ML20128E387
| ML20128E387 | |
| Person / Time | |
|---|---|
| Site: | Summer  |
| Issue date: | 01/31/1993 |
| From: | Esposito J WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19303F266 | List: |
| References | |
| SG-93-01-013, SG-93-1-13, WCAP-13523, NUDOCS 9302100411 | |
| Download: ML20128E387 (380) | |
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%"dSTtGOUSECLASS3 SG 93-01013 j 4~ WCAP-13523 4- .I Q
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V. C. Summer - j Steam Generator interim Tube Plugging Criteria for .
Indications at Tube Support Plates
.j e
January 1993
- .3 m
i.
Approved by:
J. . Esposp6,~ Manager.
)S am_ Generator Technology & Engineering.
' .j i.
1 I
'O 1993 Westinghor.se Electric Corporation--
' A!! Rights Reserved
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s v
Table of Contents -
Section Il11e Eggt-
1.0 INTRODUCTION
11 2.0 CCNCLISCNS 21 3.0 SUPPORT PLATE REGION PULLED TUBE DATABASE (3/4" TUBING) 31 3.1 Introduction & Definitions 31 3.2 V. C. Summer Corrosion Degradation at Support Plates 32 3.3 Plant R 1 Corrosion Degradation 3-3 3.4 Plan: E-4 Corrosion Degradation 3 10 3.5 Plant B 1 Corrosion Degradation 3 10 4.0 LABORATORY SPECIMEN PREPARATION & TESTING 41 4.1 Preparation of Specimens 41 4.2 Nondestructive Examination (NDE) Results 43 4.3 Leak Rate Testing - 43 4.4 Burst Testing 4-4 4.5 . Destructive Examination 4 4.6 Model Boller Data Base Summary- 47-5.0 NON-DESTRUCTIVE EXAMINATION (NDE) 51 5.1 Eddy Current Voltage Normalization for APC _ 5-1
- - 5.2 Eddy Current Data Analysis Guidelines 52 5.3 Voltage Trends for EDM Slots 52 5.4 Voltage Renormalization for Alternate Caliorations 52-5.5 Plant R 1 Pulled Tube Data 5-3 5.6 Belgian Pulled Tube Data 5-5 5.7 NDE Uncertainties for V. C. Summer 57 6.0 PULLED TUBE AND FIELD DATA EVALUATION 6-1 6.1 Utilization of Field Data in Tube Repair Limits 61 6.2 Summary of Pulled Tube Data Base 61 6.3 Operating Plant Leakt.ge Data for ODSCO at TSPs :62 6.4 Voltage Renormalization for Alternate Ca!!brations 62 6.5 Tensile Property Considerations 6-3 6.6 Evaluation of Plant R 1 Pulled Tubes 6-3 6.7 Evaluation of Plant E-4 Data 6-5 6.8 Evaluation of Plant B 1 Pulled Tubes 6-6 6.9 Growth Rate Trends 6-6 6.10 Summary of Pulled Tube Test Results 6-7 7.0 ' GUIDELINES FOR ACCIDENT CONDITlON ANALYSES 7-1
- 7.1 Limiting Accident Condition 7-1 7.2 Event Sequence Probabilities 7 4_-
7.3 NUREG-0844 Analysis 7-4 7.4 SLB Analysis Guidelines 75 i
Table of Contents (Continued)
Section Il!!g Ea;gt 8.0 ACCIDENTCONDITIONCONSIDERATIONS 81 8.1 Tube Deformation Under Combined LOCA + SSE 81 8.2 Tube Maps / Summary Tables for Potentially Affected Tubes 8-5 8.3 Effect of Combined Accident Conditions on Tube Burst Capability 86 9.0 V.C. SUMMER INSPECTION RESULTS 9-1 9.1 October 1991 Inspection 91 9.2 Prior inspection 92 9.3 Growth in Voltage Amplitude 9-3 9.4 Influence of TSP Location 9-4 9.5 RPC Inspection Resuits 94 10.0 BURST PRESSURE CORRELATION 10 1 10.1 Introduction 10 1 10.2 Data Base for Burst Pressure Correlation (3/4" Tubing) 10 1 10.3 Burst Pressure vs. Voltage Correlation 10-2 11.0 SLB LEAK RATE CORRELATION 11 1 11.1 Introduction _ _
11 1 11.2 Data Base for SLB Leak Rate Correlation (3/4" Tubing) 11 1 11.3 Leak Rate Threshold Assessment 11 2 11.4 Probability of SLB Leakage vs. Bobbin Voltage 11 6 .
11.5 General Trends for SLB Leak Rate Correlation 11-9 11.6 SLB Leak Rate vs. Voltago Correlation 11 10 12.0 V. C. SUMMER IPC EVALUATION 12 1 12.1 Introduction 12-1 i 12.2 V. C. Summer Interim Plugging Criteria (IPC) 12-1 12.3 Equivalent V. C. Summer APC Repair Limit 12 2 12.4 Projected EOC Voltages 12 3 12.5 Tube Burst Margin Assessment 12-6
, 12.6 SLB Leak Rate Analyses 12-7 12.7 Operating Leakage Limit 12-9 12.8 Conclusions 12-10 APPENDICES:
A NDE Data Acquisition end Analysis Guidelines A-1 B Regression Analysis Methodology B1 C Adjustment Procedure for SLB Le9 Rates . C 1-e li
Section 1 NTRODUCTION This report provides the technical basis for interim tube plugging criteria for outside diameter stress corrosion cracking (ODSCC) at tube support plate (TSP) Intersections
_ in the V. C. Summer steam generators (S/G) for the 1993 04 operating cycle. The -
recommended plugging criteria are based upon bobbin coil inspection voltage amplitude which is correlated wi'.h tube burst capability and leakage potential. The recommended ,
criteria are demonstrated to meet the guidelines of Regulatory Guide (R.G.) 1.121. * '
The tube plugging criteria are based upon the conservative assumptions that the tube to TSP crevices are open (negilglbie crevice deposits or TSP corrosion) and that the TSPs are displaced under accident conditions. The ODSCO existing within the TSPs is thus assumed to be free span degradation under accident conditions and the principal requirement for tube plugging considerations is to provide margins against tube burst per R.G.1.121. The open crevice assumption leads to maximum leak rates compared to packed crevices and also maximizes the potential for TSP displacements under accident conditions. Laboratory tests performed with incipient denting or dented tube intersection show no leakage or very smallleakage such that leakage even under steam line break (SLB) conditions would be negligibia, it has been demonstrated for 51 Series steam generators, using Plant A 1 as an-example, that if the crevices are packed as a consequence of TSP corrosion or if small tube to TSP gaps are present, TSP displacements under accident conditions are minimal -
. such that tube burst would be prevented by the presence of the TSPs. TSP displacement analyses under SLB loads were also performed for the open crevice assumptiv.) with the further conservative assumption of zero friction at the tubo to TSP intersections and at the TSP wedge to wrapper interaction. The wedges are installed in the TSP to wrapper gaps to align the TSPs for tubing of the S.G.s. While the TSP wedges are pressed into the gap during manufacturing, the forces are not known and thus the preload or friction force at the TSP to wrapper interface is not known. It is reasonable to expect that the friction forces at the TSP to wrapper interface would significantly reduce TSP oisplacements under accident conditions. However, the analytical results based upon the open crevice /zero friction assumptions indicate the potential for TSP displacements under SLB conditions such that prevention of tube rupture cannot be assumed for the 51 Series S/Gs with the applied ardytical assumptions. Analysis of TSP displacement under SLB conditions has not bean performed for Model D-3 steam generators as present in Summer. Therefore the regt trements for tube burst margins assuming free span degradation have been applie j to develop the tube plugging criteria for V. C. Summer S/Gs.
The plugging criteria were developed from testing of laboratory induced ODSCC specimens, extensive examination of pulled tuber from operating S/Gs and field experience for leakage due to Indications at TSPs. The recommended criterla represent conservative limits based upon Electric Power Research Institute (EPRI) and industry supported development programs that are continuing toward fur *her refinement of the .
plugging criteria. The significant database currently available permits use of burst pressures at the lower 95% confidence bound as the basis for the tube plugging limits.
The burst pressure database used here is derived from 3/4 inch diameter tog. For 1-1 s
3/4 inch alameter tubing, this is believed to be more conservative than combining the 3/4 inch and 7/8 inch databases, implementation of the tube plugging criteria is supplemented by 100% bobbin coil Inspection requirements at TSP elevations having ODSCC indications, reduced operating .
leakage requirements, inspection guidelines to provide consistency in the voltage normalization and rotating pancake coil (RPC) inspection requirements for the larger indications left in service to characterize the principal degradation mechanism as ODSCC. In addition, it is required that potential SLB leakage be calculated for tubes with TSP Indications left in service to demonstrate that the cumulative leakage at end of cycle cx>nditions is less than the allowable limit of 1 gpm. -
Two tubes were pulled from V. C. Summer steam generators in October.1988.
Devructive examination of these tubes showed axial ODSCO to be present et the first TSP intersection (lowest in elevation) in the hot leg. The degradation was similar to that observed at several other plants and consistent with the data base used to develop repair limits in this report.
To provide the technical bases for tube plugging due to ODSCC at TSPs, the following activities have been performed as documented in this report:
o Review of V. C. Summer and other plant pulled tube destructive examinations Section 3 o Preparation of cracked test specimens, the;r non destructive examination (NDE), leak .
rate testing, burst testing, and destructive examination - Section 4 o NDE inspection and data analysis guidelines, voltage trews for EDM (electrodischarge .
4 machining) slots, voltage normalizations for pulled tubes and overall NDE uncertainties -
Section 5 ,
o Review of industry wide pulled tube examination results: eddy current data, burst pressure, primary to secondary leakage, and voltage growth rate trends - Section 6 o Guidelines for accident condition analyses relating to burst probability and primary to secondary leakage - Section 7 o Structural evaluations of combined accident conditions (LOCA + SSE) - Section 8 o Review of V. C. Summer eddy current inspection results including historical growth rate data - Section 9 o Burst pressure correlation to relate the NDE parameters (bobbin voltage) to burst strength of tubing - Section 10 o Leak rate correlation to relate the bobbin coil voltage to leak rate under SLB conditions - ~
Section 11 o Integration of the inspection, leak rate and burst test results to develop ihe interim tube .
repair limits for V. C. Summer - Section 12.
The overall summary and conclusions for this report are described in Section 2.
12
7 Section 2 -
CONCLUSIONS This report documents the technical support for a V. C. Summer interim plugging criteria (IPC) of 1.0 volt for ODSCC indications at TSPs. The data base of pulled tube and model boiler specimens used in the evaluation of the IPC are described in this report. This data base is used to develop correlations relating burst pressure to bobbin voltage and SLB teak rate to bobbin voltage. These correlations are used in the tube integrity assessment to demonstrate V. C.
Summer IPC margins against Reg. Guide 1.121 guidelines for tube plugging limits.
Overall Conclusions -
The following requirements for the V. C. Summer IPC very conservatively satisfy Reg. Guide 1.121 guidelines for tube integrity:
Tube Repair Basis o Bobbin coil indications having flaw voltages greater than 1.0 volt and confirmed as flawr. by RPC inspection shall be repaired, o Tubes with bobbin collindications >1,0 volt may be repaired as an attemative to RPC inspection. Bobbin coil indications having flaw voltages greater than 2.2 volts shall be repaired independent of RPC confirmation of a flaw.
- o Projected leakage for a postulated steam line break (SLB) event at no of cycle (EOC) condalons shall be lesa than 1.0 gpm for the most limiting Sta Ecbbin coll ,
flaw indications inspected by RPC and found to have no RPC Indication do not need to be included in the leakage analyses.
o Tubes identified as subject to significant deformation at a TSP elevation under a posulated LOCA + SSE event shall be excluded from application of the IPC at that-TSP location.
Inspection Requirements o The inspection shallinclude 100% bobuir coil Inspection of all hot leg intsrsections and cold leg intersections down to the lowest TSP for which the IPC is to be applied, o Bobbin coil flaw indications above 1.0 volt and below 2.2 volts shall be inspected by i- RPC to evaluate for detectable RPC Indications and to support ODSCC as the degradation mechanism, unless the tube is to be plugged or sleeved.
. o Eddy current analysis guidelines shall be consistent with guidelines given Appendix A, o An RPC sampling program of at least 100 TSP intersections will be performed -
emphasizing intersections with greater than 5 volt (bobbin coll) dents and including some intersections with artifact bobbin indications or indications with unusual phase angles.
2- 1
Operating Leak Rate Limit o The normal operating leak rate requiring plant shutdown shall be limited to 150 gpd per S/G.
Summarv of Conclusions o The V, C. Summer pulled tubes (1988 outage) show that the crack morphology for indications at TSPs can be described as axial ODSCC within the TSP length and with negligible volumetric IGA involvement. The indications are consistent with the data base supporting the repair limits of this report.
o Recommended correlations of bobbin voltage to burst pressure and to SLB leakage, as well as attemate correlations for sensitivity analyses, are developed in this report.
These correlations form the basis for determining margins for burst and leakage as summarized below.
4 o At EOC, burst pressure capability (expressed as margin ratios relative to 34PNO and APSLB) is expected to have ratios of about 1.19 relative to 3APNO at 90% cumulative probability levels and about 1.54 relative to APSLB at 99% cumulative probability levels. A burst pressure margin ratio of 1.35 relative to 3aP NO for V, C. Summer at BOC conditions is comparable to typical values of 1.4 for plants with 7/8 inch diameter tubing with an IPC repalt limit of 1.0 volt. Thus the two tubing sizes can be considered to have comparable margins for an IPC repair limit of 1.0 volt.
o Potential SLB leakage at EOC conditions is expected to be well below the 1.0 gpa allowable limit as supported by both Monte Carlo and determlnistic evaluations including sensblvity analyses. ,
o The maximum EOC bobbin voltage indication resulting from indications at 1.0 volt is projected to be about 2.97 volts for 200 BOC Indications and up to 3.08 volts for 500 BOC Indications, o The maximum EOC voltage of about 3 volts following implementation of the IPC repair limits is comparable to the maximum voltage of 2.8 volts found in the 1991 Inspection following implementation of the 40% depth repair limits. Thus, the limiting tube for structuralintegrity considerations is essentially independent of IPC or 40% depth repair limits, o The operating leak rate limit of 150 gpd implemented with the IPC satisfies R.G.
1.121 guidelines for leak before break. This limit provides for plant shutdown prior to reaching critical crack lengths for SLB conditions at a 95% confidence level on leak rates and for 3AP conditions at less than nominal leak rates.
22
7 Setion 3 -
SUPPORT PLATE REGION PULLED TUBE DATABASE (3/4 INCH TUBING) 3.1 Introduction & Definitions The following provides summsty information regarding OD originated corrosion at support plate crevice regions of Alloy 600 tubing pulled from steam generators at various plants including V. C. Summer The data is presented in support of the development of tube plugging criteria for Summer. First, pulled tube data from V C. Summer are reviewed followed by data from other .
plants. All available (to Westinghouso) tube pull data at tube support plate locations from other plants with 3/4 inch diameter tubing are summarized.
The type of Intergranular corrosion with regard to crack morphology and density (number, length, depth) of cracks can influence the structural integrity of the tube and the eddy current ,
response of the indications. To support the tube repair criteria, the emphasis for destructive .
examination is placed upon characterizing the morphology (SCC, IGA involvement), the number of cracks, and characterlzation of the largest crack networks with regard to length, depth and remaining ligaments between cracks. These crack details support interpretation of structural -
parameters such as leak rates and burst pressure, crack length and depth, and of eddy current parameters such as measured voltage with the goal of enhancing structural and eddy current evaluations of tube degradation in selective cases, such as the Plant R 1 pulled tubes, the pulled tube evaluations included leak rate measurements,in addition to the more standard burst pressure measurements, for further support of the Integrity and plugging limit evaluations.
o Before the support plate region corrosion degradation can be adequately described, some key corrosion morphology terms need to be defined, intergranular corrosion morphology can vary from IGA to SCC to combinations of the two. IGA (Intergranular Attack) is defined as a three -
dimensional corrosion degradation which occurs along grain boundaries. The radial dimension has a relatively constant value when viewed from different axial and circumferential coordinates. IGA can occur in isolated patches or as extensive networks which may encompass the entire circumferential dimension within the concentrating crevice. Figure 31 provides a sketch of these IGA morphologies. As defined by Westinghouse, the width of the corrosion should he equal to or greater than the depth of the cerrosion for the degradation to be classified as lGA.
The growth of IGA is relatively stress independent. IGSCC (Intergranular Stress Corrosion Cracking) is defined as a two-dimensional corrosion degradation of grain boundarles that is strongly stress dependent. IGSCC is typically observed in the axial-radial plane in steam generator tubing, but can occur in the circumferential-radial plane or in combinations of the two planes. The IGSCC can occur as a single two dimensional crack, or it can occur with branches -
coming off the main plane. Figure 3-2 provides a sketch of these IGSCC morphologies. Both of the IGSCC variations can occur with minor to major components of IGA. The IGA component can occur simply as an IGA base with SCC protruding through the IGA base or the SCO plane may have a semi-three dimensional characteristic. Figure 3 3 provides a sketch of some of the -
. morphologies possible with combinations of IGSCC and IGA. - Based on laboratory corrosion tests, it is believed that the latter, SCC protrusions with significant IGA aspects, grow at rates similar to that of. SCC, as opposed to the slower rates usually associated with IGA. When 1GSCC and IGA are both present, the IGSCC will penetrate throughwall first and provide the leak path.
To provide a semi quantitative way of characterizing the amount of IGA associated with a gisen crack, the depth of the crack is divided by the width of the IGA as measured at the mid-depth of 31
_ _3
the crack, creating a ratlo D/W. Depth is the llrs ar dimension measured from the OD surface of the tube through the tube wall, i.e., along the tube radius. Width (for an axial crack) is measured along tube circumference, i.e. perpendicular to the depth. In the current context (for D/W rat!c' the width is measured at the mid depth of the crack,i.e., half the distance between the tube UD and the deepest point of the crack. Three arbitrary DN' categories were created:
minor (D/W greater than 20) (all or most PWSCC would be included in this category if it were being considered in this analysis); moderate (D/W betyteen 3 and 20); and significant (D/W less than 3) where for a given crack with a D/W of 1 or less, the morphology is that of patch IGA.
The density of vacking can vary from one single large crack (usually a macrocrack composed of many microcracks which nucleated along a line that has only a very small width and which ther g*ew together by intergranular corrosion) to hundreds of very short microcracks that may have partially linked together to form dozens of larger macrocracks. Note that in cases where a very high density of cracks are present (usually axial cracks) and where these cracks also have significant IGA components, then the outer surface of the tube (crack origin surface) can form regions with effective three dimensional IGA. Axlal deformations of the tube may then cause circumferential openings on the outer surface of the tube within the three dimensional network of l gat these networks are sometimes mistakenly referred to as circumferential cracks. The axial cracks, however, will still be the deeper and the dominant degradation, as compared to IGA.
Recognizing all of the gradations between IGA and IGSCC can be difficult. In addition to observing patch IGA, cellular IGA / SCC has been recently recognized, in cellular IGA / SCC, the cell walls have IGSCC to IGA characteristics while the interiors of the cells have nondegraded metal. The cells are usually equiaxial and are typically 25 to 50 mils in diameter. The cell walls (with intergranular corrosion) are typically 3 to 10 grains (1 to 4 mlls) thick. The thickness and shape of the cell walls do not change substantially with radial depth. Visual examinations or ,
limited combinations of axial and transverse metallography will not readily distinguish cellular IGA / SCC from extensive and closely spaced axiallGSCC with circumferentialledges linking axial microcracks, especially if moderate to signlficant IGA components exist in association with the cracking Radial metallography is required to definitively recognize cellular IGA / SCC Cellular IGA / SCC can cover relatively large regions of a support plate crevice (a large fraction of a tube quadrant within the crevice region). Figure 3-4 shows an example of cellular IGA / SCC from Plant L.
A given support plate region can have intergranular corrosion that ranges from IGA through individual IGSCC without IGA components.
3.2 V. C. Summer Corrosion Degradation at Support Plates Segments of two steam generator tubes were rJted from the V. C. Summer plant in October 1988. These were destructively examined by Babcock & Wilcox (B&W) to characterize the degradation. The results are documented in EPRI (Electric Power Research Institute) Report NP 6998-SD dated October 1990. The pulled tubes were located at R36C67 and R35070. The ,
intersections at the flow distribution baffle and at the first TSP were inspected.
Two SEM (scanning electron microscopy) micrographs of the degradation in tube R36C67 at the -
first TSP in the hot leg are shown in Figure 3-5. It may be noted that the degradation is ODSCC with some IGA aomponents. The axial macrocracks are made up of smaller microcracks. The axial orientation of the cracks was confirmed by metallography. A longitudinal section through 3- 2
the area adjacent to the defects is shown in Figure 3 6. This shows that there was no significant through wall penetration. The axial cracks penetrated 14% through the nominal wall thickness.
No general attack or wastage was noted. A transverse view of the tube at the first TSP region is 1 shown in Figure 3-7.
The first TSP region of tube R35C70 showed somewhat similar results. SEM micrographs of this tube at the first TSP are shown in Figure 3 8. Axlal ODSCC with IGA involvement is visible.
Although metallographile examination was not performed on this tube, no general attack or wastage of the tube in the crevice region was observed.
The flow distribution baffle region of tube R36C67 showed no IGA or crack like degradation.
There were several pit type indications which contained some etched grain boundaries, as illustrated in Figure 3 9. The pits were not visually detectable and were present all around the circumference in small numbers. Optical metallography was not performed on the tube at this location.
The flow distribution baffle region of tube R35070 showed axial ODSCC with minor IGA. The area of degradation extended about half the circumference of the tube. An SEM micrograph of the tube at this location is shown in Figure 310. No general attack or wastage was observed.
3.3 Plant R-1 Corrosion Degradation For olant R 1, three tubes with six intersections were pulled in 1992 and five tubes with nine intersections were removed in 1991 or earlier. This section describes the results of the destructive examinations for these tubes. ;
3.3.1 Plant R 11992 Pulled Tubes
~
Three hot leg steam generator tube segments from Plant R 1 (tubes R7-C71 and R9 076 from S/G C and tube R9-C91 from S/G D) were examined in 1992 to provide supporting data for the development of alternate plugging criteria specific to support plate crevice corrosion. The first, second and third support plate crevice regions of each tube were nondestructively examined.
Subsequently, elevated temperature leak testing and room temperature burst testing were -
conducted on the second and third support plate crevice regions of each tube. The burst tested specimens were then destructively examined using metallographic and SEM fractographic techniques. The following provides a brief summary of the more significant observations.
NDE Results A summary of the field and laboratory NDE results is given in Section 5. While OD origin indications were observed at the second and third support plate crevice regions of each tube by the various eddy current examinations performed, none were found within the first support plate (flow distribution baffle location) crevice region. There was good agreement between the
, field and laboratory eddy current results, but a significant (qcrease in the bobbin probe signal voltage was noted in three instances in going from the field to the laboratory data. These increases are probably related to the effects of the tube pulling stresses on the corrosion crack networks. The eddy current data suggested that corrosion was present in the form of axial crack +
within the crevice regions. Laboratory UT data suggested that a larger number of axial Indications were present.
3-3 1
N' _
l
Leak and Burst Testina The second and third support plate crevice region of each tube was leak tested at elevated temperature and pressure. This test is capable of accurately measuring very low levels of leakage. None of the crevice regions of Tubes R7 071 or R9 076 leaked at normal operating ,
conditions (1500 psl differential pressure) or at steam line break conditions (2650 psi differential pressure). The third support plate crevice region (SP3) of Tube R9 091 did not -
leak at normal operating conditions, but did develop a very small leak (0.023 liters per hour) at steam line break conditions. Post leak inspection (pressurlzing a leak specimen held in a glass container filled with water and then looking for the source of gas bubbles) verified that the leak occurred within the crevice region and not at a fitting. SP2 of Tube R9 C91 developed a 0.032 liter per hour leak at normal operating conditions and a 3.33 liter per hour leak at steam line break conditions. Post-leak test inspection, found that the crevice region had two leak location 3.
Both leak locations had very small (-0.08 inch) throughwall corrosion. Extensive destructive
- examination was required to locate inroughwall corrosion at SP3 of R9-C91. Neither the burst crack nor the next largest macrocrack had throughwall corrosion. The only throughwall corrosion (0.016 inch in length) was located in a crack network adjacent to the second largest macrocrack.
Room temperature burst tests were conducted on the second and third support plate crevice regions of each tube at a pressurization rate of 1000 psi per second. Results of the burst tests are presented in Table 31. All burst tests resulted in bulging or fishmouthing and tearing of the crack opening, as shown by the ductility, burdength and burst width data of Table 31. All but SP2 of Tube R9-C91 developed simple axial burst openings which were centered within the crevice regions. SP2 of Tube R9-C91 also hed an axial burst opening centered within the crevice region, but it was complex in shape it appeared to have formed from three closely spaced and interconnected axial openings. Another interesting observation was that the burst opening of SP3 of Tube R9 C91 occurred at a different location than the leak location. All six burst specimens had similar burst pressures that ranged from 5,400 psi to 7,700 psi. SP3 of *~
Tube R7 C71 had data recorder problems that prevented knowing the true burst pressure. It is known that its burst pressure exceeded 5,450 psl.
The upper figure of Figure 311 shows the recording of burst pressure versus time for R9-C76, SP2 as an example of a normal and successful burst test. It is seen that the recorder responds quickly to the pressure changes. The water pressure is controlled to follow the drawn target line. However, due to the small volume of the overall system, the initial water pump _
strokes produce spikes in the pressure, although these spikes are negligible at moderate
- pressures, The recorder follows these spikes closely.11is also seen that the r6 corder immediately follows the complete drop in pressure following the burst test.
4 The burst history of SP3 of tube R7-071 is complex;it was burst tested twice. The first time, a Swagelok fitting leak produced a sms!! leak that prevented the internal pressure from exceeding 5,450 psi The middle part of Figure 311 shows the pretsure versus time curve for this run.
Note that the recorder normally followed the initial pressure spikes, but with no tube burst -
there was no rapid drop in pressure. 'Instead, there was gradual drop-off in pressure as the '
pump tried to maintian pressure with the small fitting leak. After the initial attempt at bursting the tube,it was noted that the tube appeared normal and that a few drops of water were leaking from the lower Swagelok fitting when several hundred psi of water pressure was applied.' The -
Swagelok fitting was tightened and the test was repeated During the second run, the recorder was malfunctioning such that the recorder was moving in a sluggish, nonresponsive manner in -
the vertical (pressure) axis. The bottom figure of Figure 3-11 shows the pressure curve.-
34
First, note that no initial pressure spikes were recorded. Second, note that the post-burst test pressure drop-off was slow, even though a large burst opening was subsequently observed.
Finally, note that the maximum pressure recorded was lower than the initial burst run, but the tube only had plastic deformation during the second run. After adjustment, no subsequent
. recorder malfunction occurred. it is suspected that the recorder was set at the calibration setting (filter in recorder to facilitate calibration) and not returned to the record setting for the final burst test. Again,it is concluded that the burst test results for SP3 of tube R7 C71 are low and should be excludsd in any final data base of burst test results.
Free span portions of the three tubes had room temperature burst pressures that ranged from 11,400 to 11,900 psl. Room temperature tensile properties of free span sections are also pret.ented in Table 31.
Destructh/e Examination Results The leak and burst fracture faces were opened for SEM fractographic examinations. Table 3 2 presents a summary of the fractographic data. The burst openings occurred in axlat macrocracks that were composed of numerous axially oriented microcracks that were confined to a relatively narrow axial band. Most of the microcracks had Interconnected during plant operation since most of the microcrack ledges, separating the individual microcracks, had only intergranular features. However, the macrocracks also had two to four ledges with dimple rupture features, Indicating that the metal between these microcracks tore during burst testing. The thickness of these tom ledges ranged from 0.001 Inch to 0.024 inch. Most of these torn ledges occurred nearer the macrocrack tips than the mid macrocrack regions. The macrocracks were confined to the crevice regions and were typically 0.4 inch long with a maximum length of 0.57 inch. The maximum crack depths ranged from 77% to 100% throughwall while the average macrocrack depths ranged from 47% to 83% throughwall.
Figures 3-12 through 317 present sketches of the crack distribution found by visual (30X stereoscope) examinations of the post-burst tested specimens and by subsequent destructive examinations. The sketches show the locations where cracks were fcund and their overall appearance, not the exact number of cracks or their detailed morphology. Due to the complexities of the observed crack ne' works, radial metallography,in addition to the more standard transverse and axial metallography, was frequently used to provide an overall understanding of the intergranular carroslon morphology. In radial metallography, small sections of the tube (typically 0.5 by 0.5 inch) are flattened, mounted with the OD surface facing upwards and then progressively ground, polished, etched and viewed from the OD surface towards the ID surface. From the metallographic examinations,it was concluded that the dominant corrosion morphology was axiallntergranular stress corrosion cracking (IGSCC) with various amounts of intergranular cellular corrosion (ICC) being associated with the axial IGSCC. In some instances only axial IGSCC was present. In other instances, ICC appeared to dominate locally. However, with progressive grinding, it was shown that axial IGSCC was deeper than the associated 100, even though the 100 was deep (typically 37% or less). At the locations where ICC had been present, only axial IGSCC remained after grinding. This feature of cellular corrosion
. with partial depth cellular patterns and deeper penetration by axial IGSCO has been found in all pulled tubes with cellular corrosion, including Plant E-4 Indications with more extensive cellular corrosion than found in domestic pulled tubes.
Figure 318 provides an example of both axial IGSCC and ICC as revealed by radial metallography. The low magnification (16X) photomicrographic montage is from a portion of the mid crevice region of SP3 of Tube R9 C76 at a depth of 26% throughwall Predominantly 3 -5
4 axial IGSCC is seen to the left of the photo and predominantly 100 is observed to the right of the section. Figure 319 provides higher magnification (100X) photomicrographs of the 100 region. j All of the support plate crevice regions examined were remarkably similar in the types ,, l corrosion found and in the corrosion distr lbution. Overall crack densities were low to moderate -l (typical crack densities ranged from 4 to 25 cracks over 360* at a glvon elevation), However, j the cracks were not uniformly dlstributed and local crack densities were significantly higher. -
Crack densities would range from 48 to 144 cracks over 360* If local area data were incorrectly extrapolated over the tube circumference. All of the individual corrosion cracks found had only minor to moderate IGA components (D/W ratlos ranged from 7 to 31). No surface IGA, i.e., that Independent of corrosion cracks, was observed.
For R9 C91 of SP2, the leak locations were comprised of very short (~0.08") throughwall corrosion penetrations at the burst crack and at a secondary crack. For SP3 of R9-C91, which showed no leakage at normal operating conditions and a very small (0.0231/hr) leak at SLB conditions, neither the burst crack nor the largest secondary macrocrack (Table 3 2) were i found to have throughwall corrosion. The cellular pattern at the edges of the secondary crack at about 180' of Figure 317 was also examined and found to have no throughwall corrosion.
Throughwall corrosion of very short length (0.016") was found in the cracking pattern adjacent
- to the secondary crack. Following leak and burst testing, the throughwall penetration at the tube ID was about 0.044"long. Figure 3-20 shows fractography and a sketch for the associated crack. The short length of deep corrosion accounts for the very small leak rate, as well as the modest bobbin coil voltage (1,13 volts) for this intersection.
From Table 3 2 for the 3rd TSP of R9 C91,it can be noted that the burst crack has a shorter length and smaller average depth than the largest secondary crack. This occurs as patches of cellular SCC at the ends of the secondary crack are included in the total crack length -
1 measurement. It is expected that the tortuous nature of the crack path for the secondary crack made it somewhat stronger than the burst location.
Concludons The examined second and third support plate crevice regions of the pulled tubes had combinations 4 of axially oriented IGSCC and 100,- The corrosion was of OD origin and was always confined to within the crevice region. Usually, the corrosion was centered within the crevice region and did not extend to the support plate crevice edge locations. Overall crack densities were low to moderate, but with the non uniform nature of the crack distributions, local crack densities were high. Only minor to moderate IGA components were found in association with the IGSCC and no -
surface IGA was observed.
The field and laboratory eddy current inspections accurately described the presence of axial cracking. Destructive examinations showed that the most significant cracking occurred where the eddy current Indications were located. Laboratory UT inspection results suggesW the presence of an even larger number of axial cracks. Destructive examinations showed that the UT ~
call of a larger number of axial cracks than the eddy current data suggested was correct.
Crack networks in two of the crevice regions developed smallleaks during the leak testing that ,
was conducted prior to burst testing. The maximum leak rate at elevated temperature and pressure was 0.032 liters per hour at normal operating conditions and 3.33 liters per hour at .
steam line break conditions. The presence of intergranular stress corrosion cracking at the 3- 6 s
support plate locations did not reduce the burst pressures of these corroded regions by more than a factor of 2.2. This is well above the safety limitations. The burst pressures ranged from 5,400 psi to 7,700 psi at locations where corrosion macrocracks ranged from 0.35 inch to 0.57 inch long. The maximum depth of corrosion ranged from 77% to 100% throughwall.
3.3.2 Plant R 11991 Pulled Tubes This section describes the crack morphology and burst / leak rate measurements for tubes pulled from Plant R 1 in 1991. The burst test data are described to assess whether the measured pressures are representative of a buret or a more limited crack operdng causing leakage, A completed burst test is characterized by fishmouth opening of the crack, bulging of the tube and/or tearing at the edges of the corrosion crack as found for the 1992 pulled tubes (Table 3-2). In general, the burst test opens up the entire macrocrack length or a very large frection of the corrosion crack, it is shown that the 1991 burst tests resulted in either minor crack opening (not representative of a complete burst) or moderate openings stillless than expected for a complete burst. For essentially undegraded tube sections which had fishmouthed and bulged ruptures, the resulting burst pressures were 15% or more lower than found for other tests of the same or typical 3/4 inch diameter tubing. Upon review of the data, the EPRI APC Committee .
concluded that the burst data are not reliable and should not be included in the APC database.
The following describes the destructive examination results for the 1991 Plant R-1 pulled tubes:
Tube R50112. TSP 3 Plant R 1 pulled tube RSC112. TSP 3 had a field bobbin voltage Indication of 1.82 volts and a post-pull bobbin voltage of 5.06 volts. By destructive exam, the maximum corrosion depth (at 1 of 23 axial grinds) was 97% in the laboratory, the indication was found to leak at about 500 psi while pressurlzing for a burst test. The tube section was then leak tested at prototypic conditions before further burst testing and found to have leak rates of 0.078 and 0.56 t/hr. at normal operating and SLB pressure differential, respectively. A bladder was then inserted to continue burst testing. The " burst" pressure measured was 4,150 psi.
Figures 3 21 to 3-23 show, respectively, the post burst test crLek, a' map of OD crack indications and the crack depth vs length of the macrocracks that opened during leak and burst testing. Figure 6-3 also shows the location and length (~0.19 inch) of throughwall crack opening following the burst test. From Figure 3-21, it is seen that two post-burst crack openings are separated by a ligament. The lengths of the two throughwall penetrations are about 0.11 inch and 0.08 inch. These lengths are typical of Individual crack initiation rites '
(sometimes called microcracks). Even the end to end opened crack length of about 0.19 inch is much less than the 0.45 inch throughwall crack expected for a burst pressure of 4,150 psi. A completed burst test is characterized by fishmouth opening of the crack and/or ter2 ring at the edges of the corrosion crack. This burst test shows neither of these burst features and did not open up eltner of the macrocracks, . It is concluded that the " burst" test is an incomplete test, it is postulated that a slow pressurization rate permitted the bladder to enter the microcracks as they opened and caused the bladder to tear which terminated the test. As a consequence, the.
