for 0,7>1' without including the einttic plastic effects. 4-3 STABILITY ANALYSIS
~ The approach used to determine stability of a crack is based on a procedure similar to that developed in Reference (1_p_) . It is assumed that for a fixed displacement loading, the sum of the displacement 4 changes -at the cracked section, which can be separated into the elastic part and the plastic part, should ~ue equal to zero. Carrying through with the mathematics, we find T,pp = F2 (0)L,gg/R + F2 (0)J,ppE/(o*R) (4-5)
F2 (0) = 2Fj /n F,(0) = (cosp-2 sin 0)/2F3 L,gg = EI/[K], [K] = min. stiffness at hinge and Fj is given by Equation (4-4).
' Note that Equation (4-5) depends upon the geometric configuration as well as the boundary conditions of the piping system and J,pp.
Page 4-5 4-413E J-T DIAGRAM For a given piping system and material, the pipe diameter, wall thickness and flow stress are known. Then, for any given pair of 6 and 9,,, Fj can te calculated using Equation (4-4) and J can be found from Equation (4-3). The- rotation caused by the plastic hinge at the crack section, 0,,, depends upon the interaction between the various segments of the piping and the boundary conditions imposed. The T,pp values corresponding to the computed J,pp are obtained vic Equation (4-5). In this Equation, the only quantity that is unknown is L,f g, the ef fective length of the piping. Since actual piping systems are typically 3-dimensional structures, it is not always easy to compute the effective length (of an equivalent straight pipe) by simple analysis methods. . Hence, a clastic plastic finite element analysis of the piping system'is performed using JTPIPE (21) from which the elastic compliance is computed to determine L,gg. For -the particular loading, J,pp is computed based on the structural response. The computed (applied) values, J,pp and T,pp, are then used to generate the applied curve shown in Figure 2-1. Note that the applied J-T curve shown does not account for crack growth, that is, 20 is assumed constant throughout. The error resulting from this approximation is small and allows conservative conclusions to be drawn from the analysis, i l [
e
<C 2
1 E 1 m O g
~
5 S 4 N h n N. g O' h
\ ^
By' ?> NE i m R\/ i V i E a J C y -,n--y- mmuy,- m n e- - -Iy 2 ow
- w" w '" *=<- #I'w
7 -
-h l'
CRACKE D r- YCr
')
(Mc)r > (Ivic)P j
>el i k /
~ FULLY P L AST I c
'igum 4-2 Fully Plastic Bending of a piP o with Axial Load I
l. 1 i I
1.0 ~, . . . -
's X= a/t
's %
O g O 0.8 -
\ '+-
\ \s 0.6 -
s\
\
j g g X=0.1 9, y0 -
, x 0.4 _ x = 0.9 s _
\
\
X=0.5 g \
\ \
~ ~
* \
AXtAL LOADING
- - - -PURE BEND LOADING g g
\ \
I I I ' '\ O i O O.2 0.4 0.6 0.8 1.0 ! Figure 4-3 Stability of Part-through Crack Under Fully Plastic Bending
5 Section S
' RECIRCULATION SYSTEM, RESULTS AND DISCUSSION The safety of the recirculation system piping at the Vermont Yankee Nuclear Power Station focuses on the postulated existence of large cracks in the piping. The evaluation of this system begins with a thorough . description- thereof including the code stresses, pipe geometry and operating pressures as well as the appropriate isometrics.
The application of the criteria used for the analysis of the recirculation system piping is based on that of Section 1-3. The approach used is described in the following Sections. The reference to
- Criterion 1) through 5) in the following Sections refers to those defined in Section 1-3. The material properties used in the following Sections were developed in Appendix B according to Criterion 4.
5-1 RECIRCULATION LOOP PIPING SYSTEM The recirculation system provides a continuous flow of coolant through the RPY in order to achieve heat transfer rates greater than that
.possible by natural convection. The system is composed of 2 similar loops which are referred to as loops A and B. In this study, only loop A is considered. Loops A and B are the same except for the RHR return line attached to the loop A suction line.
_ _ _ - . _ . - - _- . . - _ _. - _ . . _ . _ _ _ _ - _ . . - . . - -~ _ _
Page 5-2 1[he portion of the recirculation loop piping system that is of interest in this study is limited to the the suction, discharge, header and riser piping portions of the loop. The RHR piping, etc. is not considered except that it is included in the structural idealization as the stiffness of the RHR piping effects the crack stability calculations. 5-1.1 System Description Following the isometric view of loop A, shown in Figure 5-1, the system can be readily explained. Flow from the RPV is via the 281n suction line at a pressure of 1040 psig under normal conditions. The normal operating temperature for the suction line and the rest of the system is-528F. The coolant flows through the suction line to the pump and exits at a pressure of 1130 psig under normal conditions. The discharge line is also 28 in and provides coolant to the 22in header. From the header, coolant is distributed to the RPV through 5-12 in risers located at 30' intervals. The location of existing anchors, cupports and pipe-whip restraints are shown in Figure 5-2. Locations for the snubbers and hangers are omitted because their ef fect is neglected in this study. The 5 riser nozzles and the suction nozzle are the anchor points for this system. Quasi-static vertical movement of the pump is not inhibited, but large vertical displacements are limited. The loop A header is connected to the loop B header. For structural idealization, an elastic spring is included to properly include the stiffness of loop B in the loop A analysis.
Page 5-3 5-1.2 Piping Code Structural Analysis Because a stress analysis had already been performed by General Electric (11), as part of the design of the NSSS, it was not necessary to perf onn another. The GE stress analysis (GESA) results (21) used herein are 'taken- directly from the stress report. The leak rate computation required by Section 1-3,' Criterion 1 uses normal operating stresses. But, the normal stresses in the GESA were not directly applicable to this
-analysis. Thus, a conservative approach was taken. For purposes of computing' leak rates, the portion of the stress due to dead weight and thermal effects was neglected and only the pressure term was used.
For the crack stability calculations of Section 1-3, Criterion 2, Level D stresses are required. These could be taken directly from the GESA report, but with some judgement. The GESA Level D stresses are based on the resolved moments about 3 principal axes and an assumption that SSE = 20PE. Because one term is a torsional component, it does not contribute to circumferential crack extension. Thus, it can be removed for computing J,pp. The pressure stress used corresponds to the maximum pressure during the bonding transient. The stress terms result from dead weight, thermal and SSE. The stresses used are given in Table 5-1 and their compohents are given in Table 5-2. For computational simplicity, the maximum value of the Level D stress along any line segment is used in lieu of point by point documentation. This approach tends to be conservative but greatly simplifies the comprehension of the analysis.
Page 5-4 Table 5-1 also lists the pipe section properties used for this study along with a tabulation of operating and design limits on pressure and temperature. Minimum wall thicknesses are used for the crack stability calculations which follow. 5-2 LEAK DETECIABILITY Criterion 2 of Section 1-3 requires the demonstration of the stability of _ e a crack that has a length equal to that which would result in detectable leakage rate as determined under Criterion 1. For this analysis, rates of I and 10 gpa, under normal operating loads, were selected as being representative of a leak that is detectable using existing sensors. 5-2.1 Circumferential Flaws i The leakage rate computation is conservatively based on normal operating stresses that-result from.the suction side operating pressure (1,040 psi) component (21) alone. As the suction side has a lower operating pressure than the discharge side, the cracks sizes computed will be longer than what would actually exist on the discharge side. The dead weight plus thermal components of stress were conservatively ignored. No dynamic
-Ioads are used in developing the normal stresses. It is noted that the lower the stress, the lower the leak rate, and the longer the crack must b's in order to have a detectable leak. Leakage rates were computed for a series of crack sizes based on the computed operating stresses. Crack lengths ranging from 6.3 to 12.9 inches corresponding to rates of leakage between 1 and 10 gpm. The results are shown in Figure 5-3 and Table 5-3.
l l
Pogc 5-5 1 l 1 l 5-2.2 Longitudinal Flaws. ' The leak rate for longitudinal flaws was computed using a hoop stress again conservatively based on a normal operating pressure of 1,040 pal (11) . The range of flaw sizes considered, ranged from 3.8 inches for the .1 spa leak rate to 7.6 inches for the 10 gpm rate as shown in Figure 5-4 and Table 5-3. 5-3 CRACE STABILITY: PRESSURE + THERMAL + SSE LOADS This assessment of crack stability relies on the small scale yleiding (ssy) theories discussed in Section 3 and Criterion 2 of Section 1-3. 5-3.1 Circumferential Flaws The solution of Equations (3-1) through (3-7) for circumferential flaws was obtained using the computer program, "0YCJT"(12), which performed the ; necessary iterations on K to obtain the plastic zone corrected K values. From the E(a+ry) values, the appropriate J,pp estimates were determined. This evaluation was performed using the pressure plus thermal plus SSE stresses (11) and the results are included in Appendix D. s Crack lengths corresponding to the lengths that cause leak rates of 1 and 10 gpm plus 2t were considered. For the 1 gpm cases, J,pp<J y, a:suming Jy, = 1300 for SMAW welds (or field welds). This insures stability and compliance with Criterion 2. This conclusion also holds for a 10 gpm
Page 5-6 rate for all cases except the riser. If SAW (or shop welds) are assumed, then Jy ,=500, and a small amount of crack extension might occur. For the levels of J,pp computed, only small amounts of crack extension would occur and T,pp is small (<7). Thus, no crack instability is indicated ice any location. Refer to the results in Appendix D. 5-3.2 Longitudinal Flaws Crack stability, as evidenced by J<J ,yand J<Jpggg, was checked using the hoop stress at the. pipe wall mid plane. Upon substituting the appropriate crack lengths (2c(10 spm) plus 2t) and stresses into Equation ( 3-8) , we find, for the 10 gym crack, that the maximum value of the
. plastic zone corrected value of J,pp = 340in-1b/in*, which is much less than J, y for either weld type, thereby insuring crack stability and no crack extension. Having satisfied the fracture toughness criterion, a check for a plastic zone instability failure (PZIF) was made following the methods of Yasquez and Paris ( H). Jp,gg was computed using the relation -of Equation (3-10) and it was found, that J,pp<Jpggg thereby satisfying the PZIF criterion. The results are included in Appendix D.
These computations were made using the "PZIF"( M) computer code.
. 5-4 CRACK STABILITY, UPPER BOUND LOADS In this section, the methods used were based on large scale yielding (1sy) theories and structural ductility. These satisfy Criterion 3 of Section 1-3.
b._.
t"o
]
Page 5-7
~5-4.1 Applied Loads Tho' intent of the Criterion 3 of Section 1-3 is to insure that brittle behavior of the piping system does not occur. This is accomplished by - demonstration of structural ductility (lg) in the presence of cracks. To prove this, the cracked section of the pipe must be capable of absorbing large amounts of energy.
It is important to consider the maximum load that can be applied to a structure within the intent of current laws, namely,10CFR50 App. A, Criterion 2(1). Criterion 2 and other Criteria of 10CFR$0, App. A require that the uncertainty in predicting the magnitude of loads resulting from natural phenomena, such as seismic events, must be lacluded in design of the plant. Thus, it is prudent to use conservative assessments of the magnitude of maximum loads. To do this, postulated "appe r-bound" loads, based on a structural ductility approach (1Q) to essess crack stability, are used for the analysis. As a result of this approach, the methods being used herein are, in essence, demonstrating that the recirculation systes piping is safe under extreme accident conditions. 5-4.2 Crack Stability Calculations Using Criterion 3 of Section 1-3 and the applied loads determined in accord with Reference IQ and Section 5-4.1, the stability of several crack sizes was examined. 11e analysis was performed using the JITIPE(11) program and the J,pp value was computed using the foregoing s
. .. n. s load assumptions. JTPIPE has several analysts options. The option.
selected accounts for the interaction of the piping with surroun'ing The material property values used were obtained from structure. Reference 29 as discussed in Appendix B. A temperature of 550F w.is assumed at every crack location. The stability of circumferential cracks having lengths of 20 - 60 and 120' are considered under the application of the upper-bound loads described in Section 5-4.1.