- burst" test is not considered reliable and is not included in the voltage / burst correlation data
. base.
A threefold increase in eddy current bobbin voltaga and the appearance of leakage at a pressure of 500 pslin a post-pull test raises questions of damage to tube R50112 prior to leak rate testing.
3- 7
__ __ _ __ __ J
_.__ ___ _ _. _ _ .m._. _ _ .
! l 1
- i i !
t i and the suitability of including this leakage data in leak rate . bobbin voltage. correlations. The j measured crack depth of Flwre 3 23, as obtained from the metallography of the successive axlat - ;
i grinds of Figure 3 22, was used to estimate the pressure at which fracture of the remaining l crack depth ligaments would be expected. The estimated pressure at which ligament fracture and l thus leakage would be expected is about 3300 psi or many times higher than the observed -
pressure of 500 psl. This Indicates the tube was damaged prior to leak rate testing and should ,
l
. not be included in the general leak rate database. The measured SLB leak rate for this Indication i of 0.56 liter /hr would be expected to have a throughwall erack length of ebout 0.1 inch or - -
j l larger. 1 R10089. TSP 3 l This indication had a bobbin voltage of 1.48 volts which increased to 3.31 volts in the post pull >
Inspection. Figure 3 24 shows the burst crack opening aher the burst test and after cutting and bendino of the tube to open the matcrocrack. It is seen that only minor crack opening has i
j occurred and the opened length is very short (0.1 inch). After bending to open the macrocrack, a
! large part of the macrocrack is not throughwall. Figure 3 25 shows that the burst opening represents only part of the approximately 0.37 inch macrocrack length which had a maximum
- depth of 75%. The measured burst pressure of 5000 psils approximately the expecteo burst 1 pressure for a 0.37 inch throughwall crack and much less than expected for an average depth equal to the maximum 75% depth.- it is concluded that the burst test did not result in a complete ,
4 burst. Therefore, this indication is not incuded in the burst pressure data base, i '
R7047. TSP 3 .;
i
!. The 3rd TSP Intersection of R7047 had a 1,57 volt Indication that increased to 4.13 volts in the ;
j post pullinspection. Figure 3 26 shows the burst crack opening for this indication and Figure .
!- 3 27 shows the crack map. The macrocrack associated with the burst opening is about 0.43 1 inch long with a maximum depth of ateout 87% Figure 3 28 shows the crack depth vs length
~
which Indicates an average depth of about 60 to 65% The crack length having depths greater than 70% is <0.1 inch. = The burst pressure for a 0.43 inch long crack with an average depth of -
65% would be expected to be at least 6900 psl compared to the measured 5800 psl. From .
Figure 3 26,it is seen that the burst test resulted in only a minor crack opening of about 0.05 inch which again Indicates an incomplete burst test, and the data point was excluded from the databas3.
R20046. TSPs 2 and 3 Both Intersections of tube R20C46 burst just above the TSP elevation at hand held grinding tool _-
marks. These marks were applied in t_he laboratory for location purposes. Since the burst 3
~
pressures are associated with the grinding marks outside the TSP rather than the degradation within the TSP, these indications are not included in the database.
R10089. TSP 2 -
No detectable bobbin Indication was found in either the field or post !pull inspection for the 2nd ;'
TSP intersection of R10069. The destructive exam also shows no measurable degradation at this <
TSP Intersection.' The burst opening ls centered at the TSP indication and shows a ductile, fishmouth rupture typical of bursts for indications with modest degradation. The measured :
. burst pressure was 9,400 psl.
34 8 a-
+
lilihud r e - g Neir,=vur
- me**a,e.Lr .,a 1,,,,-r,g,,,4--_> gy. yn m -T.q,4rm','I,m u gmebe --a..--m-pg
- g *,-v+'4 gmv V- w Q ga'r'*=F Ft # *n NW C 1D % Sm we W6+ nF-" "-
W
To evaluate the potential need to adjust the measured burst pressure for this type of indication, an undegraded freespan piece of tube R7047 was burst by Westinghouse for comparison with burst of the Plant R.1 freespan tubing as part of the destructive examination program. The Westinghouse test yloided a burst pressure of 11,100 psi which is sitnllar to that found for undegraded model boiler tubing. The freespan burst pressures during the destructive exam program were in the range of 9,400 to 9,900 psi or about 12% lower than the Westinghouse tests. Historically, burst pressures for undegraded 3/4 inch tubing have been in the range of 10,600 to 12,000 psl. The low burst pressures obtained during the destructive exam tests tend to indicate a potential systematic problem in the time frame of these tests. Based on these results, the burst test is not considered reliable and is not included in the database.
B50112. TSP 2 The 2nd TSP of R50112 was called NDD in the field evaluation,0.48 volts by reevaluation of the field data and 0.25 volt for the post pull evaluation. The tube burst at 9,700 psl above the TSP location and thus should correspond to an undegraded tube burst pressure. The expected range of -
burst pressures for undegraded tubing is 10,60012,000 psi. The low measured burst pressure indicates an unreliable data point.
R10CG. TSP 2 The 2no TSP of R1000 had a 1.46 volt bobbin indication which increased to 2,07 volts in the post pullinspection. The burst crack opening is shown in the upper part of Figure 3 29 at two magnifications. The crack opening is about 0.33 Inch long with minor bulging or tearing. Figure 3 30 shows the OD crack map and associated depths. The macrocrack that opened in the burst test is about 0.33 inch long with a maximum depth of 72%. The expected burst pressure for a
. 0.33 inch long crack conservatively sssuming an average depth of 72% would be about 7260 psi or significantly in excess of the measured 6,000 psl. It is concluded that the reported burst pressure underestimates a complete burst by at least 15% and the data point has too many i uncertalntles for including in the APC database.
R10C8. TSP 3 The 3d TSP of R1006 had a 1.31 volt bobbin indication which increased to 5.34 volts in the post pull examination. Figure 3 29 shows the burst crack opening for this indication. The burst opening length is about 0.38 inch with a maximum depth of 85% Similar to the 2nd TSP
, for this tube, a minimum increase of 15% in the measured burst pressure of 4,850 psl would be appropriate for this indication and the data point is not included in the APC database.
Crack Mornholo.gy
, Figures 3 32,3 25,3 27 and 3 30 show available OD crack maps and associated maximum depths found in the tube examination. These figures also show regions on the tube which were characterized in the destructive examination as IGA. The IGA depth was generally negligible
(<5%). However, the 3rd TSP of R7C47 was identified as having very local IGA depths up to the 5175% range as shown in Figure 3 27. The IGA characterization used to define the OD crack maps is not known. A review of the metallography data indicates negligible volumetric IGA Involvement. The Plant R 1 pulled tube crack morphology can be classified as multiple ODSCC with minor IGA.
39
l 3.4 Plant E 4 Corrosion Degradation -
1 Steam generator tubes at support plate crevice regions in the European Plant E 4 have developed
> cellular IGA / SCC. The cellular IGA /SCO is localized in the crevice region such that rnost of the crevice region is free of corrosion. The crevice regions had moderate crack densities, moderate l lGA components associated with Individual major cracks, and no significant IGA Independent ,
cracking. Burst tests conducted produced the expected axlal opening through complex mixtures of axial, circumferential and oblique cracks. For the more strongly affected areas, while the cracking remained multi-directional,there was a predominance of axial cracking. Figures 315 and 310 provide radial section photomicrographs through two of the more strongly 4
affected areas showing cellular IGA / SCC at Plant E 4.
3.5 Plant B 1 Corrosion Degradation A description of the corrosion found at TSP 5 of Plant B 1 is provided below. This region is
- singled out for two reasons. First of all, it has through wall corrosion. Secondly, the tube had a
- small reglon believed to have cellular IGA /SCO.
OD origin, axlally orientated,intergranular stress corros'.on cracks were observed confined .
entirely withln the fifth support plate crevice region on the hot leg side of tube R4 061 from
- Steam Generator C of Plant B 1. Six axial macrocracks were observed around the circumference. The largest of these was examined by SEM fractography without any metallography. The macrocrack was 0.4 inch long and through wall for 0.01 Inch. However, the crack was nearly (effectively) through wall for 0.1 inch. The macrocrack was composed of seven Individual microcracks that had mostly grown together by intergranular corrosion (the 4
separating ledges had Intergranular features that ranged from 40 to 90% of the length of the ,
ledges). Since no metallography was performed on the axial cracks, it is not possible to definitively describe the axial crack morphology at this location. At the eighth support plate region of the same tube, metallography showed that the morphology was that of SCO with a crack depth to IGA width ratio (D/W) of 15. Figure 3 31 summarizes the crack distribution and -
morphology data for the fifth support plate crevice region.
In addition to the OD origin axlat macrocracks observed at the fifth support plate region, one location adjacent to the burst crack had five intergranular circumferential cracks. The 1 maximum penetration observed for the circumferential cracking was 46% through wall. The morphology of the circumferential cracking was more that of IGA patches than of SCO. In addition to the 5 main circumferential cracks, the region had numerous smaller cracks aligned in both the axial and circumferential directions providing a crazed appearance. See Figure 3 32. This crazed degradation is now recognized as probably being cellular IGA /SCO.
Previously the crazed pattern was thought to represent only shallow IGA type degradation that completely disappeared a short distance below the surface. Figure 3 33 provides micrographs of relevant cracks showing the morphology of axial and circumferential cracks. As stated above, the axial cracks had a morphology of IGSCC with a moderate D/W ratio of 15 while the -
circumferential cracking had a morphology more like that of IGA, with a D/W ratio of 1. ,
Field eddy current bobbin probe inspection (in June 1989, just prior to the tube pull) of the fifth support plate crevice region produced a 1.9 volt,74% deep indication in the 550/100 kHz - -
differentialmix.
3 10
f-Table 3-1
- Leak Test and Room Temperature Burst and Tensile Test Results on Plant R-1 SG Tubing .
i Swet Pressure tuttility twret Largest turet Uldtin Tenelle 9.21 Y$ Tenelle U15 Tensite Elens. ;
Location Leek Rete (tthr.)-
- 8-*
(1 Die.) (lnettes) (laches) (pol) (pol) (%) f t
g p RT CT1 F2 . ne teak ust; 14.8 1.054 9.319 j ne leek R S [
RT-CT1 SP3 ne leek WSC; 11.3 S.964 9.MT sv leek me :
aT-CT1 rs 3B.9 1.958 9.37T 53,988 102,300 35.8 [
t 1 af-CT6 w2 no teek moC: 9.9 e.742 e.21e [
no leek as t
e9-CT6 3P3 no leek usC; 18.1 4.728 - 9.2 55 no teek 'Es i
i'
~
E9-CT4 rs 48.8 1.498 S.3er 55,148 106,10e 31.3 k .
R9-C95 SP2 8.852 meC; 13.9 8.611 8.264 -
- 2.=
i 39;C91 ar3 no teak eeC; 11.2 0.861 8.249 i S.825 me af-C91 FS 3T.9 1.729 8.3TZ 53,900 103,ess 29.1 Cement 38.3 1.M5 S.315 52,588 te3,4ee 3T.4 t
j 4 - see - eiwort piste location; rs - fee e,an tenettern
- uet = leek rete et normat aparating easiditlene; Et
- leek rete et seems time treek conditlese
- Tenelle specimen felted at pleg eram pesothly temering the elementless celus.
+ swet pressure is emmpacted to be higher then recorded for thle epoclean. The racer *r era oss easted to be restricted km envenant ef ter thle hasret test such that the era one eleggleh le eswesent. After ed)ustment, no eleller heibowler wee ested.
tihe receeder era eswemmet ese shoorved to be merest for ett other specisano.3 ~ .
Table 3 2
. SEM Fractography Data on Corrosion Present on Plant R 1 SG Tubes
.f M
Location Mar. Deoth Ave. Deoth Macrocrack Ductile Licaments >f (Y, depth) (7. depth) .enath (in.) (number / width, in.)
~ -
R7 C71 SP2 Q (burstcrack)
R7 C71 SP3 (burst crack)
R9 C76 SP2 (burstcrack)
R9-C76 SP3 (burst crack)
R9-C91 SP2 (leakarea&
burstcrack)
~
R9 C91 SP2 (secondary leakarea) .
R9 C91 SP3 (burstcrack)
R9-C91 SP3 (secondary crack)
R9-C91 SP3 (leak area) _
-l g
4 knus 3 12
4 i,
k i
- tube 00 I
tube 10 a
Patch IGA l
9 tube 00 i'
i .
i l
' ~
tube ID I
i Unifom IGA Figure 3-1 Patch-and u transverse niform-IGA morphology as observed in a
-made from a longitudinaltube section. - (A siellar observat1on wou action.)-
St .13
_. . - ~ . ._=..
i 1
i i .
- - tube OD i
1 i
i ,_
- tube ID Branch SCC i
1
- Simple SCC 1
i -
, transverse section schematic 4
i j
tube 00 tube 10
) longitudinal section schematic Figure 3-2 Schematic of simple IGSCC and branch IGSCC. Note that b~ ranch and simple IGSCC are not distinguishable from a longitudinal metallographic section. From a longitudinal section, they also look similar to IGA (See Flijure 4.3).
3 14 ,
tube 00 IGA with IGSCC/!GA gwith
' IGA IG w th IGSCC - IO/ )
(D/h' )
(0/W > 10)
/ tube ID transverse section schematic
'~
-. _., tube 0D tube ID longitudinal section schematic t
Figure 3-3 Schematic of !GA with !GSCC fingers and !GA with IGA
.- fingers. Note that neither of the above variations can be distinguished from a'iongitudinal section. ,
3 15
i i
i
- y ** ; - ; * * ' - v. . .
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?.
8 42
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- 7 g' N.
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4 Mlls Deep . 11 Mils Deep,. , Antal Direction 21 Mlls Deep Mag II.5X Figure 34 Photeetcrographs of radial metallography performed c- a region with axial and circumfereattal degradatles on tube R16-CJ4, support plate 1. Cellular IGA was found with little change to the cell
, shape and cell well thickness at depths of 4.11 and 21 alls below l the 00 surface. IIste that the cut secties was flattened, i preferentially opentag the circumferential well of the cells. .}
l
l i
e d
i i
l
~~~~. .
. , l l ,
t
- g #* l i 's, r s* l j 's, .. , ,9<
- i
! 't %'- l
. 9 :a
.- , -\
l i
Figure 3 5. SEM Micrographs of Tube R36C67 at the First TSP, Axial ODSCC is readily seen.
3 17 I
l L______.._. -____ __ _ _ . . _ , . _ _ _ . . _ . _ _ _ _ _ _ _ _ _______ ___ _ ______ __________ _
i 6 -
Y '
I i55um . j-- 4 25um Figure 3 6. .
Longitudinal Metallographs of Tube R36C67 at the First TSP, Shallow axial ODSCC is seen along with some IGA.
3 18
l l
i i
l I
T. , ,
., A
\
. s
' s. ,*
j
\ .
Y
) '
i 1
a '
V i ,
[ l lI - 1 0 Grn '
l l 25um i
j .
i Figwe 3 7. Transverse Metallographs of Tube R36C67 at the First TSP, Shallow i axial ODSCC is seen along whh some IGA. .!
3 19 i l
'. _ . _ . . - _ . . . . _ . . . - _ . _ _ - . _ . - . , _ ~ , . _ , - . . . - , ,-,_.-_....._.--_,-...-...--.-...__,,--_,.-....~..1,._.-
i I e
-s
- a. *
. ,. . r.
, '4 s
- l
{
.y .
t sN % -
N, 4* ~. . . .
' -, ., s
% '. ~
%, ,\,.,
- L.- -
x . .
w- N , 4 .w
%(
3- ~ -
., N ,
, ~
s -, ~
s Y, * '
'- ( % x T ./
s s . '. ,s v x ,9 mR e,
~. -
4 , .
,'- - k ;' h i ., -
c' q- Q 3,
8
~ ~
.s,.
Km -
~c -'
../
Figure 3 8. SEM Micrographs of Tube R35Cio at the First TSP. IGA and Axial ODSCC are seen.
3 20 i
I l
1 I
I,
\
r
. . . , c y -
- i q' u..7 .
g.-
e l f '
\ (- .
~ '
~~. ,
l -
~
l
.M s . -
,.( ~,
~
~ %; '
r y ' . .*~ l N*s _ *-
%'st
- .k '
X'. ' ,- ,
A
., , y - 4 m / . f
~,
...- m w os ,
x ' ..,
- M '*M . -
~,
s ,
V, . .e .
1 i n 1
l i
l t
i Figure 3 9, SEM Micrographs of Tube R36C67 at the Flow Distribution Baffle, L
3 21- t i
~ } : b s '? ;'j .'~ ~ f 4
' ;p s
, . . 4 ~
' ' '??
,2:- sf"
- ~
.p.T c ,5,
.:. . ~. g.. ,
[h.,' \ g $.-.
.' ~
'5
, w
.n 'i %
.I. ,,
- s. -
.s w . 4-g N gg 1
Figure 310. SEM Micrographs of Tube R35C70 at the Flow Distrbution Baffle.
IGAand 00 SCC are seen.
3 22
Figure 311 -
Norrnal Burst Pressure Curve and R7 071, SP2 with Fitting Leak and with Recorder Malfunction 10/27/92 Burst Test R9 C76 SP 2 (381) Foll350' 1,. . , 1_..
o ,_.
8 ._.......
- c. -
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r fyr '
-mr 1 .
4 +. .
6 f.
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,5 tt
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- p. .,m t
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4Q .. . . .
- r. s.,. .+J
~ .-
a m
mun ..
4 ,
t l
__. . . . . . ..... . . .a .
Time (sec) 10/27/92 Burst Test R7-C71 SP3 (48) Foll Location (190')
Pun # 1 I y
._.6 x ._,y.......-..y'..
- c. -_,
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}. ___.
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f . (p .._ , m.
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,f L-R
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3 i.
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L -
.@ L ,.
3;
' d' ' '
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. 1 . .. .. .:,.
p
. Time sec .
10/27/92 Burst Test R7 C71 SP3 (48) Foil Location (180')
Run #2 c ., ,.mni .. .
Li IOi3';1b,,,a
- =%
_ . -- .._.. _ ...... .. .:t_ .. . iii8a,...
..7
- g. - .,
.. .. . .. , .. .. . .. . i .' $u=1 DOC
_, _fpq1. ____ . .q,y ,
i g # .;&J. ., #. JW
,h.._%++
- z. .
JL..fA,,4
.. .,1
=
g.... _-H.. +. - q t@ $
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i ' .v. y 3 ,
=
. d, .
7
+.y
.( .i.-- y
.. .y,r e,.:
[.h 7 4lh {- .-j .
w
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$, ,,, j$
,m s-e .. .
4 f e ;
1.- ,.,., .
h h fh .4, .
- er,r
- S -
t f ei{ n s ,
.rar y
.itt
- l l t y~
rp, g g i
?!:r r
- Time (sec) 3 23
1 i
l 1, 5 '
- 1. 0 - . Sp yop
- g
.c j
- 4 8
j l 1 a \ \/ .
.2 I l
' l j[ I 'J {
s f f tu ;i4 l
'hO D
4
- 0. 5 - 1d - SP BtAtom -
0 , , ,
l O' 90* 18P 270*- 360' Circumferential Position Idegrees) :
i 4
4 '
Figure 312. Sketch of the crack distribution found at the second support plate . i'
' crevice region of tube R7 C71 from Plant R 1. Included is the -- -
, location of the burst test fracture opening. The OD origin intergranular corrosion was confined to the support plate crevios region, including - l that found on the burst fracture face, i I
3 24-
i i
l.
s 1.3
- 1. 0 - - Sp rop -
g
.c q
- 8 '
1 3:
4 y }f I- k
- h .
, 1
- j, m , u ni -l <
O. 5 - 1 - SP Bottoni 0 , , ,
0*' W 1RP 210' 360' i circumferenilal PosHlon (degrees)'
l Figure 313. Sketch of the crack distribution found at the third support plate
. crevice region of tube R7 071 from Plant R 1.- Included is tho'-
location of the burst test fracture opening. The OD origin intergranular corrosion was confined to the support plate crevice region, including ,
that found on the burst fracture face.
3 2'
- x. -. - - . . . - - ~ . . , . . . , , ,
f
- 1. 25
- 1. 0 -
. Sp g O
U '
ji l h
1$
r -m,
.E
}
i l lI
~
' ]
O. 25 - ~
- SP Bottom 0 ,
O' 90* 110 ' , 210' '360' Circumferential Position (degrees)
Figure 314. Skets of the crack distriteion found at the second support plate crevc' e reolon of tube R9-076 from Plant R 1. Included is the -
location of the burst test fracture opening. The OD origin Intergranular corrosion was confined to the support plate cr6vice .
c rogion, including that found on the burst fracture face.
3 26
- l. 25
- 1. 0 -
- SP iop
'I ,
- p. v e r ._,
- I- } t a -
- 9 e I 9 '
.e t I
I, 1, '(b a}
}
m ll }'
0.5 .
- SP Bottom 0 . .. ,
P 90' lar 210' 3gP Circumferential Posilbn (degrees)
Figure 315. Sketch of the crack distribution found at the third support plate j .- crevice region of tube R9 076 fem Plant R 1. Included is the location of the burst test fracture opening.. The OD origin L Intergranular corrosion was confined to the support plate crevice l
region, including that found on the burst fracture face.
3 27
- 1. 25
- 1. 0 - - SP Top
_. \,
E W
'l I,
/
l
\u\ '
h a ,
s' 1
- 0. 25 - NA' - SP Bottom 0 , , ,
D' 90' 1B' 2iO*. 360*
Circumferential Poshn (degrees)
I 4-Figure 316. Sketch of the crack distrioution found at the second support plate crevice region of tube R9 C91 from Plant R 1. Included is the ..
location of the burst test fracture opening. The 00 ori0 in .
Intergranular corrosion was confined to the support plate crevice region. Including that found on the burst fracture face.
3 28
~
i 1 i i
t
, 1. 25 i
- 1. 0 - - - SP Top '
]
10
-lh \
Yf '/
8 M
J \/i
=
c ,a
,u .
J2 y . u a (
- SP Bottom
. 0. 'd.
.- 0 , , ,
90* 18)* . 270'- 360*
D' _
I Circumferential Position (degreesi 4
Sketch of the crack distribution found at the second support plate
. Figure 317.
crevice region of tube R9-C91 from Plant R.1. Included is the location of the burst test fracture opening. The 00 orhin intergranular corrosion was confined to the support plate crevice .
region, including that found on the burst fracture face. -
3 29
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l Figure S 18. Radial metallography of the mid crevice region of SP3 of tube R9.C76 from Plant R 1, near 300* (from 225' to 340'). The .
Iow magnification (16X) montage shows predomiaantly axial IGSCC.
to the left from 225' to -300* and predominantly ICC to the right from 300' to 340'.
3 30
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Higher magnitcation (100X) photomicrographs of the 100 shown in Fgure 318.
3 31
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l Opened Crack k 0.016' a *'
Thioughwall
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.' // Crack, urface '
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/ l l // / / f,sll N Polished from 00 Estimated Crack Shape Figure 3-20. Through Wall Corrosion for SP3 of R9-C91 from Plant R 1.
3 32
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3 33 i
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81 f ost Burst Throughwall a
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Floure 3 23. Plant R-1 Pulled Tube RS C112 SP 3: Crack Depth Profile 3 35
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3-36
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- SPECIMEN 7-47-4B-2E
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Figure 3-25. Plant R 1 Pulted Tube R10-C69, SP3: Incremental Grind and Polish Results.
.3-37
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e Figure 3 26.
Plant R 1 Pulled Tuta R7C47, SP3: Crack Opening After Burst Test.
3-38
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l l Figure 3-27. Plant R-1 Pulled Tube R7 C47, SP3: Incremental Grind and Polish Results.
3-39 l
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Figure 3-28. Plant R 1 Tube R7C47, SP 3: Crack Depth Profile 3-40 t
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3-41 1-----
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! Figure 3-30, ' Plant R 1 Pulled Tube R10-C6, SP2: incremental Grind and Polish Results.
3 42
. _ . - - _ _ _ _ .. . _ - . _ . . . . ~ . _ -- . . _ _ _
4 l .
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f, 10 3
Sketch of Eurst Cenek i
- Macrocrack; Length'= 0.4 inch '
Throughwall Length = 0.01 inch I
i i- Number of Microcracks . 7 (all: ligaments have predominantly ,
J intergranular- features) - _
Morphology .' IG3CC with some IGA aspects (circumferential cracking 4
- has more.!GA characteristics) l . ,
1 b
- i 1
0.75 inches - - SP top
. I I
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17'08 08 908-- 1808
(-- $katch of Crack Distribution -
p
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L-Figure 3-31.
Description of OD corrosion at the fifth support plate crevice region of tube R4 C61 from Plant B-1.
3 43-q
- a l1
Axial Direction 4
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i t Figure 3 32.
Network of small. rnostly circumferentially oriented OD surface cracks
( observed in the fifth support plate region of tube R4-C61 from Plant B-1.
I I
3-44 i
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3-45
Section 4 LABORATORY SPECIMEN PREPARATION AND TESTING 4.1 Preparation of Specimens Cracked tube specimens were produced in the Forest Hills Single Tube Model Boller test facil!ty.
The facility consisted of thirteen pressure vessels in which a forced flow primary system transfered heat to a natural circulation secondary system. Appropriate test specimens were placed around a single heat transfer tube to simulate steam generator tube support plates. The tests were conducted in two boiler configurations, shown schematically in Figures 41 and 4 2.
The majority of the tests were conducted in the vertically oriented boilers shown in Figure 41, in which four support plates were typically mounted on the tube. A few tests were conducted in horizontally mounted bollers, shown in Figure 4 2 Because the' <as no steam space in the horizontal boilers, seven support plates could be mounted on the ., eat transfer tube. Since capillary forces, rather than gravity forces, dictate the flow pattem in packed tube support plate crevices, the tube orientation should have little effect on the kinetics of the corrosion processes.
The thermal-hydraulic specifications utilized in the test are presented in Table 41. As -
Indicated, the temperatures are representative of those found in PWR steam generators, and the heat flux is typical of that found on the hot leg side of the steam generator. The tests utilized 3/4 inch (1.9 cm) 0.D. mill annealed alloy 600 tubing from heat NX7368. The tubing was manufactured by the Plymouth Tubing Co. to Westinghouse specifications. The chemical and- -
physical properties of the tubing are presented in Table 4-2.
The cracks were produced in what is termed the reference cracking chemistry, consisting of either 600 ppb (1X) or 6 ppm (10X) sodlum as sodium carbonato in the makeup tank.
Typically a test was initiated with the 1X chemistry, and if a through wallleak was not identified after 30 days of operation, the 10X chemistry was applied. The occurrence of primary to secondary leakage was determined by monitoring the boilers for lithium, which would ordinarily only be present in the primary system. Because of hideout in the crevices, the boiler sodium concentration was typically between 50 and 75% of the makeup tank concentration.' Hydrazine and ammonia were also added to the makeup tanks for oxygen and pH control, respectively.
A summary of the test pieces which were subsequently leak and burst tested is presented in Table -
4 3. Two groups of tests are listed; the EPRI test pieces were prepared under (funded by) this program, while the Spanish test pieces were fabricated for a group of Spanish utilities.
Permission from the utilities has been obtained to use the results of these tests in other applications. The only difference between the two groups of tests is that the crevices were packed with different sludge formulations. As in most previous model boiler test programs, the EPRI tests used what is termed simulated plant sludge while the Spanish tests used a formulation more representative of that typically found in steam generators in Spanish plants. As indicated in Table 4-4, the only difference between the two formulations is that magnetite has replaced the
- metallic copper content in the simulated plant sludge.
As outlined in Table 4 3, three means of packing the tube support plMe crevices were utilized.
in the fritted configuration, loose sludge was vibratotily packed into the crevice and then held in place with alloy 600 porous frits placed over both ends of the crevice. In this configuration, cracks were typically produced near the interface between the sludge and the frits. In some -
cases, multiple cracks were produced at both ends of the crevice.
41 - LI I
The dual consolidated configuration consisted of two studge regions, in wh'lch the outer region contained chromic oxide, while the inner region coritained either Simulated plant or Spanish sludge. The regions had the following dimensions, with the distances given in millimeters:
a Chromic oxide -
A 19.0 Simulated Plant Sludge > 6.4 1 r J L 3.8 i k V 4-- 4.1--> 3.2 4-4.1->
1 11.4 > .
' ~ '
The two region sludge configuration was specified in order to limit cracking to the small inner region, containing an oxidizing sludge. Chromic oxide is nonoxidizing, and previous testing had found that accelerated corrosion is less likely to occur in its presence.- The outer region provided thermal insulation for the inner region, so that the temperature in the inner region was sufficiently high to produce accelerated corrosion. The two sludge regions were baked onto the tube using a mixture consisting of 5% sodium hydroxide,2.5% sodium sulfate, and 0.8% sodium silicate. The support plates were then mounted on the tube over the sludge and held in place with externally mounted set screws. Since corrosion should be confined to the inner reglon, this =
configuration was intended to produce short, individual cracks.
The mechanically consolidated sludge configuration was fabricated by mechanically compacting -
sludge within a tube support plate simulant, drilling a hole in the sludge for the tube, and then sliding the tube through the hole until positioned properly. This configuration was used because I
relatively low voltage indications had been produced in previous tests using this configuration.
As indicated in Table 4-3, there was considarable variation in the time taken for a crack to be produced in a given test piece. In general, cracking'was produced in shorter time spans with this ,
heat of material (NX7368) than for the heats used in similar tests performed with 7/8 inch =
diameter tubes. Cracks were typically produced most rapidly with the fritted configuration and .
L most slowly with the dual consolidated configuration, although a few cracks were produced very -
quickly with the dual consolidated configuration. Details of crack networks produced in the model boiler specimens are presented in Section 4.4.
L l
42 l
4.2 Non-destructive Examinatioa (NDE) Results The model boiler specimens were eddy current tested in the laboratory using both the bobbin coil probe and the RPC probe. The bobbin coll voltages were measured in accordance with the analysis guidellnes used for the EPRI APC program. In the case of the 3/4 inch diameter tubing, the bobbin coil results reported here are for the 550/130 kHz mix frequency. The reference calibration was performed with the 20% holes in the ASME standard set to 2.75 volts in the differential mix channel. The RPC test results are for the 550 kHz freaquency with the volt 1ge normalization of 20 volts for the 0.5 inch long through wall EDM (ele 0tric discharge machining) slot in the ASME calibration standard. Table 4 5 presents a summary of the NDE data for the model boiler specimens. The 3/4 inch tubing analysis guidelines used in the evaluation of the laboratory specimens is consistent with the V. C. Summer analysis guidelines describedin Appendix A.
- 4.3 Leak Rate Testing
- The objective of the leak rate tests is to determine the relationship between eddy current characteristics and the leak rates of tubes with stress corrosion cracks. Leak rates at normal operating pressure differentials and under steam line break conditions are both of interest, since leakage limits are imposed under both circumstances. The SLB leak rate data are used to develop 4 a formulation between leak rate and bobbin coil voltage.
Crevice condition is an important factor. Tightly packed or dented crevices er9 expected to significantly impede leakage through cracked tubes. Since denting is readily 9ble by non destructive means while crevice gaps cannot be readily assessed, the empha sced upon f open crevices and dented crevices as the limittag cases.
,_ Leak testing of cracked tubes is accomplished as follows. The ends of the tube arev Jg welded.
One end has a fitting for a supply of ilthlated (2 ppm LI), borated (1200 ppm B) and hydrogenated (1 psla) water to the tube inner diameter. The specimen is placed in an autoclave and brough: to a temperature of 616aF and a pressure of 2250 psi. The pressure on the outer l
diameter is brought to 750 psi. A back pressure regulator on the secondary side maintains the 750 psi pressure. Any leakage from the primary side of the tube tends to increase the secondary pressure because of the superheated conditions. The back pressure regulator then opens, the fluid is released, condensed, collected and measured as a function of time.' This provides the
. measured leak rate. The cooling coil is located prior to the back pressure regulator to prevent overheating and to provide good pressure control. Typical leakage duration is one hour unless .
leak rate is excessive and overheating of the back pressure regulator occurs. Pressure is .
controlled on the primary side of the tube by continuous pumping against another back pressure '
regulator set at 2250 psl. The bypass fluid from this regulator is returned to the makeup tank.
l To simulate steam line break conditions the primary pressure is increased to 3000 psi by a simple adjustment of the back pressure regulator and secondary sida is vented within one to .
, three minutes to a pressure of 350 psl The pressure differential across the tube is thus 2650 psi. Temperature fluctuations settle out in several minutes and the leakage test period lasts for I- approximately 30 minutes.
j l
A summary of leak test results is provided in Table 4-5. Leak rates at normal operating
[ pressure differential and at steam line break conditions were obtained for all specimens. The . (
l steam line break conditions increased the leak rates by about a factor of three compared to 1
4-3
1 normal operating conditions. More variation in this factor can be expected.-Prolonged leak rate testing under operating conditions is expected to lead to lower rates. The increase in the leak j, rate upon transition to accident conditions then becomes more variable.
4.4 Burst Testing Given the assumption that significant' support plate displacements cannot'be excluded under--
accident conditions, burst tests of tubes with stress corrosion cracks are conducted in the free =
- span condition and burst pressure is correlated with bobbin coil voltage. This burst pressure correlation is then applied to determine the voltage amplitude that satisfies the guldelines of Reg.
- Guide 1.121 for tube burst margins.
Burst tests were conducted using an air driven differential piston _ water pump at room -
, temperature. Pressure was recorded as a function of time on an X Y plotter. Seal!ag was
! accomplished by use of a soft plastic bladder. Burst tests of tubes with stress corrosion cracks were done in the free span condition. No foil reenforcement of the sealing bladders was used -
since the crack location which was to dominate the burst behavior was not always readily; apparent. Some of the maximum openings developed during burst testing were not sufficient to -
cause extensive crack tearing and thus represent lower bounds to the burst pressures.' The L openings were large enough in all cases to lead to large leakage.- Burst test results are =
cummarized in Table 4-5. ,
4.5 Destructive Examination I 4.5.1 Objectives The objective of this task is to characterize the size, shape, and morphology of the laboratory . '
j created corrosion in alloy 600 tube specimens which have been leak rate and burst tested. The -
- crack morphology is also to be compared generally to the corrosion morphology observed in tubes pulled from operating power plant steam generators. A summary of the available results--
( for 3/4 inch OD specimens is presented in this section.
i 4.5.2 Examination Methods c lJ Examination methods include visual examinations, macrophotography, light microscopy and/or SEM (scanning electron microscopy) examinations, SEM fractography, and metallography. A i- number of model boiler test specimens were selected for destructive examinations.- Most of these
- were leak and burst tested.
f The specimens were initially. examined visually and with a low power microscope. The burst
- opening and visible cracks around the circumference of the tube.within the tube support plate
' intersection were visually examined and their location in relation to the burst crack noted.~
(When the crack networks were particularly complex, such as_when circumferential- __ _
L components were strongly present, photographs of the crack networks were taken and included in ;
- this report for more complete documentation of the data.)- The major burst crack was thcn L 1 l opened for fractographic observations including crack surface morphologyJcrack length, and .- .
crack depth using SEM." One metallographic cross section of each tube specimen was selected
! containing the majority of secondary cracks within the tube support plate region. The location of.
the cracks within this metallographic cross section was noted, the cracks measured as to their-4 4-
I depth and a crack was photographed to show the typical crack morphology: Note that the on.
metanographic section through each specimen will provide the secondary crack distribution at that location. Secondary cracks at other elevations would not be recorded unless the burst test happened to open the cecondary cracks sufficiently for visual examination to record their location.