Page 5-9 The piping system was idealized for analysis using the JTPIPE code (21). Structural details were taken from appropriate drawings (25-27). Nodal locations are shown in Figure 5-5 and the elements corresponding to crack location are shown in Figure 5-6. The results of the analysis are shown in Figure 5-7 and 5-8 for circumferential crack lengths, 20 = 60 and 120' respectively. The J-T stability diagram approach due to Paris (6) was used. The material data is based on the J-modified method developed by Ernst (11). The total J,pp is composed of displacement controlled loads plus inertial loads following Equation 2-6. The latter were computed using a structural ductility approach (10). It is concluded that the most critical locations, which correspond to elements 74, 75, 81 and 82, can tolerate large defects and safisfy the ductility criteria for Equation 2-5. , i e
-- -- _-.---.----,-,,,,..-,-----,,--..,,,.w..-., , . . , , , , , , , , , , , , , . , , , - - , , . , , . , . . - - - - , . . , , , , , --,,___y,,,,.,,e.,-,,,. a. - - .
Page 5-10 Table 5-1 Section Properties and Maximum Stresses Line Dis t wal P ' Lev (in) (in (p!if* pdes)* (psi (psi ** Suction 28.17 1.151 1040. 1148. 19751. Discharge 28.34 1.23 5 1130. 1233. 18007. Hender 21.88 0.976 1130. 1233. 19413. Riser 12.75 0.6 87 1130, 1233. 19413.
- Temperatures: Top,,=528F; Tde s=57 5F
** Maximum at any point along line Table 5-2 Summary of GESA Results Line pdes F,13,g MB MC M eff (psi) (kip) (in-kip) (in-kip) (In-kip)
Section 1148. 0* 2376. 1371. 2743. Discharge 1233. 20.1 1770. 2938, 3403. Hender 1233, 16.1 1343. 1802. 2248. Riser 1233. 0* 1436. 788, 1638. F,13,3 = Fdw + Fth + 2Fobe if,g g = / MB+MC 3Mi"IMdw + Mth
- 2Mobe 3 , i=B,C 1
1
- Compressive Load Conservatively Ignored.
- J f
f
, - - . v - - - - . --
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- lable 5-3 J,pp for Leak Rates of 1.0 and 10.0 gpm Leak Rate Line Crack Crack Length, 2c J (sym) .
Orientation (inches) in.IEEin* 1.0 Riser LONGIE DINAL 3.8 97 1.0- Header LONGIEDINAL 4.1 116
- 1. 0 Discharge LONGIEDINAL 4.6 155 1.0 Saction LONGIE DINAL 4.3 13 8 10.0 Riser LONGIE DINAL 5.9 277 10.0 Header LONGITUDINAL 6.8 340 10.0 Discharge IANGIE DINAL 7.6 333 10.O Suction LONGIW DINAL 7.2 300 1.0 Riser CIRCUMFERENTIAL 6.3 540 1.0 Header CIRCUHFEP.ENTIAL 7.0 110 1.0 Discharge CIRCUMFERENTIAL 7.6 85 1.0 Suction CIRCUMFERENTIAL 7.2 65 10.0 Riser CIRCUMFERENTIAL 9.9 1700 10.0 Eesder CIRCUMFERENTIAL 11.7 260 10.0 Discharge CIRCUMFERENTIAL 12.9 1 90 10.0 Section CIRCUMFERENTIAL 12.3 151
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- 100.00.) A HEADER 1040 22.00 0.976 5.1 0.0
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f: rw no. 4171 PROJECT: VERM0lli TAlmEE SYtt T I TLE P (rcil D (in! ' (it.) tih (ksi) c) RISER 104D 12.DD D.687 8. G 100.00;. A !!EADER 104D 22.00 0.976 10.6
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VERMONT YRNKEE RECIRC - NODES
. PLOT REFERS TO RUil110 - 4102 m 2s is nuw no . vio3 3
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5 1 0 8 Figure 5-6 Crack Locations at Connection Elements
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_ jii : (.- Section 6
SUMMARY
AND CONCLUSIONS r The Vermont Yankee recirculation loop piping was analyzed using
' structural ductility methods ( "s da") in addition to the conventional lenk-before-break ("Ibb") approach. Integrity of the piping containing
. postulated flaws was demonstrated for both of these.
4 Using the 1bb approach, it was.shown that cracks, having lengths l wh'ich would result la readily detectable leaks, were stable under Code
- - loads (Level D'or faulted conditions). Stability was shown for both longitiidinal and circumferential cracks.
The sde ' approach also demonstrated lategrity. For the sde analysis, apper-bounds on local plus global (or Anertial) loads were used for caroumferential cracks havIng lengths of 60 and 120*. As a conservatism,
'It _ was assumed that all snub'bers and hangers were not fatact. Integrity was shown to salst with ample margins of safety.
,s 1 4
4 This' analysis demonstrated that the sde approach la valid for BWR systems. It was previously demonstrated for a PWR system (1Q). It further proved that it could be used to evaluate flaws found during 151. This conclusion is limited however, because the boundary conditions play an important role in J,pp and T,pp, and they are plant specific. In summary, it was concluded that the Vermont Yankee recirculation loop piplag has large margins of safety against fa!!ure due to the presence of flaws. (
\
p Section 7 REFERENCES
- 1. Title 10-Chapter 1, Code of Federal Regulations, United States Nuclear Regulatory Commissloa, Part 50, Appendix A, General Design Criteria for Maclear Power Plants (Abbreviated 10CFR50, App. A.)
2 USNRC Letter to Consumers Power d6ted 12 December,1981. 3 ASNE Boller and Pressure Vessel Code, Section 3, 1974 (and subsequent ed).
- 4. Rise, J.R., Jour. Appd. Nach., Vol. 3 5, 196 8, pp. 379-3 88.
5 Butchinson, J.W. and Paris , P.C., " Stability Analysis of J-Controlled Crack Growth", ASTM STP-668,1979, pp. 37-64.
- 6. Paris, P.C., "A Nethod of Application of Elastic-Plastic Fracture Neobaales to Nuclear Vessel Analysis", USNRC Report NUREG-0744, App. A, April 1981.
- 7. Paris, P.C., et al., "A Treatment of the Subject of Tearing Instability".
NUREG-0311 Aug.1977.
- 8. Paris, P.C. and Tada, N., " Farther Results on the Subject of Tearing Instability - I ", NUREG/CR-1220, Vol. I, Jan.1980s and Zahoor, A. and Paris, P.C., "Further Results on the Subject of Tearing Instability -
II", NUREG/CR-1220, Vol. II Jan.1980.
- 9. Tada, M. and Paris, P.C., " Tea ring Instability Analysis Ha ndbo ok ",
NUREG/CR-1221, Jan.1980. 20 Johnson, R.E., et.a1., " Resolution of the Reactor Vessel Naterials Toughness safety Issue", USNRC Report NUREG-0744, April 1981. 11 Paris, P.C., Tada, M. and Baldini, S.E., 'Tra ct ure Proof Des i g n ", in "CSNI Specialists Neeting on Plastic Tearing Instability", USNRC Report, NUREG/ CP-0010, Jan.1980.
- 12. Zahoor, A. and Kanninen, N.F., "A Plastic Fracture Mechanics Prediction of Fracture Instability in a Circumferentially Cracked Pipe in Hcnding",
July 1990, ASNE Paper 80-WA/PVP-3 Accepted for Publication in the ASME J. of P. Vessel and Technology,1981.
- 13. Gedas, J.P. and Joyce, J. A., " Degraded Pipe Esperimental Program". IISST Review - VIRG Hesting, 23 + 24 July 1980. Silver Spring, MD.
14 "A Critical Review of Nethods of Alleviating the Requirement to Postulate Gul!!otine Breaks," Paris, P.C. and Cotter, K . ll . , CSNI Specialista Neoting on Leak-before-break, Monterey, CA. Septembct, 1983.
- 15. Folias, E.S., "A Cirw om f e ren t ia l Crack ir t Pressurized Cylindrical She!!, Int! . J. cf Fracture Nect anica, Vol . 3. pp.1-11,1967.
Page 7-2 x 16 Erdogan, F. and DeLule, F.,
" Ductile Fracture of Pipes -and Cylindrical Containers- with a ' Circumferential Flaw", ASME J. of Pressure Vessel Technology, .,Vol.103, May 1981, pp.160-168.
- 17. Vasquez, J. and Paris, P.C., *The AppIlcation of Instability.
the Plastic Zone Criterlon to Pressure Vessel Failure", Journadas Metalurgicas,,.Sociedad Artentina de Metales, Cordoba, Argentina. Nov.1970,9also la NUREG/CP-0010, Sept.1979.)
- 18. "LERATE". USNRC computer code for determining leakage rates through cracks in nuclear piping and pressure vessels'. Transmitted by M. Boyle, USNRC/SEP to FPDC .on Sept. 9, 1982,
- 19. Tada, M., Paris, 'P.C. and Gamb'l e , , R., " Stability Analysis of Circunferential Cracks .-In Reactor Piping Systems", NUREG/CR-083 8, June 1979. ; '
Iv -
- 20. Zahoor, A., Monthly Progress Report, EPRI Project T11'8-9-1.
w -
- 21. "JTPIPE", A - Finit e Element Program for Computing the Stability of
- Circusif erential Cracks in Np~itg, .Ver's' ion: 2, Level 3, FPDC, St. Louis , M0
~(PROPRIETARY). , _-
~
- 22. "0YCfY", A Computer Program for, Computing Values of Circumferential Cracks in Pipes, Version .1, Level 5, FPDC, St. Louis, J,pp for MU.
- 23. "PZIF", ;A Computer Program for Computing Values of the J and J
~ Longi tudinal Cracks ' irs Pipes, Version 1, Level 3 5.FPDC, N Louis # hb.for
- 24. GE St ress Analysis Report 22426L5 6/10/73: reviewed by KHC at Vt. Yankee
' 06/26/84.
N
- 25. Recirc'ulation Loop Piping, YAEC Drawing 5920-26 8, Rev. 9, 6/ 83. ,
26 Recirculation Loop Pipe-Whip Rostraints, YAEC Drawing 5920-424, Rev. 5. ,
- 27. RBR Piping,.Ebssco' rawingv G-191210, Rev. 9 and G-191211, Rev. 11.
'28 A Miscellaneous info $mation obtained by K11C during June 26, 1984 trip to Vt. Yankee; includes Materidis,- Operating Tempertures and Pressures.
- 29. ; " Development of Material Proprty Data for the Tearing Stability Analysis of -the Indian Point 3 Primary Coolant System", FPDC Report 83-104, Rev. 1, September 28, 1983 (PROPRIETARY).
= '4 0. "Sammary.of'the Tearing Stability Analysis of the Indian Point 3 Primary .
-Conlant Systen", K.H. Cotter and Paul C. Paris, FPDC Report 83 -7 5, Rev.1, May 4,1984 (PROPRIETARY) .
L
' 31. "Ma t e rini > Resistance and Instability Beyond J-Controlled Growth",
H. A. : Ernst in'EI r s t ic-Pla s t i_c - Fra c t ur e : Second Symposium, ASDf STP 803, C. F. Shih and J. P. Gudas , Eds . , AS'IN, 1983. l3 2. The Ap~ plication of Fracture Proof Design Methods L ing Tearing Instab~111ty Theory to Nuclear Piping Postulating Circumferential Through Yall Cracks " Paris, P.C. and Tada, H., USNRC Report NUREG/CR-3464, 6 Sept. ,19 83 ~. 4
, , , . , - +-- - , , , -
----n- .v ,. , . - - --e, r- ,- - - , -r ,
y. . ... ,e 9 I ' ) . 1 I APPENDIX A r s , j l c ). JTPIPE 3 3 A FINITE ELE!!EhT PROGRAM FOR C0!!PUTING ~ PIPING SYSTDI CRACE STABILITY PARAMETERS r o l t i h s l. } i k-i a run _ , _ , _ _ . _
l'a ge A-2 C0hT&TS A-I INTRODUCTION A-2 APPROACH A-3 ANALYSIS AND IDEALIZATION OF THE STRUCI11RE A-3.1 ELEMENT TO STRUCTURAL MATRICES A-3.2 BOUNDARY CONDITIONS
'A-3.3 COMPLIANCE COMPUTATION AT CRACE SECTION A-4 PROGRAM ORGANIZATION A-4.1 NODAL POIhT AND ELEMEhT DATA INPUT A-4.2 ASSEMBLAGE OF STIFFNESS MATRIX A-4.3 COMPLIANCE CALCULATIONS A-4.4 COMPUTATION OF J,pp A-4.5 COMPUTATION OF T,pp 1
, - - - - . , , ,-,-n, -- -- - , , . , -,., --
,-y r
r Page A-3
' A-1 INTRODUCTION t~ 12 NUREG/CR-0838 Tada, et al., applied tearing modulus stability concepts to a s3103ted. nuclear reactor piping system geometry and concluded that the piping system was " fracture proof"; that is, unstable ductile crack extension was shown to be unlikely. This was a maj or breakthrough for the inelastic fracture
. mechanics analysis of piping. However, in Tada's analysis, the piping system was idaclized as a straight beam with simple boundary conditions and the vaine of J,pp w:s specified. In general, the geometry and the boundary conditions of a nuclear pipics system are complicated. To extend the application of Tada's approach to cat:01 piping systems, it became nec e s s a ry that a finite element program be d;valoped to overcome the structural compicxities of typical piping systems and to
- s:rp;te the value of J,pp for the case of interest. The JTPIPE program was devaloped for that purpose.