, 4.5.3 Destructive Exam: nation Results Tube 5901 The crevice region of tube 590-1 showed only three axial cracks, two of which were through-wall. The longest of the two through wall cracks caused the burst opening. At the tube burst opening, the macrocrack (composed of one microcrack) was 0.275 inch long at the OD and 0.21 inch long at the ID. The crack morphology was IGSCC. A metallographic cross section capturing the three axial cracks is shown by a sketch in Figure 4 3. The crack morphology is shown in a photomicrograph In this figure. The shape of the burst crack and its morphology is described in Figure 4-4 together with the OD crack distribution found in this tube.
Tube 590 2 The crevice region of tube 590-2 had large numbers of axial and circumferential cracks. The cracking was concentrated on one quadrant of the tube's circumference. Photographs of the tube following burst testing are shown in Figures 4 5 and 4-6. The burst fracture occurred in a highly irregular fashion dictated by the axlal and circumferential tube degradation. The burst opening was formed by at least five small cracks which joined partial circumferential cracks to form the irregular overall crack pattern, The macrocrack length due to corrosion measured 0.38 inch at the OD surface and it was through-wall for 0.30 inch. The microcracks and their
~
ligaments had intergranular ligaments and the morphology of the burst crack was that of IGSCC (Figure 4-7) A metallographic cross section through the region with the highest crack density showed a crack distribution as sketched in Figure 4 8. A photomicrograph of two typical secondary cracks is also shown in this figure. They suggest that the cracking is primarily (GSCC with some IGA contributions. Figure 4 9 provides a summary of the overall crack distribution and. summary information regarding the burst crack.
3 Tube 590-3 Rupture in tube 590-3 occurred from a single axial OD origin crack confined to the crevice region. The macrocrack was 0.31 inch long and was through-wall for a length of 0.27 inch.
Only one microcrack could obviously be observed on the macrocrack. The morphology was that of IGSCC. Figure 410 provides summary data regarding the corrosion observed on tube 590-3.
Tube 591 1
- Burst in tube 591 1 occurred from a single, relatively small axial crack which was 0.24 inch long on the OD and 0.18 inch long on the ID. _ While two small axial secondary cracks were observed away from the burst near the bottom of the crevice region, no secondary cracks were observed near the burst opening. However, a metallographic cross'section through the center of the burst opening revealed two additional axial secondary cracks which were located away from the burst._ The location of these cracks in relationship to the burst opening is indicated by a sketch in Figure 4-11. A photomicrograph of one of the secondary cracks is also shown. All-4-5 l
cracks had a morphology of IGSCC. The shape of the main crack and the distribution of cracks are depicted in Figure 412.
Tube 5912 The burst fractu'e in tube 5912 occurred in an area of the crevice region where many small but deep axial cracks were concentrated. The burst created a macrocrack which was 0.21 inch '
long on the OD and it was formed by four smaller microcracks. The crack was through wall for a 1 length of 0.03 inch. The ligaments forming the macrocrack all had ductile features and the morphology of the cracking was that of IGSCC. Figure 413 shows the crack distribution observed by metallography in a circumferential cut through the lower region of the crevice where the crack density was highest. A photomicrograph of one of the cracks is also shown. A sketch describing the shape of the burst macrocrack, as well as the overall distribution of '
secondary cracks within the crevice region as observed by visual examination, is shown in 4 Figure 414.
Tube 591-4 A group of small, deep, OD origin, axial cracks, concentrated in one region of tube 5914 within the crevice region, caused the burst fracture. The irregular shape of the burst opening (Figure 415) was formed by five small microcracks which grew together by intergranular corrosion to form the macrocrack. The morphology of the macrocrack was that of IGSCC. The macrocrack crack was 0.45 inch long and through-wall for 0.35 inch. A metallographic cross section through the center of the burst crack revealed many secondary cracks of considerable depth. The cracks are depicted by a sketch in Figure 416 together with a photomicrograph of the cracking.
Summary data regarding the burst crack and the overall crack distribution are shown in Figure 4 17.
Tube 596-3 , ,
The burst fracture in tube 596-3 occurred from a group of small axial cracks of OD origin.
Four of the deep microcracks joined together during the burst test to form the burst opening macrocrack. The ligaments between the microcracks had only intergranular features and the crack morphology was that of IGSCC. The macrocrack caused by IGSCC was 0.45 inch long on the OD and was through wall for a length of 0.44 inch. A metallographic cross section through the region with the highest density of cracking revealed a crack distribut!on shown by a sketch in Figure 418. A photomicrograph in this figure shows the crack morphology of two of the secondary cracks. A summary of the burst crack data and of the overall crack distribution within the crevice region is shown in Figure 419.
4.5.4 Comparison with Pulled Tube Crack Morphology Mest cf the support plate cracking on pulled steam generator tubes was OD origin, intergranular.
stress corrosion cracking that was axially orientated. Large macrocracks were frequently -
present and were composed of numerous short microcracks (typically < 0.1 inches long) - .
separated by ledges or ligaments. The ledges could have either latergranular or dimple rupture -
features depending on whether or not the microcracks had grown tt gether during plant operation.
Most cracks had minimal to moderate IGA features (minor to moderate D/W ratios) in addition to the overall stress corrosion features. Even when the IGA was present in association with the cracks in significant amounts,it did not dominate over the overall SCC morphology. The numbers of cracks distributed around the circumference at a given elevation within the crevice 46
region varied from a few cracks to typically less than 100. In a few cases, ttfe number of cracks was significantly larger than this,in one case possibly approaching 500. For this situation,
< patches of IGA formed where the cracks were particularly close and the individual cracks had some IGA characteristics. Even for this situation, the axial SCC was still the dominant corrosion morphology as the IGA was typically one third to one half the depth of the IGSCC. in addition, cellular IGA / SCC was occasionstly observed confined to small areas within the crevice region.
Finally, IGA, separate and independent of SCO, has been observed. It is usually present as small isolated patches of IGA, la the few cases where more uniform IGA has been observed, it is q typically shallow and intermittently distributed within support plate crevice regions.
4 The model boller corrosion observed in this investigation was similar to that observed within -
typleal pulled tube support plate crevice locations. Most corrosion was axially orientated IGSCC with negligible to moderate IGA aspects (minor to moderate D/W ratios) in association with the cracking. Some of the model boiler specimens had cracking with almost puro IGSCO,i.e., with no obvious IGA aspects (D/W ratios of 50 or higher), more similar to PWSCC than to the typical OD IGSCC observed within support plate crevice corrosion on pulled tubes. IGA independent of the cracking was not observed in the model boiler specimens. The numbers of cracks at a given 4 elevation was typically less than 20, slmllar to that observed in many of the pulled tubes.
However, only one model boiler specimen had a moderate crack density and none had high crack
. densities as have been occasionally observed in plants. A number of the model boller specimens from the second set of tests conducted in 1991, however, did have very complex crack networks
, that frequently had circumferential cracking in association with the predominant axial cracking.
Some of the complex crack networks may have had cellular IGA / SCC components similar to that occasionally observed in pulled tubes.
4.5.5 Conclusions from Specimen Destructive Examinations
! It is concludcd that the laboratory generated corrMion cracks have the same basic features as
_ support plate crevice corrosion from pulled tubes. The laboratory created specimens freQ! sally had somewhat lower crack densities, but individual cracks usually had similar IGA aspects 4 (minor to moderate D/W ratios). IGA independent of IGSCC was not observed in the rnodel boller e specimens as was sometimes observed in puhed tubes. The observed differences in corrosion morphology between the model boiler speamens and the pulled tubes is not believed to be i
significant. ,
4 4.6 Model Boiler Data Base Summary As described in the above subsections, model boiler specimens have been fabricated and tested to augment the pulled tube database at support plate intersections. 53 laboratory specimens have been prepared using 3/4 inch OD tubing. The specimens were subjected to eddy current
. examination. Degradation at simulated tube support plate Intersections have ranged from NDD to 65 volts in bobbin coil amplitude. All of these specimens have been burst tested, with the ..
results disp %yed in Table 4-5. Specimens with significant degradation (41) have also been leak
. tested. Further, several of the samples were destructively examined to determine degradation characterists and crack morphology. The currently available maximum and through wall crack length data obtained for many of these specimens from the destructive examinadons are listed in Table 4-5. The model boiler database is combined with the pulled tube database and the total used for determining leak rate and burst correlations.
47
. . , - . _ . . _ . _ ~ . - . . - - . . - . . - . ... . . -.-- .
- g. -
y
.-- Table 4' 1~
- Model Boller Thermal and Hydraulic Specifications - ,
[.
F 4,--
Primary loop temperature 327'C (620*F) -
1 .
L Primary loop pressure - 13.8 MPa (2000 psi)c i Primary boller inlet temperature '324'C A 3*C (615'F i S*F) .
[ Primary boiler outlet temperature l 313*C A 3'O (595'F 15'F)
- Secondary Tsat at 6,1 MPa (900 psi) 278'O A 3'O (532*F A S*F)
Steam bleed ~ 0.1f 0.2 t/ day 1 i 3
Blowdown 8 cm3 /min (continuous)-__ '.
I
.16.28 x'104 kcal/m2 hr-
- - Nominal heat flux i
j . (60,000 Blu/ft2.hr) -
i 1 t r
- l
- - . .s p ,
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n
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4 4
4
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48
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- J Table 4 2 :
l.
Chemical and Physical Properties of Tub!ng Material (NX7368) ; .
Chemical Composition (Weight %).
Ni 76.21 Cr 14.87 Fe . 7.98 0 0.04 Mn O.41
- SI 0.30 ,
Cu 0,15 Cb 0.04-Physical Properties
- Ultimate Strength'(KSI)- 109.4 - (744 mPa)
Yield Strength (KSI) 54.2- (368 mPa) .
% Elongation- 37.0.-
Hardness ' 83.- (Rockwell B) - ,
W
'I P
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l
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l
- - , . .-: ,-e ,,,r, , p
Table 4 3 Model Boller Test Spec l men Summary 3/4" Diameter Tubing Specimen Crevice Days in
- A Group Confiouration ,,,,Jmt 590 1 EPRI Frit 8 5?42 EPRI Fait 15 590 3 EPRI Frit 15 590-4 EPRI Frit 19 591 1 EPRI Frit 8 591 2 EPRI Frit 10 591 3 EPRI Frit 21 591-4 EPRI Frit 10 592 1 EPRI Mech. Cons. 138 692 2 EPRI Mech. Cons. 138 592-3 EPPI Mech, Cons, 138 592 4 EPRI Mech. Cons, 138 592 5 EPRI Mech. Cons, 138 592 6 EPRI Mech. Cons, 138 592 7 EPRI Mech. Cons. 138 593-1 EPRI Dual Cons. 133 593 2 EPRI Dual Cons, 133 593 3 EPRI Dual Cons. 133 593 4 EPRI Dual Cons. 133 594 1 EPRI Dual Cons. 85 ,
595 1 EPRI Dual Cons. 34 595-2 EPRI Dual Cons. 84
- 595-3 EPRI Dual Cons, 84 595 4 EPRI Dual Cons. 113 596-1 EPRI Dual Cons. 48 596-2 EPRI Dual Cons. 10 596-3 EPRI Dual Cons. 5 596-4 EPRI Dual Cons. 48 597-1 EPRI Dual Cons. 133 597-2 EPRI Dual Cons. 133 597 3 EPRI Dual Cons. 133 597-4 EPRi Dual Cons. 133 5981 EPRI Mech. Cons. 27 598 2 EPRI Mech. Cons.27-598 3 EPRI Mech. Cons, 27-598-4 EPRI Mech. Cons. 46 .
603 1 EPRI Frit 34 603-2 EPRI Frit 34 603-3 EPRI Frit 34 603 4 EPRI Frit 34 (Continued on next page) 4-10
Table 4 3 (Continued)
Model Boller Test Specimen Summary 3/4" Diameter Tubing Specimen - Crevice Daysin 1D Group Confiauration Test 604 1 EPRI Frit 14 604 2 EPRI Frit 7 604 3 EPRI Frit 22 604-4 EPRI Frit 22 600 1 Spanish Dual Cons. 10 ,
600 2 Spanish Dual Cons. 14 600-3 Spanish Dual Cons. 38 601 1- Spanish Frit 12 601 2 Spanish Frit 12 601 3 Spanish Frit 17 601 4 Spanish Frit 17 601 5 Spanish Frit 17 601 6 Spanish Frit 17 c
w 9
s 4
11
2
- Table 4 4 ;
1 --
j Composition of Sluoge used for Crevice Packing j; Welaht %
Simulated -
Plant. Spanish -
Constituent Studae - Studae Magnetite 59.7 92.2.
P Copper 32.5 1
i= Cupric Oxide . 4.5 4.5 -
j . Nickel Oxide 2.'1 - 2.1 i Chromic Oxide. 1.2 1.2 '
t I
t se*
E -
_4 h'
5 -
E s
i 1.
E i
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4 W W
P 4
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Table 4 5 Leak Rate & Burst Test Results for 3/4 inch OD Laboratory Specimens
- Preliminary .
Bobbin Burst Destructive Exam.
Amplitude tank Rate (1/hr) - Pressure Lanath(Inch)
& Soactmen (vehs) N.O.AP SLB AP '(oan Maximum Thruwall'
_ 9 l
(Continued on next page)
- I
- When crack is not throughwall, maximum depth of penetration is shown.
l'- 4 13 i
L ,
j
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Table 4 5 (Continued) i i
Leak Rate & Burst Test Results for 3/4 Inch OD Laboratory Specimens
' Preliminary Bobbin . Burst Destructive Exam. -
, Amplitude Leak Rate Whr) Pressure - Lenoth (inch)
& Snecimen (volta) N.O.AP SLB AP (oan Maximum Thruwati'
{ ,
4 4
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- When crack is not throughwall, maximum depth of penetration is shown. .
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Figure 4-3. Sketch of a metallographic cross section through the crevice region of tube 590-1. The burst cra$ and two secondary crads were observed. A photomicrograph of a secondary crack is also shown. The crack morphology is that of IGSCC. Mag.100X 1
4 17
~
l 1
a o
00 '
- N- ' -
s N ' !
l '
i .
10 :--
i Sketch of Burst Crack f.. 1 Macrocrack Length . 0.275 inch
- Throughwall L.ength . 0.21 inch a
Number of Microcracks . 3 1
1 1-Morphology .
l 1 IGSCC 1 1 i
l i
i 0.75 inches - .
, -. SP top-
- tt
- 3 O.6 inches -
t i
b l 0.2 inches -
, 0.0 inches - -
SP bottom-i 1
1800~ 2700- 00 900 1800 Sketch of Crack Distribution .
Figure 4 4.
Summary of burst crack observation and the overall crack distribution at .
, the crevice region of tube 5901.
r
' 4-18 4
4
l
's i
)* .' * ,
s l
h ,
g' -
b .
t I i
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4 l
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t
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e .
e * ;
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u i
i i
1 Figure 4 5. Photographs of the burst opening in tube 590 2 showing <:xlal and circumferential cracking, l 4 19 l
1 4
a
,r .' ,, . , ,4 I,Ik- Nk
... . . - m.
I 4R C
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, . -s -
y, .
j '. '
,{ '
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10x Figure 4 6, Photographs away from the burst opening in tube 590 2 showing axial and !
circumferential Cracking.
4 20 i
.-v-=h'" -# * " * * * . . _ _ - - . . - - ~ ~ - . . - - - - - '
e l
I I
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- i I 4 l . . a.
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a t
t
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- l. Figure 4 7. Photomicrograph of the burst crack in tube 590 2 showing IGSCC morphology, r
4 21 I
4m_..-- . - , . . . - , _ . - - - , . --...-.,,,..v.%,,m...wew-,--...--..., ..w- - - . . .,,- . - - - . - - , . . e .-e-,.-... ._ ..
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Crack A .
Figure 4 8. Sketch of a metallographic cross section through the crevice region of tube 590 2. The burst crack and a number of secondary cracks were observed.
l A photomicrograph of two secondary cracks is also shown. The crack .
I morphology is that of IGSCC with some IGA contrbution. Mag 100X 4 22
l
! OD
\ '.' y\f'kU ,
~
10 ..
Sketch of Burst Crack-Macrocrack Length = 0.38 inch .
^
Throughwall Length = 0.30 inch Number of Microcracks - -5 (ligaments have intergranular features)
Morphology - IGSCC -
4 0.75 inches - - SP top h;I l,I 0.6 inches - )
Ig 0.2 inches - ,
0.0 inches - . - SP. bottom 0 00 1800 180 2700 900 Sketch'of Crack Distribution Figure 4 9. ' Summary of burst crack observations and the overall crad. distribution at the crevice region of tube 590 2.
4 23
3 l i
4 OD f,',,\
3 4
s
\s ] >
1 -
10 1
Sketch of Burst Crack e
i Macrocrack Length = 0.31 inch ihtoughwall Length = 0.27 inch ,
Numbar of Microcracks = 1 4
Morphology = IGScC i .
{ 0.75 inches - - SP top 0.6 inches - ; .
1' 4
f O.2 inches -
1 0.0 inches - - SP bottom
~
'1800- 2700 00 900 1800 '
Sketch of Crack Distribution Fgure 410. Summary of burst crack oburvations and the overall crs& distrt>ution at the crevice region of tube 590-3.
f 4 24
~
100%
1 78% (A) l l
l 60% I l
i
+
l l
l i
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I 1
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'.. i. ; 4 !., gu, . - %. f, d.h L . ', s. J , i i Crack A l
Figure 411. Sketch of a metallographic cross section through the crevice region of tube l*
591 1, The burst crack and two secondary cracks on one quarter of the circumference were observed. A photomicrograph of a secondary crack is ;
also shown. The crack morphology is that of IGSCC. Mag.100X 4 25
00
,}
)
10 .'
Sketch of Burst Crack Macrocrack Length = 0.24 inch Throughwall length = 0.18 inch ,
Number of Microcracks = 1 Horphology = IGSCC 0.75 inches - -
SP top
,1 -
0.6 inches - ,
l 0.2 inches -
0.0 inches ' - - SP bottom L
l 1800 2700 0' 908 180' l Sketch of Crack Distribution Figure 412. Summary of burst crack observations and the overall crack distrbution at -
the crevice region of tube 591+1.
4 26 l.
100%
1
, r l 100% d 1
l I
/
85%
/ ,s 80% (A) 85%
... h ?** '
- e. .
.5 .
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- s ..o<2:.. . .~. . l+ .c a .; . .. , ,;.:.$6 . d,7:v . >> < : . . r. e e. ~ .
Crack A Figure 413. Sketch of a metallographic cross section through the crevice region of tube 5912. The burst crack and a number of secondary cracks around the circumference were observed. A photomicrograph of two secondary cracks -
is also shown. The crack morphology is that of IGSCC. Mag.100X 4 27
~
10 - - - -
Sketch of Burst Crack -
Macrocrack Length = 0.21 inch
, Throughwall length . 0.03 inch Number of Microcracks = 4(ligamentshave !
ductile features)
Morphology = IGSCC i
4 0.75 inches - -
SP top
\
)
4 0,6 inches -
i 1
0.2 inches - o i
0.0 inches - -))'d 4 - SP bottom 0
180 2700 00 900 1800 Sketch of Crack Distribution _
- Fyure 414. Summary of burst crack observations and the overall crad distrbution at '
the crevice region of tube 5912, 4 28
mmm--mam--w-----_.=me----=_ - - - -
-w_w- = = - -w------w--- - - - - . w-w .wmm l
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e V's . . . y* . or w' -
E *
- ., a w [- E i I
i l
l l
l .
, t Figure 415. Photographs of the burst opening in tube 5914.
l 4*29 l
. - - . - . . - . _ - . . _ . . . . . . . . - _ _ - - - - - - - - . - - - - . - - . - . . - _ _ _ . _ _ = _. . . - . - - -
i i
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Crack A Figure 416.
Sketch of a metallographic cross section through the crevice region of tube 5914. The burst crack and a number of secondary cracks around the circumference were observed. A photomicrograph of the burst crack and a ,
secondary crack is also shown. The crack morphology is that of IGSCC .
Mag.100X 4 30 i
._- _ - - _ - _ _ _ . _ _ _ _ _ _ - _ _ _ _ _ _ . , _ _ _ . _ _ _ _ . - - _ - - . - . , ~ . . - v _ - , . . . . . . . . _ _ _ . -
^
[ [ \, , -
. io I
Sketch of Burst Crack
. Macrocrack Length = 0.45 inch
. Throughwall length . 0.35 inch i Number of Microcracks a 5 (ligaments.have intergranular features)
Morphology = IGSCC 0.75 inches - - SP top 0.6 inches -
i 0.2 inenes - ,
)
0.0 inches - I k'4lI - SP bottom 1800 2700 00 900 1800 Sketch of Crack Distribution-Figure 417. Summary of burst crack observations and the overall crad distrbution at -
the crevice region of tube 5914.
4 31
P 4
5 100% 30% .
~
i 30% (A) i .
i j
100%
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Crack A Figure 418.
Skeich of a metallographic cross section through the crevice region of tube 596 3. The burst crack and a number of secondary cracks in one quadrant -
, of the circumference were observed. A photomicrograph of two secondary
- cracks is also shown. The crack morphology is that of IGSCC Mag.100X i
4 32
00 g'
! , )
'N' \ ,
\ 'Y q .
40 - - . -
. .s Sketch of Burst Crack Hacrocrack Length =
i 0.45 inch ;
I
] Throughwall Length = 0.44 inch
! Number of Microcracks = 4-(11gaments have intergranularfeatures)
- Morphology
- IGSCC i
i
- _ 0.75 inches - -
SP top
. 0.6 inches -
1 1
1(-
0.2 inches - '), I I
0.0 inches - ,
- SP bottom 2'
1800 2700 00 '900 180'
. Sketch of Crack Distribution-l' Figure 419. Summary of burst crack observatbns and the overall crad distribution at l- - the crevice region of tube 596 3.
-4 33 l-
-"- -'-"-- - - W .meM @.MM.sMM f -
--'A4--
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.- e 4+ w w, e,we e w per- .ww m,-.ww ,%.c m ,,, W % ws ,, ,,w w e n, A- 13,...me v .etera carr+r r e **. c e m = e 5 m r e w-
Section 5 -
NON DESTRUCTIVE EXAMINATION (NDE) 5.1 Eddy Current Voltage Normalizatlon for APC Normalization of observed support plate ODSCC signal amplitudes is performed to permit direct comparison of the voltage levels associated with field measurements with the laboratory calibration used to Joln pulled tube and model boiler signal amplitudes, in cases where field data is collected using different voltage calibrations or different frequencies, conversion factors are developed which permit the field data for tubes and ASME standards taken at different voltage normalizations to be integrated into the overall database for voltage, burst pressures and leakage correlations.
The existing data base for pulled tube and model boller samples includes amplitude measurements which are referenced to a common voltage calibration for both 3/4 inch diameter 0.043 inch wall thickness and 7/8 inch diameter 0.050 inch wall thickness tubing. Specifically the bobbin coil reference calibrations for 3/4 inch x 0.043 inch tubing are:
4.00 volts at 550 kHz for 4 x 20% ASME holes, and 2.75 volts in the $50/130 kHz support plate suppression mix output, also for 4 x 20%
ASME holes The frequency and voltage normalizations applied for the 3/4 inch diameter APC data base, including all model boiler specimens, were developed based on scaling from the 7/8 inch diameter practices (400/100 kHz frequency mix). The 3/4 inch and 7/8 inch tube diameters are geometrically similar, that is, all linear dimensions are scaled by the same factor. Eddy current probe dimensions are scaled by the same factor applied to the 0.720 inch probe diameter for 7/8 Inch tubing to obtain =0.615 inch, which has been rounded to a 0.610 inch probe diameter for 3/4 inch tubing. The probe excitation frequency is inversely proportional to the square of the tubing thickness. This applies to the 550/130 kHz and 400/100 kHz frequencies used for 3/4 inch and 7/8 inch tubing, respectively. Thus the bobbin coll dimensions and test frequencies are intended to yleid rJmilar responses for the 3/4 inch and 7/8 inch tubing.-
However, the ASME calibration standard holes are not scaled and other probe characteristics such as coil size are not scaled so that the resulting 3/4 inch and 7/8 inch tubing eddy current measurements are not direedy comparable.
For the above bobbin coil voltage normalization, the probes tested (Echoram, Zetec) in the laboratory yleided both 4.0 volts at 550 kHz and 2.75 volts at 550/130 kHz. However, testing of other probes with different frequency sensitivity can yield a different ratlo between 550 and 550/130 kHz normalizations. For probes yleiding tiie laboratory ratio be' ween 550 and -
550/130 kHz, the voltage normalization at 550 kHz to 4.0 volts is preferred as it is less
. sensitive to small analyst variations in setting up the mix. However, the voltage normalization to 2.75 volts is more genera'ly app!! cable to different probes and should be used for probes
' differing from the laboratory ratio by more than about 5%. That is, if the voltage is normalizaed to 4.0 volts for the 20% ASME hole at 550 kHz and is outside the range (5%) of 2.6 to 2.9 volts when carried over to the 550/130 kHz mix, the bobbin voltage normalization to 2.75 volts for the mix should be used for the data evaluation. For V. C. Summer voltage measurements the 2.75v mix normalization was the basis employed for sizing the support plate indications.
51 j
5.2 Eddy Current Data Analysis Guldelines The generalinspection protocol for bobbin probe EC testing specified that data be collected at four frequencies 050 kHz,400 kHz,130 kHz and 35 kHz. For V. C. Summer, this has been used in prior inspections and hence data at oddy current frequencies consistent with the APC .
database is directly available without any renormalization. The eddy current inspection guidelines to be app!Ied for the upcoming 1993 outage are described in Appendix A. Thus the
- V. C. Summer support plate amplitude measurements used for repalt lirnit disposition will be taken as described in Appendix A of this report. in a fashion consistent with the " Appendix A" guidelines presented in the APC submittals for Plant A and Plant D.
5.3 Voltage Trends for EDM Slots in order to anticipate the behavior (bobbin amplitude response) of cracks, EDM slots of varying depth and length were prepared for 3/4 inch tubing. As with the 7/8 inch data for EDM slots, the NDE measurements were made according to the EPRI study guldelines on which the Appendix A guldelines from the Plant A submittal were based. For 3/4 inch tubing, the support plate mix (550/130 kHz) data obtained using a 610 mil bobbin probe were evaluated to determine the .
peak to peak voltage values for each notch. These data are displayed in Figure 51. The trends apparent in these data are virtually identical to those collected with a 720 mil probe from the 400 kHz/100 kHz mix channel for 7/8 inch tubing (Figure 5 2).
Laboreleo has performed an extensive study on the voltage response for machined slots in both 3/4 and 7/8 Inch diameter tubing. Machined defect depths and lengths were accurately measured to minimize the uncertalnties associated with manufacturing tolerances. The data were then correlated by an empirical formulation wnich was found to show excellent agreement between -
calculated and measured voltages. Use of the empirical correlation permits development of data trends and ratlos that minimize the influence of depth or length tolerances. Figure 5-3 shows the ratio of voltage amplitudes for 7/8 inch tubing relative to 3/4 inch tubing as a function of simulated crack length. The crack lengths in this figure are adjusted to equal structural strength -
(3/4" tengths increased by a factor of 1.17) rather than equal 3/4" and 7/8" lengths. The voltages ratio of 7/8 to 3/4 inch tubing sizes is seen to be significantly different between 4
throughwall cracks and partial depth cracks. For partial depth indications, the voltage ratlo has little dependence on crack length while throughwall cracks show an increasing voltage ratio as crack length decreases. A constant ratio, which would simpilfy assessments of combined 7/8 and 3/4 inch diameter data ls not expected based on the Figure 51 data.
5.4 Voltage Renormalization for Alternate Calibrations To increase the supporting data base,11is necessary to renormalize available data to the calibration values used in this report. For data on 3/4 inch diameter tubing, voltage renormalization has been obtained by applying a normalization for the ASME 20% holes of 4.0 volts in the 550 kHz channel and 2.75 volts in the 550/100 kHz mix. The APC voltage - .
normalization and data analysis guidelines have been discussed in Sections 5.1 and 5.2. The voltage renormalization for the Plant R 1 and Belgian pulled tubes are described in Sections 5.5 and 5.6. The Plant R 1 renormalization from 400/100 kHz mix to the 550/130 kHz mix was obtained from post pult laboratory measurements and is shown in Figure 5 4 The Belgian renormalization was obtained by direct measurements of field indications as shown in Figure 52
.- - . - - - = - .. - --.
5 5. As discussed in Section 5.6, the Belglan voltages have been increased by a factor of 1.7 for cross calibration of the Belglan ASME standard to the reference laboratory standard.
5.5 Plant R 1 Pulled Tube Data 5.5.1 Plant R 11992 Pulled Tubes Three
- ubes with two TSP Intersections per tube were pulled from Catawba 1 in 1992. These six TSP Intersections have had post puillaboratory NDE, leak rate testing at operating and St.B oonditions, burst testing and destructive examination. This section describes the NDE evaluation of the pulled tube test results for including the data in the APC database. l The field and laboratory evaluations of the pulled tube NDE data are chown in Table 51. The data were collected at the APC 550/130 kHz frequencies and voltage normalization. The field and laboratory reevaluation of the indications are generally !n good agreement although R9091, TSP 2 has a significantly larger voltage for the laboratory evaluation. The field and laboratory bobbin evaluations for this Indication are shown in Figure 5 0. The laboratory evaluation is based on maximum peak to peak voltage for the indication based on guidelines for APC voltage analysis. The field evaluatlan utilized the maximum depth flaw indicaiion for the voltage which results in a somewhat smaller voltage. For consistency with the overall APC database, the laboratory revaluation of the field data are used for the voltage amplitudes. The ASME standards used in the field data collection were cross calibrated to the reference laboratory standard. The cross calibration corrections are given in Table 51 and applied to the evaluated data to obtain the final voltages as given in Table 51, The post pull NDE data are also shown in Table 51 The post pull bobbin and RPC voltages increased significantly (factors of 2 3) over the pre pullvalues for R9076, TSP 3 and both TSP Intersections of R9091. It is suspected that the increased voltages are associated with crack ligament tearing from the tube pulled operations.
The field and laboratory re evaluation of the field data tapes show good ngreement for bobbin voltages as seen in Table 51. The largest voltage difference was fourd for R9091, TSP 2. The field and laboratory bobbin voltage c <aluations (Mix 1:5) for this indio2 tion are shown in Figure 5 6. It is seen that the laboratory eva!.iation applies the peak to pcak voltage guldeline of Appendix A, while the field evaluation measured the voltage for the deepest phase angle rather than total flaw peak to-peak.
5.5.2 Plant R 11991 Tube Pulls Tubes which were pulled from Plant R 1 steam generators in 1991 and in earl er Insp6ctions were field examined using the 400/100 kHz mix, with the bobbin probe, calibrated on the basis of a carbon steel support simulator (ring) on an ASME standard tube yielding 5.0 volts at 400
. kHz. To include this Information in the 3/4 inch tubing databate, the field EC data was '
recalibrated to the 2.75 volt APC normalization for the 4 X 20% hole on the ASME standard-
- In addition, the post pull data provided by B&W (Babcock & Wilcox) on Plant R 1 tubes was used to develop renormalization ratios from the field to the APC normalization. Post pull laboratory data were obtained for 550/130 kHz mix witn the APC normalization at 2.75 volts (Table 5-2). The post pull voltages are much hlgher than pre pull voltages and thus are not 53
used to support the APC development. However, the post pull data are used to develop the ,
conversion f actors for renormalizing the 400/100 kHz field data to the 550/130 kHz :
normalization. Using the correlation of the 550/130 kHz mix to the 400/100 kHz mix when both evaluations are independently normalized to 2.75 volts for tne 20% ASME hole (Figure 5 4), one obtains: .
APC volts (550/130 kHz) - 1.094*(400/100 kHz volts) + 0.143 (52)
The pre pull Plant R 1 voltages were converted to the APC normalization using this equatlon.
TSP voltage measurements by the field normalization of 5.0 volts yleids the voltage adjustn ent factor given in Table 5 2 for obtaining the 20% ASME hole normalization (2.75 volts for ,
400/100 kHz mix). This adjustment factor is applied to the field evaluation with TSP '
normalization as shown in the field evaluation columns of Table 5 3 to obtain the field voltages for the 400/100 kHz mix normalized to 2.75 volts for the 20% ASME hoh. The Westinghouse -
' evaluation for the 400/100 kHz mix is also shown in Table 5 3. The agreement is generally ,
- better than 15% between the field and Westinghouse evaluations.
i The voltage renormalization f actors were then applied to the pre pull 400/100 kHz voltages of .
Table 5 3 to obtain the APC normalization voltages also given in Table 5 3. The Westinghouse evaluated voltages are used for the APC development although differences from the field -
evaluation are small. Also shown in Table 5 3 are the Westinghouse evaluated RPC voltages based on evaluation of the available field data at 300 kHz with normalization to 20 volts for a 0.5 -
inch long EDM notch. The field RPC voltages were normalized to 10 volts for the ASME holes and are not directly comparable to the APC voltage normalization.
Comparison of the pre pull voltages of Table 5 3 with the post pull voltages of Table 5 2 shows that the bobbin voltages for the larger voltage indications increased by f actors of 1.5 to 4 as a result of the tube pull operations. The four largest voltage indications, which show increases of factors of -2.4 to ~4.4, are associated with the lowest four burst pressures for the Plant R.1 pulled tubes.
4 5.5 Belglan Pulled Tube Data The 3/4" tube database is significantly expanded by inclusion of tube pull and burst test data produced by Laboreleo from Plant E 4 in Belgium, in support of the industry effort to develop alternate plugging criterla for support plate ODSCO, Laborelee has collected field data using both Belgian and APC voltage calibrations on U.S. testing equipment (MlZ 18) as well as Belglan -
equipment; this data has included several pulled tubes among -57 Indications evaluated.
For six of eight pulled tubes for which burst and/or leak rate data are available, the Plant E-4 -.
eddy current data were collected for the APC voltage normalization as well as the Belgian voltage normalization. For the APC 550/130 kHz data. Zetec equipment was used to obtain the data. For the Belgian 300 kHz data, both Zetec and Belglan equipment were used to obtain the data. In 54
l addition the Delgian ACME calibration standards were cross calibrated to the reference !
laboratory standard. Table 5 4 Summarlzes the NDE results for the Plant E 4 pulled tubes. l Two tubes, R19035 and R26047, pulled in 1991 were inspected only with the Belglan equipment at 300 kHz. Voltage renormalization for these indications is described Mter in this section.
l The Zetec bobbin coli data tapes for 550/130 kHz were independently evaluated by Westinghouse 4 for 53 TSP Intersections. Figure 5 7 shows the correlation between the Westinghouse and Belglan (Laborelec) voltage evaluations. It is seen that both evaluations are in excellent agreement so that olther Westinghouse or Belglan data analyses may be used for APC applications. ;
Where Westinghouse evaluations are available, they are used as the reference voltage amp!!tudes to enhance genere! consistency with the other APO data. Where not available, the Laboreleo evaluations are used. Table 5 4 includes the Belglan and Westinghouse evaluations for the pulled tubes.
Evalustions of the Belgian ASME calibration data were performed to assess potential differences between Belglan and domestic ASME calibration standards and probes. Table 5 5 tummarizes results of a Westinghouse assessment of the Belglan field data for the Belglan ASME standard and compares the results with values obtained using domectic ttandards and probes. For 550/130 kHz results normalized to 2.75 volts for the 20% ASME hole,it is seen that the Belglan standard / probe leads to lower voltages for the remaining holes including the Belgian 41.25 mm throughwall holes. This leads to a ratlo of about 9.30 for the Echoram probe versus about 5.16 for the Belgian standard / probe when the 550/130 kHz APC normalization is compared to the 300 kHz Belgian normalization. To further evaluate th's difference, a domestle ASME l calibration standard and an Echoram probe were provided to Laboreleo for both 3/4 inch and 7/8
- inch tubing. Cross callbration results for the 3/4 inch U.S. transfer standard are t"wn in Table 5 5 for which the transfer standard yleided 2.68 volts for the 20% hole corresponding to 2.75 volts for the laboratory standard.