This Appendix summarizes the capabilities of the current version of the JTPIPE etap:ter program. The detailed theory and the numerical techniques used in JTPIPE cre Ect presented in this Appendix.
. The piping systems to be analyzed with JTPIPE can be modeled by combinations of fear dif ferent types of finite elements. The four element types are:
a) 3-d straight beam element b) 3-d curved bcam element c) Flexibic connection element d) Special element
Page A-4 A-2 APPROACll Th3 program determines the elastic compliance of the piping system at specified lec:tions for use in the crack stability analysis. The location of the maximum cenpliance is also identified. The computed compliance values are then used to determine principal stiffnesses at each location to be analyzed. From the minimum
' stiffness at each location, the Legg/R is determined. The Legg/R data is stored 1 fcr post-processing.
Using the_ aforementioned Legg/R data, J,pp and T,pp are computed using Equations (3-3) and -(3-5) -for each postulated crack location in another program. These lattar values are tabulated for a reries of ci rc nm fe rential through-wall cracks having included angles of 60 to 300 degrees in 60 degree increments. Alt e rna t ely, sp elfic angles can be selected. All J vs. T data is saved and later utilized for etcputer ~ plotting the stability diagram where corresponding material- resistance in
'th) torm of Jest vs. Tmat is also-included.
A-3 ANALYSIS AND IDEALIZATION OF THE STRUCI'URE In this section, a brief. description of the method of idealization of the straiture i s -- presented. The direct stiffness method is used to analyze the i stractural systems. ' A-3'.1 Formulation of Structural Matrices A piping system .is basically a three dimensional frame. It can be idealized as a stabar of discrete beam (straight or curved) elements, flexible connection
1%ge A-5 sicaants and special elements. The beam elements are two node elements with six degrses lof freedom at each node. The stiffness matrices of the elements are 12 x
'12 'cymmetrics1 matrices which can be directly formulated f rom beam theory. After 'the- trans formation from the local element , coordinate system to the global ocardinate system, the total system stiffness matrix can be formed by direct cddition of the element matrices according to the index of the degree of freedom.
It ccn be expressed in the following manner: N s- (m) K K gj (A-1) 3j=}1 m= (m) whns K3 ; is the stiffness matrix component of the total system, K 33 is the stiffness matrix component of the m th element and N is the total number of 01saants in the. system. The external force can be expressed in the form: F 1=fK33
- Uj (A-2)
J
-whars F; is the external force applied at the i th degree of freedom and Uj
~
is the displacement at the jth degree of freedom. A-3.2 Bourdary Conditions Ta simplify the programing problems associated with the specific displacements on
~ths boundary, a spring that is very stif f in compr.rison with the structure, is assumed to connect the bounda ry nodal point to a fixed point. If the applied utdal displace =cnt component is zero, the node will be restrained by the stiff
l'a g e A- 6 l l 1 spring. If a non-zero displacement component is specified, it can be replaced by that nodal point. The equivalent force is as equivalent force applied at evaluated by the specified displacement applied on the stiff spring with the Since the spring is much stiffer than the system structure stiffness ignored. ;
)
str:cture, the error introduced is negligible. l These elements may have G:p elements are included as a feature of the program. tny one of the principal directions. Displacements limits can be specified in either the 11,17 or 12 directions. A-3.3 Compliance Computation At Cracked Section in a piping In the stability analysis of a through-wall c i rcumfe rentia l crack system, the rotational compliance at the pipe cracked section is required for the the fact co putation of the applied tearing modulus, T,pp. This is because of the pipe idealized as a plastic hinge. The that the cracked section of is flexural rigidity rstational compliance at the pipe cracked section is due to the cf two elastic piping sections joined by the hinged section. Fr a the total system stif fness, including the boundary conditions, as formulated in Se c t i on A-3 .1 an d A-3 . 2, the rotational compliance at the pipe cracked section section. ecn be obtained by applying unit moments on opposite sides of the hinged Thi principal rotational compliance at that section and the maximum rotational secpliance of the selected locations in the piping system are both calculated.
Pnge A- 7 A-4 PROGRAM ORGANIZATION {sf Ile computation process in the JTPIPE program is basically divided into five distinct phases plus post processing. A-4.1 Nodal Point And Element Data Input In _ this phase, the control .information and nodal point geometry data are input and c;dal points are generated by the program as required. The indices of the degrees
. cf freedom at each nodal point are established. The element data are input and clenent groups generated, the element connection arrays and the element coordinate te nsformation matrices are calculated and all element information is sto,ed r in a L diss file for use in the second and third phases.
A-4.2 Assemblage Of System Stif fness Matrix JTPIPE uses a compacted storage scheme in which the system stiffness matrix is stered as a one-dimensional array. In the second phase, the index of the storage is sstablished, then the system stif fness matrix is essembled and modified to cotisfy the boundary conditions. A-4.3 Compliance Calcula tions 13 the third phase, the locations of the postulated crack locations desired for
- th 2 compliance computation, are input. The rotational compliances and minimum stiffnesses at each cracked nodal point is calculated based on the response of the structure to the irposed load. The status of gap elements (open or closed) are L_
l'a ge A- 8 tnksn into account at this point. Next, the Legp/R are calculated and stored for post processing. A-4.4 Computation of J,pp or an input value for rotation
.J epp can be specified by an input vaine such as J Ic at the cracked section. Alternately, J,pp can be determined from the response of tha structure. This latter method is the preferred approach but involves c nniderably longer computer run times. -A-4.5 Computation of T,pp ' Finally,'a post processor is used to compute T,pp for specified crack sizes and crack rotations. The data is displayed in tabular form and is stored on a disk fer subsequent post processing: namely, the generation of J vs. T diagrams.
^^ J
L 6 t APPENDIX B { MATERIAL PROPERTY DATA i x 4 s k i 1 3
l s=Jie qJ ' YERMONT YANKEE
\ RECEC BPJNG -
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COMMITTEE CORREbrummmu
-Keep ASME Codes and Standards Department Informed- '
l l l Attachment #1 AGENDA Task Group on Pi p ing Fl aw Anal ysi s San Antonio, Texas April 23,1984 -___=====- -
==========_=--
Chairman Approval of Minutes / Agenda / Elect Secretary John Landes EPRI/W Weldment JR Tests and Data Summary Mike Vasseleros NRC/DTNSRDC JR Curves From Pi pes and Data Surrenary Fred Copeland Effect of Weld Procedures on Initiation and Toughness Douglas Norris
-R;und Robi n QA Cal cul at i ons Summary All o Resolution of Differences
. Low Toughness Weldment Issue - New Calcul ations Sensitivity of Instability Load to Yi eld Stress Asao Okamoto o
Har Mehta o Consideration of Secondary Loads /New Results New Results Ron Gamble and Akram Zahoor o Fred Copeland o New Results Joe Bloom o New Results Loads Issue Other Speakers on the Low Toughness We l dmen t/ Secondary 82td [arl3 All Conclusions and Recommendations __====.---- -__====================______=====
=======_=========-
Douglas Norris Techn i c al Support Document Recommendations Other Issues Recommendations for Changes to Table IWB-3641-2 Fred Simonen o Gery Wilkowski o Ligament Collapse o Other New Issues b
- a. ..
Afrsl 23 f/174 9tDuhN 3pln L2h Attachment e3 CONTENTS OF PACKAGE
~A. Tables
- 1. Weldment Types
- 2. Table of Properties-B. Data Curves
- 1. ' Stress-Strain 304 Base, TIG RT, 5500F
- 2. J D
vs. Aa R Curves All Data
- 3. J m
vs. Aa R Curves All Data C. Comparison Plots - R Curves, 5500F
- 1. Base vs. Weld
- 2. All Welds
- 3. CL vs. CR Direction (Compact vs. Bend Bar)
k WELDMENT TYPES. SS Type Type Practice Source Identification , 304 Submerged Arc Shop J.A. Jones SA (A) 304 TIG Automatic J.A. Jones TIG (B) l-Field 304 Shielded Metal Manual. J.A,. Jones SMAW (C) - Arc Field l 316 Submerged Arc ? Shop Battelle SW (E) ; (Nine Mile Point) i i 304 Base 4B (B) 4CB (C) 316 Base 6B (E) ; 4 l i'
Test Y.S. . T.So CVN. In Ib da/dJ (from JD)
.in;2 :
psi - Material Code Temp. ksi- -ksi ft-Ibs SA - WM A 750F 50.4 87.0- 49/81 580 '25500. . A 5500F 36.0 61.8 .46/109 556 7600 i-A 75 F Bend Bar 570 21950 l A 5500F Bend Bar 360 11600 .! . SA - HAZ A 750F -- -- -- 5200 69600
- A 750F -- -- --
4360 54400 0 304 - BM 4B 75 F 38.2 81.0 239 5000 67900 4B 5500F 23.1 61.3 221 4000 37900 TIG - WM B 750F 68.9 90.5 140 2314 81400 0 B .550 F 53.9 63.4 239 4480 33000 0 TIG - HAZ B 75 F -- -- -- 3700 50400 0 B 75 F -- -- -- 6000 65500 0 . 316 - BM E 550 F 33.2 72.7 239 4000 38400 l 0 WM E 75 F 60.0 91.8 34/65 690 21100 l WM E 5500F 40.8 70.3 35/96 650 9400 ; ilAZ E 750F -- -- -- 1700 57800 l HAZ E 5500F -- -- -- 4000 19800 0 304 - BF C 75 F 42.3 86.2 239 5900 60200 0 C 550 F 25.3 61.5 228 4580 22800 , WM C 750F 62.6 87.-8 71 1530 26800 d WM C 5500F 46.9 61.4 84 990 14500 0 '- HAZ C 75 F -- -- -- 1900 61700 HAZ C 5500F -- -- -- 4200 25100
Attachment #4 HRC PIPING MATERIALS PROGRAM M. G. VASSILAROS R. A. HAYS DINSRDC J. P. GUDAS o STAINLESS STEEL COMPACT DATA CF8A WELD + BASE PLATE l 30!! WELD + BASE PLATE o 4 INCH DIAMETER WELDED 304 STAINLESS STEEL PIPE , CIRCUMFERENTIAL THOUGH FLAW LOAD VERSUS DEFLECTION .