F Laborelec evaluated the differences between Belglan and U.S. standards / probes using the Belglan
! manufactured laboratory ASME standard, the U.S. mcnufactured transfer standard cross
- calibrated to the reference laboratory standard, a Belglan probe snd an Echoram probe. Holes in
! the Belgian ASME standard were obtained by EDM while the U.S. standards are drilled. Holes i sizes for the 20% ASME and 1.25 mm holes were measured and found to be in close agreement such as to minimize tolerance effects on the measurements. The manufacta,ing process for the standards and the influence of probe design were separately evaluated. Results of the Laborelec evaluation are given in Table 5 6. Results for the manufacturing process compare measurements for the domestic drilled hole standards with the Belgian EDM hole standards.
-Voltages were normalized using the Belgian standard as applied for field measurements and then compared to voltage measurements for the U.S. drilled standard. From Table 5 6,it is seen that ratios of EDM to drilled hole voltages for throughwall holes (1.25 mm holes) are approximately unity for 3/4 inch tubing and 1.061.10 for 7/8 inch tubing. However, for the 20% ASME holes, the EDM/ drilled ratio is 1.763 for 3/4 inch tubing,550/130 kHz and 1.447 for 7/8 -
Inch tubing,400/100 kHz. The 550/130 kHz ratio factor of 1.763 represents a direct
" adjustment factor to the Plant E 4 pulled tube voltage measurements obtained at this mix with a .
Belglan ASME standard. To further check this ratio, additional Belglan standards were cross calibrated against each other and found to be in excellent agreement (within tight EDM
. tolerances). For example, the field ASME standard used for the Plant E 4 measurements was cross calibrated against the Belgian laboratory standard and found to have a cross calibration factor of 0.986.
55 l:
I
Laborelee also evaluated the influence of probe design on ratios between the same $1mulated defect and between frequencies using the U.S. calibration standard. Results are also given in j Table 5 6. For 3/4 inch tubing, the differences between probes are small (-5%) for 20%
holes and even smaller for throughwall holes. These differences are typleal of 1
probe to probe variations of < a same manufacturer or between domestic manufacturers and can .
be ignored as a correction fo APC data. These variations are included in the APC database for model boller and pulled tube data for which the small probe to probe differences contribute to the spread of the burst and leak rate data. For 7/8 inch tubing, the differences between Echoram and Labore!co probes at 240 kHz are more significant ( 20%) while small at 400/100 kHz. The factors for adjusting 240 kHz to 400/100 kHz are continuing to be evaluated including plant data obtained by Laborelec.
i Based on the above results, the not adjustment factor of the Plant E 4 550/130 kHz data for cross calibration of ASME standards is obtained as follows:
Cross calibration of 1.aborelee laboratory 1.763 standard to U.S. transfer standard.
Cross calibration of Plant E-4 standard 0.980 to Belgian laboratory standard.
Cross calibration of U.S. transfer standard MIS to reference AFC laboratory standard.
Net cross calibration factor 1.70 Thus the Plant E 4 pulled tube voltages measured at 550/130 kHz with the Belgian ASME -
standard need to be increased by a factor of 1.70, The adjusted or reference voltages for inclusion in the APC database are given in the last column of Table 5-4.
9
^
The Plant E 4 field measurements at 550/130 kHz (APC normalization) and at 300 kHz
. (Belgian normalization) can be applied to obtain a general correlation for renormalization of Belglan dala (3/4 inch tubing) to the APC normalization. Figure 5 5 shows the correlation obtained for the voltage renormalization based on Zetec equipment for the 550/130 kHz data (adjusted for cross calibration of ASME standards) and Belglan equipment for the 300 kHz data.
When Zetec equipment is used for both measurements, the correlation has less spread than that of Figure 5 5. The slope of 8.39 for large voltages is slmllar to the ratio obtained for throughwall defects in Table 5 5 while the low voltage slope of 4.2 was feathered into the 1
correlation based on consistency with the 20 40% ASME hele data of Table 5 5. The renormalization factors thus increase with increasing voltagh The correlation of Figure 5 5 is -
to be applied for incorporating Belgian data at 300 kHz into tie APC database in Table 5 4, pulled tube voltages for two intersections (R19C35, R26C47) are renormalized using the correlation of Figure 5 5.
^
5.7 NDE Uncertalntles for V. C. Summer 5.7.1 General Approach for APC -
The usualIndustry practice with respect to NDE uncertainty is based on the adequacy of a sizing model which relates the measured NDE parameter (e 0. depth from phase angle or amplitude for 56
EO testing) to the true value as determined from metallographic examinellorrof representative specimens, actual or simulated, it has been shown that unique interpretations of depth from bobbin amplitude signals are not to be expected and that depth as measured from phase angle is l not an adequate predictor of the structural capability of a tube. The need to relate measured NJE parameters to structural adequacy has resulted in the subject amp!Itude (voltage) based relationship with burst pressure as a predictor of structural adequacy. This approach is based on the relationship between amp!!tude and volume of tubing affected by degradation, a well founded dependency which predicts that as the tube condition becomes more extensively degraded the EO signal response in volts becomes larger; concurrently the more extensively degraded the tube becomes, the less capable the tube becomes with respect to the internal pressure it can withstand before burst.
Thus for NDE uncertalnty the focus la placed on standards and measurement repeatability. Since all the measursments must be referenced to a known condition, the Industry practice of using ASME standards is the comerstone of the APO practice. To minimize effects of the variability of standards, each particular ASME tubing standard used to calibrate the field NDE responses is cross calibrated to the ASME standard used in the EPRI laboratory study. Thus each standard is constralned to produce measurements which are directly comparable to those produced from each plant using the same size tubing. To assess the effects of probe construction differences on amplitude measurements, the EPRI study compared bobbin probes manufactured by Zetec and Echoram, finding them essentla!!y equivalent for the purpose. Additionally Westinghouse has compared the responses of a number (12) of production probes built by Echoram_on the same standard. li was found that the variability of the responses in the support plate mix channel was less than 5%. Eddy current system . cabilng, instrumentation, etc. variability arising from nolse is of the order of 0.1 volt at the calibration used for field measurernents; this is essentially negilglble compared to other sources of error for applications to plugging limits of the order of ,
one or more volts.
Special concern attends measurement variability arising from wear of the probes' centering devices. Excessive play may result in off center positioning of the probe relative to the flaws which affect the EO response. Thus a new probe with design centering produces the proper response, while the same probe with worn centering devices may lean away from the flaw or toward it producing smaller or larger amplitude responses.- To reduce thls variability, limits _ ,
are placed on the usage of an otherwise electrically sound probe; each probe is required to give <
amplitudes no greater than 115% different at any time from the new probe responses to four identical,100% deep holes staggered axlally on a standard tube (" probe wear standard").
Periodio measurement of the probe wear standard identifies when the probe centering is inadequate and replacement is required. A maximum allowance of 15% is provided in the NDE uncertainty.
Data analysis guidelines for voltage measurements are provided in EO sizing guidelines,in Appendix A. It has been found through experlence at Plants A 2 and L that, when given a common orientation to specific measurement guidelines, the variability arising from analysts' differences are reduced to less than 110% (90% cumulative probability). As expected, this
- l. uncertalnty is larger at low voltages; this results from the lower signal to noise ratio. As the S/N value increases, measurement variability diminishes with the result that for APO plugging -
limit voltages the overall average is a conservative correction.
To conservatively represent responses to a given morphological condition, the measurements :
used for the voltage / burst pressure correlation are taken from the field, pre pullinspection data. It has been observed on many occasions that there are unpredictable differences in 1:
57 1 -. - . . . .
amplitudes of flaw signals between pre pull and post pull inspection data. This results from mechanical deformation of the tube, such as elongation, denting, scratching, etc. which occurs in the process of removal.
The contributions to the NDE uncer'ainty at the 90% cumulative probabil;ty are calculated for .
each of the error sources. These sources are treated as independent variables and combined as a square root sum of squares (SRSS) to obtain the net NDE uncertainty. This value is then applied in the calculation of the tube plugging voltage limit. For probabilistic SLB leak rate evaluation, the cumulative probability distribution or a normal distribution of NDE uncertainty-is utilized.
5.7.2 V. C. Summer NDE Uncertainties i The EC uncertainty consists of the EC analyst variablilty and the probe wear contribution. For the 1993 V. C. Summer inspection, these are developed as described below:
EC Analyst Variability The most extensive evaluation for the EC analyst uncertrainty was performed at Plant L Figures 5 8 and 5 9 show the indications and analyst uncertainty from the Plant L study. At 90%
cumulative probability, the EC analyst uncertainty is 10%. The uncertainty in percent represents the voltage difference from Figure 5 9 divided by the mean voltage of 1.41 volts. An upper limit of 20% on the analyst variability uncertainty results from plant specific guidelines for resolving voltage differences between analysts.
For V. C. Summer, the 1991 Indications and the associated 1990 indications for developing growth were reanalyzed using guldelines consistent with APC requirements as described in -
Section 5.2. Similar guidelines will be applied in the 1993 Inspection, as described in Appendix A. Henceit reasonable to apply the Plant L EC analyst uncertainty for the 1993 V. C. Summer '
Indications. Thus the EC analyst variability can be represented as the cumulative distribution function of Figure 5 0. This uncertainty has been applied for the V. C. Summer analyses.
Probe Wear Uncertalnty Figures 510 and 511 show the database on voltage sensitMty to probe weat. For plants implementing the probe wear standard, the voltage variability of Figure 511 is obtained from the Figure 510 data by including all data to 20 mil radial wear for the Echoram probe and to 5 mils for the Zetec probe. The resulting probe wear uncertainty has a standard deviation of 7%,
based on the most limiting (Echoram probe in vertical tube) standard deviation of Figure 511 dMded by the average voltage measurement of Figure 510.
The probe wear standard will be implemented during the 1993 V. C. Summer inspection. Since Zetec probes will be used for the 1993 V. C. Summer inspection, the data for the Zetec probe (bottom figure) of Figure 511 are applicable to V. C. Summer. Mockup tests with the probe wear standard have shown that at 0.0075 inch wear, the wear standard requires probe ,
replacement for 90% of the tests and only the data up to 5 mils wear was used for the EC uncertainty of 7% With the probe wear standard, the probe wear uncertainty is cut off at 15%
by the probe wear replacement requirement. -
58
Combined EC Uncertaintv .
The probe wear and EC analyst NDE uncertaintles can be considered to be independent variables.
For Monte Carlo analyses to obtain EOC voltages, separate distributions can be used and independently sampled for the two contributions to the NDE uncertainty, For deterministic analyses of tube integrity, the EC uncertainties at 90% and 99% cumulative probability are required. The independent uncertainties can be combined as square root sum of squares
. (SRSS). The uncertainty distributions are a 7% standard deviation for probe wear with the distribution cut off at 15% and the analyst variability of Figure 511. At 90% cumulative probability, the combined uncertainty is 14% based on 9% for probe wear and 10% for analyst variability The combined standard deviation is 11%. The analyst variability is cut off at 20%
based on resolution of large voltage differences between analysts and the probe wear cutoff is 15% based on the probe wear standard. The combined uncertainty is 25%, which can be applied for 99% cumulative probability analyses.
1 4
I 4
i e
4 59
. - - . . . - - o , ,w'. ,. .- w - - -.r .. .
.',_ . . , , .y-
1 i
Table 5-1 >
Eddy Cwront Data for Plant R-11992 Puned Tebes i
LabEvaluationot Field Data Post pus Evaluation Field Evaluation Caillwation Bobtnn Coil IFC Bobinn Coil FFC Bobben Cosi FFC ,
Tube TSP Correction Volts & Volts (1) Volts Volts (2) Depth Volts 0.610V 0.630V Depth Volts R7C71 2- 1.033 0.96 87 % 0.14 0.8 0.83 DI 0.14 0.7 0.64 74 % 0.4 SGC 3 50415 (4) 1.70 90 % 0.42 1.9 1.96 DI O.42 0.8-2.4 (5) 0.8-2.4 60 % 0.9 R9C76 2 1.033 0.95 49 % 0.44 1.26 1.3 57 % 0.25 1.5 1.8 60 % 0.4 SGC 3 1.40 97 % 0.36 1.45 1.5 90% 0.31 4.2 4.5 46 % 1.12 50415 (4)
R9C91 2 0.995 2.81 88 % 2.4 3.56 3.54 90% 2.44 7.7 8.8 81% 3.7 SGD 3 50418 (4) 1.12 80 % 0.36 1.14 1.13 80% 0.43 2.7 3.4 68 % 1.5 NOTES:
- 1. APC nonnehramon. Fald evaluesson normalization to ASaEE throughwell holes rather then 0.5 inch slot yisided fWM:: volleges about a factor of luo smeRor. '
- 2. sabann voas cease <subraeed to set =ence immoresory standant Values usW for APC appEcahon.
- 3. Voltages for 0.810 and 0.830 inch diernecer bobbin cod pW
- 4. AsasE standeed nwnhor used 1m cheen bobtun date. *
- 5. Range of voltages given as we5 defined New vohage not obtained in bobtnn le 5-10
. - o
Table 5-2 [
Voltsgo Adjustment Factors to Obt In APC Hon...llaatton for 55N130 kHz Mix i
?
i Post Pull 550/130 kHz [
Factor for Adjusting Field and 400/100 kHz Datp(4)
ISE,, Norm. to 20% ASME Norm. 400/100 kHz 550/130 kHz Iuba ISE TSP Volts 0) Artinetmant Faetar(2) . yggg yggg . !
t R5C112 2 6.92 1.38 0.25(5) 0.37 3 4.44- 5.06 R1006 2 6.4 1.28 - 1.82 2.07-3 4J7 5.M
-R10C69 2 6.4(3) 1.28 - NDD.
3 2.92 3.31-R20C46 2 6.04 1.21- 0.59 0.82 3 0.75 .1.04 R7C47 2 7.8 1.56 - ~
, . 3 3.65 4.13 .
blQles:
- 1. Westinghouse measure of standard TSP volts when 20% ABME volts set at 2.75 volts.
- 3. Adequata TSP not available on standard. Assumed same as tube R10C6.
- 4. B&W evaluations of post pull data.
- 5. The 40W100 kHz data were renormalizoo to 2.75 volts for the 20% ASME hole. -
5 11 ,
.-~_-__.__._.u_.---_-__
-A_.-,-.~._.-.._.__--_ .-_-- .--.-_w._.---._ - _ , _ - - - - - _ u -_.w_- . _ . _ - - . _ - - - _ - - - _ _ _ _ _ _ - . _ . . _ . _ . _ . - - . - _ - - - -_____-_-_.____.a__.__.-__ _. _ _ . . - - - -
T,tble53 .
Fleid and Westinghouse Evaluations of Plant R 1 Prea. pull Voltages .
4 .
ll 4
i
, Flau Evahintbn Wantinahnuam Evaheathn -
)
) 40CV100 kHz Mir $50/130 kHz 400/100 kHz 550/130 kHz i
- 20% Hole l
20% Hole 20% Hole 20% Hole RaC i Lha ISP. TSP Norm. ASME Norm. ASME Norm. ASME Norm. ASME Norm. 2011g(1) i i
R50112 2 NDD ~
0.31 0.48(2) 3 1.15 1.59 1.88(2) 1.53 1,82 1.30 R1006 2 0.82 1.05 1.29 1.20 1.46 0.93 3 0.77 0.99 1.23 1.07 1.31 1.20 R10C69 2 NDD ~ ;-
NDD' ~
3 0.93 1.19 1.45 1.22 1.48 0.97 R20046 2 0.31 0.38 0.56 '
0.25 0.42 3 0.40 0.48 0.67 0.59 0.79 R7C47 2 0.33 0.40 0.58 0.34 0.51
! 3 0.80 1.25 1.51 1.30 1.57 1.40
- 9 blatas:
1.
RPC volts at 300 kHz normalized to 20 volts for 0.5* EDM notch.
- 2. OtWained from 400/100 kHz evaluation using Equation 5 2.-
i 9
k
(:
5 12-4
,, , . _ , - . . . - ~. y-, . . , , y.,.
.._ , . . . , , - , ..m.<, . ' . ,c, - -. - . . . _ ,
t Table 5-4 .
Bohlan and Westinghouse Evaluations of Plant E-4 Eddy Current Data
[ , Balalan Field Evaluation _ Westinahauam Evaluation Reference Zetac Fouin. Dalalan Faulo- (3) Voltage,s
~
h M S50/130 kH2(1) 300 kHz(2) , , 100 kH2(2) 550/130 kHr(1) 3QQ,M jg(2) 11Q(,13q,hSg (Volts) (Volts) (Volts) (Depth) (Volts) (Depth) (Volts)
R26C34 3 4.95 1.17 1.33 71 % 5.03 70% 1.11 8.55(7)
R16C31 2 5.75 1.43 1.27 65 % 5.85 6fr 4 1.32 9.55(7) l 3 9.30 2.02 2.25 70% 9.25 72 % 1.95 15.7(7)
R40C47 2 0.17 0.09 NDD(C) - 9.17 41 % 0.09 0.29(7)
R45C54 2 ' 9.57 2.29 2.25 65 % 9.53 69 % 2.21 16.2(7) l 3 0.47 0.22 0.20(d) (6) - 0.83 53 % 0.16 1.41 (7)
R47C66 2 9.28 2.26 2.12 72 % 9.39 69 % 2.13 16.0 (7) -
- 3 1,39 0.53 0.52 40%(4) 1.37 38%- 0.51 2.33(7)
.. 4 0.18 0.41 (7) i R33C96 2 3.57 0.96 1.07 60 % 3.54 67% 0.88 6.02 (7) l l
j R19C35 2 2.27 17.8(8) l
! R26C47- 2 1.54 11.7(8)
HQ!as'
- 1) Voltages e arf % kHz mix normalized to 2.75 volts on 20% ASME hole.
- 2) Voltage i E Mil r, tz normalized to 2.0 volts on 4 throi.1pma8.1.25 mm (0.049*) holes.
- 3) Amplituu rae Ape,; 'nessured by automated signal analysis.
- 4) Manual conse.;on fcw small si9 nal to noise ratio.
l- 5) Signal below automated detection threshold.
- 6) No depth measurement for insuff'cient signal to noise ratio.
l 7) Voltages in parentheses corrected for crosa calibration factor of 1.70 between Belgian made 4ME standard and the reference laboratory standard _
- 8) Voltages renormalized from 300 kHz to 550/130 kHz using conelation of Figure 5-5.
5-13 i
Table 5-5 Ratio of U.S. 550/130 kHz to Bel 94en 300 kHz Frequency ASE ASME 4 TW Borden 4 TW 125 mm Probe Frobe Type 100' Ext. Standard 0.20 0.40 0.60 0.80 1.00 0.03 0.03 0.05 (kHz) l 0.05 550/130 Zetec Mon eag. No Lab 2.75 2.62 4.13 4.54 4.17 550 3.93 3.14 18.10 4.17 4.17 3.50 300 0.89 15.70 0.65 0.69 0.66 0.52 2.00 550/130 Echoram Mag.-blas Yes Lab 2.75 2.51 4.15 4.81 4.44 550 3.86 3.01 4.18 4.
300 3.72 0.68 0.51 0.62 0 0.51 2.00 550/130 Echoram Meg.-blas Yes Transfer 2.68 3.18 4.23 5.w 5.51 550 3.82 18.60 3.82 4.Si 5.22 300 4.69 10.10 0.62 0.60 0.65 0.72 0 63 2.00 550/130 Belgian Yee Belgian 2.75 2.50 550 3.17 3.63 3.09 3.57 2.91 10.32 3.31 3.49 2.90 300 0.94 0.72 9.58 0.79 0.77 0.59 5 2.00 Ratio 550r130 (APC) to 300 (Belgian)
Zetec Non-meg. I No Lab 3.09 4.03 5.99 6.88 8.02 9.0,5 Echoram Mag.-bias Yse Lab 4.04 4.92 6.69 7.76 8.71 Echoram Mag.-blas Yes Transfer 4.32 5.30 6.50 7.82 8.75 9.30 Belgian Yes Belgian 2.93 3.47 4.01 4.71 5.24 5.16 5-14 1
0 8 9 9
, Table 5 .
Laborelec Results for Renormellaation of Belgian to U.S. Volta 3 3/4" Tubina Volts- 7/8" Tubina-Volts '
j, item 300 KHz 550/130 KH 240 KHz 400/100 KHz .
i A. Manufacturina Proces's: 4 TW 1.25 m Holes ,
! U. S. Drilled 1.99 10.58 1.88- 9.26 P Belgian'EDM 2.00 -10.56 2.00 10,172 .
l Ratio EDM/ Drilled 1.005 0.998 1.064 -1.098. -;
i Manufacturino Frocess: 4 20% ASME Holes i
L U.LS. Drilled- 0.63- .l.56~
0.71 1.90:
l Belgian EDM 0.94 2.75 0.94- -2.75~
- Ratio EDM/Orilled 1.492- 1.763- 1.324 1,447 j B. Influence of Probe Desian
- U. S. ASME Std.
J
! 20% Holes (U.S. Standard) 4 l Echoram 0.61(1)'2.76(2) --o,94(1) 2.77(2)
Laborelec 0.64 2.74 0.76- :2.74 j'
4 Ratio APC to Belgian Normalization at 20% Depth l . _. o Echoram 4.52 2.95 -*
- o Laborelec 4.28 -3.611
- 4 TW 1.'25 m Holes (U.S. Standard) ,
Echoram 2.00 18.52 2.00 .11.10 l Laborelec- 2.01 '.18.77 2.00 13.11 Ratio APC to-Belgian' Normalization at 100% Depth o Echoram 9.26 :5.55-
[
-o Laborelec- 9. 3 4 :.
6.56'-
i
-Notes: 1) Normalized to-2.0" volts for-4-TW-1.25 mm holes.
- 2) _ Normalized to'2.75' volts-for 4-20% ASME drilled holes-
, (APC Calibration).
, 5-15
+
r- l i
b Figure 51 Bobbin Coll Voltage Dependence on Slot Length and Depth + 0.75" Tubing E
0.75 INCH OD 43 MIL THICK _ALLOYeoo TUBE-
- 100. '
g 5- i m.
a en .
E
' d'
> 10: .
p :
, A A -.
g .
j-- .
1 -X a a-1- E O D o
[ .
a
- a * *
. O .
O 810 INCH DIAMETER BOB 8W CO!L AT 550/130 kh:: -
' 0.1 . i- . i i i i i 0,1 .
O 0.2 - 0.3 0. 4 - 0.5 = 0.6 - <
0.7 i
0,8 .- 0.9 L l
- AXIAL SLOT LENGTH, INCH .-
i 4-
. E. THRUWALL & 80% CEEP -- I ~ $0% CEE* - '
D, 50% C55? K 40% C55? ,
+-
3 W
9 :
e 5 18 i
M +
Figure 5-2 Bobbin Coll Voltage Dependence on Slot Length and Depth . 0.875" Tubing 100.,
=
. J B
- X E Rectangular Axial Siet t
. X E g ,
o
$ *E
> 10- '
W g ~ 3L
- g L'y.
a f 7 Tapered Axial Slet- Tute Wall l
E Z
- 2 E h1 o
3 - O O O O O .O m -
O
- l. :
C7:D 'NCM C'.AMETIM BCBBIN CO.t AT 4CIy1C $2 1
0.1. . . . . . . . . .
0 0.1 - 0.2 0.3 0.4 0.5 0.6 0.7 0.8- 0.9 -1 SLOT LENGTH, INCH E THRU WAU. A 80% DEE? . Z 80% DEE?
O 50% DEEP X TAPERED 9
-5 17
d 1.111 -
i 1.094 - -
4 1
i 1.677 -
- a f
I a2n I'h y 1.080 --
EI w > 100 %
w $ -
3 3 1.043 --
80 %
1.026 --
- j. ~
60 % T :
4
, 1.000 l l l l l l l l l l_ l 0 2 4 5 6 8 10 12 14 16- 18 20 0.2" 0.4' Crook Length (mm)
Equivalence factor = ratio of signal amphtudes between 74* 00 and 3/4" 00 tubes with axial flaws of same depth and prpu.al kngth (equal stuctural strength) -
Figure 5-3 Ratio of Voltage Amplitudes for 7/8 inch Tubing RelatWe 2 3/4 inch Tubind as a Function of Simulated Crack Length l.
5-18 i .
Correlation 550/130 to 400/100 KHz mix 1 6
5- Unear Regression Slope 1.094
.g Intercopt 0.143 y 4 -
R Squared 0.999 I
- ( Std Dev 0.012)
!3-B f2-
.- 1-0 - - . . .
0 1 2 3. -
4 5 6 Volts 40W100 KHz mix--
Figure 5-4. Correlation Between Bobbin Voltages at 550/130 and 400/100 kHz from Plant R 1 PulledTube Data 5 19
Evaluation of 1992 Voltage Ind. at TSPs 30 E .
W c 25- f 3 No. Data Points 45 EC equipment .
550/130 - MlZ18
- 2& 300 - Belgian Q )
E I T "'
2 15- S6 m = 4.2
{ intercept = 0 ange 0 0.3 V m ,
f r
<g 1&
at 300 KHz . , ; Linear Regression E t J. -
. slope e.39 g ,, Intercept -1.27
- 5- " Std Dev. 0.36- -
R Squared 0.927_-
h '
4 0
\ c ,
)
0 0,5 1 1.5 2 2.5 3 300 KHz, Belgian Norm., Belgian Eval.
Fgure 5 5. Correlation Between Bobbin Voltages at 550/130 and 300 kHz from Plant E-4 Data 5 to l
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5e21
4 Evaluation of 1992 Voltage Ind at TSPs '
r d 16-No. Data Points 53 '
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Figure 5-7. Compariso'n of Bobbin Voltages at 550/130 kHz Between Westinghouse and Belgian Evaluations 5-22
PLANT L MEAN OF ALL SIX ANALYSTS 120 1 1 ", __: : : :
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VOLTM3ES l
l Figure 5-8. Distribution of Voltage Indications Used for EC Anahst Variability Evaluatio'n
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VOLTAGE DIFFERENCES Figure 5-9. Distribution of Volta 0e Differences Between IndMdual Analysts and Mean Values '
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Fgure 510.' Bobbin Coll Amplitude Dependence on Probe Wear 5 25 l
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i l
Section 6 -
PULLED TUBE AND FIELD DATA EVALUATION This section identif'es the field experience data from operating steam generators that are utilized in the development of tube plugging criteria for ODSCO at TSPs. The field data utilized include-pulled tube examination results including tubes pulled from V. C. Summer during 1988 and occurrences of tube leakage for ODSCC Indications at support plates. Emphasis for the pulled tube data are placed on bobbin coll voltages, burst pressures and leak rate measurements.
6.1 Utilization of Field Data in Tube Repair Limits Operating steam generator experience represents the preferred source of data for the plugging criteria. Since the available operating data are insufficient to fully define plugging criteria, data developed from laboratory induced ODSCC specimens are used to supplement the field data base.
The field data utilized for the plugging criteria are identified in this section.
The overall approach to the tube plugging criteria is based upon establishing that R.G.1.121 guidelines are satisfied. It is conservatively assumed that the tube to TSP crevices are open and that the TSPs are displaced under accident conditions such that the ODSCC generated within the TSPs becomes free span degradation under accident conditions. Under these assumptions, preventing excessive leakage and tube burst under St.9 conditions is required for plant operability. Tube rupture under normal operating conditk ns is prevented by the constraint provided by the drilled hole TSPs with small tube to TSP clearances (typleally ~16 mil
. diametral clearance for open crevices). For the plugging criteria, however, the R.G.1.121 guidelines for burst margins of 3 times normal operating pressure differentials are applied to define the structural requirements against tube rupture.
In add! tion to providing margins against tube burst, it is necessary to limit SLB leakage to acceptable levels based on FSAR evaluations for radiological consequences under accident -
conditions. Thus SLB leakage models are required for the plugging criterla in addition to tube -
burst data.
Based on the above considerations and the plugging criteria objective of relating tube integrity to NDE measurements, the primary data requirements for the plugging criteria are the correlation of burst pressure capabhlty and SLB leak rates with bobbin coil voltage. For plant operational '
considerations,11is desirable to minimize the potential for operating leakage to avoid forced outages. The field data indicate very low leakage potential at normal operating conditions for -
ODSCC at TSPs even at voltage amplitudes much higher than the plugging limits. European operating experience at much higher voltages than the APC repWr limits of this report have shown negligible operating leakage.
6.2 Summary of Pulled Tube Pata Base -
. The available pulled tube data base for ODSCC at TSPs in Westinghouse steam generators is summarized in Table 6-1 for both 3/4 and 7/8 inch diameter tubing. The number of 7/8 inch pulled tubes is provided as a general comparison with the 3/4 inch data and is not utilized in the =
3/4 inch evaluation of this report. Both tubing sizes have a comparable number of pulled tube l
6-1
intersections although the 7/8 inch tubing has more tube burst data. None of the pulled tubes have been reported as leakers during plant operation. The field eddy current data for all pulled tubes were reviewed for voltage normalization consistent with the standard adopted (see Section 5.1) for the plugging criteria development.
Operating plant leakage experience for ODSCO at TSPs is summarized in Section 6.3. Evaluations of the 3/4 inch diameter, pulled tube burst and leak iate data are given in Sections 6.6 to 6.8.
V. C. Summer pulled tube examination results are described in Section 3.2. The :esults support ODSCC as the dominant degradation mechanism, although the indications were not burst tested and are too small for tube leakage considerations. The most extensive leak rate and burst test data for 3/4 inch diameter pulled tubes are from Plants R 1 and E-4 as described in Sections 6.6 and 6.7.
6.3 Operating Plant Leakage Data for ODSCC at TSPs Table 6-2 summarizes the available information on three suspected tube leaks (3/4 inch -
tubing) attributable to ODSCC at TSPs in operating steam generators. These leakers occurred in !
European plants with two of the suspected leakers occurring at one plant in the same operating cycle, in the latter case, five tubes including the two with Indications at TSPs were suspected of contributing to the operating leakage. Leakage for the two indications at TSPs was obtained by a fluoresceln leak test as no dripping was detected at 500 psi secondary side pressure.
For the Plant B 1 leakage indication, other tubes also contributed to the approximately 63 gpd totalleak rate. Helium leak tests identified other tubes leaking due to PWSCC indications. Using relative helium leak rates as a guide, it was judged that the leak rate for the ODSCC indication was less than 10 gpd. These leakage events indicate that limited operating leakage can occur for .
Indications above about 7.7 volts. No leakage at V. C. Summer has been found that could be -
attributable to ODSCC at TSPs. .
6.4 Voltage Renormalization for A!temate Calibrations To increase the supporting data base, it is necessary to renormallze available data to the calibration values tsed in this report. For data on 3/4 inch diameter tubing, voltage .
renormalization has been obtained by applying a normalization for the ASME 20% holes oi4.0 volts in the 550 kHz channel and 2.75 volts in the 550/100 kHz mix. .The APC voltage normalization and data analysis guidelines have been discussed in Sections 5.1 and 5.2. The voltage renormalization for the Plant R 1 and Belgian pulled tubes are described in Section 5.
Plant R 1 tubes pulled in 1992 were obtained at the APC voltage normalization and 550/130 kHz. The Plant R 11991 pulled tube renormalization from 400/100 kHz mix to the 550/130 kHz mix was obtained from post-pulilaboratory measurements and is shown in Figure 5-4.
Most of the Belgian pulled tubes with burst and leak rate measurements were pulled in 1992 and eddy current data were obtained at the APC voltage normalization and 550/130 kHz. The Belgian
- renormali:ation for the remining indications was obtained by direct measurements of field '
indications as shown in Figure 5-5.
62
6.5 Tensile Property Considerations -
The 3/4 inch diameter model boiler specimens have above average tensile properties while the pulled tube data have both higher and lower tensile properties than average values. The tensile property differences between model boiler and pulled tube data are greater for 3/4 inch tubing than found for 7/8 inch tubing. The 3/4 inch model boiler tubing had above average (6%)
material properties while the 7/8 inch model boiler tubing had properties slightly below average. For the 3/4 inch tubing APC development, all model boiler and pulled tube burst pressure data are renormalized to approximate average tensile properties (150 ksi for Sy+Su) for 3/4 inch tubing as described in U.6 Ation.
Tubing manufacturing data have been utilized to develop mean tensile properties together with the standard deviation and lower 95% tolerance limit at room temperature and 650 0F. These data are given in Table 6-3. Also given in the table are the values for (Sy + Su). An Sy+Su value of 150 ksi (twice the flow stress) at room temperature (WCAP 12522) is used to normalize the measured burst pressures for the model boiler and pulled tube data. The ratio of the 95/95 Lower Tolerance Limit (LTL) flow stress at 650*F to a 75 ksi flow stress at room temperature is utilized to adjust the voltage / burst correlation obtained at room temperature to obtain the operating temperature LTL correlation.
Table 6 3 also includes the tensile properties for the 3/4 inch model boiler specimens and for each of the available pulled tubes. Since burst pressures are proportional to the flow stress, the measured burst pressures are normalized to approximate mean properties by the ratio of the tubing mean (Sy + SU) of 150 ksi (flow stress of 75 ksi) at room temperature to the tube specific (Sy + Su) given in Table 6-3.
6.6 Evaluation of Plant R.1 Pulled Tubes This section describes the evaluation of Plant R-1 tubes pulled in 1992 and in 1991. Bobbin voltages for the Plant R 1 pulled tubes are discussed in Section 5. The burst and leak rate measurements are evaluated and summarized in this section.
6.6.1 Tubes Pulled from Plant R 1 in 1992 Three tubes with six intorsections were pulled from Plant R 1 in 1992. Allintersections were burst tested at room temperature and leak tested at operating conditions as described in this section. Destructive examination results cre given in Section 3.3.1 and NDE data in Section 5.5.
Bobbin coil voltages were measured using the APC normalization at 550/130 kHz and ASME standards were cross calibrated to the reference laboratory standard.
Table 6-4 summarizes the 1992 pulled tube data. Burst pressures were obtained for 5
- intersections. As described in Section 3.3.1, the burst pressure for TSP 3 of R9071 was not reliably obtained due to a malfunction of the pressure recorder during the burst test. All six c intersections were leak tested at operating temperature conditions and the two intersections of R9C91 were found to have small leaks at SLB conditions. The 3rd TSP Intersee.en had a very small 0.023 liter /hr (0.0001 gpm) leak rate. The only throughwall corrosion found through -
. extensive destructive examination of this intersection had a 0.016 inch throughwalllength which had opened to about 0.044 inch following leak and burst testing. This indication had a bobbin voltage of 1.13 volts and represents the lowest voltage found to date for a throughwall 63
crack as well as the lowest voltage indication with measurable leakage, although the leak rate is .
negligibly small.
. TSP 3 of R9C76 and between TSP 1 and TSP 2 of R9091 had significant (0.9% and 1.7%,
respectively) tube elongation resulting from tube pulling operations. The tube pull report .
shows the highest pull force of 2000 lbs occurred as the second TSP Intersection of R9091 3
entered the secondary face of the tubesheet. A maximum pull force of.1500 lbs was applied to
- the other tubes. These results support an expectation that the increases in post pull voltages are a conecquence of damage (ligament tearing) from the tube pulling operations. Ligament tvanng would likely have the greatest influence on measured leak rates at normal operating conditions, influence on SLB leakage is also possible but less conclusive as the related ligament tearing may have occurred at the SLB pressure differentials. For R9076, with measurable SLB leakage, throughwall corrosion was found by destructive examination and the leak rates are included in the database although tube pulling damage may have influenced the leak rates.