MECHANICAL PROPERTIES OF STAINLESS STEEL BASE METAL Al4D WELI) IEMP Y.S. U.T.S. % ELONG. ("F) (KSI) (KSI) (2 IN.) % R.A. RT 36 6 89 0 68 77 TYPE 304 1 STAINLESS BASE METAL 300 27 0 71 8 54 77 550 22 0 69 9 50 72 l l RI 67 4 88 7 38 65 I I WElo ' 300 51 6 69 1 29 59 RANSVERSE) 550 49 0 65 6 25 55 CF8A RT 43 81 57 67 STAINLESS STEEL BASE flETAL 300 29 67 45 70 550 45 78 40 52 RT 64 85 48 46 WELn 300 48 73 33 59 (LONGlTUDINAL) /
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RECIRC PlPING Naberial Test Teg (aF) me stress (xsi) Jg ('$O . 5A- WM Eso @8. 9 557o l 2o4- BM sro @,2 t/ coo TIG-wM SCO SE lo @EO 316-BM Fro 63.0 YoOO 31G-WM cr0 SElo (So 304- BM sro 43.f 15~80 304-wM sro 51,2 990 soy samttss sso dio.O 6ASE MET 8L M sThtuss 550 57.3 - TMf45 VERSE WRD CF9 A STAWLEss 5fC 63,6 LogtTLADICAL (AJEth L L. - - - - - - - -
APPENDIX C USNRC ALTERNATIVE CRITERIA i l
umr - memmemmunersumarasmuursumsm - -w -._ muman- - x USNRC ALTERNATIVE SAFETY ASSESSMENT FOR SELECTED IIIGH ENERGY PIPE BREAE LOCATIONS AT SEP FACILITIES 6 This assessment is required only if a LWR high energy piping system (i.e., 275 psi or higher) is being considered. It is only required if a postulated double ended pipe break would impair safe system shutdown by pipe whip (lacking pipe whip constraints) consequences, or by the consequences of the implied leakage or its jet action. The following guidance is for a safety assessment that may be permitted as an alt e rna t ive to other system modifications or alterations for locations where the mitigation of the consequences of high energy pipe break (or Icakage) have been shown to be impractical. Guidance for Alternate Safety Assessment The suggested guidance are as follows: A. Detectability Requirements Provide a leak detection system to detect through-cracks of a length of twice the wall thickness for minimum flow rates associated with normal (Level A) ASME B+PV Code operating condition. Both circumferential and longitudinal cracks must be considered for all critical break or leak locations. Methods for estimation of crack opening areas are attached in Appendix 2(not included). Surface roughness of the creck should be considered. B. Integrity Requirements (1) Loads for Which Level D is Specified (a) Show that circumferential or longitudinal through-cracks of four wall thicknesses in length subjected to maximum Level D loading conditions do not exhibit substantial monotonic loading crack growth (e.g., staying below J yc or Eye by plastic zone corrected linear-elastic fracture mechanics methods or a suitabic alternative. For 4t flaws that are calculated to be greater than Eye or J ye, consideration will be given to: (1) flaw growth a rguments , (2) postulation of small flaw sizes than 4t if justified by leak detection sensitivity. Also assure that local or general plastic instability does not occur for these loading conditions and crack sizes. (b) Under conditions in "B.(1)", show that the flow through the crack and the action of the jet through the crack will not impa ir safe shutdown of the system. Acceptable methodology for the estimation of crack opening area for a circumferential through crack in a pipe in tension and bending and for longitudinal cracks subject to internal pressure are attached. (2) Extreme Conditions to Preclude a Double-Ended Pipe Dreak Using clastic plastic fracture mechanics or suitable alternative, show that c i rc um f e ren t i al through-cracks will remain stable for local fully plastic large-deformation bending conditions under the following additional conditions: (a) Fully plastic bending of the cracked section is to be assumed, unless other load limiting local conditions (such as elbow collapse) dictate maximum bending loads, for all critical locations. C Letter to Consumer's Power dated 12/12/81.
sI11'l;RNA11VE CI:11 Eld A l'n ge C-3 (b) Assume all system anchors are effective. To simplify the analysis, supports may conservatively be considered inoperative. If supports are included, consideration should be given to the adequacy of the support to resist large loads. (c) Other "as built" displacement limits or constraints may be assumed as especially justified (such as displacement limits of a pipe running through a hole in a suf ficiently strong concrete wall or floor, etc.). (d) Assume a through-crack size of 4t or 90 total circumferential length, whichever is greater, or a larger crack only if especially justified. (e) Assume large deformations means deformations proceeding to "as built" displacement limits or other especially justified limits. (3) Material Properties Cons e rva tive material properties should be used in the analyses. Sufficient justification must be provided for the properties, both weldment and base metal, used in the analyses. C. Suberitical Crack Development Consideration should be given to the types of subcritical cracks which may be developed at all locations associated with this type of analysis. From prior experience and/or direct analysis, it should be shown that: (1) There is a positive tendency to develop through-wall cracks. (2) If there is a tendency to develop long surface cracks in addition to through-wall cracks, then it should be further demonstrated that the long surface crack will remain suf ficiently shallow. D. Augmented Inservice Inspection Piping system locations for which corrective measures are not practicable should be inspected volumetrically in accordance with ASME Code, Section XI for a Class I system regardless of actual system classification.
DRAFT 11/29/83 STEP-WISE APPROACH, LEAK-BEFORE-BREAK (LBB) ANALYSIS
- 1. De' scribe the line(s) for which LBB is to be applied.
- a. Provide a discussion to support a conclusion that this line or lines is(are) very unlikely to experience stress corrosion cracking or excessive loads such as might occur from thermal
. - or mechanical-low and high cycle fatigue or a water-hammer.
- b. Identify the types of materials and materials specifications used for base metal, weldments and safe-ends and provide the materials properties including the J-R curve used in analyses, long term effects such as thermal aging and other limitations such as limits to valid data (e.g. , maximum J, maximum crack growth).
- c. Specify the type and magnitude of the loads applied (forces, bending and torsional moments), their source (s) and method of combination. Identify the location (s) at which the highest stresses coincident with poorest material properties occurs for base materials, weldments and safe-ends. For geometrically complex lines or systems, it may be necessary to analyze several locations to assure that the more vulnerable locations are identified. At this location or these locations, postulate a crack that may be missed during fabrication and preservice inspections or would be permitted by code, whichever is larger. Demonstrate by fatigue analysis that the track will not grow significantly during service.
s l
- 2. Postulate leakage size crack (s).
l
- a. Even though Step 1 should demonstrate that a leaking pipe is unlikely, postulate a through-wall crack at the selected location (s). The l size of the crack should be large enough so that the leakage is assured of detection with adequate margin using the minimum installed leak detection capability when the pipe (s) is(are) subjected to normal operational loads. If auxiliary leak detection systems are relied on, they should be described.
- b. Further, assuming that a safe shutdown earthquake (SSE) occurs prior to detection of the leak, demonstrate that the postulated leakage crack is stable under normal plus SSE loads for long periods of time; that is, crack growth if any is minimal during an earthquake.
- 3. Determine crack size margin by comparing leakage size crack to critical size crack. Using normal plus SSE loads, demonstrate that there is adequate margin between the leakage size crack and the critical size crack to account for the uncertainties inherent in the analyses and leak detection capability. In some cases, a limit-load analysis may suffice for this purpose, however, an elastic plastic fracture mechanics (tearing instability) analysis is preferable.
hh m M
' ? NL i
- 4. Determine margin in terms of applied loads by a crack stability analysis.
Demonstrate that the leakage size that crack (s) will not experience unstable crack growth even if larger loads (larger than design loads) are applied. Demonstrate that crack growth is stable and the final crack size is limited such that a double-ended pipe break will not occur. l . - 1 i l NOTE: Steps 1 through 4 are illustrated in the attached figure.
3 GENERAL DISCUSSION The preceding analytical steps assume that circumferentially oriented postulated cracks are limiting. If this is not the case, then the analyses described in Steps 1 through 4 should also include the postu-lation of axial cracks and/or elbow cracks. Also if applied moments are quite low and axial forces dominate, it may be necessary to consider relatively long part-through-wall cracks in Step 1 and demonstrate that they are unlikely to result in unstable axial or elbow splits or a double-ended pipe break. In general, the LBB approach does not rely on crack detection by inservice inspections (ISI). If,.however, conclusions reached via the LBB analyses are marginal, then augmented ISI at potentially vulnerable locations may be necessary. Positive conclusions reached via the LBB approach, will allow the removal of or non-installation of protective devices such as pipe whip restraints and jet impingement shields and thus obtain the benefits in both cost and man-rem saved as well as other safety benefits. If it can be demonstrated that large pipes will not fail catastrophically, then reconsideration can be given to design requirements for other safety systems such as containment and emergency core cooling systems. The latter approach, however, will involve consideration of other systems, component or operator failures affecting the design requirements of these systems and which must be addressed in any request for reconsideration. i m...
^ LEGAL /ADMlt4ISTRATIVE CONSIDERATI0t45 The utilization of LBB technology to demonstrate that protective devices are not required and possibly that other safety system design require-ments can be relaxed will require an exemption from the current t4RC regulations, particularly GDC-4 and/or the definition of a LOCA. To justify such an exemption, applicants or licensees should provide a sufficient basis for such exemptions until the regulations are modified, including the cost and man rem benefits to be accrued at a specific facility versus any potential additional risks that might occur. The i4RC, in the meantime, will initiate rule-making activities to remedy the' situation in the long run.
The elimination of large LOCA loads can also affect the future design requirements for support systems. This aspect is under consideration by the staff. For all facilities currently operating or under construc-tion, appifcants and licensees should retain the present design requirements. Similarly, the current requirement tc postulate specific intermediate break locations in various lines is affected by LBB and is being reconsidered by the staff. t,
'\ t
(
STEP-WISE APPROACH, LEAK-BEFORE-BREAK ANALYSIS 5-o Select highest stress, poorest material properties _ location in pipe under consideration. o Postulate crack that may be missed during fabrica-tion and preservice inspections or would be permitted ;; by Code, whichever is larger. -_ o Demonstrate by analysis that crack will not grow Postulated significantly during service either by fatigue, - j-Fabrication corrosion or impact forces (water-hammer). 7, law -
^
=
I , pt/ "!/ t 4 o Demonstrate that even if crack propagated through wall that: ~
~
Leakage through crack is significantly greater than minimum leak detection capability under - normal operating loads so that detection of ' Postulated crack is assured and _r Leakage - Crack - even if undetected prior to an earthquake, crack - is stable under normal plus earthquake loads _, (growth, if any, is minimal for long periods of time).
,y ' y,%?""^. 5 ; s -
o Demonstrate margin via crack sizes 1 S. Compare leakage crack size to critical crack size - under normal plus earthquake loads. " Demonstrate that there is adequate margin to - Critical account for uncertainties inherent in analyses Crack and leak detection. . Size t r E os 1. L'"si t (. 4 o Demonstrate margin via loads =$ Demonstrate that leakage size cracks will not experience unstable crack growth even if larger 13 1
._- loads are applied and that final crack size is '
limited (that is, a double-ended pipe break will Stable Crack J-Growth Under n t occur. Larger Loads
}
4 (.. N
\
- g APPENDIX D
..CRACE STABILIIT CALCULATIONS mL N
?
Y e. I 5 e ? k, s ? t a,.._
.+g.-- ,._.__. __
VYNPS RECIRCULATION SYSTEM PIPING 12" RISER IANGITUDINAL CRAG STABILITY cosc o o * * " " * * *
- CAS E = 1 " * " " * * " " * " " * "
LONGlTUDINAL CRACK LEAK RATE, LEVEL A Leak Rate = 0.1 gpm Soper 2c 8.823 2.300 LONGITUDINAL CRACE STABILITY, LEVEL D LOADS Sleak = 8823. psi Shoop = 10191. psi PIPE OD = 12.750 in THICENESS = 0.687 in Sflow = 45000. psi CRACE RY Ceff J Jpzif LENG1H,IN IN IN-LB/ IN"2 1 3.67 0.000 0.18370E+01 0. 433 55E+02 0. 40483 E+ 03 2 3.67 0.097 0.193 40E+01 0. 4 8188E+02 0.41570E+03 3 3.67 0.108 0.19448E+01 0.4 87 50E+02 0.41691E+03 4 3.67 0.109 0.19461E+01 0. 48816 E+02 0.4170SE+03 5 3.67 0.109 0.19462E+01 0.4 8824E+02 0.41707E+03
""* CONVERGENCE ACHIEVED *""
ococoo********** CASE = 2 ********************* LONGIIUDINAL CRACE LEAK RATE, LEVEL A Leak Rate = 1.0 gpm Soper 2c
- 8. 823 3.800 LONGI 1EDINAL CPJLCE STABILITY, LEVEL D LOADS Sleak = 8823, psi Shoop = 10191. psi PIPE OD = 12.750 in THICENESS = 0.6 87 in Sflow = 45000. psi CRACE RY Ceff J Jpzif LENGTE,IN IN IN-LB/ IN" 2 1 5.17 0.000 0.25870E+01 0.80136E+02 0. 4 83 55E+03 .
2 5.17 0.179 0.27663 E+01 0.94150E+02 0.51108E+03 ! 3 5.17 0.211 0.27 97 6 E+ 01 0.96756E+02 0.51589E+03 4 5.17 0.216 0.2 803 4 E+ 01 0.97245E+02 0. 5167 8E+ 03
$ 5.17 0.218 0.28045E+01 0.97337E+02 0.51695E+03
""* CONVERGENCE ACHIEVED *"' I
cocc************ CASE =3 ********************* LONGI'lTDINAL CRACE LEAK RATE, LEVEL A Leak Rate = 10.0 gpm Soper 2c 8.823 5.900 LONGITUDINAL CRACE STABILI'IT, LEVEL D LOALS Sleak = 8823. psi Shoop = 10191. psi PIPE'0D = 12.750 in THICENESS = 0.6 87 in Sflow = 45000, psi CRAct RY Ceff J Jpzif LENG'IH,IN IN IN-LB/IN**2 1- 7.27 0.000 0.36370E+01 0.1852 8E+ 03 0.64376E+03 2 7.27 0.414 0.40515E+01 0.243 84E+03 0.70650E+03 3 7.27 0.545 0. 4182 5E+01 0.26461E+03 0.72630E+03 4 7.27 0.592 0.42289E+01 0.27225E+03 0.73331E+03 5 7.27 0.609 0.42460E+01 0.27509E+03 0.73 589E+03 6 7.27 0.615 0.42524E+01 0.27616E+03 0.73 6 86 E+03 7 7.27 0.618 0.42547 E+01 0.27656E+03 0.73722E+03 8 7.27. 0.619 0.42556E+01 0.27671E+03 0.7373 5E+03
***** CONVERGENCE ACHIEVED *****
VYNPS RECIRCULATION SYSTEM PIPING 22" HEADER LONGIEDINAL CRAG STABILITY cmcoco e e e e e n e e CASE = 1 * '
- m "'
LONGITUDINAL CRAG LEAK RAIE, LEVEL A Leak Rate = 0.1 gym Soper 2c 10.997 2.400 LONGIEDINAL CRAG STABILITY, LEVEL D LOADS Sleak = 10997. psi Shoop = 12573. psi PIPE OD = 21.879 in *IHIGNESS = 0.976 in Sflow = 45000. psi CRAG RY Ceff J Jpzif LENGTH,IN IN IN-LB/ IN"2
'I 4.35 0.000 0.21760E+01 0.60 87 8E+02 0.55501E+03 2 4.35 0.136 0.23122E+01 0.67E31E+02 0.57092E+03 3 4.35 0.152 0.23 277E+01 0.6 8660E+02 0.57272E+03 4 4.35 0.154 0.23 296 E+01 0.6 8760E+02 0. 57293 E+ 03 5 4.35 0.154 U.23298E+01 0.6 8772E+02 0. 57 296 E+03
"* CONVERGENCE ACHIEVED "**
- c0cescoo * * """ CASE = 2 * """""" m " m LONGIEDINAL CRAG LEAK RATE, LEVEL A Leak Rate = 1.0 gpm Soper 2c 10.997 4.100 LONGIEDINAL CRAG STABILI77, LEVEL D LOADS Sicak = 10997, psi Shoop = 12573. psi PIPE OD = 21.879 in 'DIIGNESS = 0.976 in Sflow = 45000, psi CRAG RY Ceff J Jpzif LENG'HI,IN IN IN-LB/IN**2 2 6.05 0.000 0.30260E+01 0.11441E+03 0.65128E+03 2 6.05 0.256 0.32819E+01 0.11625E+03 0.63 840E+03 3 6.05 0.260 0.3 2 860E+ 01 0.11659E+03 0.63 90 5E+ 03 4 6.05 0.261 0.32868E+01 0.11665E+03 0. 63 916 E403
* * * *
- CONVr't0ENCE ACHIEVED " m
lccocosce * * * " * *
- CASE = 3 """"""? """ *
- LONGITUDINAL CRACE LEAK RATE, LEVEL A Lesk Rate = 10.0 gpm Soper 2c 10.997 6.800 LONGITUDINAL CRACE STABILITY, LEVEL D LOADS Sleak = 10997. psi Shoop = 12573. psi PIPE OD = 21.879 in IIIICENESS = 0.976 in Sflow = 45000. psi CRACE- RY Ceff J Jpzif LENGIH,IN IN IN-LB/ IN"2 1 8.75 0.000 0.43760E+01 0.2293 4E+03 0.80736E+03
-2 .8.75 0.513 0.48890E+01 0.3 007 6E+03 0. 8857 8E+03 3- ~ 8.7 5 0.673 0.50488E+01 0.3 2566E+03 0.91012E+03 4 8.75. 0.728 0.51045E+01 0.33 46 5E+03 0. 91860E+03 5 '8.75 0.749 0. 5,1246 E+01 0.337 93 E'03 0.92166E+03
(. 8.75 0.756 0.51319E+01 0.33 914E+03 0. 9227 8E+03 7 8.75 0.759 0.51346E+01 0.33 95 8E+03 0.92319E+03 8 8.75 0.760 0. 513 56 E+ 01 0.33 97 5E+03 0.92334E+03 + "H* CONVERGENCE ACHIEVED "*" _, y- -.4 -,- . . - . , - , - - - - - - . - -, . , , , m-- , - - , . + ,r - c- - - - - - , e- - - - -
VYNPS RECIRCULATION SYSTEM PIPING 28" DISCHARGE LONGITUDINAL CRACK STABILITY ceococco " * " *
- CASE = 1 """""""" m "
LONGITUDINAL CRACK LEAK RATE, LEVEL A Leak Rate = 0.1 gpm Soper 2c 11.295 2.600 LONGITUDINAL CRACK STABILITY, LEVEL D LOADS Sleak = 11295. psi Shoop = 12899. psi PIPE OD = 28.337 in 7EICENESS = 1.235 in Sflow = 45000. psi CRACK RY Ceff J Jpzif LENGIE,IN IN IN-LB/ IN" 2 1 5.07 0.000 0.253 50E+01 0.69792E+02 0.68024E+03 2 5.07 0.156 0.26 911E+01 0.77222E+02 0.6 9942E+03 3 5.07 0.173 0.27077E+01 0.7 80 46E+02 0.70142E+03 4 5.07 0.175 0.27096E+01 0.7 813 8E+02 0.70164E+03 5 5.07 0.175 0.27098E+01 0.7 814 8E+02 0.70167E+03
""* CONVERGENCE AGIEVED m" y,0cc oo c * * * * * *
- CAS E = 2 " * " " " " " * " " "
- LONGIIUDINAL CRACE LEAK RATE, LEVEL A Leak Rate = 1.0 gpm Soper 2c 11.295 4.600 LONGITVDINAL CRACE STABILITY, LEVII D LOADS Sleak = 11295. psi Shoop = 12899. psi PIPE LD = 28.337 in TIIIGNESS = 1.23 5 in Sflow = 45000. psi CRACK RY Ceff J Jpzif LENGT11,IN IN IN-LB/ IN' 2 1 7.07 0.000 0.3 53 50E+01 0.127 93 E+03 0.7 9606E +03 2 7.07 0.286 0.3 8212E+01 0.14977E+03 0.82807E+03 3 7.07 0.335 0.3 8700E+ 01 0.15375E+03 0. 83 3 55E+ 03 4 7.07 0.344 0.3 87 89E+01 0.15449E+03 0. 83 455E+03 5 7.07 0.346 0.3 8 80 6E+01 0.15463 E+ 03 0. 83 473 E+03 "m CONVERGENCE ACHIEVED * **
oc co o0c c * * * * * * *
- CAS E = 3 * * * * * * * * * * * * * * * * * * * *
- l LONGITUDINAL CRACK LEAK RATE, LEVEL A Leak Rate = 10.0 gpm Soper 2c 11.295 7.600 LONGITUDINAL CRACK STABILITY, LEVEL D LOADS Sleak = 11295. psi Shoop = 12899. psi
-PIPE OD = 28.337 in THICENESS = 1.235 in Sflow = 45000. psi CRACK RY Ceff J Jpzif LENGTH,IN IN IN-LB/IN**2 1 10.07 0.000 0.50350E+01 0.23 899E+03 0. 946 53 E+03 2 10.07 0.53 5 0.556 96 E+01 0.3 0446 E+03 0.10287 E+04 10.07 0 .6 81 0. 57161E+01 0.32425E+03 0.10 511E+04 3
4 10.07 0.725 0. 57 603 E+01 0.33 03 9E+03 0.10579E+04 5 10.07 0.73 9 0.57741E+01 0.33 231E+03 0.10600E+04
.6 10.07 0 .7 43 0. 577 84E+01 0.33291E+03 0.10607 E+ 04 7 10.07 0.745 0. 377 97 E+01 0.33310E+03 0.10609E+04
***** CONVERGENCE ACHIEVED *****
I i t
, . . . ~ . .
VYNPS RECIRCULATION SYSTEM PIPING 28" SUCTION IANGI'IUDINAL CRACE STABILITY
- c:Co * * * * " * * * * *
- CAS E = 1 " " " " * " " " " " "
LONGIIUDINAL CRACE LEAK RATE, LEVF' A Leak Rate = 0.1 Fpm Soper 2c 11.188 2.400 LONGITUDINAL CRACE STABILITY, LEVEL D LOADS Sleak = 11188. psi Shoop = 12888. psi PIPE OD = 28.169 in IIIICENESS = 1.151 in Sflow = 45000, psi CRACK RY Ceff J Jpzif LENGTH,1N IN IN-LB/ IN" 2 1 4.70 0.000 0.23 510E+01 0.63019E+02 0.6 43 94E+03 2 4.70 0.141 _0.24920E+01 0.69405E+02 0.6617 2E+03 3 4.70 0.155 0.25063E+01 0.7007 7 E+02 0.66348E+03 4 4.70 0.1 57 0.2507 8E+01 0.70148E+02 0.66367E+03 5 4.70 0.157 0.25079E+01 0.70156E+02 0.66368E+03
""* CONVERGENCE ACHIEVED *""
occc e e e e e n o u n CASE = 2 """
- LONGITUDINAL CRACK LEAE RATE, LEVEL A Leak Rate = 1.0 gpm Soper 2c 11.188 4.300 LONGIIVDINAL CRACE STABILITY, LEVEL D LOADS Sleak = 11188. psi Shoop = 12888. psi PIPE OD = 28.169 in IIIICENESS = 1.1 51 in Sflow = 45000. psi CRACE RY Ceff J Jpzif LENGIll,IN IN IN-LB/IN**2 1 6.60 0.000 0.33010E401 0.11565E+03 0.7 5542E+03 2 6.60 0.259 0.3 5597 E+01 0.13440E+03 0.7 843 5E+03 3 6.60 0.301 0.36016E401 0.13764E+03 0.7 8 904E+ 03 4 6.60 0.308 0.36089E+01 0.13 821E+03 0.7 8985E+03 5 6.60 0.309 0.36102E+01 0.13 831E+03 0.7 8999E403
*"' CONVERGENCE ACHIEVED *"'
c oc o c co * * * * * * * *
- CAS E = 3 * * * * * * * * * * * * * * * * * * * *
- LONGITUDINAL CRACK LEAK RATE, LEVEL A Leak Rate = 10.0 gpm Soper 2c 11.188 7.200 LONGIIIIDINAL CRACE STABILITY, LEVEL D LOADS Sleak = 11188. psi Shoop = 12888. psi PIPE OD = 28.169 in -IIIICENESS = 1.151 in Sflow = 45000. psi CRACK RY CeIf J Jpz1f LENGIII,IN IN IN-LB/IN**2 1 9.50 0.000 0.47510E+01 0.21864E403 0.89660E+03
'2 9.50 0.489 0.52401E+01 C.27633E+03 0.97185E+03 3 9.50 0.618 0. 53 691E+01 0.293 05E+03 0.99164E+03 4 9.50- 0.656 0.54065E+01 0.29802E+03 0.9973 8E+03 5 9.50 0.667 0.54177E+01 0.29951E+03 0.99908E+03 6 9.50 0.670 0.54210E+01 0.29996E+03 0.99959E+03 7 9.50 0.671 0.54220E+01 0.3 0009E+03 0.99974E+03
* * * *
- C011 VERGENCE ACHI EVE's ""
- 1
. ., , ., - . . . _ . . _ - _ _ . _ _ . . _ . _ . . . , . . . _ _ ._ _ _ . ~ _ . _ _ _ . _ _ . . _ . . . _ . _ _ . . .