The crack morphology was found to be principally axial ODSCC with some local patches of -
cellular corrosion. Based on progressive radial (into tube wall) grinding, the cellular corrosion 4
was found to about 37% depth with deeper penetrations showing only axlal cracks. This pattern of partial depth cellular cracking with deeper axlal cracks has been found at all cellular -
Indications examined by radial metallography, including the Plant E 4 Indications with greater cellular involvement. Allindications were located entirsly within the TSP intersections.
6.6.2 Tubes Pulled from Plant R-1 in 1991 Five tubes were pulled in 1991 and earlier, with 9 TSP intersections destructively examined. -
An assessment of the burst and leak rate measurements is glven in Section 3.3,1. Upon review -
by the EPRI APC Committee,it was concluded that the data should not be included in the burst
, database and the RSC112, TSP Intersection should not be included in the leak rate data. The burst tests resulted in incomplete burst tests and lower than expected values for undegraded tubes. No throughwall corrosion was found for R50112, although leakage was found at 500 psi and it is expected that damage during tube pulling operations resulted in throughwall penetration. The NDE and destructive exam data are summarized in this section for use in probability of leakage assessments and overall data summaries.
Table 6 5 summarizes the Plant R-1,1991 pulled tube results. The NDE data are developed in
. Section 5. Four of the intersections had increases in bobbin voltages by a factor of 2 to 4 between pre-pull and post-pullinspections. None of the Intersections had throughwall-corrosion. Eight of the nine intersections showed no leakage during room temperature pressurization tests, which is consistent with the maximum corrosion depths ranging from insignificant to 85% depth. The crack morphology was found to be multiple axial ODSCC with neglialble volumetric IGA involvement. The destructive exam did not include radial metallography to examine for potential patches of cellular corrosion. All Indications ere entirely within the TSP intersection and nearly centered within'the TSP. ,
- 6.7 - Evaluation of Plant E-4 Data .
Recent (1992) tube pulls from Plant E-4 provide a major contribution to the 3/4 inch tubing burst pressure and leak rate data base. Burst and/or leak rate data were obtained for 7 tubes and 64
. w ,. ~ . . u.,u.,- _ , . . . - . .m.
12 TSP intersections. NDE data for 10 of the intersections on 5 tubes were abtained at the APC 550/130 kHz mix and voltage normalization. The eddy current data were cbtained to the Belglan and APC voltage normalizations to provide the basis, as described in Section 5.6, to convert prior and future Belglan data to the APC data base, la Section 5.6, the results of cross calibration of
, Belgian (EDM holes) and domestic (drilled holes) ASME calibration standards are discussed. The cross calibration factor of.1.7 was applied to the Belglan data fur APC applications.
Leak rate and burst test measurements were performed on the Plant E 4 pulled tubes as summarized in Table 6-6. These data include free span burst and leak rate measurements for bobbin voltages up to -17 volts, which are higher than obtained for other 3/4 !nch pulied tubes.
The Plant E 4 burst tests were performed with a plastic bladder and no foll reinforcement. The burst test results showed tearing, except for tube R26047, and are considered to require no adjustments to burst pressures other than the adjustment for material properties. Tube R26C47 is included as a minimum burst pressure since no tearing occurred. Free span bt ,
pressures were obtained for eight Intarsections with bobbin Indications and one NDD intersection.
The leak rate measurements are aho given in Table 6-6. This table includes tubes R19C35 and R26C47 which had been previously (1991) pulled and examined. Leak rates were measured in free span at room temperature. The Plant E 4 leak rate measurements were made at room temperature at 1450 to 1525 psig and 2400 to 2750 psig for normal operating and SLB d!fferential pressures, respectivel). Laborelee has defined an analytical procedure using measured leak rate dependence on pressure differentials to adjust the room temperature test results to prototypic temperaturer and pressure differentials. The adjustment procedure is described in Appendix C and cor; firmed against more detailed crack models in the CRACKFLO code.
applied to the measured leak ratas to obtain the adjusted leak rates given in Table 6-6. The SLB teak rates are given for a presstxe differential of 2650 psi.
The Plant E 4 pulled tubes have been found to have axlal CDSCC and cellular SCC crack morphologies. The cellular mo!phology involves larger areas than found for the Plant R 1 pulled tubes. SimHar to the Plaat R-1 morphology, the cellular SCC depth is limited; deep or throughwall indications are axlial cracks.
6.8 Evaluation of Plant B-1 Pulled Tubes Bobbin and destructive examination data are available for 16 intersections from Plant B Units 1 and 2 pulled tubes. However, only the 5th TSP intersection of R4 C61, Unit 1 was burst tested and this data point is described in this section. The bobbin data was obtained at a 550/100 kHz mix normalized to 2.75 volts for the mix at the 20% ASME hole. The 550/100 kHz mix is -
sufficiently close to the 550/130 kHz mix of the APC normalization such that no voltage .
adjustment is necessary. The pre-pull field bobbin voltage for this indication was 1,91 volts and the maximum depth was 74%. The post-pull bobbin data was 2.33 volts and 80% depth.
, Tube R4C61 at the 5th TSP was burst tested with no bladder and inside a TSP simulant (0.75 Inch long,0.016 inch diametral gap). No leakage was detected (by loss at pressure) until the crack opened to a large leak rate and loss of pressure at 6750 psi. The initial crack was found by
- - destructive exam to be 0.40 inch long with a 0.01 inch long throughwall penetrationc Given the throughwall penetration and that feak rates were not measured with significant accuracy, this indication is not used in *e APC leak rate database. The post-burst crack had minor opening of the crack faces with negligible tearing at the edges of the crack. The maximum change in tube 65
diameter as a result of the burst test was 1.3% OD or about 0.010 inch which is less than the 0.016 inch diametral clearance in the simulated TSP. Thus there is no apparent influence of the TSP on the leak / burst test such that the data point can be used as a lower bound to the burst pressure.
ODSCC with negligible IGA involvement. Short circumferential branch Indications show more -
IGA invotvement at the faces of the cracks, The largest axial macrocrack was examined by SEM fractography and found to be 0.4 inch long with 0.01 inch throughwall penetration. The crack was nearly throughwall for a 0.1 inch length. Seven Individual microcracks compris1r:g the macrocrack had mostly grown together by corrosion with only partially uncorroded ligaments remaining. The maxirnum depth found in the circumferential branching cracks was 46%
throughwall.
6.9 Growth Rate Trends For implementation of alternate plugging criteria (APC) in the range of 2.5 4.0 volt repair .
limits, growth rate dependence on BOC voltage amplitude becomes important to establish the repair limits. This results as current domestic plugging limit!. result in little data in the h,gher range of voltage amplitudes near the APC repair limits. For the V. C. Summer interim plugging.
criteria (IPC) limit of 1,0 volt, the growth rate data developed in Section 9.3 do not require any .
extrapolation to higher BOC voltages, it may be noted that the BOC voltages for V.'C. Summer 4
over the last cycle exceeded the 0.0 to 1.0 volt range, in several cases. _It is desirable to compare the V. C. Summer average growth rate trends with other domestic plants and with European plants. This comparison is provided to show that the V. C. Summer growth data are comparable to -
other domestic plants for percentage growth with a trend for percentage growth to decrease with increasing BOC voltages. French and Belgian plants, which operate with higher voltage indications in service due to differences in ph gging criteria, tend to show less dependonce of percentage growth on BOC amplitudes, Available French data (Plant H-1) indicate percent growth rates nearly independent of initial amplitude (WCAP-12871). Belgian growth data from Plant E 4 have,not been evaluated for percentage growth although the trends appear similar to the French units. For Plant E 4 BOC '
amplitudes in the range of about 0.5 (0.1 volt Belgian) to about 3.7 volts (0.6 volt Belgian), the.
average growth increase in amplitude was about 3.5 volts (0.57 volt Belgian). No strong trend
. of growth dependence on initial amplitude was found although a linear fit to the broad scatter oi growth data indicate a trend for the change in voltage to increase with amplitude. ' Overall, the European plants operate with higher voltage amplitudes in service and with trends toward higher growth rates than domestic plants.
The V. C. Summer percentage growth trends (developed in Section 9.3) are compared with other
' domestic plants and French Plant H-1 in Figure 6-1. This figure shows that the V. C. Summer growth rates are comparable in magnitude and dependence on BOC amplitude with other domestic
- plants. Plant R 1 has 3/4 inch diameter tubing, while the other domestic plants shown in Figure 6-1 have 7/8 inch diameter tubing'.
. ' ,l I
'l 66 l
6.10 Summary of Pulled Tube Test Results Based on the above evaluations, the 3!4 inch tube diameter, pulled tube data for application in tube burst and leak rate correlations is summarized in Table 6 7. The Plant R 1 data includes 5 burst values and 6 SLB leak rate values (two with >0 leakage). The Belgian Plant E 4 dats provides 6 burst data points and 11 SLB leak rate data points (8 points with >0 leakage).- Leak-rates given in Table 6-7 include adjustments to operating temperatures and a SLB pressure differential of 2050 pst. Burst pressures are Ol ven as measured and as adjusted to Sy+Su-150 ksi, based on the tube dependent tensile properties given in Table 6 3.
The overall pulled tube database having bobbin voltages and destructive examination depths for 3/4 inch tubing is shown in Figure 6-2. For comparisons, the equivalent 7/8 inch data is shown in Figure 6 3 and the 3/4 inch and 7/8 inch data are combined in Figure 6 4. A voltage reduction factor of 1.36 (Section 5.5) was applied to the 7/8 inch data for comparison with the .
I 3/4 inch data. Overall, the data sets are comparable in size and general trends toward higher voltages at increasing depth. The European pulled tube data show a number of pulled tubes in the 10 to 30 volt range.
i 67
Table 61 ,
Number of Pulled Tubes with NDE and Destructive Exam Data Number of intersections Number Burst Leak Tested - DestructNe ElaD1 of Tubes . Tasted MM Eram -- .
3/4 Inch Pulled Tube Data Base Summary R1 8- 5 (10)(1) 2 4 (8)(2) -15 E-4 9 8 8 3 14 B1 1 1 0 0 (1)(2)- S B2 3 0 0 0 11 C2 2 0 0 0 4 Totals 23 14 10 7 (9)(2) 49 7/8 inch Pulled Tube Data Base Summary A1 1 0 0 0 1 A2 4 3 2 1 4 .
D2 7 7 0 2 (5)(2) - 15 L 8 21 0 0 (22)(2) 23 P1 2 2 0 0 (3)(2)_ 3 J1 9 1 2 3- 13 Totals 31 -- 34 4 6 (30)(2) ' 59 80181:
- 1) - Number in parentheses represents number of additional pressurization tests -
performed without complete burst or successful burst test. Data not included in data base, -
- 2) Number in parentheses represents room temperature pressure tests performed _ .,
with no identified leakage at pressure differentials exceeding SLB conditions.
One additional Plant R 1 tube was leak tested but throughwall penetration is likely the result of tube pulling and results are not included in data base.
68
Table 6 2 -
Fleid Experience: Suspected Tube Leakage for ODSCC at TSPs(1) hhhin Cell giant !nnoection vehn (3/4" Tubinal DgD1h Comments g
B.1: Outage following R22C58 suspected leak ,
E 4: Outage following R11087 suspected leak R17C58 Outage following suspected leak j-I Notes:
2)- Field voltages of 1.4 and 4.2 volts, as given in parentheses,' were obtained at 300 kHz with Belgian normalization. Voltages converted to 3/4 inch tubing '-
normalization of this report utilizing Figure 5-5.
Sg-6 9'
~
Table 6 3 Tensile Strength Proporties for 3/4 inch Diameter Tubing Source of Tubina Sv.Yleid Strenath-Knl Su.Ultim. Stranath-Kal - ' Sv+So.Ksl Room Temo. ECOE Room Temp. $E0E - Boom Tomo. SECE - .
Tubing Manufacturing Data:
Mean . 53.05 45.78 101.29 97.35 154.34 143.13 Standard Deviatio.n 4.86 3.91 4.22 3.97- 8.28 7.13 ;
I Lower 95% Tolerance 44.55- 38.95. 93.92 90.40 139.85 130.65-Model Bollac Samples 54.2 -- 109.4- - 163.6 -
i 1
Plant R 1 Pulled Tebes R7C71 . 53.9 -- 102.3 -- 156.2 --
R9076 55.1 -- 104.1 - 159.2 - -
R9C91 53.9 -- 103.6 - 157.5 --
l Plant E 4 Pulled Tubes R26C34 53 49 100- -93 - 153 142 R16C31' - 60 59' 112 108: -172 167- ,
j R40047 46 46 101- 101 147 147 R45054' 54 44 97 22 151 136 R47C66 51 40- 97 91 148 131 .
>- R33C96* 54 44 97- '92 151~ 136 R26047 -148 Plant B-1 Pulled Tubes R4061 52.0 - 101.0 - 153.0 - --
{
O 4
-6 10
l Table 6 ,
Plant R 11992 Pulled Tube Burst' Pressures and Leak Rates 4
1 Bobbin Coil RPC Destructive Exam Leak Rate (1,hr) - _ Burst Row / Col M Mgha Qggth yelta Mar. Depth L80001 Norm. On. SLR Pressure
. - (in.) (psi) -
~
R7C71 2
~0 SGC 3 R9C76 2 SGC 3
- R9C91 2 SGC 3 Notes:
- 1) Average depth given in parentheses. -
- 2) Throughwall depth for ~0.08* In burst crack and also in a second crad.
- 3) Burst crack depth and length.
- 4) Throughwall length of 0.016" found in a crack away from burst crad.
- 5) Burst test not reliable due to pressure recorder malfunction.
e 6 11 1
a.. a. ~ u . sus- ,,~e- .. > . , , as a - a ..na a +..a x..a:... ---. -
4
, Table 6 5 -
Summary of Plant R 11991 Pulled Tube Results Westinghouna Field D4 ta Eval Lab N=*ructive Fram Mn RPC Post, Pull Max. Burst i =ak Rats (Ltd
.IddL ISE V.sha Danth Vohn B.C. Volta Osalb Laosth .anL Norm.Oo. Sia .
- _ g RSC112 2 3
4 e
i R10C6 2 3
e R10C69 2 3
- MO6 2 -
3
. R7C47 3 .
Notes: 1, NA = Not Reliable.
- 2. Evaluation indications crack opening for leakage may have resulted during tube pull,
- 3. No leak identifed during room temperature pressurization tests, i
i
- s e
i 6-12
1 J
Table 8 - -
Plant E 4 Pulled Tube Burst Pressures and Leak Rates
, Bobbin Actual Laak Ratan (thr) Maasured h ISE- bglg(1) Mar. Death Norm.Op.(1305 pan SLB (2650 ps!) Burst Pressures ,
I
- -g R26C34 3 R16C31 2 3
7 R45C54- 2 3
R47C66 2 3
4 l R33C96 2 L
R19C35 2 R26C47 2 Notes: 1. Belglan voltage developed in Section 5.
- 3. Leak rates measured at room temperature with variable pressure differentials and adjusted to operating tergeratures and reference pressure differentials in Appendix C. - Measured leak rates are given in parentheses with adjusted leak rates for APC applications given without parentheses.
.I ..
l=
l 6-13' u
?'
4
- Table 6 7 -
3/4-Inch Diameter Pulled Tube Leak Rate and Burst Presst;re Measurements e
i' N tw _RPC peatnzee Exam Laak RaleflM(1) Burst Pressure (2) '
ElAut RowCol ISg -yggg(3)D58b 2011 Max. Depth Laggb(4) NormalOper. SLE. Ma3L. -AL'. ;
(in.) (psi) (psi)
, _g_ ,;
1 B1 R4C61 5 E-4 R26C34 3
. R16C31 2 3-F R45C54 2 3
! R47C66 2 ,
- 3 4
l R33C96 2 R19C35 2 l R26C47 2 f R1 R7C71 2
.3 R9C76 2 3
. R9C91 2 3
i
, 1.- Leak rates at operating temperature and preeauro differentials of 1300 pel for normal operation and
~L 2650 pai for SLB conditions based on adjustments given in Appendix C.- .
- 2. Measured (Meas.) burst pressure and bur.4 preesuie adjusted (Adj.) to 150 kol for Sy+Su at room temperature.
l 3. Votage normalization for 560/130KHz to 2.75 volts on 20% ASME holes.
- 4. Crack network length for burst crack with through wall crack length given in parentheses.
- 5. Measurement no reliebie (N.R.).
- 6. Leak rates measured at room temperature conditions and analytmally adjusted to operating conditions.
- 7. Not measured at 550/130 KHz. Voltage renormalized from 300 KHz data.'
8, Leak rate at SLB condsons -ed with 0.016 inch throughwaR penetration at a cred location '
separated from the burst crack. .
- 9.' Minimum burst pressure, as no duaile tearing extension c,! burst crack was found after burst test.
~ 10. 00 corrosion extended additional 0.16' above the top of the TSP as microcracks separated by ligaments with -
individual microcred depths in range of 3% (farthest above TSP) to 27% (nearest TSP).-
6 14- ,
80 3
.. ._ -;_ . , U.S. VOLTAGE NORNLIZATiche 70 "
3 o 60 -
0
$ C*
w 50 -
o g_ g ,___
__; g. . o
. __i f,
.O 6 30 -
w e
8 20 -
> Hi- a i y ,
10 - .
l] '
O O 1 2 3 4- 5 6 INITIAL BOBBIN AMPLITUDE, VOLTS o V.C. SUMMER V PLANT R 1. E PLANT A 1 0 - PLANT A 2 0;. PLANT H 1 Figure 61; Average Percent Voltage Growth Rates br V. C. Summer, Plant A, Plant H 1 -
and Plant R 1. V. C. Summer and Plant R 1 have 3/4 inch tubes.-
6*'15
7..
i
.,t t
k
- r
- ~ g e t
I r
r i
t t
'F I
.t
. .i L
i t
- i t
I
.I 1
Figure 6 2. ' 3/4 inch Pulled Tube Data: Bobbin Col Voltage versus :
Maximum Depth from Destructive Examination i 6 16; 4
, ., .[. .[.---s $ .,.... y , % + ,y, ...r'9
i i
1 1
l 4
1
- g 4
I_
1 1
1 a
i l
I J i
,1 ,
i i 4
r
} ,
4 L I \ {
-i e
'. t t
4 i
k t
p 9
0 ,
Y
' f
-?
e F
-3
'~ '
4
&-. 3 3
' Figure 6 3. 7/8 inch Pulled Tube Data: Bobbin Coil Voltage versus - ,.
~
- Maximum Depth from Destructive Examination
- , l l 1 i 6 .: 17,. ;,4 ., ; . .
+
- r. ;;
J
-O__. g.
p y -
,,,~ ,i r a -
^,w-, n , - - , .~-,,v, - , r , , , -
r i
t t
h 9 ,r r
f i
i .
I i
I i
8 4
t L , .*
Figure 6 4, - 3/4 Ind and 7/8 Inch Pulled Tube Data: Bobbin Coll Voltage . .
versus Maximum Depth from Destructive Examination 6.-18
I Section 7 f GUIDEUNES FOR ACCIDENT CONDITION ANALYSES This section develops guidelines for the pressure differential to be used in SLB leak rate ,
and tube burst analyses and the confidence levels or probabilltles to be applied for the :
analyses. Probabilistle safety analyses results are used to assess the frequencies assoclated with the sequence of actions that result in a given primary.to secondary pressure differential for an SLB event. The results of NUREG 0844 analyses together with the event sequence frequencies are used to define guidelines for the confidence L levels to be applied to SLB leak rate analyses and for an acceptable tube rupture probability.
7.1 Limiting Accident Condition The most Ilmiting accident condition for contalnment bypass (potential radiation release to the public) is a steam line break (SLB) In the piping between the containment and the main steamilne isolation valve (MSIV). The SLB ovent also envelopes the feedline break (FLB) event with regard to sequence frequencies as obtained from probabilistic safety analyses.
For most Westinghouse plants, a value of 2600 psid bounds the maximum pressure differential during a limiting secondary system pipe break. This value is determined by applying a maximum uncertainty of 4% of the safety valve setpoint, which includes allowances for setpoint uncertainty, valve accumulation and setpoint shift (setpoint shift of 1% maximum is applied to plants with pressurizer loop seals). For V. C.
Summer, since there is no pressurizer loop seal, the maximum primary pressure would be limited to 2575 psla, including 3% allowance for uncertainty and valve accumulation. Therefore, with atmospheric pressure on the secondary side, the limiting primary to secondary pressure differential for V. C. Summer would be 2560 psla.
Historically, Westinghouse has conservatively bounded the FLB AP at 2650 psl and, for simplicity in analyses, applied this AP also for the SLB ovent. To obtain an upper bound FLB AP. a long term reactor coolant system (RCS) heatup event was postulated with the pressurizer assumed to go water solid during the event. The bounding RCS pressure was obtained assumlng water tellef through the pressurizer safety valves at 110% of design pressure (~2750 psla). The secondary pressure in the faulted SG was 100 psia to >
obtain a 2650 psid across the SG tubes.
A large AP during an SLB is dependent on assuming no operator action to terminate safety injection and failure of the PORVs to open. As described in Section 7.2 below, these assumptions lead to a low probability (-3x10 6) for primary pressure exceeding the PORV relief pressure of 2350 psi and hence a SG tube AP of 2335 psiis the appropriate
,. AP for the SLB analyses, in the case of the FLB, for the everit to become of concern for atmospheric release, it also requires that the check valve in the feedline fail to close; hence the frequency would be further reduced. Typical variations of pressure with time for the SLB and FLB events are described below.
7e 1
_ - - - - ~ - - - - _ - . - . . - - -
, Steam Line Break (SLB) Event Deeerlotion Figure 71 shows a typleal dependence of the primary pressure as a function of time in an SLB event. The primary pressure variation of Figure 71 assumes no operator action on safety injection and that PORVs fall to open. The primary pressure and SG tube .
AP initlally decrease below the normal operating pressure differential (APNO)I0f about the initial 3 minutes of the event. Safety injection then increases primary
- pressure until the plant operator acts to terminate safety injection. If a reduction in safety injection is not initiated within about 20 25 minutes, the primary pressure reaches the PORV pressure setting of 2350 psl. Only if the PORVs fall to open would the i pressure continue to increase to 2575 psi. Thus the typical SLB event would have SG tube AP decrease and tube leakage would not increase above normal operation for about the first 6 mlnutes of the event. Thereafter, tube leakage would increase above that for normal operation as dependent on operator action to terminate safety injection with
- maximum leakage at the PORV rellet pressure of 2350 psl.
1
] Feodline Break IFLBt Event Deterloilga
(
During a postulated feedline break event, it is assumed that the feedwater pipe suffers a '
double ended guillotine break. Whether the break occurs inside or outside of the containment building, the feedwater isolation valves, located outside of the containment, will perform their intended function. Pipe t'reaks upstream of the isolation valve would be modeled as a loss of normal fW9ter since the feedwater isolation valve would restrict the backflow of secondas wpft reedwater line check valves located Inside 4
containment would prevent steat:QM w lth subsequent containment bypass.
Containment bypass can occur only asswung failure of the check valve (to close) and an
- FLB between the isolation va!ve and the containment. The following event description relates to such a scenarlo, i The Model D3 steam generators at V.C. Summer incorporate flow restrictors in the feedwater nozzle inlet. The effective area of the restriction is 0.223 ft2 , and blowdown rates through the break would be much less than for the steam line break event, where a
( larger (1.4 ft2) restrictor Is utillzed in the steam nozzle.' Additionally, the installation of the Model D3 steam generator preheater inlet modification will add to the overall resistance of secondary fluid reversing through the broken feedline.
- At approximately 22 seconds into the event, a turbine trip signal will cause isolation of the main steam lines. Secondary system pressure will rise to the steam generator PORV ,
(steam dump valve) setpoint of 1107 psi, coincident with the Inltlalincrease in
_ primary system pressure. Secondary system PORV operation results in a lower secondary pressure, and therefore a slight rise in AP across the steam generator tubes.
During the initial stages of the accident, primary to secondary pressure differential would not be expected to exceed normal operating values. At about 400 seconds, the ,
faulted loop pressure is assumed to have dropped to atmospheric (assuming a break on -
S/G side of the check valve within containment or failure of the check valve to close) and - .
4he primary system pressure has achieved the PORV setpoint, resulting in a maximum -
AP of 2335 psl.- .
For V.C. Summer, Figure 7 2 shows reactor coolant system pressure and steam pressure as a function of time for the FLB event where offsite power is available. For the 72
-- -e---w- - - N * - d -- y a r
FSAR analysis, loss of PORV function was assumed, and RCS system pressure is eventually elevated to approximately 2500 psi (RCS safety valve setpoint). An initial rise in primary system pressure to approximately 2300 to 2350 psi occurs, and is due to turbine trip. Upon reactor trip, primary system pressure drops to a minimum (for t
the first portion of the event) to approximately 2000 psl at about 120 seconds, and then escalates due to coolant expansion caused by the reduction in heat transfer capability in the steam generators. For the case where offsite power is available, credit for PORV operation can be assumed, in which case RCS pressure will be limited to approximately 2350 pW.
t In the case where offsite power is not available, pressurizer pressure peaks at about 2500 psl. Th6 pressurizer water volume increases due to coolant expansion from the reduced heat transfer capability of the steam generators but does not cause the pressurizer to go water solid. At approximately 1200 seconds, decay heat generation decreases to less than the emergency feedwater heat removal capability, and RCS temperatures and pressures begin to decrease. For the case where offsite power is available, RCS pressure and temperature do not begin to drop until approximately 3300 seconds, due to increased heat input from the reactor coolant pumps.
For considerations of tube rupture for Indications at TSPs in a feedline break event,it can be noted that tube burst would be prevented by the constraint of the TSPs. The SLB event has a rapid loss of SG pressure, resulting in pressure drops across the TSPs and potential TSP displacement to uncover the Indications formed at TSPs under normal operating conditions. The steam pressure changes in an FLB are slow and result In minimal loads to displace the TSPs. Potential offsite dose consequences for the feedline -
break are bounded by the steam line break event. Offsite dose is incurred due to steam rollef from the non faulted loops. Upon implementation of the Interim plugging ;
criteria, the reduction in allowable primary to secondary leakage from 500 gpd to 150 opd leads to lower offsite doses.
Primary Coolant l_enknoe Assumotions Durina Steam Line Break Used in Dose Analysis For a SLB event, negligible leakage would occur until the primary pressure increased again above the normal operation AP. The postulated limiting tube in the SG could then conceptually rupture if safety injection was not terminated. The single tube rupture for the limiting tube would tend to reduce the rate of RCS pressure increase or decrease the pressure. This effect reduces the likelihood of large cumulative leakage and multiple tube ruptures in an SLB event, in contrast, the SLB leakage analyses and burst probability analyses assume all tubes reach the PORV relief pressure of 2350 psl.
Per NUREG 0717, Safety Evaluation Report for Operation of Virgil C. Summer Nuclear Statlon, a continuous primary to secondary leak rate of 1.0 gpm was assumed for the ~
calculation of 0 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> site boundary doses and low population zone entire accident duration dose estimates. Per the V.C. Summer FSAR, primary system pressure drops upon reactor trip, and the pressure differential across the steam generator tubes is not expected to surpass the normal operating value of about 1300 ps!. Upon initiation of :
Safety injection (SI) the reactor coolant system pressure is reestablished. The normal operating AP is re established at approximately 6 minutes into the event. Primary system pressure equal to the pressurizer PORV setpoint is achieved at about 23 minutes, and is only obtalnable if the operator falls to terminate St. Therefore, actual volumes of released coolant could be reduced to as little as 16% of that assumed in NUREG 0717, to 73
as much as 100%, based upon operator action. NUREG-0717 utilizes the guidance of 100FR100, Reactor Site Criteria,in establishing the extent of the protected area. Per 10CFR100, O to 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> doses at the site boundary are limited to 25 Rom whole body, 300 Rem thyrold. These values, however, are not used as acceptable limits for Individual accident scenarlos, Sul establish maximum limits for emergency conditions. .
NUREG 0800, Standard Review Plan, establishes acceptable doses during Design Basis Accidents, based on a percentage of these limits. For application to guidelines for
- IPC/APC SLB analyses, the two hour duration for SLB leakage can occur only if the operator falls to terminate Si and the probability of reaching the allowable leakage limit must include the probability of falling to terminate SI.
The dose estimations are calculated based on primary coolant activity levels of 1.0 micro curies per gram of dose equivalent 1131. An increase in the coolant activity of 500 (lodine spike) is assumed due to the reactor trip. The 1.0 micro curie per gram value is a Technical Specification maximum limit. If this limit were reduced, dose estimates would be reduced by a nearly proportional level. Residual activity in the Jeoondary system due to primary to secondary leakage during normal operation adds to the dose estimate, in the case of the IPC, allowable primary to secondary leakage is reduced from 500 gpd to 150 opd. This reduction will have an overall effect to lower dose estimates in the event of a postulated SLB. Also, the operating history of V.C. Summer has shown that measured coolant activity levels (generally less than 0.1 micro curles per gram) are far below the allowable Technleal Specification limit. These Octors all have an overall effect of lowering actual doses compared to the calculated values. Lowering of the allowable coolant activity, such as through administration controls on operation, would permit an increase in the allowable SG leak rate as measured in gpm.
4 7.2 Event Sequence Probabilities Probabilistic Safety Analysis (PSA) results can be used to assess the probability of SLB event sequences leading to various pressure differentlais across the SG tubes.
V. C. Summer PSA results for an SLB event are given in Table 71 and compared with the assessment given in NUREG 0844. The highest probability sequence (1.8x10 3) includes the operators terminating safety injection, which would result in a AP <2335 psid. As indicated in Figure 71, the SLB pressure differentiat inillally decreases and operator action can be initiated before primary pressure reaches the PORV set point.
The time in the event at which operator action is taken to terminate safety injection determines whether or not the 6P reaches 2335 psld, if the operator falls to terminate safety injection (Sequence 2), the SLB AP remains at 2335 psid but at a much lower probability of about 3.3x10 6/RY. If the PORV falls to open (Sequence 3), the 4P could increase to 2560 psid at an even lower probability of about 2.6x10'7/RY.
7.3 NUREG.0844 Analysis ,
From Table 71,it can be noted that NUREG 0844 conservatively assumed a 10 3/RY probability of a MSLB with a AP of 2000 psid. No operator action was assumed. This is -
a factor of about 3x103higher than the V. C. Summer estimate for obtaining a AP of 2335 psid.
7- 4
The conditlonal probability of a tube rupture at 2000 psid was estimated is 2.5x10 2 in NUREG 0844. Thus, the probability of an SLB event with a consequential tube rupture in NUREG-0844 was 2.5x10 0/RY. This rupture probability is a factor of 10 greater than the V. C. Summer Sequence 2 probability of reaching or maintaining a AP of
- 2335 psid, not considering the probability of a consequential tube rupture event. The NUREG 0844 analyses result in a 2.5x10 8 probability of a core melt in combination with the SLB event and consequential single tube rupture. ThIs core melt frequency is a f actor of 1000 greater than what is calculated for V. C. Summer.
The NUREG 0844 evaluatlons led the NRC staff to conclude that the increment in risk associated with a single tube rupture event is a small fraction of the accident and latent fatality risks to which the general public is routinely exposed (i.e., the associated risk of early and latent fatalities were found to be zero and 1.1x10-5/RY, respectively).
This report Indicates that the 2.5x10 5 probability for an SLB with a tube rupture represents a negligible incremental risk to the public. This probability is greater than the Sequence 2 frequency of 3.3x10 6/RY for a pressure differential of 2335 psid ,
For APC applications, SLB leakage is limited to satisfy a small fraction of 10CFR100 dose over a 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> time frame. Applying average SLB leak rate analyses at 2335 psid yleids a probability of <10 5 of exceeding a small fraction of 10CFR100 limits. This probability is comparable to the probability of a single steam generator tube rupture found in NUREG 0844, which as noted above,is concluded by the NRC staff to result in smallincremental risk to the public. Consequently, it is judged that the use of a 95%
confidence limit on the mean regression fit to SLB leak rate is adequate to determine SLB leakage. Application of the 95% confidence level on the mean regression fit to SLB leak rate yleid a probability of satisfying a small fraction of 100FR100 limits comparable to that which was determined to be acceptably low for the large leakage associated with a single tube rupture event as discussed in NUREG 0844.
7.4 SLB Analysis Guidelines 7.4.1 SLB Pressure Differential Based on the low probability (<10 5) for SLB event sequences leading to a AP of 2335 psid, a 2335 pstd AP is adequate for the SLB leak rate and tube burst analyses in support of alternate plugging criterla. The probability for exceeding SLB leak rate limits or obtaining a tube rupture will be Ngher for the 2335 psid of Sequence 2 of Table 71 than for the 2560 psid of Sequence 3. This results as the differences in leak rates between 2335 and 2560 psid are about 25% while the difference in sequence probabilities is a factor of 10. The contribution of Sequence 3 to leak rate can be ignored as it would contribute only about 10% to the totalleak rate.
7.4.2 Guldelines for Tube Burst Analyses Although the NUREG 0844 analyses show smallincremental risk to the public for a
. ' single tube rupture at e ,nditional probability of 2.5x10 2/RY , the analyses supporting implementation of APC require significant confidence levels to support the NUREG 0844 probabilities. Maintaining the NUREG 0844 tube burst probability in combination with the V. C. Summer Sequence 2 probabilities provides a net probability L
7- 5 l'
i=i
"'- e ==- g a=
~,
h i
of a combined SLB and tube rupture event of a factor of about 10 3 lower than
- NUREG 0844. Thus a 2.5x10*/RY rupture probability represents a conservative
- goal for V. C. Summer. For deterministic analyses of SLB butst margins, the analyses shou;d be performed at a plus 99% uncertainty level. When Monte Carlo analyses are performed for tube burst, the prodleted burst probability should be less than .,
2.5x10 2/RY. The 2.5x10 2 robability p is app!Ied per fuel cycle, which can be 18 months rather than per year. Applying the 2.5x10 2 to APC applications permits the *'
102V.C. Summer margin to cover other potential causes of a rupture such as looso parts, etc.
In applying the R.G.1.121 guldelines for tube repair critoria, tube burst capability at three times normal operating pressure differentials (34Pyo) is found to be most limiting for V. C. Sumrner. Satisfying the 3APno requirement provides margins against burst at SLB conditions. The tube burst correlation is evaluated at the lower 95% '
l predletion interval to provide a high confidence of burst capabllity exceeding 34Pyo.
l The V. C. Summer SLB tube burst margin analyses of Section 12 show that an Indication :
i at the repair limit based on 3APN o (at the 95% interval) results in a much lower probability of burst at SLB conditions than the ~10 2guldeline established above.
7.4.3 Confidence Levels for SLB Leak Rate Analyses The NUREG 0844 analyses and the low Sequence probabilities for V. C. Summer Indicate that a single consequential tube rupture event does not result in significant incremental tir.k to the public. For V. C. Summer, the Sequence 2 probability for reaching or maintaining 2336 psid in an SLB eventis about 3x10*u/RY. Safety injection could be .
terminated by the operators prior to reaching the PORV relief pressure leading to 2335 psid across the tubes, in eithe' case, Fafety injection would be terminated in less than the 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> period used for dosi ana!yses of the leak rate at the $lte boundary. Thus, .
. assuming a constant leak rate at 233G psid is conservative for the leak rate analyses to satisfy the allowable leak limit developed for the o to 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> dose at the site boundary.
The probability of 3x10 6/RY, which assumcs operator failure to terminate Si, for reaching or maintaining 2335 psid is thus also conservative for the O to 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> dose estimate. The use of average values for SLB leak rate analyses at an SLB pressure differential of 2335 psid result in <10 5 probability of radiological consequences which exceed a small fraction of 10CFRt00 limits. These consequences at a 10 5 probability are conservatively bounded by the NUREG 0844 analyses of the consequences of a single tube rupture. Therefore, it is judged that a 95% confidence limit on the mean regression fit for SLB leak rate analyses provides adequate conservatism for event hequence probabilities comparable to the NUREG 0844 SLB
, consequential tube rupture event probability, in this report, the V. C. Summer SLB leak rate calculations are more conservatively performed using deterministic analyses starting with a BOC voltage value and adding .
allowances at 95% cumulative probability for NDE uncertalnty and growth (average per cycle) to obtain a corresponding EOC voltage value. The SLB teak rate is obtained by evaluating the SLB probability of leakage and leak rate correlations at +95% confidence ,
on the mean regression fit at the EOC voltage value, 1
76 l
Table 71 SLB Event Sequence Frequencies and Resulting Pressure Differentials V. C. Summer NUREG.0844 Seauence_1 Soauence 2 Seauence 3 Foquence SLB.1x10 3/RY SLB Outside SLB Outside SLB Outside initiator Containment Containment Containment 1,8x10'3/RY 1.8X10 3/RY 1.8x10 3/RY-Action i IM Operators Operators Fall Operators Fall Terminate Si to Terminate Si to Terminate SI 0.998 2.0x10 3 2.0x10 3 Action 2 tM IM Pressurizer - Pressurtzer PORV Relleves PORV Falls to Pressure Rolleve 0.929 Pressure (1) 7.1x10 2 Total 1X10 3/RY 1.8x10 3/RY 3.3x10 6/RY 2.6x10 7/RY Sequence Frequency Pressure Assumed <2335 psid 233'i psid(2) 2560 psid(3)
Differential 2600 psid Across Tubes Notes:
. 3) Pressurizer safety valve set pressure plus 3% for valve accumulation less atmospheric pressure in SG.-
7 7-
-J
Figure 71 1 RCS Pressure During Steam Line Break r
2600 -
2400 2200 .
$ t vi o.2000
, 7 W
e .
@1800 M PORVS OPEN AT . .
$. Pr = 2360 PSIA :
o- 1600 -
vi o
x .
1400 7 1200
/, :
TIME TO RETURN TO AP(NO) , ,
1000 ;
n 0 500 ~1000- .-1500 2000--
l TIME, SECONDS
-i p
4 p
I 78-
?
. . , , . . . , , . _ . - . , , , . - _ _ _-. ,_. ~ .. _ . . . , , , . - . - - ,. __.s
t Figure 7 2 ,
Main Feedline Rupture With Offslie Power Pressurizer and Steam Generator Pressures vs. Time SOUTH CAROUNA ELECTsuC & GA5 CO. l VIRGIL C. SUMMER NUCLEAR STATION I Mein feedhne RWpture With Offsete Power Preflurifer PffIlute vl. Time E
{ titi. <
26B0. <
m gagg, . lWW b '
Y u tatt. ' fff$Ps*
W 2t28. <
g i8ee. .
3pC 331 , 332 335 Igd T!MC ISCCI SOUTH CAR 0uMA ELECTIUC & gal CO.
VIRGIL C.$UMMER NUCLEAR STATION Mom Feedline Rupture With Offstte Power steem Generetor Pressure vs. Time 5
1548 1 Intact 1258. '
m ri..
! 75.. <
s W $88. '
d 4
g 258. ' '
Faulted w
338 ist -sp2 ist led flMC 15CCI
- 79
h 9
4 4
- 4 e
i 9
9 s
_ ,_m __ . _ _ . _ - - --a---d-------^--"- '----' '" ---
Section 8 -
ACCIDENTCONDITION CONSIDERATIONS This section deals with accident condition loadings and thelt effects relative to the alternate plugging criteria. More specifically, this section considers the Safe Shutdown Earthquake (SSE), Loss of Coolant Accident (LOCA), Foodline Break (FLB), and Steamline Break (SLB)
- events. These events are considered both separately and in combination.
8.1 Tube Deformation Under Combined LOCA + SSE For the combined SSE + LOCA loading condition, the potential exists for yleiding of the tube support plate in the vicinity of the wedge groups, accompanied by deformation of tubes and subsequent loss of flow area and a postulated in leakage. Tube deformation alone, although it impacts the steam generator cooling capability following a LOCA, is small and the increase in PCT L is acceptable. Consequent in leakage, however, may occur if axlat cracks are present and propagate throughwall as tube deformation occurs. This deformation may also lead to opening of pre existing tight through wall cracks, resulting in primary to secondary leakage during the SSE + LOCA event, with consequent in leakage following the event, in leakage is a potential concern, as a small amount of leanage may cause an unacceptablo increase in the core PCT, 8.1.1 SSE Analysis Selsmicloads result from motion of the ground during an earthquake.The SSE excitation of the steam generators is defined in the form of acceleration response spectra at the steam generator supports.To perform the non linear time history analysis,it is necessary to convert the response spectrum input into acceleration time history input. Acceleration time histories for the nonlinear analysis are synthesized from El Centro Earthquake motions, using a frequency suppression / raising technique, such that the resulting spectrum in each of the orthogonal axes closely envelopes the original specified spectrum in the corresponding axis. The resulting three orthogonal time histories of the earthquake are then applied slmultaaeously at each steam generator support to perform the analysis. Plant specific response spectra are used to calculate the TSP loads for V. C. Summer.
The analysis is performed using the WECAN finite element computer program. The mathematical model consists of three dimensionallumped mass, beam, and pipe elements as well as general matrix input to represent the specific steam generator piping stiffnesses. The TSP shell interaction is represented by a rotating, concentric gap spring dynamic element, using impact damping to account for energy dissipation at these locations. The mathematical model with selected node numbers is shown in Figure 81. The primary loop piping and the lower column support stiffnesses are input as 6x6 matrices. The upper and lower lateral support restraints are represented by compression-only (single acting) spring elements, with the shell flexib!!!ty included for the upper support stiffnesses. At the lower elevation, the support
,, structure is connected to a relatively rigid channel head foot combination. In modeling the tube bundle Internals to shell interface, the TSP local shell stiffness, obtained from detailed finite element analyses, is also included. The local shell stiffness at the top TSP location is higher than
- at lower TSP locations because of its proximity to the upper lateral supports.
The tube bundle geometry is shown in Figure 8 2, with the tube baffles and support plates identified. The flow battes (3,4,6,7,9,10) are typically not included in the seismic model, 81 i
k f
both due to the difficulty in representing them accurately in the model, and also because it is ,
conservative in terms of tube stresses to exclude them from the model, if the baffles were included in the inodel,it is anticipated that contact impact loads for the lower plates would be distributed among the various p'ates and baffles resulting in reduced loads. However, it is difficult to estimate these loads, due to the flexible nature of the partition plate whl0h forms one
- of the support members for these plates. For the flow baffles, it is concluded to be conservative
, to use the loads developed for the support plates (2,5,8).
1; Results for V. C. Summer show the flow distribution baffle (plate 1 in Figure 8 2) to not l experlence seismlo impact loads. Thus, it is judged that there will not be any tubes at the flow distribution baffle location that are potentially susceptible to in leakage.
For reasons that will be discussed later, tube deformation calculations are performed for three
8.1.2 LOCA Analysis LOCA loads are developed as a result of transient flow, and temperature and pressure fluctuations
- following a postulated primary coolant pipe break. Based on the prior qualification of the V. C.
Summer steam generators for leak before break requirements for the primary piping, the limiting LOCA event is either the accumulator line break or the pressurtzer surge line break.
However, bounding LOCA load calculations for V. C. Summer for the accumulator or pressurizer
, surge lines are not available. As a conservative approximation, the available LOCA loads for the
- primary piping breaks are used to bound the smaller pipe breaks. The large pipe break loads
- have been shown for other model steam generators to be several times larger than the smaller i
pipe breaks, and thus,it is judged that these loads form a conservative basis for the small pipe .
1 breaks for V. C. Summer. ,
! As a result of a LOCA event, the steam generator tubing is subjected to the following toads: -
, 1) Primary fluid rarefaction wave loads.
- 2) Steam generator shaking loads due to the coolant loop motion.
- 3) External hydrostatic pressure loads as the primary side blows down to atmospheric pressure.
- 4) Bending stresses resulting from bow of the tubesheet due to the secondary to primary pressure drop.
- 5) Bending of the tube due to differential thermal expansion between the tubesheet and first .
tube support plate following the drop in primary fluid temperature.
- 6) Axially induced loads resulting from differential thermal expansion between the tubes ,
and tie rods / spacers due to the tube being tight in the first TSP, and the reduction in primary fluid temperature. (Based on available data, the majority of intersections are considered to be tight. Because the majority of the intersections are tight, the TSP will - .
respond with the tubes, and the resulting loads on the tubes are judged to be small for this loading.)
82
Loading mechanisms (3) through (5) above are not an issue since they are a non-cyclic loading condition and will not result in crack growth, and/or result in a compressive membrane loading on the tube that is beneficialln terms of negating cycIle bending stresses that could result in crack growth.
~
8.1.2.1 LOCA Rarefaction Wave Analysis The principal tube loading during a LOCA is caused by the raref action wave in the primary fluid.
This wave initiates at the postulated break location and travels around the tube U bends. A differential pressure is created across the two legs of the tube which causes an in plane horizontal motion of the U bend. This differential pressure, in tum, induces significant lateral loads on the tubes. The pressure time histories input to the structural analysis are obtained from translent thermal hydraulic (T/H) analyses, using the MULTIFLEX computer code. A break opening time of 1.0 msee of full flow area, simulating an instantaneous double ended rupture is assumed to obtain conservative hydraullo loads. The fluid structure interaction effect due to the flexibility of the divider plate between the inlet and outlet plenums of the primary chamber is included in the analysis. Pressure time histories are calculated for two tube radil, Identified as the average and maximum radius tubes. A plot showing the tube representation in the T/H model is provided in Figure 8 3. Typical primary pressure time histories following a LOCA are shown in Figure 8 4 for nodes 815 (in Figure 8-3) on the cold leg of the largest radius U bend. For-the structural evaluation, the tube loads result from the hot to cold leg AP. Plots showing the hot to cold leg AP for the maximum and average radius tubes are provided in Figures 8 5 and B 6, respeetively.
For the rarefaction wave induced loadings, the predominant motion of the U bends is in the plane of the U Bend. Thus, the Individual tube motions are not coupled by the anti vibration bars.
Also, only the U bend region is subjected to high bending stresses. Therefore, the structural analysis is performed using single tube models limited to the U bend and the straight leg region over the top two TSP's. The node hnd element numbering for a typical single tube modells shown in Figure 8 7.
The tube inodel consists of three dimensional stralght and curved pipe elements. The mass inertla is input as effective material density and includes the weight of the tube, weight of the primary fluid inside the tube and the hydrodynamic mass effects of the secondary fluid. Damping coefficients are defined to realize a maximum damping of 4% at the lowest and highest significant frequencles of the structure.To account for the varying nature of the tube / TSP Interface with increasing tube deflection, three sets of boundary conditions are considered. For the first case, the tube is assumed to be laterally supported at the TSP, but is tree to rotate. This is designated as the ' continuous' condition, in reference to the fact that the finite e'.ement model for this case models the tube down to the second TSP location. As the tube is loaded,it moves laterally and rotates within the TSP. After a finite amount of rotation, the tube will become wedged within the TSP and will no longer be able to rotate. The second set of boundary conditions, Werefore, considers the tube to be fixed at the top TSP location, and is referred to as the ' fixed" case.
Continued tube loading causes the tube to yleid in bending at the top TSP and eventually a plastic -
, hingo develops. This represents the third set of boundary conditions, and is referred to as the
'ptnned" case.
For the average radius tube, only the continuous case is analyzed. Results for the continuous case analysis indicate that both the tube rotations and moments at the TSP nodes are small compared to those required to cause the locking in or plastic hinge, respectively, at the support locations.
Since the main objective in analyzing the average radius tube is to determine the maximum 8+3 o
reaction load on the TSP due to the overall response of the tube bundle, the continuous configuration is the most appropriate for the average radius tube analysis. In addition to the pressure induced bending loads, the raref action wave analys1s also includes the membrane stresses due to the primary to secondary AP.
Each of the dynamic solutions results in a force time history acting on the TSP. These time histories show that the peak responses do not occur at the same time during the transient.
However, it is assumed for this analysis that the maximum reaction forces occur simultaneously. .
Using these t ;ults, a TSP load corresponding to the overall bundle is then calculated. A summary of the resulting TSP forces is provided in Section 8.1.3.
8.1.2.2 LOCA Shaking Loads Concurrent with the rarefacilon wave loading during a LOCA, the tube bundle is subjected to additional bending loads due to the shaking of the steam generator caused by th9 break hydraulics and reactor coolant loop motion. However, the resulting TSP loads from this motion are small compared to those due to the rarefaction wave Induced motion.
To obtain the t X .nduced hydraulle forcing functions, a dynamic blowdown analysts is performed to ot on the system hydraulle forcing functions assuming an instantaneous (1.0 msec break opening ume) double. ended gulliottne break. The hydraulle forcing functions are then applied, along with the displacement time history of the reactor pressure vessel (obtained from a separate reactor vessel blowdown analysis), to a system structural model, which includes the steam generator, the reactor coolant pump and the primary piping. This analysis yleids the time history displacements of the steam generator at its upper lateral and lower support nodes. These time history displacements formulate the forcing functions for obtainlng the tube stresses due to LOCA shaking of the steam generator. .
To evaluate the steam generator response to LOCA shaking loads, the computer code WECAN is used. The model used is similar to the one used for the selsmic analysis, discussed previously.
- The steam generator support elements are removed, however, because the LOCA system model accounts for thelt influence on the steam generator response, input to the WECAN modells in the form of acceleration time histories at the tube /tubesheet interface. These accelerations are obtained by differentiation of the system m; del displacement time histories at this location.
Acceleration time histories for all six degrees c! freedom are used. Past experience has shown that LOCA shaking loads are small when compared to LOCA rarefaction loads. For this analysis, these loads are obtained from the results of a prior analysis for a Model D steam generator.
8.1.3 Combined Plate Loads A summary of the resulting LOCA and seismic loads is provided in Table 81. In combining loads, the LOCA shaking and LOCA rarefaction loads are combined algebraically, while LOCA and SSE loads are combined using the square root of the sum of the squares.The TSP loads, which are reacted by wedge groups located at their periphery, are divided into three groups based on the wedge group arrangements for the plates. The number of wedge groups varies in number, size, "
and orientation among the various plates. The wedge group wentation for plates 2 and 5 is shown -
In Figure 8 8, and for plates 11 14 in Figure 8 9. The wsdge group sizes and locations for the remaining plates are combinations of plates 2 and 5. Typically, the wedge gioups are ,
symmetrical about the centerline of the bundle, and also hot leg to cold leg. TSPs 2 and 5 are .
two cases where hot to cold leg symmetry does not exist.
84
Relative to wedge group size for the V, C. Summer steam generators, the wedge groups are comprised of either two or three wedges, each two inches wide. Thus, the overall wedge group size is either 4 or 6 inches. The wedge group size is important, because it affects the local distribution of load into the neighboring tubes.
In reacting the load among the various wedge groups, a cosine distribution is assumed among the wedges that are loaded. Typleally, only half of the wedge groups are loaded at any given time, in
Selsmic loads on the other hand are random, and can act in any direction. Calculations are performed to determine load factors for the various plates, grouping the TSP by commonality of their wedge group locations. The load f actors are not a function of the wedge group size, only of location.
Applying these load factors to the overall TSP loads in Table 81, loads for each of the wedge ,
groups are determined. A summary of the Individual wedge loads is provided in Table 8 2.
8.1.4 Tube Deformation in estimating the number of deformed tubes, the results of TSP crush tests for Model D steam generators are used. The deformation criteria for estabilshing a tube as being susceptible to ,
in leakage has been defined to be 0.030 inch. in reporting the crush test results, tube deformations were reported foi various deformation magnitudes. This is the smallest deformation reported. Although test data is not available for leak rate as a function of tube deformation, it is judged that deformation levels of this magnitude will not result in significant in leakage.
Using the crush test data, a correlation is developed between elastle plate load and the number of tubes that would have a deformation of 0.030 inch or greater. it is this correlation, summarized
- In Table 8 3, that 19 used to approximate the number of affected tubes. Summaries of the number of potentially affected tubes for each of the wedge groups are provided in Tables 8 4 through 8 7.
8.2 Tube Maps / Summary Tables for Potentially Affected Tubes V. C. Summer is a three loop plant. The layout of the plant is such that all three loops have
- left hand' steam generators.The left hand designation refers to the orientation of the nozzles -
and manways on the channel head. For the purpose of this analysit. *left hand
- units are defined -
to be those loops where the primary fluid flows from the reactor to the steam generator to the pump and back to the reactor vessel in a counter clockwise direction. The left hand designation affects the location of the nozzles and manways, and the manner in which the columns are numbered for tube identification purposes. The reference configuration used in identifying wedge locations is shown in Figure 810. As shown in the figure, the nozzle and tube column 1 are
, located at 0.
Maps showing the locallon of the potentially susceptible tubes are provided in Figures 811
- through 8 21. Identification of the potentially cusceptible tubes is based on the crush test results. The wedge / tube configurations considered in the tests are not identical to those for the V. C. Summer steam generators. As such,it is not possible to identify exactly the tubes that might be !!miting at each wedge group. Thus, due to the uncertainties involved, there are more tubes 85
Identified at each wedge group as being limiting than estimated in the calculations.
Tabular summaries of the tubes that are potentially susceptible to oo!! apse and subsequent in leakage are summarlzed in Tables 8 8 through 813. Finally. Table 814 provides an index of the applicable tables and figures identifying the potentially susceptible tubes for each TSP. ,
8.3 Effect of Combined Accident Conditions on Tube Burst Capability .
Since the tube support plates provide lateral support against tube deformation that may occur during postulated accident conditions, tube bending stress is Induced at the TSP intersections.
This bending stress is distributed around the circumference of the tube cross section, with tension on one side and compression on the Cther side, and is oriented in the axial (along the tube axis) direction. Axlal cracks distributed around the circumference will therefore e?!her experience tenslon stress that tends to close the crack or compressive stress that tends to open the crack. Tlie compressive stress has the potential to then reduce the burst capability of the cracked tube due to opening of the crack.
l e
la o
86 l
_ . . . . . ~ . . . . . . . - . . . - . -.. - - -=
Table 81 Summary of LOCA Plus Seismic TSP Loade V. C. Summer Steam Genersters Steam Generator Inlet greek i
I a
N 1
iL a
f O 9 e
6
+
+
4 9
+
87 L -
4 Table 8 2 Summary of TSP Wedge Leeds , , -
V. C. Summer Steam Generstns""~ ~
. . .- :n .
..4.
f l
8 4
)
i 4
6 h
4 9
9 i
e f
I .
(
W .
I l
I l
1 4
w l-S. 8 r
., , ' , - , , , ,, - -,ev a,. -- g. n -,- , , - . , - - - - - , ,.U,, ,, -
Table 84 '
Number of Tubes with AD = 0.030 inch Versus Lead V. C. Summer. Model D Steam Generetor O
. E 6
4 4
e 4 N 4
4 49
9 Table 84 Number of Tubes with AD > 0.030 inch V. C. Summer Steam Generetore Steam Generefor inlet Broek TSPs 2 and 8
~'
.s .
4 8
e
+ 8 e
~
d 8 10
4 Table 8 5 ,
Number of Tubes with AD > 0.030 inch .
V. C. Summer Steam Generetore Steam Generator inlet Break
- TSPs 3,4, 5, 6, 7, 9,10 O
uns em G
9
- 4 i
W 8 11' yN-7 m*%mg W W f W F-WT*rM $ *w - e r ?*- -' w
Table 64 Number of Tubes with AD > 0,030 indh V. C. Summer Steam Generetors Steam Generatorinlet Break TSPs it 12.13 4
. 8_
\
4 4
4 W
M 4
en 0
8 12
____m_ - - - - - - - ~ ^ - ' - - - - ' - - - ^ - ' - - ' -
i Table 8 7 3 Number of Tubes with AD > 0.030 indh V. C. Summer Steam Generatore Steam Generator inlet Break -
TSP 14
\
l I
- a l
l A
O 4 b
b EO W
! t.
- i 8+13
. j 1
e Table 8-8 :
i Summary of Tubes holuded fromIPC l Y. C. Summer Steam Generatore i TSP 2 i
)
- 1 i
b i
d j i-d i
n s
2 4
I i
i F
h 4-e L
0 k
d, 1
i i e a
t 4
t~
t h
k I
.h 3
l 4 s
- 1. .
n I
J F - 6 dp o
i i
, 8 14 4
i y
.,m,.,d..:_._.,_w,, .. ,, , +, U..., y , -. - ,a.-- ,, g , ..,w.. r ._,v h m ,+ . ,,,.r... , .-.- , - y,
Table 8 9 Summary of Tubes Excluded fromlP9 V. C. Summer Steam Generators TSP 5-C
= .,3 e
4 E
<m e 4
6 8+15
i Table 810 Summary of Tubes Excluded from iPC'~ " :
V. C. Summer Steam Generators -
TSP 8 8
- 4 4
a 4
W e
8 16
d r
Table 611-
- Summary of Tubes Excluded fromlPC V. C. Summer Steam Generators TSPs 3,4,4,7,9,10 ;
49 m
- ~
s 8 '
e i
1 9
4 e
i un 4
.e. .
t 8 17
J d
. Table 812 Summary of Tubee Excluded fromik 4
V. C. Suh,mer Steam Generators i TSP 11,12,13
- i
-t I
8'. I
- Y
)-
4 4
i 1
'S l
Ib .
s i
3 t
8 18
.1
Table 813 ,
Summary of Tubes Excluded fromiPC V. C. Summer Steam Generators ~- -
TSP 14 e.
W, g'}
a
-k e
DD e
.?*
8 19
i Table 814 e Summary of Tables and Figures for TSP Row / Column identification I
i 1- .
i i
4 i Summary Tube Map l
ISP. Iahias Elautas I 2 88 811, 812 I
3 8 11- 8 17 4 8 11 8 17
- 5 89 813 to 818 '
6 8 11 8 17 7 8 11 8 17 8 8 10 811, 817 ,
, 9 8 11 8 17 10 ~ 811 8 17, i 11 8 12 818,819 l 12 8 12 818, 819 13- 8 12 818,819 14 8 13 820,821 l
5 e
l l
1 I
'8 20 i l
I
._ _ _. _ _ . . . _.- - .. .. _.. _ . _ . . ~ . _ . - . _ _ . _ . _ . _ _ _ . _ __ _ - . .
I Table 8-15 Combined Bending and Intemal Pressure Burst Tests on Tubes with Through Wall Slots
?
,, 8 0
k 4
e M
?
_.4 _
. d.
3 i.
3 21 i
li
- },
I a ,
S F
t . .
. Fgure 81. Seismic Model Representation of Steam Generator r
3 22-
. +n. - - , ~ -
i i
i .
i-k , ,
e l* , l 1
l 4
1% ,,
i IT ,
l l
l 12 l
. 11 1 1
. 10 - s 9 s
. 8m - '*
<- 7- s o 6- - -
m
\ w Sm ( '
+
- 4. - '
2 3 m
x 1m
.m l
~
ah '
l 5 '
o ; i
,4cm.Pm m m m DONHOT540WN Figure 8-2. Tube Bundle Geometry l
8 23
a' 4
9 4
Figure 8 3. Thermal / Hydraulic LOCA Tube Model 8 24
_- . .. . .. . .- -. . - . =. . . . - - . - . ~ . . . -- . .. -_
. - - - . . . . , .-- t i
4 4
7 a
- i i
4 I
s
?
g I-1 b
i 1
l, Figure 8-4. LOCA Pressure Time Histories, Maximum Radus Tube Nodes 815 :
8 25
-+-
5 -
3-
- a-i.
i .
1 i d
t j
t d
c .
t *
?
I i
4 4-a i
- i i- .
4 L
e- -
i i
f -
J'.
l _..
h b
i 1
i' t
i Fgure 8 5. LOCA Pressure Time History, Hot to-Cold Leg Pressure Differential,; .
j '. Maximum Radius Tube f :.-
8 66 w
a
4 e
4 '
- a 4
i o.
t k
2 i
i i
1 4
i 2.
1
- l
- Figure 8-6; LOCA Pressure Time History. Hot-to _ Cold Leg Pressure Differential, -
, ~
f
- Average Radus Tube = ,
-8 27 .
1 1 '
f a 4 t ,
a e
Figure 8-7. Structural LOCA Tube Model 8-28 i
1
_ _a 4
1 -
Figure 8-8. Wedge Group Orientation, Plates 2 and 5 -
8 29
a Figure 8 9. Wedge Grot'p O.dentation, Plates 11 14 8 30
90 i
Quadrant 2 i l Quadrant i Hot 1.og i-DMder Plate !
Y o
l M 0 o
180 -
j -
o Column 114 '! Coiwan 1 l
Quadrant 3 Cold Leg Quadrarit 4
?
o-270
- Figure 810. Reference Configuration for Tube identification Looking Down on Steam Generator -
Left Hand Unit 8 31-
.. a.
4 Figure B-11. Tubes Excluded from IPC, TSPs 2H,8H - Quadrant 1 -
8 32
1 i
i J
4> a Figure 8-12. Tubes Excluded from IPC, TSP 2C - Quadrant 4 -
8 33
., a 4
Figure 813. Tubes Excluded from IPC, TSP SH . Quadrant 1 8 34 i
e a
4 I
1 1
Figure 814. Tubes Excluded from IPC, TSP 5H - Quadrant 2 i
I 8- 35
a l
I, ;
i i
j l Figure 8-15, Tubes Excluded from IPC, TSP SC - Quadrant 3 .
L 8 36
. .,a-i
- Figure 8-16. Tubes Excluded from IPC, TSP SC - Quadrant 4 -
8 37 l
l l
a
,i - .
Figure 817. Tubes Excluded from IPC, TSPs 3,4,6,7,80,9,10 - Quadrant 4 '
8 38
i a
I i
9 Figure 8-18. Tubes Excluded from IPC, TSPs 11,12,13 - Quadrant 1 8 39
3 r
\
\ .
Figure 819. Tubes Excluded from IPC, TSPs 11,12,13 - Quadrant 2 .
8 40
i l
- a j o
4 O
Num &20. Tubes Excluded from IPC, TSP 14. Quadrant 1 8-41
.,a A
J F%ure 8-21. Tubes Exduded from IPC, TSP 14 - Quadrant 2 8 42
Pr; P/2 4- 2" c 3" > 4-- 2" ->
l if 1 f I
\_
O,__ / \ , )
J k
[ 1 J k 2
P/2 P/2
- Figure 8-22. Extemally Applied Bending Load and Locations of Through Wall Slots 8 43
Ja. 424--a.J 6*2AM 4 6544MLwSW +44Me++444 pa=M+2b-2 -e-44 4a--E m15- d a h M 4 A M."* rehhr s' i- h+h+. AdJ.-- A--Mm aM* 4ma h 1. m. G4.M ds=* A m 5. a4 2-m .Ja6 M h4dM4)BAMs + .&d 5+=**
$'#9 I
4 s r
4 I
1 a
t I
e.:
1 1
s P
I
,e b
6 h
6 Y
4 i
' +t f
1
-t
.1 I
t
\
e \
p 4
P
- e b
6 6
h 6
8
')i ih t
i.
7
. g ..
+
h e
g
.' f
, , .. .. .., .. - - . . ..n + .-- - ~., w- r - - - - -r w ,-c--, ~ ~ ~ - - *'- c e r"**'- ^#'" ' ' ' ' ' '
Section 9 V. C. SUMMER INSPECTION RESULTS 9.1 October 1991 Inspection A scheduled inspection of the V. C. Summer steam generators was conducted during the refueling outage in October, iv91. All tubes in service were inspected by bobbin ooll eddy current tests.
131 Indications at the tube support plate (TSP) locations were reported from the bobbin coll tests and analysis conducted during the outage. These indications (reported during the 1991 inspection) were reevaluated with the primary objective of estimating voltage growth rates during the last cycle.
This reevaluation was conducted in the laboratory and particular attention was paid to obtalning accurate voltage amplitudes needed to estimate growth in Indication voltage. The voltage amplitudes of TSP Indications discussed below are for the 550/130 kHz mix channel with the amplitude of the 20% holes in the ASME standards normalized to 2.75 volts. This is the same frequency mix and normalization used in the APC (allemate plugging criterla) database and hence renorrnalization of the NDE data was not required.- For the indicatlons reevaluated from the 1991 Inspection, eddy current (EC) data tapes from the prior inspection (April 1990) were also evaluated so that voltage growths could be determined for the operating cycle. This reevaluation is the source of the results described in this section,i.e., no reference was made to the data analysis performed in the field during the outages. The following is a summary of the results relating to ODSCC Indications at TSP locations.
The tube support plate indicatlons from the 1991 inspection were distributed among the three steam generators as follows. There were 19 Indications in steam generator A,72 in S/G B, and
~
40 in S/G C. Allindications were in the hot leg Figure 91 displays the frequency distribution of TSP Indications by support plate locations. The upper figure shows the number of observations at a TSP while the lower figure shows the percent of all TSP Indications falling at a given support plate location. Location 1H is the flow distribution baffle. Figure 9 2 shows a schematic of the support plates in the tube bundle and their designations. Besides the flow distribution baffle, there are seven tube suppuri plates in the hot leg, designated 2H,5H,8H, 11H,12H,13H and 14H. They are numbered from bottom to top,14H being the highest. It may be noted from Figure 91 that over half of the indications were at the lowest TSP in the hot leg (2H) and a third in the next two TSPs combined. There were no indications in the two uppermost TSPs (13H and 14H). Thus the TSP Indications were located predominantly in the lower support plates in the hot leg. This is consistent with observations at other plants evaluated by Westinghouse for APC application. Four (4) of the previously reported Indications (1 in S/G B and 3 in S/G C) were found to be NDD (no detectable degradation) upon reevaluation.
In general, the TSP Indications at V. C. Summer had low amplitudes. A frequency distribution of the 1991 voltage amplitudes from the TSP Indications is displayed in Figure 9 3. The upper
, ends of the voltage ranges are shown on the X axis scale. Thus the first bar at 0.2 volts
- represents indications of amplitudes between 0 and 0.2 volts. -The cumulative percent frequency is also plotted and is shown as a curve. 97% of the Indications had amplitudes of less than 2 volts (550/130 kHz mix channel). There were four (4) Indications between 2 and 3 volts in amplitade and none above 3 volts. A histogram of the bobbin voltages in each S/G is shown in Figure 9 4. As in Figure 9 3, the X-axis scale displays the upper ends of the voltage ranges.
91
t 9.2 Pilor inspection ,
The reevaluation included a review of th9 eddy current inspection results from the April 1990 inspection conducted at V. C. Summer. The objective of the review was to assess the progression **
of ODSCO in the Summer S/Os. The 131 Indications from the 1991 Inspection were traced back to this prior inspection to assess whether they may be observed in light of the 1991 data and if so, to obtain the bobbin signal evaluation. The results of this review is used to obtain growth '
rates of the TSP Indications.
The bobbin signal amplitudes will depend on the calibrMion standards used during each of the inspections. The ASME standard for flaw depths specules a tolerance of A 3 mils or 20%
whichever is t kvas the diametric tolerance is A 10 mils. Hotvever, measurements on plant tubing e - t 41 the diametral variation on the tubing is typically only a few mils. This leads to a pot w... diametrio variation of A 16% and A 5%, respectively, for the 100% and 20% deep ASWlE holes assumlng that the machined standard is within the specifications. The engineering drawing of the calibration standard used during an Inspection provides the exact as-built (measured) depths for the machined flaws; but does not give the as built diameters of the flaws. Furthor, the accuracy of depth for the through wall hole is absolute. Thus while the standard was well suited for depth estimates with hlgh degree of confidence, the signal amplitude estimates tend to be less reliable due to the variation in hole diameters in the standard. This uncertainty is inherent in the voltage results from past inspections.
The bobbin voltage amplitudes of the TSP Indications during the 1990 Inspection obtained from the reevaluation ranged from 0.05 to 1.47 volts with an average value of 0.66 volts. Growth in -
amplitude for the last cycle was determined from the data. Growth estimates were calculated only for the cases where bobbin signal voltage data were available for both the inspections;l.e., -
no assumdlon about the signalvoltage for the prior year was made if a flaw indication was not available. As a result, growth estimates for the 1990 91 cycle were made for only 87 of the 131 Indications (five Indications were not detectable,i.e., NDD, during the reevaluation of the 1990 Inspection data and 39 were not in the population of intersections tested during 1990).
Figure 9 5 shows a plot of the growth in arnplitude from 1990 to 1991 as a function of the 1990 amplitude for the TSP Indications in all S/Gs. It may be noted that the amplitude growth ranged from about 0.5 to 42,0 volts. The negative growths (and possibly some of the high positive growth values) result from the uncertainties in the eddy current inspection and data evaluation discussed above. Overall, the distribution of amplitudes and their growths appear consistent with experience with data from certain other plants.
Figure 0 6 shows a frequency distribution of voltage growths in all S/Gs during the 1990 91 cycle. The upper ends of the ranges are displayed on the X axis. The cumulative frequency distribution in percent is also displayed in the figure, as an 'S' curve, it may be noted that most of the indications had growth rates between 0.0 and 0.5 volt per cycle.
The average growth in amplitude for indications in each S/G was calculated for the last cycle ' ,
This Is displayed in Table 91 along with the overall average which includes all S/Gs. The number of Indications used in the calculation of the average is also shown in the table. Some of the variation in the growth rates is attributable to the uncertainty in the voltage Indications .
from prior inspections. The overall average growth rates of TSP Indications during the 1990 91 cycle was 0.29 volt. The standard deviation associated with the overall average growth in amplitude during the last cycle was 0.37 volt. As discussed before, part of the scatter 92 l
results from the uncertainty in the eddy current test results and the data evaluation and the remaining from the variabl!!ty in growth between Indications. Overall, it may be noted that the average amplitude growth is low, being in the neighborhood of 0.3 volt per cycle.
/ 9.3 Growth in Voltage Amplitude As discussed in Section 9.2, the growth in amplitudo per fuel cycle ranged from 0.5 to +2.0 volts. Experience with the data from certain European plants indicates that the percent growth in amplitude tends to be stable,i.e., independent of amplitude (however, this is not supported by i data from certain domestic units including V. C. Summer). - The European data suggests that the small amplitude indications grow by smaller voltages and that large amplitudo signals are more likely to grow by a larger amplitude during the subsequent cycle.- Percent growth rate is calculated by taking the ratio of the growth in amplitude during an operating cycle to the amplitude of the signal during the prior inspection. Figuro 9 7 shows a plot of the percent growth in amplitude vs beginning of cycle (BOC) bobbin voltage for the 1990-91 operating cycle. Two Indications with beginning of cycle amplitudes less than 0.1 volt are not shown in this figure since they lie beyond the displayed range of the y axis, it may be noted that the high percent growth rates are observed only at low amplitudes. Two factors contribute to this: 1) at ;
low signal amplitudes, the uncertalnty in the signal analysis may be higher, and 2) since the BOC amplitude is in the denominator in the percent growth calculation, the percent growth value is magnified for the low amplitude indication. -
The average percent growth rate of all the TSP Indications during the last cycle was calculated for Summer from the average voltage growth rate and the average BOC amplitude. This is displayed in Table 9 2. The average growth rate of TSP Indication voltage was 44% per cycle.