h' e,
~
Rnotti YAlal;EE RCCIRC ' LIf 4E C P A Cl? STHBILITY
's F ISER , 12.:n)- STRESSES FROrl GE/SAR'22A2615
~
M4:: 4147. Fakial. : O. 11 applied : .O.16380E+07 Poper : 1479, ps
; s ;, ,. i n 1. = 0. Sbending : 21979. Smem : 5774. psi ' PIPE _ODL:. 12.750' THICKf4ESS : O.687 SfIow : 70000. psi ALFA _ = 2. ELAS MOD =0.256E+O8 Jic : 4500. in-lb/in*s2 '. C R A C K : - LE Al' .- ARE A L/Dh J T
$El4GT H,114 I!4 *
- 2 Il4-L B/ Il4*:t 2
'f. 2 7 - 0.002 0.le434E+03 0.57827E+02 0.56075E+00 l3,59 O.005= 0.11929E+03 O.91456E+02 0.62746E+00 2.51 0.009 0.86565E+02 0.12921E+03 0.70634E+00 L
3.12 0.015. O.66878E+02 0.17178E+03 0.79722E+00 3.72 0.023 0.53767E+02 0.21985E+03 0.90027E+00 4.31 0.032 0.44452E+02 0.27412E+03 0.10159E+01 4J88- -0.042 0.37533E+02 0.33530E+03 0.11445E+01 g;as 0.055 0.32227E+02 0.40416E+03 0.12666E+01
;g,93- 0.070 0.28059E+02 0.48147E+03 g,gg 0.14429E+C1 O.056 0.24725E+02 0.56SO7E+03 0.16139E+01 7.03 'O.104 O.22021E+02 O.66479E+03 O.10004E+01 7.52 -0.124 0.19802E+02 0.77254E+03 0.20029E401 7,99 O.145 O.17966E+02 O_.89225E+03 O.22222E+01 P
i,aa- O.167 O.1643SE+02. O.10249E+04 O.24591E+01 E.E5 O.190 - 0.15160E+02 0.11715E+04 0.27141E+01 9,29 0.214 0.14089E+02 0.13330E+04 0.296COE+01 9,g1 O.237 O.13191C+02 O.15107E+04 O . 3 2 61 E.E + 01
.9,9a O.260 0.12440E+02 0.17056E+04 0.35955E+01 10.2a O.282 .O.11816E+02 0.19188E+04 0.39306E+01 10'50 0.303 0.11304E+02 0.21517E+04 0.42876E+01
.. g o , 7 3 0.321 0.20891E+02 0.24055E+04 0.46.E71E+01
. 3 p,9 3 - O.337 O.10569E+02 O.26814E+04 O.50701E+01 11'.'03 0.349 O.10?31E+02 0.29809E+04 0.54571E+01 11.'14 .O.357 O.10175E+02 O.33053E+04 O.59491C+01
'11.19 0.361 0.10099E+02 0.36560E+04 0.64166E401 3g,39 O.361 O.10104E4OI O.4034EE404 0,69306E401
=en* E;:CEEDED RY INN SI2E : h 2.YC o4 **
1 l
.rCF;10NT -Yolli:EE REC IRC . L IIIC CR ACk SinD1 LILY tHEAbly , 22 in) STRESSES FROl1 GE/SAR 22A2615
$4 L: [4144 Mapp!ied : 0.22480E+07 Fayi&l ':' 16110. Poper : 1479. ps sexgat : 251.- Sbending : 7010. Smem : 7448. psi PIPE OD = 21.879 THICKNESS = 0.976 SfIow : 70000. psi ALFA : 2. ELAS MOD =0.256E+O8 Jic : 4500. in-lb/in g2
' CRACK LEAK AREA L/ Die J T
$ENGTH,IN IN**2 IN-LE/INu2
'1.91 : 0 ~. 0 0 3 0.34731E+03 0.22230E+02 0.15171E-OO 2.86 0.006 0.22469E+03 0.35273E+02 0.16932E+00 3.01 O'. 0 1 1 -O.16304E+03 O.49829E+02 O.18891E+00 4.70 0.018 .O.12589E+03 0.66059E+02 0.21045E+00 5.-70 0.026 0.10107E+03 0.84121E+02 0.23397E-OO
'6.64 _O.037 0.83350E+02 0.10418E+03 0.25948E+00 7.58 0.050 .O.70095E+02 0.12639E+03 0.28703E-OO S.51. 0.066 0.59842E+02 0.15093E+03 0.31668E-OO
' 9, 4 4 - -0.085 0.51707E+02 0.17797E+03 0.34847E+00 710.36' O.107 0.45121E+02 0.20768E+03 0.38247E+00 111.27 O.133 0.39704E+02 O.24025E+03 O.41873E+00
' '12,18 0.162 0.35191E+02. O 27587E+03 0.45732E+00
, . 13, 09 - 0.195 0.31390E+02 0.31472E+03 0.49831E-CO U .13,99' .O.232 0.28159E+02 0.35701E+03 0.54175E+00
- 14.88 0.274 0.25392E+02 0.40293E+03 0.58771E+00
! 15.77 0.32O_ O.23OOSE+02 0.45269E+03 0.63626E-OO
?i6.64L 'O.371 O.20935E+02 O.SO651E+03 O.6874SE-OO L 17. 51 - .O.42G O.19129E+02 0.56461E+03 0.7413EE+00
-10.33 0.489 O.17546E+02 O.62720E+03 0.79809E+00 19.23 O.555 0.'16154E+02 O.69452E+03 O.85765E+00 420.07 O . 6 F.8 0.14923E+02 0.76681E+03 0.92012E+00 30;91- 0.706 O.13833E+02 0.84430C+03 0.98558E+00 21,-73 0.739 0.12862C+02 0.92724E+03 0.10541E+01 4
2.E5 0.878 0.11997E+02 0.10159E+04 0.11257E+01
" 2'3 . 3 5 - 0.972 0.11222E+02 0.11105E+04 0.12OOSE+01 ja,la 1.071 0.10527E+02 0.12113E+04 0.12786E-01 24,92 1.176 0.9902SE+01 O.13186E+04 O.13600E-01 gg,c9 1.285 O.93399E+01 O.14326E+04 O.14447E-01
,26;43- 1.399 O.88315E+01 O.15537E+04 O.15329E-01 27._19 1.518 O.83716E+01 O.16821E+04 O.16247E+01 17.02. 1.640_ O.79547C+0i O.19181E+04 0.17199E+01 20.63 1.7CC O.75762C+01 O.196.'OE+04 O.18189E-01
.g9,33 11.895 0.72320E+01 0.21141C+04 0.19215E+01
' 9. 01_ 2.027 O.C9166E+01 0.- 2 27 4 7 E + 04 O.20260E,01
< 0.se - 2 .~_1 6 2 O.66330E+01 O.24442E+04 0.21382E-01
--21.33 2.295 O.C3723E401 0.26227E+04 0. 2 2 5 2 4 E- 01 21.96 2.435 0.61342C+01 0.28107C+04 0.23705E-01
,; 2, 5 3 2.574 0.59165E+01 0.3OO85E+04 0.24927E-01 33.17 2 . 7 '. 2 0.57174E+01 0.32164E+04 0.26159E-C1
;3,75 2.351 0.55352E+01 0.34347E+04 0.27494E-01
- 34.31 2.93G .O.53655E 01 0.36536C+04 0.2&G4CE-01 3a,er ~.124 3 0.0215?E-01 O.39040E+04 O.3023CE-01 as,jg 's.257 0.50762E*01 0.41557C+04 0.31662E-C1
((I3 . - re .c, 2.388 0.49485L+01 0.44192E404 0.33140L+01 I: 36.'7 2.I16 0.JG316C401
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. 4.44 EXCEEDED L I MIT 11011EllT
- 4 4 e 4-
w VEF!!^ ~7 4*;n'CE FECIRC LillE CR ACR STHBILITY tSuctic . 28 in) ~ STRESSES FPOt1 GE/SAR 22A2615 t UH;: 4146 F&xial : O. Mapplied : 0.2743OE+07 Poper : 1375. psi garial. .: .O. Sbending : 4326. Smem : 7396. psi PIPE OD :' .28.169 THICKNESS : 1.151 Sflow : 70000, psi ALFA -: 2. ELAS ItOD =0.256E+08 Jic : 4500. in-lbrinat2 CF.ACK LEAM AREA L/Dh J T IN4"2 IN-LE/Ilie 4 2 SE14GTH IN 2,27J O.CO3 0.44364E+03 0.16370E+0; O.94784E-01 3.~41 0.007 0.28728E+03 0.26021E+02 0.10580E+00 4,54 0.012 0.20873E+03 0.36700E+02 0.11779E+00 5.67- 0.019 0.16143E+03 0.48740E+02 0.13074E400 6.79 0.029 0.12983E+03 0.61993E+02 0.14466C+00 7.fl2 ' O.041 0.1072SE+03 0.76634E+02 0.15955E+00
'9,04 0. C ES. 0.90352E+02 0.92757E+02 0.17544E+00 c10.1G' O.073 0.77259E+02 0.11046E+03 0.19233E+00 11127_- C.~?? O.66850E+02 0.12934E+03 0.21025E+00 ,12.39 <
0.117 0.58406E+02 0.15099E+03 0.22922E+00 13.49 O.145 0.51443E+02' O.17402E+03 0.24927E+00 1'4,60 c.1 7 0.45625E+02 0.19904E+03 0.27040E+00 15.70 0.213 0.40710E+02 0.22614E+03 0.29265E+00 16,80 0.2?4 0.36520E+02 0.25544E+02 0.31603E+00 17.89 0.?SO . O.32918E+02 0.28705E+03 0.34058E+00 18.08 0.25; O.29801E+02 0.3210EE+03 0.36630E+00
._ 2 0. 0 7 - .O.409 0.27088E+02 0.35764E+03 0.39324E+00 21.14 0.473 0.24712E+02 0.39686E+03 0.42139E400 22.22 0.543 0.22623F+02 0.43385E+03 0.45080E400 23.29' O.619 0.20778E+02 0.48373E+03 0.48143E+00 24.35 0.703 0.19142E+02 0.53163E+03 0.51345E+00 25.41 0 ~94 0.17686E+02 0.58267E+03 0.54675E+00
- 26. 4 E 0 '.43 0.16385E+02 0.63690E+02 0.5813GE400 27.51 . ? ?9
. 0.1522OC+02 0.69470E+03 0.61727E400
.28.75 1.113 0.14174E+02 0.75594E+03 0.65475E+00 29 re- 1.I}6 0.13231E+02 0.32085E+03 0.69354E+00 30.o.1 1.:47 0.123SOE+02 0.88957E+03 0.73375E*00 21.62 1*X O.11610E+02 0.96223E+03 0.77541E+00 32.63 1.s54 0.10911E+02 0.10390E+04 0.G1855E+00
,33.'64 1.!!O O.10276C+02 0.11199E*04 0.86318E+00 34.66 1.F7E O.96977E401 0.12053E+04 0.90933E+00 35.62 2.142 0.91C96C+01 0.12951E+04 0.95701E+00 G.6t 2.130 0.86867C+01 0.13G97C+04 0.10063E401 27.57 ~ 2.12v O.82444E+01 0.149?OE-04 0.10571E401 29.53' 2. 19' O.73387E+01 0.15932E+04 0.11095E+01
-30.48 2.?I5 0.74G58E+01 0.17028E+04 0.11E36E401 40.42 3.179 0.71226E+01 0.1E175E+04 0.12193E+01
-41.?5- 3.1i1 0.CSO62E401 -
O.19377E+04 0.1276CE401 42.20 1 !90 0.65142E+01 0.20634E+0a 0.13357E401 42.19- 2.EIC O.62443E+01 0.21950E*04 0.13964E+01
'44.09 4 :i9 0.5994GC+01 0.23225E+04 0.14599E401 44.9E 4.:15 0.57C21E+01 0.24761E+04 e. 15232C+01 45.25 J.f"2 0. 55 4 F.4 E + 01 0.2625?E+04 ". 15992E+01
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' *** ** . E'/CEEDED L IM I T MOMEf 4T * * * ** ,s 4
e
ATTACHMENT A VERHONT YANKEE I&E BULLETIN 83-02 EXAMINATION PROGRAM a.
L f.f f _ lW A. ' f i ; Q [ j (g j y l l/ i f 4 \,.[~-[ .). ; :.i.[ l ? f flh? ; %F 0 kb% ATTACHMENT A VERMONT YANKEE I&E BULLETIN 83-02 EXAMINATION PROGRAM UTILIZED DURING THE 1983 AUCMENTED ISI PROGRAM Ultrasonic Examination Technique I&E Bulletin 83-02 requires that we demonstrate the effectiveness of the detection capability of the ultrasonic examination technique to be used for examining weld joints in our recirculation system piping. The bulletin also establishes provisions for demonstration tests to be performed at the EPRI-NDE Center in accordance with specific criteria. This includes equipment / procedure similarity, personnel participation, pipe sample size, acceptance criteria, demonstration time limit, and procedures review. On March 11, 1983, Vermont Yankee and its contractor, Magnaflux, successfully
. passed the demonstration. A copy of the form used to document this demonstration is provided as Figure A-1.