It was noted before that the percent growth rate of amplitude is lower at higher BOC amplitudes ,
(see Figure 9 7). This observation is typleal of other plant data evaluated by Westinghouse to date. Thus an overall average of percent growth rate may have an upward bias due to the small amplitude signals. To assess the impact of this factor, average percent growth rates were calculated for two different BOC amplitude ranges above and below 0.75 volt. The results are displayed in Table 9 2. For BOC amplitudes below 0.75 volt, the average percent growth rate for the 1990 91 cycle was 77% whereas the average for indications with BOC amplitudes
, equaling or exceeding 0.75 volt was 16*/J Tnis ditterence must be noted as additional conservatism in the development of the plugging critoria. The standard deviations associated-with the averages are also listed in Table 9 2.
Growth rate projection for Cycle 8 (1993 94) is made from the Cycle 6 results based on the relative lengths of the operating cycles in effective full power days (EFPD). The Cycle 6 operation lasted 427 EFPDs whereas the planned operating duration for Cycle 8 Is 444 EFPDs.
This is an increase of only 4%. Never the less, the gowth rates calculated for Cycle 6 were factored up by the ratio of 444/427 to estimate the projected growth for Cycle 8. The results are shown in Table 9 2. The corresponding frequency distributions are shown in Figuro 9 8.
1 This growth distribution is used in the Monte Carlo analyses for projecting the EOC voltage distribution of TSP Indications (Section 12).
9 93
9.4 Influence of TSP Location Initiation and progression of ODSCC could be affected by, among other factors, the temperature of the local environment at the tube elevation. Since the primary coolant temperature decreases as it flows up through the tube on the hot leg side, the local tube wall temperature decreases with
necessary to assess the growth rates by location and determine if any dependence exists.
The average growth in amplitude was calculated for each of the support plate locations in the V. C.
Summer data. The results are presented in Table 9 3. The number of growth values used in calculating the average in each case are also shown in the table. The average percent growth rates shown in the table were calculated from the average voltage growths and the average BOC amplitudes. Only three support plates (2H, SH and 8H) had 5 or more indications with amplitude growth values. These are the three lowest TSPs in the hot leg, not including the flow distribution baffle. There is no apparent correlation between amplitude growth rates of Indications and TSP elevation, if at all, the growth rate is slightly higher at the two upper TSPs (5H and 8H) than that at the lowest TSP (2H), a trend opposite to that expected on the basis of temperature dependence. This is true of both the absolute growths as well as the percent growths in amplitude. The absolute growth is the same for TSPs SH and 8H (0.35 and 0.34 volt, respectively). Therefore, the variation in growth between support plates appears to be random rather than systematic. Since the indications at the first TSP dominate the data, the averages for the cycle are quite close to the conesponding averages for the first TSP. Thus the growth rate data from the last V. C. Summer cycle 6 (1990-91) suggests no variation in growth rateh between the TSP locations.
9.5 RPC Inspection Results ,
During the 1991 V. C. Summer outage, RPC inspection was performed at TSP Indications in several tubes in each steam generator. In general, RPC test results confirmed bobbin coil indicationa. The review of the RPC data suggests that the TSP signals are due to axial ODSCO.
They are comprised of both single and multiple axlat indications. Most of the RPC traces could be easily interpreted. However, interpretation of the signalis difficult in a few cases as a result of noise or other complexity in the RPC tcsults.
A sample of the RPC traces from the three steam generators are shown in Figures 9 9 through 911. RPC traces in Figure 9 9 are from S/G-A. They are from tubes R40067, R42C48 R27C99, and R14C56 and had bobbin amplitudes of 1.05,0.52,1.11, and 1.41 volts from the laboratory reevaluation. Three of the traces are from the lowest TSP (2H) and the fourth trace (R14C56)is from the second lowest (5H), The two upper traces show multiple axial .
Indications whereas the two lower traces are single t clalindications. These are similar to Indications found at other plants with ODSCO.
RPC traces from S/G B are shown in Figure 910. They are from tubes R10C44, R10C94, R90103, and R46C71. Bobbin coil voltages obtained from the laboratory reevaluation of these indications were 1.24,1.02,1.01, and 0.75 volts, respectively. Three of the traces are from ,
the lowest TSP (2H) and the fourth trace (R46C71) is from the second lowest TSP (5H). The two traces in the upper section of the figure show multiple axlalind'eations. The two lower traces are of single axialind' cations.
l l 9- 4 u- w m m. a yw m -
y-
Figure 911 shows RPC traces from the C steam generator. These are from-tubes R39091, RBC27 R37C40, and R21084. Laboratory reevaluation of the EC data from the 1991 inspection had shown their bobbin 0011 amplitudes to be 1.23,1.18,1.73, and 1.40 volts, respectively.
The top traces in the figuie are from the lowest support plate (2H) while the lower traces are from the plate 6H. The trace in the upper left, from tube R39091, shows multiple axlal Indications. The other three traces show single axialindications. 7 These are examples of the RPC traces from the V. C Summerinspection of 1991. Revlow of the overall RPC data suggests that the support plate indications are axial ODSCO signals, f
4 a
S 95
Table 91 .
Volfa0e Growth Per Cycle for V. C. Summer S/Gs -
1990 91 Cycle ,
800 Indication Vettana Voltaam Growth nar Cvela SLQ Number 0) Average Std. Dev. Number U) Averana Std. Dev.
A 17 0.60 0.25 17 0.29 0.30 B 43 0.71 0.33 43 0.30 . 0.34 C 27 0.63 0.27 -27 0.26 0.44 All 87 (2) 0.66 0.30- 87 (2) 0.29 0.37 Notes
.' 1. Number of Indications included in the calculation of the statistics (average and std.
deviation).
- 2. This total differs from (is lower than) the 131 Indications from the 1991 Inspection '
primarily because many of the TSP Intersections with Indications in 1991 were not tested during the 1990 inspection and a few others from the 1990 inspection were '
not detectable (NDD).
t 96
Table 9 2
. Percent Voltage Growth Per Cycle for V. C. Summer S/Gs (U 1990 91 Cycle Cycle 8 Average Percent Projected Number of (2) BOC Voltage Growth / Cycle Growth Growth (d)
Indicatient Vohace Avernae Std. Dev. / Cycle Volt'evelo Entito voltage ra'nge 87 (3) 0.66 0.29 0.37 43.9 0.30 55 0.47 0.36 0.30 76.6 0.37 VBOC < 0.75 volt VBOC a 0.75 volt 32 1.00 0.16 0.44 16.0 0.17 blcle1
- 1. Percent voltage growth per cycle determined as (VEOC VBOC)/ VBOC
- 100. .
- 2. Number of indications in the calculation of the growth statistics.
- 3. This total differs from (is lower than) the 131 Indications from the 1991 inspection primarily because many of the TSP intersections with indications in 1991 were not tested during the 1990 inspection and a few others from the 1990 inspection were not detectable (NDD).
4 97
Table 9 3 ,
Voltage Growth Per Cycle by TSP Location 1990 91 Cycle Average Average Average -
TSP Number of U) BOC Voltage Percent Loention Indientions Voltaan Growth Growth 8H 5 0.48 0.34 7i%
SH 8 0.67 0.35 52 %
2H 67 0.69 0.26 38 %
All 87(2) 0.66 0.29 44 %
HQ11t1
- 1. Number of Indications in the calculation of the growth statistics. .
- 2. This total differs from (is lower than) the 131 Indications from the 1991 Inspection .
primarily because many of the TSP intersections with indications in 1991 were not tested during the 1990 inspection and a few others from the 1990 inspection were not detectable (NDD). The totallncludes 2,1 and 2 indications at TSPs 1H,11H, and -
12 H respectively.
I 98
80
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60
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~ LOCATION OF NEATIONS e
~ Figure 91. Frequency Distrbution of TSP Indications by Location (1991 inspection) :
99
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Figure 9 2. TSP Locations and Designations in the V. C. Summer S/G Tube Bundle '
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, 0.4 - 0.8 1.2 1.6 2 2.4 2.8 1991 BOBBIN AMPUTUDE, VOLTS Figure 9-3. Frequency and Cumulative Percentage Distribution of Bobbin Amplitudes:
(TSP indications from the 1991 Inspection) .
. 9 11
14 12- ~~~.;~~~~~~~~~ -- -- * ~ ~ ~ ~ ~ ~ ~ ~ " * ~ ~ . ~ ~ ~ ~ ~ ~ ~ ~ . ~ ~ ~ . ~ ~ ~ -
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4 t
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- 1990 BOBBIN AMPLITUDE, VOLTS g ?
. Figure 9 5. Growth in Bobbin' Amplitudes During Cycle 6 (1990-91) _
,9 13-
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,R 0 0.5 0.3 0.1 0.1 0.3 0.5 0.7 0.9 1.2 1.6 2.1 0.4 0.2 0 0.2 0.4 0.0 C8 1 1.4 1.8 GROWTH IN BOBBIN AMPUR.lDE, VOLTS Figure 9 6, Histogram and Cumulative Probability Distribution of Voltage Growth During Cycle 6 (1990-91) 9 14 t
400 2R o a 8,_! 300- ~ ~ " " " " " " " " * " " " " " " " " " " " " " " " " " " " " " " " " " " " " " " " " " " -
O Z D
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7 200- ......................O............................................................................
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.a o o> ao 100 , , , , , , i 0 0.2 0.4 - 0.6 0.8 1 1.2 1.4 1.6 1990 BOBBIN AMPUTUDE, VOLTS Figure 9 7. Percent Growth in Bobbin Amplitudes During Cycle 6 (1990 91)
O 15
_ _ _ . , - _ _ -m .2 _m_. . . . n__ g ._ - - _ _
n V. C. SU MMEFl TSP IN DICATIONS .
PROJECTED VOLTAGE GROWTH DISTRIBUTIJN 100 .
00 ........................... ................, ....................................................
w g Qoe ...e..e..e........ ....e'.ee.ee.............ee......e.e...e..e.......n.....ee.....................
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0 0.6 1 1,5 2 2,5 BOBBIN AMPLITUDE GROWTH RATE, VOLTG Figure 9 8, Cumulative Frequency Distribution of Growth Rate Projection for Cycle 8 i
9 16 l
l' .-
= _- _ . . _ - _ _ _ _ _. - _ _ - -_
9
) f'
/
j .45 /
~
0.48 2H+0.44 2H+4.35
-0.48 36e 4 .63 360 a) R4 C67. Location 2H b) R42 C48, Location 2H
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,y
. 45 4 // ,
8.54 j% j
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2H+4.37 5H+0.33
-0.48 36g c) R27C99, Location 2H d) R14 C56, Location SH Figure 9 9. Example RPC T races from Hot Log Support Plate locations in S/G-A (1991 Inspection) 9 17
~f l97
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/
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8.51 -e.ss 0
'd m e.4, 2N4.39
- 8.64 360 8.4s e a) R8s0044, Locatbn 2H b) R10 C94. %tbn 2H 9
4 .
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8.79 k
//,
0.48 < '< 5N*0.33 2H+0.31
't.56 360 4 .79 360 c) R9 C103, btbn 2H d) R46 C71, Locatbn 5H I* 9~10. Exarnple RPC Tra s rotn 9 18.
l l
s.se 2H+4.41 .
4.62 Me' *4.M M4 I i
a) R39 C91. Location 2H b) R8 C27, location 2H I
e.ss
-e.53 ase e.se e c) R37 C40. Location 5H d) R21 C64, Location SH Figure 911. Example RPC Traces from Hot Leg Support Plate Locations in S/G C (1991 Inspection) 9 19 y ww-. , - - . > wevs-~-op- '~w4 vw-- ' - - + - -
Section 10 -
BURST PRESSURE CORRELATION 10.1 Introduction This section utillzes the model boiler (Section 4) and pulled tube (Section 6) data to develop a correlation of burst pressure vs. bobbin voltage. A preliminary correlation was reported in WCAP 13494 (in support of Plant R.1 IPC application). Subsequent to that submittal, significant progress has been made in the development of the burst correlation. A thorough statistical procedure was implemented as described below. All normalization of eddy current data has been reviewed in detall and resolved by the EPRI APC Committee. The 19g1 pulled tube data from Plant R 1 were not included due to the .
large uncertainty in the burst test results. Thus the data, methodology, and correlation provided in this section are no longer considered preliminary and hence provide a high degree of confidence In the results.
10.2 Data Base for Burst Pressure Correlation (3/4 inch Tubing)
The database used for the development of the burst correlation (burst strength vs.
bobbin 0011 voltage amplitude) for 3/4 inch diameter tubing is derived from model boiler specimens and pulled tubes. All of the data were derivei from Alloy 600 tubing with 3/4 inch OD and 0.043 Inch nominal wall thickness. The model boller test results for 3/4 inch tubing are described in detall in Section 4. All reported bobbin coll measurements on model boller specimens are for 550/130 kHz mix frequency with the 20% holes in the reference ASME standard normallzed to 2.75 volts.
The pulled tube data included in the database are obtained from Plants B 1, E 4 and R 1. The bobbin voltages from Plant B 1 were measured at a frequency mix essentially the same as the model boiler specimens. The bobbin data from Plant E 4 pulled tubes were dominantly (10 of 12 Indications) measured at 550/130 kHz mix with the APO voltage mormalization, and two indications were measured at 300 kHz using the Belglan NDE procedures. This data has been normalized to correspond to the APC database using the process described in Section 5.6. The principal adjustment to the direct field measurements resulted from cross-calibration of the Belgian ASME standard to the reference laboratory standard. Conversion of the Belgian 300 kHz data to the APC 550/130 kHz mix is obtained from direct field measurements obtained with both voltage normalizations. This process involved significant and detailed efforts by two independent parties (Laboreleo and Westinghouse) and thorough review by the EPRI APC Committee such that there is negligible uncertainty in the normalization procedure. The bobbin data from the 1992 pulled tubes from Plant R 1 were obtained using a procedure equivalent to that used for the model boiler database (field /J ME standards were a calibrated against the reference standard). The 1991 pulled tube data from Plant R 1 were not included due to the large uncertainty in the burst test results. The pulled tube '
results are described in Sections 6.6,6.7 and 6.8.
The burst pressures of all the room temperature data are normalized to a reference flow stress of 75 ksi. This value is close to the 77 ksi mean flow stress for the mill annealed '
Alloy 600 tubing at room temperature (WCAP 12522). The resulting burst pressure 10- 1
--nue- - - - - _ _ - - -
database is summarized in Table 101 for both the model boiler specimens and the pulled tubes. The database includes 47 rr "iel boller spec! mens and 14 pulled tube intersections.
10.3 Burst Pressure vs. Voltage Correlation ,
The bobbin 0o11 voltage amplitude and burst pressure data of Table 101 were used to determine a correlation between burst pressure and bobbin voltage amplitude. This is *!
not to say that a " formal" functional relationship,in the sense of one variable being dependent on the other, exists between the variables since the burst pressure is not caused by the bobbin voltage and vice versa. The burst pressure and bobbin voltage variables considered are mainly functions of a third variable,l.e., the crack morphology. While the variation in crack morphologies is essentially infinite, suitable descriptions can be effected based on the depth, average depth, profile description, etc.
However, the characterization of the morphology is not essential to this analysis since a relationship is being independently established between two offspring varlables.
Although the correlation analysis does not estab!!sh a causal relationship tatween the variables,it does, however, estab!!sh a " working" relationship that can be employed for ,
the prediction of one variable from the other, The data considered are shown on Figure 101 along with the results of three correlation analyses which will be discussed in later sections.
The analysis performed considered the scale factors for the coordinate system to be employed,i.e., logarithmic versus linear, the detection and treatment of outilers, the order of the regression equation, the potential for measurement errors in the variables, and the evaluation of the residuals following the development of a relation by least squares regression analysis. -
In summary, It was concluded that the optimum linear, first order relation could be -
achieved by considering the burst pressure relative to the common logarithm (base 10) of the bobbin amplitude voltage. For this relationship it was determined that a bobbin voltage value of 0.1 volt could be ascribed to the burst data where degradation was not detected, i.e., no detectable degradaton (NDD). This is necessary to include NDD specimens in the database since the burst pressure should be a continuous function to the point of no existing degradation. A linear, first order equation relating the burst pressure to the logarithm of the bobbin amplitude was developed. The correlation coefficient from the regression analysis was found to be significant at a >99.9% level.
Analysis of the residuals from the regression analysis indicated that they are normally -
distributed, thus verifying the assumption of normality inherent in the use of least squares regression, 10.3.1 Selection of Coordinate System in order to establish a correlation between the pairs of variables, but expected to have independent variances, the method of least squares (LS) curve fitting was employed. The ,
simplest functional form is a linear relationship of the type y - a o+ ajx (10-1) ..
where the variabtes x and y may be linear or logarithmic independently, and the 10 2
l coefficients of the relation, aoand at ate to be determined from the analysls. in addit!on, the choice of the regressor variable is not pre determined. Both variables are assumed to be sub'ect to random fluctuations which are normally distributed about the mean of the variable or the logarithm of the variable with a mean of zero and some unknown, but reasonable variance. It Is also assumed that this variance is constant, or unMorm, over the range of interest of the variables. In practice this may not be the ;
case; however, any non uniformity present would not be expected to significantly affect the analysis outcome.
Analyses were performed to determine the optimum nature of the variable scales,l.e.,
linear versus logarithmic, and the appropriate selection of the regressor varlable. It was concluded that the most meaningful correlation could be achieved by considering the log of the voltage as the regressor and the burst pressure as the response. Thus, the functional form of the correlation is Pa = a o+ aglog(V) (10-2) where Ps is the burst pressure and V is the bobbin voltage amplitude. The basis for selection of the form of the variables was based on performing least squares regression analysis on each possible combination and examining the square of the correlation coefficient (the Index of determination) for each case. The selection of the regressor variable does not affect the calculation of the Index If the calculations are performed on the transformed data. The results of the calculations are shown below.
Index of Determination, r2, for Various
, Selections of Coordinate Scales PB V Unear Logarithmic unear 28.7 % 31.0 %
Logarithmic 79.7 % 71.2 %
The result s clearly chow an advantage for treating the voltage on a logarithmic scale and the burst pressure on a linear scale Given this cholce of axes scales,it is apparent that the burst pressure should be regressed on the logarithm of the voltage amplitudes. The rationa!e for this is that the residual error bands will be reduced with this selection relative to performing the regression in the opposite direction, it is noted that the data contain some results for specimens in which there was no detectable degradation (NDD) from the non destructive examinatlon. The inclusion of this data is necessary in order to predict burst pressures for Indications with very low bobbin amp!!tudes. Two methods were considered for inclusion of the NDD data in the analysis. The first method consists simply of assigning a low bobb!n voltage ampiltude to the data. This was done for the determination of the best choice of scales for the coordinates of the plot. The value assigned was 0.1 volt for the NDD specimer s. The second method consists of modifying the prediction model to include a voltage offsef variable to be determined from the data,i.e., the prediction equation becomes Pg - a +o a3 og(V 1 + Vo) (10-3) l t
l 10 3
where Vo is the additional parameter to be determined from the data. V represents a voltage offset at which the NDD data are included in the regression 0. at V o the Since equation is now non linear in the parameters the a 3 plication of least squares techniques !
is not appropriate. However,if an assumed value s assigned to the offset term the equation is once again linear and least squares can be applied to find the values for the other two coefficients. Analysis was performed for a varlety of offset values with the .
result that the index of determination was a maximum of 79% for Vo- 0.236 volt.
This is slightly less than the value reported above for determining the best selection of
- the coordinate scales, it was thus concluded that the comp!! cation of including a voltage offset in the prediction modelis not necessary to account for the NPO specimens.
Additional analysis was performed to determine if a value less than 0.1 volt would be appropriate for NDD specimens. It was found that the Index of determination becomes a maximum if the NDD specimens are assigned a voltage level of 0.15 V. However, the improvement was significant only in the third decimal place. Since this level of amplitude was present in the data for a confirmed indication it was judged to be ineppropriate as an assigned value for NDD specimens. Likewise, reduction of the assigned value was considered inappropriate since it would reduce the fit of the burst regression curve artificially.
An additonal consideration for the analysis of the data was to increase the order of the prediction equation. This would allow for the assignment of a lower value for the NDD specimens. Under this consideration the model would be Pa - a +o ai og(V) l + a2[ log (V)]2 (10 4)
The results of this analysis are shown on Figure 101. The results Indicate no improvement in the index of determination and the introduction of a second order term provides no improvement in the model. Thus, a linear (first order) model was retained for the analysis. .
10.3.2 Regression Analysis for the identification of Outliers The data contained in Table 101 were analyzed to identify any potential outlying data points. The analysis was performed using the robust regression technique based on ndntmizing the median of the squares of the residuals. The method, known as the least median of squares method,is described in Appendix B. Specimen 5913 (Model Boller) was found b have a resloual to scale ratio of 3.31, Indicating that the burst pressure tesult does not conform with the rest of the data at a levelof about 99.9% Such a data point might be expected to occur in approximately 6% of repeated test sets for the same number of tests performed per set. This value is shown as a solid circle on Figure 101. The bobbin amplitude for the Indication was 14.46 V, and the burst pressure was found to be 7647 psl when normalized to a flow stress of 75 ksi.
No specific error in either the burst pressure or voltage measurements was identified for Specimen 5913. However, as seen in Table 101, this specimen has more (6)'
RPC Indications than other specimens. The bobbin voltage increases with the number of indications around the tube circumference, while the burst pressure is limited by the limiting single crack. Thus, specimens with multiple Indications (hlgh voltages relative to burst pressure) are expec'cd to contribute to the high (non conseNative) burst- ,
pressure tall of the burst / volta 0e correlation. Specimen 5913 is the most extreme, high outlier and is excluded from the correlation on this basis. In contrast, the low burst pressure outliers have a different physical basis (crack morphology). The low ,_
burst pressure outilers result from the burst crack having one or more relatively largo ductile (uncorroded) ligaments sepaiating deep microcracks that comprise the overall macrocrack. In this case, the remaining ligaments act to decrease the bobbin voltage 1o 4-
@ r m
1 proportionally and the extreme talls of the distributions could be different, Only high !
I outilers are considered for deletion from the regression analysis. The low outilers are retained to provide conservatism in the regressbn fits, it is seen by later analyses of residuals that the retained data follow a normal distribution. On this basis, the data from specimen 5913 was deleted from the analys s of the data. Confirmation of the
'. outlying nature of the burst pressure result for this specimen is also apparent from the cumulative probability plot of the residuals (Figure 10 3).
10.3.3 Error in Variables Analysis A general assumption in performing a leata squares regression analysis to establish a correlation is that both of the variables, burst pressure and bobbin voltage in this case, arc subject to random fluctuations about their respective mean values (or their re.pective log mean values). If there are significant uncertainties in the measurement of one or both of the variables the slope of the I.S correlation line will be biased. it is judged a priori that significant measurement error does not exist in the values of the burst pressure, however, no such judgement was made relative to the reported bobbin coll voltage amplitudes. Thus, it was assumed that mearurement error could present itself in one of the two variables, but not both. This means that the variance of the variable, X, with the measurement error, as estimated from the data, consists of two parts, the intrinsic variation of the variable, Ox, and the variation due to measurement error, Om* 0 0
- 02 ,o,2+ g ,2, (10 5)
This results in the expected value of the slope of the regression line, by , being E(bj)= / [1 + (an/ Ux)2) (1g.6) where represents the slope of the true relation,if any. Thus, the regression performed always underestimates the slope of the true relation, if the measurement
- error variance or the ratio of variances is known, the effect of the error on the slope of the regression line can be evaluated directly. An alternative evaluation of the potential effect can be performed based on an examination of the data per the partitioning procedure developed by Wald and subsequently improved by Bartlett. A Wald Bartlett evaluation of the data was performed and it was concluded that the presence of l
i measurement error would not have a significant effect on the slope of the correlation line, The results of the analysis are shown in Figure 10-1. The Wald-Bartlett line l
corresponds to assuming a measurement error variance en the order of 5% or 6% of the variance of the log of the bobbin voltage data.
Since the omission of correction for measurement error results in an under prediction of the slope of the correlation, predicted burst pressures for voltages below the centroid of the data willlikewise be less than a prediction based on consideration of the measurement error. For voltage values greater than the centroid of the data the correlation slightly over predicts the burst pressure. An examination of the plot Indicates that for the 3AP limit of 3996 psl, the offset in the voltage direction will be approximately zero for tha prediction curve. For the St.B AP limit of 2335 psi, the-
-- offset is on the o; der of 6% to 10%, i.e., on the order of 1 volt, under the assumption -
that the variance of the measurement error is such that the Wald Bartlett line would result from an analysis including the error. Consequently, the measurement error can
,- be acceptably ignored and the more conventional regression analysis can be applied for data analysis. 1 1
10* 5 I
l 10.3.4 Burst Pressure Correlation for 3/4 inch Diameter Tubing .
The final fit of the data is shown on Figure 10-2. The correlation line is given by PD = ao + ajlodV)i
= 7.837 2.900 log (V i) (10 7) .
where the burst pressure is measured in ksi and the bobbin amplitudo is in volts. As -
' noted in Section 10.3.1, the index of determination for this regression was 79.7%, thus, the correlation coefficient is 0.89. The estimated standard dovlation of the residuals, 1 1.e., the error of the estimate, sp, of the burst pressure was 0.96 ksi. A one sided 95%
simultaneous confidence bound. corresponding to a 90% two-sided band, was calculated for the mean burst pressure, F,, , corresponding to a specific voltage, Vl, per the following equation: i l
Es, * %
- aglogG'd * /2Fu.a...: 8r f* , , (10 0) i n i i where n is the total value corresponding number to an upper tall ofarea dataof points 10% in additoused, a and 95% Fo one 9sided g,n,2 s the F distribution prediction band for individual values of burst pressure, P;, as a function of voltage was also calculated por the following equation:
8*
1 P,,=%+a:l*V)eim .:sp 1 3. +
(10 9) s E l t<V)~VsTV)l' where to.95, n 2i s the t-distribution value corresponding to an upper tall area of 54 .
The simultaneous confidence bound and the prediction bound are shown on Figuro 10 2.
Since the burst tests were performed at room temperature (RT) conftlons, the 95% .
prediction bound wast further reduced to a level corresponding to the 95%/95% lower tolerance limit (LTL) for the material properties of the tubes. This was done by multiplying the predicted burst pressure by the ratio (0.871) of the LTL limit of flow stress at 650'F (65.3 ksi) to the reference, or normalized, value of the RT flow stress (75 ksi) used in the analysts. A second order fit of bobbin amplitude to differential
. pressure for the 95% prediction curve as adjusted by the LTL material flow stress was performed for the purpose of determining lower bound voltage amplitudes as a function of the applied pressure differentla!.
V LTL - 10 2116. o.382P . o.oo17 P^2 (10 10)
Using this result the voltage amplitude corresponding to a differential pressure of 3.996 ksi would be [ ja.g volt, and the amplitude corresponding to a differential pressure of 2.335 ksi would be [ }a.g volt, 10.3.5 Analysis of Residuals Vorification of the regression analysis was performed based on testing the correlation .
coefficient for significance. For tho analysis with 58 degrees of freedom a correlation coefficient in excess of 0.42 would be significant at a level of 99.9% Given tho .
. calculated correlation coefficient of 0.89,it is reasonable to accept the hypothesis that -
the burst pressure and bobbin amplitude are correlated.- In order to verify conformance with the assumptions inherent in performing the LS analysis, the residual values were loiG
-- _- __ __ . - - . - - - _. - . - = . _ . - _ . - . . .~ .- .
olotted against the predicted burst pressures. The results, shown on Figure 10 3, ndicated acceptable scatter, i.e., non-descript, of the residuals about a mean of zero.
Finally, a cumulatke probability plot of the ordered residuals was prepared, also shown on Figure 10 3. Since the data form a straight line it is obvious that the distribution of the residuals is normal. The solid clicle on the probability plot represents the ,
. previously detected potential cuttler. Such outliers would b9 expected to lie to the right of the cumulatNo probability line in the upper half of the plot. A description of the methodology used for preparing the normal probability plot is provided in Section 11.6.5 (Analysis of Residuals for Leak Rate Data), ,
9 e
i l
l 10 7
Table 1 -
Burst Pressure and Leak Rate Date' Base for 3/4. Itch Tubing :
, emu '
(
O t.
MAlu
-10 8 1
m.. .- - . _ _ _ , _ . _ _ - - - - - - - - - . - - _ - _ _ , . . . - _ - - . - - . _ _ . - _ - _ . _ . - . . . - . ~
- _ . . _ - . . . . . - . _ = . _ . _ . _ . _ _ _ . . . _ , _ _ . _ , _ . _ . . _ _ , . . . . _ _ _ . . _ _ _ ._._. _. _ , . __ ,_,
- :, ,i
+
"I r
Table 101--- -t (Continued)- ..i Burst Pressure and Leak Rate Data' Base for 3/4-inch Tubing ,
9 t
a 4
4 l
M
+
t l;
t-l l.
O l -
l- -
'. . M l
I l
.m .
- l. f 9 - .
l I..
f:
.p-g.-9%.---,.qq. ,s4 y'vg #- +e- -e , v gy g es,-i"4-_q _g ..e i. y 4 -- y 9
Figure 10-1: Burst Pressure vs. Bobbin Amplitude 3/4" Tubes, Comparison of 1" & 2"d Order Reeression 4
e a SPI
~
4 10 10
a- M__- -.-4,-- -@ s,--4cdE
- . 4- 4JO-,---4 _.Ah,-_%w+ aa **-4 5 'sb'i.e. ---e+ - - + -#h-r w e,4ep-< e w -M9+- 4 - 4 L Im.ma_
-l
. a I
1 .
. Figure 10 2: Burst Pressuit vs. Bobbin Volts, Final LS 4
3/4" Tubes, Model Boiler & Field Data - -, a,g ; ,
+
'I t
W i
i j
t -
l*
+
i- ,
4 .
5 ..
1 3 -- . _. _
unne -
(
k.
g
.,h L
t
!* s
. ?
I' p.
10- 11' a
4
+
._ , --- --, - w. w. , ,
9
i l Figure 10-3: Examination of Regression Residuals ,
3/4" Tubes, Model Boiler & Field Data ,,,
j l
2 e
i i
I
- l l
l
- i
-l l
I J
l I-l-
l 1
i i
I
?-
i Y
l-
! 10-12
Section 11 -
St.B LEAK RATE CORRELATION 11.1 Introduction
. This section utilizes the model boiler (Section 4) and pulled tube (Section 6) data to develop a correlation of SLB leak rate (primary to secondary leak rate under steam line break oonditions) vs. bobbin voltage. ' A preliminary correlation was reported in WCAP-13494 (in support of Plant R 1 IPC application). Subsequent to that submittal, significant progress has been made in the development of the leak rate correlation. A thorough statistical procedure was implemented as described below. All normalization of eddy current data has been reviewed in detail and resolved by the EPRI APC Committee.
Data from non-leakers are not included in the leak rate correlation. However, a separate correlation has been developed for probability of leakage using all availabile data from 3/4 inch tubing, as described below in Section 11.4. Thus, the data, methodology, and correlation provided in this sectir.n are no longer considered preliminary and hence provide a high degree of confidence in the results.
The developed probability of leakage (Section 11.4) as a function of voltage provides a statistically based estimate of the leakage threshold, Alternate leakage threshold assessments are given in Section 11.3 to estimate a threshold for significant leakage. To assess trends expected for leak rate vs. voltage correlations, a combination of analytical calculations of leakage as a function of crack length and regression fits to crack length vs. voltage data are presented in Section 11.5. These trend analyses help to guide selection of a more bounding leak rate correlation in Section 11.6.
l 11.2 Data Base for SLB Leak Rate Correlation (3/4 inch Tubing)
The database used for the development of the SLB leak rate correlation (primary to
- secondary leak rate under SLB conditions vs. bobbin coil voltage amplitude) for 3/4 inch diameter tubing is derived from model boiler specimens and pulled tubes. All of the data were derived from Alloy 600 tubing with 3/4 inch OD and 0.043 inch nominal wall l thickness. The model boiler test results for 3/4 inch tubing are described in Section 4.
All reported bobbin coil measurements on model boiler specimens are for 550/130 kHz l mix frequency with the 20% holes in the reference ASME standard normalized to 2.75 t volts.
l l The ' pulled tube data included in the database are obtained from Plants E-4 and R-1. The
( bobbin data from Plant E-4 pulled tubes were principally (10 of 12 data points) l- obtained at the APC 550/130 kHz mix and voltage normalization. These data include l cross-calibration of ASME standards to the reence laboratory standard, as described l_ in Section 5.6. This process involved signil. and detailed efforts by two independent l parties (Laborelec and Westinghouse) and review by the EPRI APC Committee such that there is negligible uncertainty in the renormalization procedure. Leak rate
. measurements on the Plant E 4 pulled tubes were performed at room temperature.
These have been adjusted to provide leak rates under operating temperature using l procedures described in Appendix C. The bobbin data from the 1992 pulled tubes from Plant R-1 were obtained using procedures equivalent to those used for the model boiler 11 1 I
database (field ASME standards were calibrated against the reference standard).The pulled tube results are described in Sections 6.6,6.7 and 6.8.
The SLB leak rate data are normalized to a primary pressure of 2350 psia and to the secondary side pressure of 15 psia (atmospheric) at operating temperature.
Renoimalization from the measured leak rate conditions for SLB leak rates (as described in Sectie 7.4) is developed in Appendix C. The resulting SLB leak rate database is summarized in Table 101 for both the model boiler speciment, and the pulled tubes.
Data from non leakers are not included in the leak rate correlation. The database includes 47 model boiler specimens and 12 pulled tube intersections.
11.3 Leak Rate Threshold Assessment As a crack is initiated and grows to certain size (in depth and length) It becomes detectable through eddy current inspection. Such a crack indication would have a signal amplitude associated with it. As the crack grows, so does its signal amplitude. When the crack becomes throughwall, it would have a significant voltage amplitude signal.
Although extremely short cracks may be construed to have small amplitudes, in practice, a throughwall corrosion crack will have some minimum signal amplitude, it is not known explicitly what minimum voltage amplitude may be associsted with such a crack.
In order for a crack to result in leakage across the tube wa!!, the crack must be throughwall. Further, for an OD initiated throughwall crack to leak, it must have some minimum length at the tube ID. This is exemplified by tubes pulled from Plants A 2 and B-1 which had throughwall cracks but did not leak during leak rate tests at SLB conditions. This is also supported by the scarcity of indications with leaks for bobbin coil amplitudes below 3 volts. Thus,it may be concluded that there would be an eddy current voltage threshold below which corrosion cracks would not cause leakage. A -
voltage threshold for SLB leakage is assessed below from the large body of available data.
11.3.1 Threshold Based on Destructive Exam Depth of Pulled Tubes Figure 11-1 is a plot of bobbin voltage versus maximum depth of cracks in pulled tubes l from operat og steam generators with 3/4 inch diameter Alloy 600 tubing. There are 45 pulled tut
- specimens from five different plants, both domestic and European. Due to the differench in crack morphology among the various pulled tube specimens, the plotted data is scattered as expected. However, an increasing trend between bobbin voltage and maximum throughwall depth from destructive examination is clearly visible.
On the semi-log plot, a linear relationship between the voltage and destructive examination depth is broadly indicated. A regression fit of the data from the partial depth cracks yields an amplitude of ~1.7 volts for a crack reaching throughwall depth, and ~1.2 volts at 90% confidence on the mean. Virtual!y all of the data from throughwall cracks fall above the 90% mean confidence fit. Since a crack must be ,
throughwall to cause leakage, it then follows that the minimum bobbin voltage threshold for leakage in 3/4 inch tubing is about 1.2 volts.