The examination methodology made use of dual element, 1.5 MHz, 45 0 and 600 shear wave search units coupled with pulse-echo ultrasonic instrumentation. The equipment was set up in a master-slave configuration, allowing maximum use of qualified examiners with minimum radiation exposure. Detection of IGSCC was based on signal characteristics and location with respect to the weld root geometry. Sizing was performed on indications in 12", 22", and 28" pipe. Although sizing was performed on the 12" pipe, all 12" welds with flaw indications, regardless of size, were overlayed. The primary method utilized for sizing ultrasonic indications of IGSCC at Vermont Yankee was the " Amplitude Drop Method" using dual element 1.5 MHz transducers having a nominal shear wave beam angle of 450 The through-wall dimension of the indication is compared to that of a 10% notch in a basic calibration block. The sweep changes corresponding to the maximum amplitude from the 10% notch and the leading and trailing ray half maximum amplitudes (6 dB drop) are recorded during the evaluation calibration. During evaluation scanning, the sweep changes are recorded for the noted indications. The recorded sweep readings are then plotted on full size sketches of the wold joint section as determined by actual field measurement. A linear relationship is maintained in comparison to the 10% notch. Through-wall dimensions are calculated to the next higher full percent and reported for engineering evaluation. No beam spread correction was applied to the depth sizing. Linear extent was plotted similarly. Linear extent was considered at an end point when the amplitude of the signal dropped to 50% of the average maximum signal for a given indication when scanned in a manner intended to determine linear extent. Beam spread correction was not used. In order to determine thi reliability of the " Amplitude Drop Technique" for the sizing of ICSCC flaws, two investigations were performed. A-1
Initially, Vermont Yankee assessed sizing capability by evaluating indications on a cracked specimen of large diameter Nine Mile Point-1 (NMP-1) pipe. Three teams measured the through-wall dimension of specified flaws. These measurements were compared to the through-wall dimension of a crack which was exposed on the edge of the block. The examiners sized the flaw between 10 and 15%. Physical measurement after liquid penetrant exam indicated a crack depth at that location of 15% through-wall. Additional confirmation of sizing accuracy was felt to be necessary; and, as a result, two areas of the same NMP-1 specimen were selected and sized by the examiner responsible for a large portion of ultrasonic examinations at Vermont Yankee. Following ultrasonic flaw sizing, two areas of the circumferential weld joint ID-SW-19-4 (MP-01 specimen), were sectioned, liquid penetrant examined and dimensioned for through-wall dimension. Selection of the areas to be sectioned was based upon indication location in an effort to minimize impact on the sample. These were not considered as maximum flaws and are instead average flaws. The results of this effort are tabulated as follows: Destructive Testing Ultrasonic Measurement Indication No, Measured % TWD Measured % Error 1 .170 12% .227 +25% 2 .150 12% .170 +121 A-2
MTT E 4 NRC IE BULLETIN 83-02 Demonstration of UT Performance Capability EPRI KDE Center Charlotte, NC Demonstration Results Date: Procedure No.: Utility: ISI Contractor: NRC Region: Demonstration Team Nembers and Levels: f 1. 2. 3. 4. 5. 6. Results: Acceptable ( ) Unacceptable ( ) Pending ( ) Basis for Failure: Crack Detection ( ) False Calls ( ) Comments: s NRC Representative Utility Representative (Signature) (Signature) cc: NDE Center NRC IE FIGURE A-1 A-3 L
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ATTActeIENT 5 7- f PROJECTION IIIAGE SCANNING TECHNIQUE INFORIIATION s e
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l ATTACHMENT B 1 l P-Scan Principle In the P-Scan technique (Projection image Scanning technique), echoes from weld defects are recorded together with their corresponding positions. Defect positions are then visualized on two projection planes: One plane parallel to the surface and another norttal to the surface, parallel to the weld. In other words, defects appear as seen from a Top View and Side View (see illustration below) . By using two projection planes a complete three-dimensional location of weld defects is obtained. g Top view ununcun 1
- !8!:i i I
Sic e view i The P-Scan display for weld inspection can be divided into 3 sections: (1) Top View - 1 Scan mm: 0000 (0000) 0125 M F;it jc& k&M?wwb fk t:J24,.dp MT*#A* M9J; Weld TOP CM 64:ih i Viewing f Dir:ction n -.
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i l COMPLETE P-SCAN IMAGE i WITII WELD INSPECTION DATA Ci +Q CD 00 -+ CD --3 iO E CD CD >^ H- C#J O CD E cg .. wO . cp o .. CD (D CD e. k s 1
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l l A recent addition to the P-Scan System used at Vermont Yankee, includes the < ability to evaluate the end view as well as the top, side, and amplitude ! displays. This view imposes the information displayed by the top and side l views at any given cross section of the base material and wold nugget, thus l aiding the examiner with additional information as to the development and ; nature of an indications. j The P-Scan System is sensitive to the input parameters entered at the start of examination. Extensive pre-exam reviews of construction conditions are necessary to assure correct parameters are established. Additional assurance are achieved by measuring 0.D. profiles and thickness gauging of the examination area. The qualified examiners are also capable of recognizing the effects of incorrect parameters and adjustments or re-examination may be necessitated. P-Scan WS-2 Scanner The WS-2 scanner provide a compact, reliable means of obtaining all necessary j positional information and served as the only scanner used with P-Scan. This I scanner was ineffective only in extremely tight configurations or difficult geometries. 1 l Iltti .it ure a,t nii . i l w.t n - ile f I<>f ( I n n. i n 5[le'i: ( ,6 iii l OI j)[,4: W e l 615 , Wlit/ I' +
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ATTACHMENT C IMPROVEMENT IN FLAW SI7.IWC CAPABILITY Page C-2 represents the results of the original flaw sizing round robin held at the NDEC on August 4, 1983. This chart indicates the need for corrective action. Page C-3 represents the improvement of flaw sizing ability after the first four workshops on flaw sizing at the NDEC during the period April / June 1984. Page C-4 represents a sub-set of flaw sizing examiners which met the then proposed acceptance criteria for flaw sizing. C-1
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O.' f INSPECTOR PERFDRMANCE OURtNG BS INDUSTRY ROU ND ROSIN TE.ST Com -
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ATTACHMENT D VERMONT YANKEE REACTOR COOLANT LEAKAGE LIMITS
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ATTACHMENT D . VERMONT YANKEE REACTOR COOLANT LEAKAGE LIMITS COOLANT LEAKAGE
- 1. -During power operation, Reactor Coolant System leakage into the primary containment shall be limited to:
- a. 5 GPM unidentified leakage when averaged over the previous 24-hour period; and Jb . 20 GPM' identified leakage when averaged over the previous 24-hour period. '
- 2. Any time the reactor is in the run mode, Reactor Coolant S stem leakage into the primary containment from unidentified sources shall be limited to:
~
- a. 2 GPM increase in unidentified leakage within the previous 24-hour period (see Note 1).
- 3. If the requirements of Item 1 cannot be met, initiate action as follows:
- a. ~With any Reactor Coolant System leakage greater than any one of the limits specified in Item 1.a or 1.b reduce the leakage rate to within the limits or be in at least hot shutdown in 12 hours and in cold shutdown in the next 24 hours.
- 4. .If the requirements of Item 2 cannot be met, initiate action as follows:
- a. With any increase in unidentified leakage of greater than or equal to 2 GPM, averaged over the previous 24-hour period, identify the source of leakage or be in at least hot shutdown in 12 hours and in cold shutdown in the next 24 hours.
.5. Both the drywell sump and air sampling systems shall be operable during power operation. From and after the date that one of theses systems is
- made or found inoperable.for any reason, reactor operation is permissible only during the succeeding 7 days.
.6. If the requirements of Item 5 cannot be met, an orderly shutdown shall be initiated and the reactor brought to a cold shutdown condition within 24 hours.
7 BOTE 1: .During the first 24 hours in the run mode following startup, the
' limits of Item 2 may be waived provided the requirements of Item 1 ,
, are met.
1 COOLANT LEAKAGE (Surveillance) Reactor Coolant System leakage shall be demonstrated to be within the limits of Items 1 and 2 by checking and logging the leakage collected in the primary z containment floor and equipment sumps at least once per 4 hours. In addition,
" the primary containment atmosphere activity shall be checked and logged at least once per 8 hours. -<
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g_.__ TABLE I VERMONT YANKEE WELD JOINT INSPECTION MATRIK 1983 - 1984
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L -- TABLE II Details Of UT Indications And Weld Joint Stresses Pipe - a/T a/T ' L/ circ
- gize Whld ISI No. (P+DW+0BE+Th) Or' ant (%) (%) (in) L/2TI R S"(4) A= Axial C=p' ire 28" 1A 0.56 #0"
- 22 22 5.0 .057 2 0.47 C 15 15 2.0 .023 ISB 0.64 C 18 18 3.0 .034 26A 0.50 C 15 20 19.0 .216 27 0.45 C 19 20 4.5 .051 61- 0.44 C 20 20 24.0 .273 59 0.43 C 20 20 18.0 .148 65A 0.49 C 23 25 15.0 .160 9A 0.43 C 20 22 5.0 .057 17B 0.43 C 20 20 7.0 .060 6 0.48 C 17 19 3.0 .034 22" 16A 0.67 C 20 20 12.0 .101 16B 0.71 C 12 12 0.8 .012 30B 0.34~ C 20 25 20.0 .300 49 0.58 C 22 23 1.5 .022 23B 0.34 C 27 27 6.0 .087 s
20" 'RHR-32-4 A > 10 NA NA
.(1) Total length of all-circumferential indications at the weld. .-(2) Weighted Average Dept of all flaws.
(3) Maximum Dept of any one flaw. (4).A110weble Stress at 5500F. E
f
.J_ . -
t f ,'. 'p-n > r s
) A
+
r t w N 4 k TABLE III
, SUIEEARY OF PREDICTED GROWTH DURING THE NEIT CYCLE OF OPERATION f
Y
+ ,.'. J 1-
, l' l' [ ~ l E I 4 4 J l'. <-1..- k, -
# 1
, g
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i + 2
- l. - _
- . f t
i l I-l. I. , [ -- l.' . l -> .c ,e >.,/ I l4-- - l.s r-I
"&F T==Mf# Newweeh-m'MW raa-erMwWw' eP er __arwTr1 ** N-4+9 r wW_ - - W9"f MNpg---- '#M N 1El-GF N W -
TABLE III Summary of Predicted Crack Crowth For A 14-Month Operatinz Period
- Circumferential Flaw Size Allowable Flaw Size Pipe Weld- Start Start Start Of Cycle End Of Cycle Size ISI No. Depth Length Depth a/t Depth a/t s/t(%) (in) 28" 1A 22 5.0 0.42 0.5 2 15 2.0 0.40 0.5 ISB 18 3.0 0.43 0.5 26A 15 19.0 0.39 0.5 27 19 4.5 0.42 0.5 61 20 24.0 0.47 0.5 59 20 18.0 0.47 0.5 65A 23 15.0 0.45 0.5 9A 20 5.0 0.44 0.5 17B 20 7.0 0.44 0.5
.6 17 3.0 0.44 0.5 e
22" 16A 20 12.0 0.43 0.5 16B 12 0.8 0.35 0.5
'30B 20: 20.0 0.47 0.5 49 22 1.5 0.47 0.5 23B 27 6.0 0.47 0.5 -
b k e
.s
.1 TABLE IV DISPOSITION OF UT INDICATIONS r
a o I
TABLE IV Disposition'of UT Indications Disposition Pipe Wald Accept For 14-Mo. Weld Size ISI No. By Analysis (I) Overlay Repair 28 1A X 2 X ISB X 26A X 27 I 61 X 59 I 65A X 9A X 178 I 6 I 22 16A X 16B X 30B X 49 X 23B H 20 RHR-32-4 X* b v
~
- Mini overlay on axial indication
a w Y 4' o TABLE V COIIPARISON OF 1983 TO 1984 REINSPECTION RESULTS (LARGE BORE PIPING) J f L k t s {r s
=
}
A v K
TABLE V Large Diameter Piping Comparison of 1983 to 1984 Inspection Results 1983 1984 A/T A/T Pipe Size Weld ISI Wo. L( ' (% TWD) L( ' (% TWD)
.28" 64 4" 10-15 No Flaw N/A 1A 38" 15 5" 22 2 -3600 10 2" 15 (inter-mittent) 9A 3600 10 5" 20 (inter-mittent) 65A 9.5" 15 15" 23 ISA 11" 15 No Flaw N/A 58 17.5" 15 No Flaw N/A 59 3" 15 13" 20 22" 16B 4.5" 10 0.8" 12 36B 12.0" 10 No Flaw N/A 30B 4.5" 15 24.0" 20 24" RNR-31 4.0" 7 No Flaw N/A Weld 1 - (1) . L - Total length of all circumferential indications.
(2) A/T - Flow depth as a percentage of wall thickness (based on weighted average depths of all flaws). m.
4-. . c I t i s TABLE VI } ). VERMONT YANKEE STRESS INFORMATTE s + 1 b-E k s i h / 5: 1 7 s. 2 s r ..
TABLE VI Ve mont Yankee Stress Information Weld Actual 1 Overlay Joint. Wall' Shrinkage Number Thickness Pressure Deadweight OBE Thermal Stress (OS) P+DW+0BE+TH+0S 8 (inches) (psi) . (psi) (psi) (psi) (psi) ;m 1A. 1.2 5954 1177 155 2122 0 .557 2 1.2 5954 635 371 917 0 .466 ISB 1.18 6053 464 2164 1887 200 .637
'26A' 1.15~ 6210 637 636 958 0 .499 27 '1.15 6210 475 182 735 0 .450 61' .1.25 5711 83 1158 325 150 .439
- 59 ~ 1.34- 5330 54 1221 389 200 .426 65A 1.29 5534 461 1149 537 700 .484 9A 1.29 5534 259 476 393: 600. .430 17B 1.27, 6023 196 227 537 250 .428 6 '1.26 6068 .173 1320 435 200 .485 16A 1.05 5614 1417 758 2303 1190 .668 16B 1.03 5718 1422 758 2909 1190 .710
'30B 1.04 5666 -N/A N/A N/A 0 .340 23B 1.09 5408 N/A N/A N/A 0 .340 49 1.09 5408 546' 230 1136 2400 .575
W, TABLE VII COMPARISON OF 1983 TO 1984 UT PROGRAM
TABLE VII Comparison of 1983 to 1984 UT Program 1983 Details 1984 Details Equipment P710B P-Scan ALN 4060 USIP 11 USL 30 (Series) P710
' Probes 450S Dual 1.5 MHz 450S 1.5 MHz 6003 Dual 1.5 MHz 520S 1.5 MHz 450S 2.25 MHz 520S 2.25 MHz RTD 700 RL 4 MHz RTD 700 RL 2 MHz WSY 70-2 WSY 70-4 520 5 MHz SLIC 40 Calibration 10% NOTCH - 6 Db 10% NOTCH - 64 Db Scan -10 Db Unlimited Positional Recording Manual Auto - MSW-2 Manual Plotting Manual P-Scan SHARP 600 Manual Personnel 1 Level III 3 Level III 4 Level II 8 Level II
' Approx. 30 Level I Approx. 25 Level I Sizina Qualification None- EPRI Program Training In-House EPRI Qualifyinz Exam 83-02 Team EPRI Individual
~Sising -Db Deep HALT PATT SPOT MOST Pull Vee
gi e pays' T - l ( l r c { j d l l 1 i- : l l t TABLE VIII
- f. . 1983 FLAW
SUMMARY
t t i r h' i L l b I ir ) F l' . 3 t'; 1 a i i s 3 i s 1 1 _.
s fN ge h '1 ~ p s. 3 .. TABLE VIII
- 1. i
, 1, 1983 Flaw Summary
~ *leid Exam Cardinal Flaw Flaw
'gimber Size' Configuration Restrictions Point? Length Depth N
X 16Ba 22" P/CRS Pipe Side only No 4.5" 10% 4 ;
- 30B" 22" P/EC ,
No 3.0" 15% s, . . 1.0" 7-1/2%
.7" 410%
x 1.0" (10%
<~ 1.0" 15%
'c 1-
~
' - 1.0" 15%
t- J' 3.6" 15%
.. I %. '
; 1.0" 10%
.I ' l' s 1.5" f15%
,. ,,' ,. s. ,~, ,
, 3.0" ---
)-
.- t.
2 .
?.
.28" P/EL No No Int. 360 <10%
y s
. s ;' _.
28" P/EL 9Ai No No Int. 360 10% 58 28" /MP Elbow Side Only Yes 1.5" 7% 5.0" 7%
.25" 7%
17.5" 15% 7.0" 15% 64- 28" P/EL " No No - 1.5" 4 15%
, ' , 1.5" 4 15%
A c. < 2.75" 415% i s 2.0" 415%
'/*^' .4.0" 415%
i 4.0" 415% 2.0" (15%
'.. \
368. ' 2 7? P/CRS Pipe Side Only No 14.3" 10% i \l: , 65A 28" ' " ' Pipe Side Only No 9.5" 15% 7)P/T 59 28" PNP /P Pipe Side Only Yes 3" 13% s 33" 9% 1A. 28" )P/EL No - No 10" 15% 38" s 15% 15A 28" P/T. Pipe Side Only Yes 5" 15%
' - 7" 15%
11" 15%
'vf.
I t - l
.A .w., .m i
i k TABLE IX 1984 FLAW
SUMMARY
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V TABLE X 1984 EXAMINATION RESTRICTION
SUMMARY
777:: TABLE I 1984 Examination Restriction Summary Wold System Wumber Size Configuration Method Restriction Recire~ 1A 28 H. Pipe / Elbow P-scan None 2 28 Elbow /V. Pipe P-scan None 4 28 Valve /H. Pipe P-scan No Scan Valve Side 5 28 Elbow / Pump P-scan No Scan Pump Side SA 28 H. Pipe / Elbow P-scan No Scan Pipe Side 625-750 mm & 1550-1590 mm No Scan Elbow Side 0-375 mm & 2000-0 mm Manual Areas not Scanned by P-scan 6 28 Pump /H. Pipe P-scan No Scan Pump Side 8 28 H. Pipe / Valve P-scan No Scen Valve Side No Scan Pipe Side 500-625 mm & 1625-1750 mm Manual No Scans 9A 28 Elbow /V. Pipe P-scan None 36B 22 Cross / Header P-scan Not Scanned Cross Side 0-375 mm 625-1250 mm 1500-1875 mm Due to Geometry Not Scanned Header Side 0-250 mm Manual Scanned 0-250 mm Header side
r-TABLE I (Continued) 1984 Examination Restriction Summary Weld Systen Number Size Configuration Method Restriction Recire 41 12 V. Pipe / Elbow P-scan None (Cont'd) Manual None 44 12 V. Pipe / Elbow P-scan Not Scanned Elbow Side 375-750 mm Manual Elbow Side Scanned 375-750 mm 47 22 Valve / Header P-scan No Scan Valve Side 48 22 Valve / Header P-scan No Scan Valve Side 49 22 Header / Valve Manual No Scan Valve Side 51A 12 Elbow /H. Pipe P-scan No Scan Elbow Side 375-500 mm 16B 22 Cross / Header P-scan No Scan Cross Side Header Side Not Scanned 0-250 mm 875-1000 mm Manual Areas Not Scanned By P-scan 17 28 V. Pipe /V. Pipe P-scan None 17A 28 V. Pipe /V. Pipe P-scan None 17B 28 V. Pipe / Elbow P-scan None 23A 22 Header /Sweepolet Hanual No Scan Sweapolet Side 23B 22 Header / Cap P-scan None 26 28 Safe End/H. Pipe P-scan No Scan Safe End Side 26A 28 H. Pipe / Elbow P-scan None 27 28 Elbow /V. Pipe P-scan None
{E - TABLE I (Continued) 1984 Examination Restriction Summary Weld System Number Size Configuration Method Restriction Recirc 30A 22 Header /Sweepolet Manual No Scan on Sweepolet (cont'd) 30B 22 Header / Cap P-scan None 36A 22 Cross / Header P-scan No Scan On Cross 98 28 V. Pipe /RHR Tee P-scan No Scan On Tee 15 28 V. Pipe /RHR Tee P-scan No Scan On Tee 15A 28 RHR Tee /V. Pipe P-scan No Scan On Tee 15B 28 V. Pipe /V. Pipe P-scan No Scan 0-125 mm Either Side HVAC Interfers Manual No Scan 15C 28 V. Pipe / Elbow P-scan No Scan 0-125 mm Manual Areas Not Scanned By P-scan 16A 22 Cross / Header P-scan No Scan Cross Side No Scan On Header Side 0-250 mm 875-1250 mm 1625-0 mm Manual Areas Not By P-scan 54A 12 Elbow /H. Pipe P-scan Elbow Side Not Scanned 135-1800 Shielded (hot spot) Manual Scanned 135-1800 plus evaluations 56 28 Elbow /Va'. ve P-scan No Scan Valve Elbow Side Not Scanned 0-125 mm 125-0 mm Permanent Interference
TABLE X (Continued) 1984 Examination Restriction Summary Weld System Number Size Configuration Method Restriction Recirc 58 28 Elbow / Pump P-scan No Sean On Pump (cont'd) 59 28 Pump /H. Pipe P-scan No Scan On Pump 61 28 Pipe / Valve P-scan No Scan On Valve 64 28 Elbow /V. Pipe P-scan None 65A 28 V. Pipe /RHR Tee P-scan No Scan On Tee 45 12 Sweepolet/V. Pipe Manual None w/ Overlay 50 12 Sweepolet/V. Pipe Manual None w/ Overlay 51 12 V. Pipe / Elbow Manual None w/ Overlay 53 12 Sweepolet/V. Pipe Manual None w/ overlay 54 12 V. Pipe / Elbow Manual None w/ Overlay 16 12 Red Cap /V. Pipe Manual None w/ Overlay 18 12 V. Pipe / Elbow Manual None w/ Overlay 20 12 Sweepolet/V. Pipe Manual None w/ Overlay 23 12 Sweepolet/V. Pipe Manual None w/ Overlay 24 12 V. Pipe / Elbow Manual None w/ Overlay 29 12 V. Pipe / Elbow Manual None w/ Overlay
TABLE X (Continued) 1984 Examination Restriction Summary Weld System Number Size Configuration Method Restriction
'Recirc 30 12 Sweepolet/V. Pipe Manual None
'(Cont'd) w/ Overlay 32 12 V. Pipe / Elbow Manual None After Overlay Build-up 35 12 V. Pipe / Elbow Hanual None w/ Overlay 36 12 Red Cap /V. Pipe Manual None w/ Overlay 42 12 Sweepolet/V. Pipe Manual None w/ Overlay RHR-30 1 24 Elbow / Tee Manual No Scan On Tee 3 24 Pipe / Elbow P-scan No Scan On Elbow Side 9 24 Pipe / Elbow P-scan Weld Inaccessible 750-1125 mm No Scan Either Side 10 24 Elbow / Pipe P-scan Wold Inaccessible 875-1125 mm No Scan Either Side RNR-31 1 24 Elbow Tee Manual No Scan On Tee RNR-32 1 20 Tee / Elbow Manual No Scan On Tee t
2 20 Elbow /V. Pipe P-scan No Scan On Elbow F1 20 Pipe /Sweepolet Manual No Scan On Sweepolet 4 2G Pipe / Elbow P-scan No Scan 1125-1375 mm Permanent Interference 5 20 Elbow / Valve Manual No Scan Valve Side 6 In. Obstruction (Penmanent Support) l'
l TABLE I (Continued) 1984 Examination Restriction Summarv Wald System Number Size Configuration Method Restriction RHR-32 6 20 Valve / Pipe P-scan No Scan Valve (cont'd) No Scan Pipe Side 500-1125 mm Permanent Intecference
.7 20 Pipe / Valve P-scan No Scan Valve Side No Scan Pipe Side 500-1125 mm '
Permanent Interference 4 20 Pipe / Elbow Manual None After overlay
---- -_ ---_-- _ -- -- _ - ------ - ----- l
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