Figure 11-2 is a similar plot for pulled tubes of 7/8 inch diameter. A similar trend as for the 3/4 inch data may be observed. The data is scattered and a linear relationship on the semi-log plot is displayed. A regression fit of the partial depth cracks suggests a 11 - 2
bobbin amplitude of about 1.6 volts from the 90% mean confidence fit for,s crack in 7/8 inch diameter fubing to reach throughwall depth. Again almost all of the pulled tube data from throughwall cracks lie above this value.
11.3.2 Threshold Based on Leakage Data A considerable amount of data exists on leak rate of ODSCC flaws in pulled tubes and
- model boiler specimens. This database includes both 3/4 inch and 7/8 inch diameter tubing. The database consists of 74 model boiler specimens and 93 pulled tube intersections. The data from 7/8 inch tubing is at 400/100 kHz differential mix with the 20% holes in tne ASME standard normalized to 2.75 volts in the mix. Similarly, the 3/4 Inch tubing data is normalized to 2.75 volts for the 20% holes in the ASME standard at the 550/130 kHz differential mix.
In order to combine the data from the 7/8 inch and 3/4 inch specimens, a conversion factor equal to the square of the diameter ratios is applied. This factor results from the fact that the ASME standard hole size is the same for 3/4 and 7/8 inch tubing. Thus, to convert the 7/8 inch data to the same basis as the 3/4 inch data, the voltage amplitudes from the 7/8 inch data (at 400/100 kHz) were divided by the factor 1.36. The results were then combined with the 3/4 inch database.
The leak rates were, for most (131 of 167) cases, the direct result of measurements in the laboratory under SLB conditions. For the other (36) cases, laboratory data on leak rate measurement was not available and the likelihood of leakage was inferred from crack morphology (throughwall depth and length) obtained from destructive examination.
The data was classified into leaking and nonleaking specimens. A frequency Stribution of voltage amplitudes (corresponding to the 3/4 inch data normalization) in each classification was determined. This is shown in Figure 11-3 as a stacked bar chart. The number of leaking specimens in each voltage range out of the total number in that range is also shown listed at the top of each bar in the figure.
The ratio of the number of leaking specimens in a voltage range (bin) to the total number of specimens in the bin was calculated from the above frequency distribution of voltage amplitudes. This result, probability of leakage, within each voltage range is plotted as a bar chart in Figure 11-4.
The above results (Figures 113 and 11-4) were developed using data from both the 7/8 and 3/4 inch diameter tubing. If the 7/8 inch data is excluded and only the SLB leak rate data from 3/4 inch tubing is used, the results are not changed significantly. This is displayed in Figures 11-5 and 116.
For the 3/4 inch tubes, the data supports no leakage under SLB conditions for indications up to 1.0 volt (bobbin data for 550/130 kHz mix with 20% hole ASME standard normalized to 2.75 volts). A low probability of leakage between 1 and 3 volts is indicated by the data, with only one leaker at 1.13 volts below the more frequent leakage
, data at >3.5 volts. The contributing factor to the low voltage leaker at 1.13 volts is discussed in Section 11.3.5.
11 - 3
11.3.3 Threshold Based on Bobbin Voltages of Non-leaking Specimens in this paragraph, a voltage threshold for SLB leakage is derived using data from non leakers. All available data for corrosion cracks with throughwall or near throughwall indications is used in this evaluation. Specimens which had crack depths of 90% or higher (from destructive examination) and did not leak during the leak test are listed in Table 11 1. It may be noted that, in the case of 3/4 inch tubing, these flaws had signal amplitudes in the range of 1.3 to 5.7 volts. This sample has a mean of 3.30 .
volts and a standard deviation of 1,47 volts. These values represent voltage Indications for the deepest cracks that do not cause leakage during SLB conditions. Even using a lower 90% confidence level, this data suggests a SLB leakage threshold of 1.3 volts it must be recognized that this estimate is very rudimentary and conservative since it is derived from non-leakers and since only a limited number of specimens were available.
11.3.4 Threshold Based on Throughwall Crack Length The threshold for SLB leakage can also be assessed by evaluating the lowest bobbin voltages resulting in leakage at SLB conditions and by evaluating the throughwall crack length generally required for measurable leakage. If the throughwall crack length associated with measurable leakage can be defined, the voltage vs. crack length relationship of Section 11.5 above can be used to assess the voltage threshold for leakaga.
The crack length method for estimating a voltage threshold for leakage provides a more physical insight into the threshold estimate.
It can be noted that significant efforts were applied in the 3/4 inch tubing model boiler specimen preparation to obtain the lowest voltage associated with leakage. In the model boilers, leakage is monitored by sensing for lithium which provides a leakage sensitivity -
of about 3x10 3 Uhr. Upon detection of any leakage, the model boilers were shutdown and the tube (typically 4 to 6 TSPs) was removed for NDE inspection. TSP Intersections with bobbin indications above about 1 volt were removed from the tube for further NDE, -
l leak and burst testing. The smallest bobbin voltage from this program having a j measurable leak rate in the leak test facility (capability to measure down to 10-3 to 10-41/hr) was 4.24 volts (No. 601 1). It is possible that a specimen at 2.79 volts (No. 595-2, throughwall crack length = 0.17 inch) was detected in the model boilers j with no measurable leakage in the leak test facility.
l As discussed above, the leakage threshold can be assessed by examining crack length data, i Table 11 2 shows specimen throughwall crack lengths for no leakage, leakage <1.0 I liter /hr and between >1.0 and 6.0 liter /hr. The largest throughwall crack which exhibited no leakage was 0.17 inch long. The smallest throughwall crack length for-which even small leakage (<1 liter /hr) was detected is 0.016 inch. This pulled tube I indication, which leaked at 0.02 liter /hr (10-4 gpm), is discussed in Section 11.3.5.
Specimen 601 1 had a 0.05 inch throughwall corrosion crack with a thin ligament suspected to have partially opened at SLB conditions. This specimen had no leakage at operating condition pressure differential and 0.33 liter /hr at SLB conditions. For this ,
type of indication (above APC repair limit), voltage is a better indicator of leakage l potential than even throughwall crack length or total crack length (0.29 inch). Overall, .
l the average TW crack data of Table 112 indicate that throughwall crack lengths >0.07 .
inch are generally required for SLB leakage and throughwah lengths >0.13 inch are generally required for leakage >1 liter /hr (0.0044 gpm). All leakers for 3/4 inch tubing in Table 112, except R9C91-TSP 3 have bobbin voltages exceeding the full APC 11 - 4 I
l
repair limit of 2.2 volts for V. C. Summer. From Figure 11 11 (discussed later) It may be noted that a corrosion crack with a throughwalliength of 0.07 inch is expected to have a bobbin amplitude of about 2.5 volts The results of Figure 11 11 Indicate a bobbin voltage of about 1.3 volts is generally required for throughwall crack penetration, with 2.5 volts corresponding to the average crack length of 0.07 inch required for leakage.
11.3.5 Voltage Threshold Considerations for SLB Leakage (3/4 inch Tubing) in the above sections, the SLB leakage threshold was evaluated from different perspectives: voltage indications required for throughwall cracks (~1.2 volts), voltage threshold from leak rate data (-13 volts), voltages of non-leakers with throughwall and near throughwall degradation (~1.3 3.3 volts), and crack lengths required for leakage (~2.5 volts). In all cases, the data show that the bobbin amplitude threshold for significant leakage (20.3 liter /hr or -10 3 gpm) across the tube wallin a 3/4 inch diameter tube is greater than about 2.5 volts.
Small leak rates below about 2.5 volts can occur; l.e., Plant R 1, tube R9091 at TSP 3 with a bobbin voltage of 1.13 volts is an example of potential leakers between about 1.0 and 2.5 volis. As discussed in Section 3.3.1, neither the burst crack nor the second largest macrocrack were found to have throughwall corrosion. The voltage response for this tube would be dominated by these macrocracks, which had maximum depths of 82%
and 91% and average depths of 47% and 49%. Throughwall corrosion was found only at another short crack (Figure 3-20) with a throughwall corrosion length of 0.016 inch, which is expected to have opened to about 0.044 inch during the SLB leak test. No leakage was measured at normal operating conditions. This is the only leakage occurrence found in both the 3/4 and 7/8 inch tube diameter databases which did not include throughwall penetration and leakage from the burst crack. If the dominant macrocrack had penetrated throughwall, the voltage response would have significantly -
increased. Leak rates for indications such as found for R9C91 at TSP 3 are inherently small, such as the 10-4 opm for this indication, due to the very small throughwall penetration and overall short crack length.
Based on the above, leak rates should be related to a voltage threshold. The threshold for zero leakage is expected to exceed 1.0 volt, while the threshold for meaningful leakage
(~0.3 liter /hr or -10 3 opm) is expected to exceed 2.5 volts. The SLB leak rate methodology and analysis for V. C. Summer is intended to be conservatively applied.
Therefore, a leakage threshold of 1.0 volt is applied for V. C. Summer. That is, all indications above 1.0 voit have a probability of leakage as developed in Section 11.4 (Figure 11 7) and a conservative leak rate (Figure 11 16) for Indications in the range of IPC applications (s3 volts).
The best fit or maximum likelihood estimate for probability of leakage given in Figure 11-7, developed below in Section 11.4, supports a 1.0 volt leakage threshold.- The
& upper 95% bound of Figure 117 extends the leakage threshold below 1.0 volt. This lower voltage threshold estimate is inconsistent with the threshold estimates developed -
above and is judged to result from the limited database in the range of about 2 to 6 volts.
In either case, both the 1.0 volt probability of leakage at +95% uncertainty (-5%) and the SLB leak rate at +95% confidence (~0.21/hr or <10-3 gpm) are very low such that' the 1.0 volt threshold can be applied with negligible error in the SLB leak rate.
11 - 5
11.4 Probability of SLB Leakage Versus Bobbin Voltage -
The above discussion and the supporting data indicate that even when the voltage amplitude of a corrosion crack exceeds the leakage threshold, there is a likelihood of no leakage. In order to quantify this fact, the probability of SLB leak 9ge as a function voltage is derived in this section.
11.4.1 Database for Probability of Leakage .
The current evaluation of probability of SLB leakage is limited to 3/4 inch diameter tubing. Hence, the database used here is limited to the 3/4 Inch specimen results given in Table 101 and Figure 11 1. Figure 11 1 includes pulled tubs results for which leak or burst tests were not performed. The data set for this analysis consists of a total of 85 pairs of test results for which leak test results are available, or for which it could be verified from destructive exam results that leakage would not occur at a pressure less than or equal to the postulated SLB differential pressure. When leak tests were not performed, the maximum crack depths were evaluated to define the indications as non leakers or leakers. An example would be for a very small bobbin amplitude wherein a large burst pressure was observed during burst testing (e.g., for a burst pressure test in which no leakage was observed to a pressure in excess of about 3100 psl at room temperature for a tube with a material flow stress of 75 ksi). The database consists of 41 model boller specimens and 25 pulled tubes indications for which SLB -
leak rates are known from direct measurement. In addition, there are 19 pulled tube specimens for which leak rate tests were not performed and whether or not they would leak was assessed on the basis of the crack morphology from destructive examination.
11.4.2 Statistical Evaluation for Leakage Probability .
An analysis was performed to establish an algebraic relationship that could be used to predict the probability of leakage during a postulated SLB as a function of bobbin amplitude voltage. Two approaches were considered for the analysis. The procedure for the first approach would be to segregate the results into a series of discrete bobbin amplitude ranges, called bins, based either on the actual voltage observed or based on the logarithm of the bobbin voltage. This would be followed by the preparation of a l cumulative histogram of the results, e.g. Figure 11-4, and the fitting of a smooth polynomial, or a cumulative normal distribution type curve through the results.' There are, however, two significant drawbacks associated with this type of an approach. The first drawback is that the shape of the plotted histogram is dependant on the number of subdivisions used to segregate the data range. The second drawback is that there would be no direct way of establishing confidence bounds on the model, although binomial limits could be calculated for each bin.
To consider the second drawback further we note that the probability of 1eakage must be -
limited to a range of 0 to 1. A correlating equation to fit the data could be achieved by employhg the method of least squares (LS) or maximum likelihood (ML) to estimate the '
- parameters of the equation. if the expression was based on correlating probabihty as a function of leakage, the upper confidence band for the resulting expression would exceed unity for some voltages. Ukewise, for lower voltages the lower confidence band would yield probabilities less than zero.
11 - 6
in general, the criteria for a good estimator are that it be unblased, consistent, efficient, and sufficient. The parameters of a correlation curve from either LS or MLE analysis will satisfy the above requirements if the requirements for performing the analysis are met. However, for the probability of leak data, the estimating curve will not be strictly consistent for a sample size less than infinitely large due to variability inherent in establishing the subdivision size. It is noted that although the objections of the second drawback might be overcome by correlating voltage to probability, the requirement for a consistent estimator might not be met. It was therefore decided, that the first analysis approach would not be pursued.
The second approach to the analysis stems from the observation that the data may be considered as samples from a dichotomous population. This means that the data are categorical, and that the number of categories is two, i.e., either no leak or leak.
However, for an analysis treating the data in this manner a normal linear correlation model of the type P(leak l volts) - ao + at f(volts) + e (11 1) where f(volts)is either volts or log (volts), is not appropriate since normal distribution errars, e, do not correspond to a zero (no leak) and one (leak) response, it is appropriate to analyze the data in this situation by non-linear regression analysis, and, logistic regression in particular, since the data is dichotomous. Letting P be the probability of leak, and considering a logarithmic scale for volts, V, the logistic expression is i
P - 1/(1+e-[ao+ at og(V) (11 2)
~
This can be rearranged as in[P/(1 P)] = ao + ai log(V) (11 3) where the logit transform is defined to be logit(P)-In[P/(1 P)]. (11 4) The model considers that there is a binomial probability of leak for each value of voltage. The objective of the analysis was to find the values of the coefficients, aoand ay, that ' best fit the test data. Since the outcomes of the leakage tests are dichotomous, the binomial distribution, not the normal distribution, describes the distribution of model errors. It would, therefore, be inappropriate to attempt an unweighted LS regression analysis based on equation (113). The appropriate method of analysis is based on the principle of maximum likelihood. The application of ML leads to estimates of the equation parameters,i.e., aoand af, that are such that the probability of obtaining the observed set of data is a maximum. The results of applying ML to the dichotomous outcomes are the likelihood equations, I 11 - 7
n E [P i - P(vi)] = 0 (11 5) i=1 and n E vi [Pi P(vi)] = 0 (11-6) , i=1 13:re, Pi is a test outcome and P(vi) is an expected outcome based on the input value of the voltage using equation (112), where vi-log (V i ). Since these equations are non-linear in the coefficients aoand ay, an iterative solution must be determined. Two evaluations of the parameters were performed, one using a commercially available statistics program with logit fitting capability, and a second based on manually trying to iterate to a solution based on weighted least squares. The purpose was to provide an independent verification of the results from the commercial program. The accepted measure of the goodness of the solution is the deviance, n D - 21 { P In[P i /P(vi)] i + (1 P )ln[(1-P;)/(1 i P(vi))] }. (11 7) i=1 The devlance is used similar to the residual, or error, sum of squares in linear . regression analysis, and is equal to the error sum of squares (SSE) for linear . regression. Fct the probability of leak evaluation P iis either zero or one, so equation
- (11-7) may be written ,
n D - 2 E { P in[P(vi)] i + (1 P )ln[(1 i P(vy] }. (11-8) 61 Both evalursr. provided similar deviance valuec and the final solution obtained was logit(Pi ) = 5.276 - 9.185 log (V), i (11 9) Asymptotic confidence limits for each individual logit(P i ) are found as logit(P)iZg o[logit(P)] i (11 10) where 100(1-a)% is the associated probability for a two-sided confidence band. , 11 - 8
i The standard error of logit(P;) is found for each voltage level as - o [logit(P )) i - 1.1, { log (Vi )}TJ y, [1, {log(y;)}TJ T (11 1j) j where V, is the estimated variance-cwarlance matrix of the parameter estimates. Letting f);- logit(P)i (11 12) the upper and lower 100(1 cr/2)% one-sided confidence limits are 1/{1 + eillii Za/2 c(Tli)l} (11 13) The upper bound values were calculated for each voltage / leak pair. The results are shown graphically in Figure 117. 11.5 General Trends for SLB Leak Rate Correlation An evaluation has been completed that provides results supporting the recommended correlation methodology employed for SLB leak rate versus bobbin voltage. The evaluation establishes leak rate versus voltage from:
, 1) Formulation of a throughwall (TW) crack length (L) versus voltage (v) correlation from available data using regression analysis {L-fj(v)).
- 2) Calculation of leak rate (O) as a function of L using the CRACKFLO computer e e.
Formulate a simplified relationship of O as a function of L through regression analysis of CRACKFLO predictions {O-f2(L}}.
- 3) Development of a correlation between O and v from the above. Substitute the formulation L-f i(v) into 0-f (L)2 to get O=f 3(v). Compare O=f 3 (v) to the direct correlations.
11.5.1 TW Crack Length vs. Voltage The bobbin voltage versus TW EDM slot data of Figures 5-1 and 5-2 are utilized to set the form of the equation to be used in the regression analysis of TW stress corrosion crack length versus bobbin voltage since this data covers a much broader range of TW length and voltage than that reflected in the crack data. The form of the equation is found by regression analysis to be y-a+b*v3+c'in(v) as shown in Figure 118. An attemate expression of the form - y-a+b*v .5ois found to be more accurate for TW slot lengths of less than 0.25 inch (Figure 11 9). A dependence of length on square root of voltage is also more consistent with eddy current theory for short, single throughwall cracks but does not adequately represent -
- voltage saturation at long crack lengths.
Regression analysis of the TW crack lengths versus bobbin voltages of a limited set of the data provided in Tab:e 10-1 using the two forms described above yields the solutions presented in . 11 - 9
Figures 11 10 and 11 11. Three data points (590-2,5981 and 6013) of Table 101 were not included as they have multiple, large voltage cracks which excessively increase the voltage compared to a single throughwall crack, in addition, specimen 6016 was not included, as the throughwall length is the sum of two cracks separated by a ligament such that the voltage and leak rate are not representative of a single throughwall crack. Both formulations give very similar results up to about 17 volts. The data included are primarily samples with single dominant cracks where the voltage is reflective of the crack thc.t in fact was throughwall. The data is somewhat scattered and limited in number. The curves do. - however, show expected trends in that the voltage is >0 (~1.3 volts) for a TW crack, that the curves rise quickly at low voltages (s3 volts) and then tend to plateau before reaching a long enough TW length for the voltage to rise quickly again due to bobbin co!! vc!! age Insenshlvity to very long cracks (21 inch). 11.5.2 Leak Rate Versus TW Crack Length The CRACKFLO code described in WCAP 12871, Rev. 2, is utilized to predict leak rate as a function of axial TW crack length for SLB condulons. For simplicity, i.e., to avoid making many specific runs on CRACKFLO, regression analysis is used to fit a subset of the CRACKFLO solutions required. Two correlations are obtained to provide an accurate representation of the CRACKFLO solution over the range of interest. Figure 11-12 provides the expression used for less than Log (0.25") and Figure 11-13 for values greater than or equal to Log (0.25"). It should be noted that comparisons of measured versus predicted values by CRACKFLO indicate that CRACKFLO tends to overpredict SLB teak rate (Figure 11 14). The two points with crack lengths <0.1 inch had " bathtub" flaws (thin ID ligament) for which tearing of the ligament is expected at SLB conditions. CRACKFLO uses only corrosion throughwall crack lengths - i can be expected to underestimate leakage when tearing of ligaments increases the effective throughwall crack length. - 11.5.3 SLB Leak Rate Versus Bobbin Voltage Combining the crack length versus voltage equations with the SLB leak rate versus crack - length equations results in the two trend curves plotted in Figure 11 15. Both trend curves reflect the TW crack length versus voltage characteristic shape of the Step 1 correlations giving consistent predictions up to about 17 volts, in the low voltacc range (=1 volt), a large variation in predicted leak rate is noted between the two trend curves of Figure 11-15. However, the expectation of a leakage threshold is confirmed. As noted above, the crack length dependence on square root of voltage (open circles in Figure 11 15) is judged to be the better correlation for <0.25 inch crack lengths (~8 volts), but there is insufficient data to clearly support the differences between the two trend curvas, 11.6 SLB Leak Rate Versus Voltage Correlation The bobbin coil and leakage data of Table 11-3 were used to determine a correlation between SLB leak rate and bobbin voltage amplitude. This is not to say that a " formal" functional relationship,in the sense of one variable being depende'it on the other, exists between the variables since the amount of leakage is not caused by the bobbin voltage and vice versa. Both of the variables considered are really mainly functions of a third variable, namely the crack morphology. Whlie the variation in crack morphologies is - essentially infinite, suitable descriptions can be effected based on the depth, average depth, profile description, etc. However, the characterization of the morphology is not essential to this analysis since a relationship can be independently established for two 11 - 1 o
offspring variables. Since both bobbin voltage and leakage are offspring variables the results of the cx>rrelation analysis do not establish a formal relationship between the variables, however, the results do estaMish a " working" relationship that can be employed for the prediction of one variable from the other. ., 11.6.1 Selection of Coordinate system in order to establish a correlation between the two variables, which are paired, but expected to have independent variances, the method of least squares (LS) curve fitting was employed. The simplest functional form is a linear relationship of the type y = ao + atx (11 14) wher; the vadables x and y rnay each be considered to be linear or logarithmic. In addition, the choice of the regressor variable is not pre deiefinined. Doth verlab!es are assumed to be subject to random fluctuations which are normally distributed about the mean of the variable or the logarithm of the variable with a mean of zero and some unknown, but reasonable variance, it is also assumed that this variance is constant, or uniform, over the range of interest of the variables. In practice this may not be the case: however, any non-uniformity present would not be expected to significantly affect the analysis outcome, and can be tested at the conclusion of the analysis. Analyses were performed to determine the optimum nature of the variable scales,l.e., linear versus logarithmic, and the appropriate selection of the regressor variable, it - was initially concluded that the most meaningful correlation could be achieved by considering the log of the leak rate as the regressor and the log of the voltage as the predicted variable. Thus, the functional form of the correlation is
- log (V)- ao+ agiog(L) (11 15) where t is the leak rato and V is the bobbin voltage. The final selection of the form of . the variable scales was based on performing least squares regression analysis on each possible combination and examining the square of the correlation coefficient for each case. A summary of the results of the calculations is provided as follows:
Index of Determination, r2, for Various Selections of Coordinate Scales Lenk Rate Scale
' Voltace Scale Linear Logarithmic Linear 21.9 % 47.1 %
Logarithmic 15.7% 57.6 % These results strongly suggest that the appropriate choice of axes scales is log-log. The number of data points used for the above evaluations was 35. The corresponding critical value of 8 for significance at a level of 99.9% is approximately 28%. Thus, the log-log regression with an 8value of 58% is significant at a level greater than 99.9%. This is also true for the linear log set of axes, however, the coefficient is signif'cantly - better for the log log set. .'~ Guidance on the appropriate choice of the regressor variable is based on the knowledge, from CRACKFLO analyses, of the approximate value of the slope of the leak rate relationship (Section 11.5). The CRACKFLO results show the slope of the log (V) on 11 .1 1
log (O) regression line more closely matches the CRACKFLO results in the lower voltage range (<8 volts). The CRACKFLO results indicate the potential for a slope change between about 8 and 10 volts. it was concluded, therefore, that both regression lines would be determined and the higher of the two for any voltage level would be taken as the upper bound leakage. As shown below, this two slope approach results in a leak rate vs. voltage correlation having the same slope trends as the more theoretically based trend , analysis using CRACKFLO, as shown by Figure 11 15. 11.6.2 Correlation Analysis and identification of Outilers - In order to determine if the parameters of the relationships were being biased by the presence of unduly influential data points, a least median of squares regression analysis was performed on the entire data set for both regression directions. The analyses were repeated for SLB AP levels of 2.650 ksi and 2.335 ksi. For the leak rate on bobbin amplitude analyses three points were identified as potential outliers for a AP of 2.650 ksi with one additional point being identified at 2.335 ksi, For the bobbin amplitude on leak rate regressions four data points were identified for a AP of 2.650 ksi with one less point being identified at 2.335 ksi. Only two points were identified that were common to all four analyses. These were model boiler specimens 5981 and 598 3. Examination of the testing program information revealed that specimen 5981 had a very large number of deep, high voltage cracks with a modest (0.37") throughwall crack length such that the specimen represents an extreme condition and would be expected to be an outi;er. Specimen 598 3 had a significantly large throughwall crack length (0.27") with very low leakage such that the leak rate measurement is questionable. Examination of the burst pressure result for the spccimen Indicated behavior typical of.the average. The residual to scale ratios from the rogression analyses were 5.77,-5.44,3.33, and 3.82 considering the bobbin amplitude to be the regressor for the first two cases and the leak rate as the re for the second two. The associated probabilities of occurrence are 1.4x10 8, gressor . 1.3x10-8,0.0004, and 0.0001 respectivel of 35 specimens are 1.4x10-7,9.5x10 0.014,7,y. The corresponding and 0.003 respectively. Sinceexpectations the in a test rate was abnormally low it was decided to treat the leak rate result as an outlier and the ' data was omitted from further analysis. All of the data and the correlation line fit with the two outliers removed are shown on Figure 11 16 and 11 17 for SLB AP's of 2.335 and 2.650 ksi respectively. Data points for which there was no leakage are not shown, and were the subject of a separate analysis (Section 11.4) to determine the probability of leakage being observed for a given bobbin amplitude. 11.6.3 Error-in Variables Analysis The potential effect of measurement errors in the regressor were discussed in the section on burst pressure relative to bobbin amplitude. it was noted that the variance, OX, of a variable, say X, with measurement error, as estimated from the data, consists of two parts, the intrinsic variation, Ox, of the variable, x, and the variation, Um, due to measurement error, m,i.e., cX2 . g,2 + o,2 (11-16) A Wald-Bartlett type of analysis was performed for each direction of regression. Considering the leak rate as the regressor resulted in a line indistinguishable from the standard regression line. Considering the bobbin amplitude as the regressor resulted in a line with a slightly larger slope than the standard fit. This was *o be expected since the - , previous use of the technique for the burst pressure correlation indicated that measurement errors for the bobbin amplitude were not significant. Since the decision was made to proceed using both regression lines (leak rate on bobbin amplitude and vice 11 - 1 2
versa), the results of these analyses become moot. This is because the two lines cross at the centroid of the data. Therefore, omitting the consideration of errors results in a larger prediction of leakage for alllevels of bobbin amplitude. 11.6.4 SLB Leak Rate Correlation for 3/4 Inch Diameter Tubing 4 The final fits of the data are shown on Figures 11 16 and 11 17 for SLB pressure differences (AP's) of 2335 psi and 2650 psi, respe#ely. The correlation lines are given by log (L)- ao + ai og(V), l (11 17) at log (V) = bo + bi log (L). (11 18) The coefficients for the above equations are shown below. Regression Coefficients for SLB Leak Rate to Bobb!n Amplitude Correlation SLB at SLB at 2.335 ksi 2.650 ksi ao -1.627 1.613 ai 2.552. 2.795' bo 0.781 0.731 b; 0.206 0.206 In addition to the least squares regression lines, the upper 95% one-sided, simultaneous confidence bound, upper 95% prediction bound, and upper 99% prediction bound were
, determined. Each is shown on Figures 11 16 and 11-17. Figure 11-17 also includes the CRACKFLO results for comparison.
The expected, or arithmetic average, leak rate, Q, corresponding to a voltage level, V, from the above expressions was also determined for the lower (s8 volt) voltage range. G!nce the regression was performed as log (O) on log (V), the regression lina represents the mean of the log of the leak rate at each level of bobbin amplitude. Thus, the expected Q for a given V is given by E{QlV) = 10a o+ajtog(v)+0.5tn(10)o^2 , where a2is the estimate variance of log (O) about the regression line, A plot of the expected leak rate is provided on Figure 11 18. Examination of the figure Indicates that below a voltage level of ~3.6 volts, the upper 95% confidence band on log (O) is conservative relative to the expected mean leak rate. The statistical analyses indicate that the correlation of voltage on leak rate is the - preferred regression fit, although both fits are conservatively applied to bound the leak rate vs. voltage for the V. C. Summer IPC. The slope of leak rate vs. voltage based on the regression analysis of voltage vs. leak rate is
. 1 1/bt = 4.9 l l
l 11 - 13
which is consistent with the trend analysis of Section 11.5 and the general dependence of leak rate on crack length. As shown in Figure 11 15 for the expected trends from CRACKFLO, the leak rate has essentially a step change from 0 to -0.01 liter /hr at the voltage threshold for throughwall penetration and then a modest slope to about 15 to 20 volts. Above about 15 25 volts, the slope would be expected to increase significantly based on Figure 11 10 as the voltage become:, less dependent on throughwall crack . lengths above 0.8 inch. However, this voltage range is well above the range of interest for APC applications and <20 volt Iridications are appropriate to develop the leak rate correlations. As noted in Section 11.3.5, a leakage threshold of 1.0 volt is applied for the V. C. Summer analyses. As developed in Section 7, a AP of 2235 psiis appropriate for SLB analyses and Figure 11 16 is used in Section 12 for leak rate analyses. 11.6.5 Analysis of Residuals As previously noted, the correlation coefficients obtained from the analyses indicate that the log log regressions at the two SLB AP's are significant at a level greater than 99.9%. Additional verification of the regression is obtained by plotting the ordered residuals on normal probability paper, if the residuals so plotted approximate a straight line it is concluded that they are normally distributed, in addition, the information on the plot may be used to determine if the mean is approximately zero. To prepare the plot the residuals are sorted in ascending order and then plotted against an ordinate percent probability value given by 100(1-0.5)/n, where n is the number of data points used in the regression and lis an index ranging from 1 to n. If the unit area under the normal curve is divided into n equal segments, it can be expected, if the distribution is normal, that one observation (residual) lies in each section. Thus, the . /h observation in order is plotted against the cumulative area to the middle of the ith section. The factor of 100 is used to convert the scale to percent probabilities. A plot of the residuals for each SLB AP is provided on Figure 11 19. For each case the residuals approximate a straight line and it is concluded that they are normally distributed. The line on each plot was Ctted by engineering judgement and does indicate that the mean of , the residuals is approximately zero. Thn calculated mean for each case is ~-0.1 1/hr. For this type of plot, outllers in the data tend to appear on the f ar left in the lower half of the residual normal plot and on the far right in the upper half, i.e., large negative and positive residual values. The outliers for the analyses are shown as solid circles on the - two plots. The additional point that appears somewhat outlying in the lower half of the plots is a value that was identified as a potential outlier in some, but not all, of the robust regression analyses. Given the results of the normal probability plots no additional information would be available from a scatter plot similar to the one used for the residuals of the burst pressure versus bobbin amplitude regression. 11 - 1 4
-Table 11 1 ,
i Non leaking 3/4 inch Specimens with Throughwall or Near Throughwall Cracks 0)
+ Spedmen Mar. Denth hhhin Amolitude.vo!ts Plant E 4 9 Plant R 1 595 2 592 6 592 7 591 2 601 Avera0s . .
Standard Deviation Note 1: Maximum depth > 90% from destructive exam. l i. t u
.:e j .- 11 15 l
V .:
l Table 112 Dependence of SLB Leakage on Throughwall(TW) Crkck Length 1W - No Lankana TW <1 liter /hr TW >1 liter /hr. & c 61/hr & Bobbin Volts TW Lenoth h Bobbin Volts TW Lenoth A Bobbin Volts TW Lenoth _,0
- Notes:
- 1. Bathtub flaw (thin OD ligament) reasonably expected to open crack at SLB AP.
9 11 --1 6-
Table 113 . SLO Leek Rate vs. Bobbin Amplitude- . Data for 3/4" x 0.043' MA Alloy 600 Tubes . g. M 9 te
.9.
6 4 11 - 1 7
+
f Figure 11'- 1: Bobbin Coll Voltage vs. Maximum Examination Depth 3/4" Pulled Tubes Data, Destructive Examination _ "'8 9 O W
+
D im GM - 11 18-
Figure 11-2: Bobbin Coil Voltage vs. Maximum Examination Depth- ,,,
- 7/ 8" Pulled Tubes Data, Destructive Examinationi _.
i e $ p. J i- o E O
- M o
11-19
, ~ .
a,g t 6 Figure 113,': Frequency Distribution of Leakers and Nonleakers (at SLB Conditions)'versus Bobbin Voltage for 3/4 inch Voltage Normalization:-
.11-20
, y - , ,4.m
-. , .. . .. ..~ . . .. .. . .- . . - . . , _ - . . . . - . . . - - ~ - . -.
-4 :
1 4 . . t i 4 a,g -
.A .
s it
- . ..t.
e i J= ,
- -i s
W 4 + 1 ? ,x t d 1 r t I ? 1 P ( s [r ?A - i.3. I i : Figure 114. Probability Distribution of SLB Laakage versus Bobbin Voltage? , i :- 0 l' r o - 11 21 g s s , 4 3 L v YC et F - ,'b e-'v.e , , .,.w , , , , ',
a,g b
~[
t Figure 115. Frequency Distribution of Leakers and Nonleakers (at SLB Conditions) versus
- Bobbin Voltage for3/4 inch Data 11 22 -
. . - - - . . ~. . . . . . .- - . . ..--, .. .. .. -. .. .
+
a,g . ; t e i r i
. ~
.-e
$' 4 I
.1
. Flours 116. Probabiiny Distribution of SLB Leakage versus Bobbin Voltage for 3/4 Inch :-
Tubing. - V 11-23 d 'T e TTW I Y E 'T *- m 4,r--+
- Y
i 1 1 4 I . .,4 J 4 s 1 i i 1 Figure 117: Probability of Lenk vs. Bobbin Amplitude , j 3/4" Tubes, Model Boller & Field Data _ ag i 1 4 Geu 3, k 0 h
~
J i 4 i 4-f 1 J I
- i. a f
t i J l i l i i j ' 4 6? I t I l-. I b 4 4 3 h enne i 4 i n E 4 l i. 4 4
~
1124 = r i
. t w M- y rw yy.,p -yrg-w-wr --
gyTN-v'~N'y e-M T- -*m*C -t't#%T= M' 'w' e Tr- eh -W'e q' y- F-Y e *vb-' w- T wp- ,-w-T -t -- g- --
-_ _- ._ . . - -. ~.
. Figure 118 ,
TW EDM Slot Length Vs Voltage-3/4&7/8" , s.b F sq' S 4 11 25
Figure 119 M ECM Sot Length Vs Voltage-3/4&7/8" a,b .
+
0 4
+
9* 9 11 26
i Figure 11 10 TW Crack Length Vs Voltage - 3/4" Tubing ,,3 0 0 9 l l 11 27 l
Figure 11 11 TW Crack Length Vs Voltage - 3/4" Tubing
,,3 4
e 4 J meme b 4 6 11 28
Figure 11 12 SLB Leck Rote Vs Crock Length-CRACKFLO ,,3 e 4 e N ensu 4 e 11 29
Figure 11 13 - SLB Leak Rate Vs Crack Length-CRACKFLO ,,3 9 6 ' 4 unma - e 11 30
- . . . . . . . . . . - . . . - . . . -.. ..- _.... - _ .- . . - . . --_ . .. .. ~ = . .
1 I i i, , i 1 l
- Figure 11 14 l Comparison Of CRACKFLO To 00 SCC Leakage
3/4' Tubing Data- a,o
?
< j i t 4 i r i l 1 E . t 1 . + t e i
- i 4-t I
i f 1. l-
~ -
2 j.- e
,- 11 31-my-- 4 we ~
Q w--ww--n -- q m e - e,yy --we-e ,. - p- yr.m.en.-sowwn & Y
.._ . - - - - - . - , . . . _ = _ - _ - . . . . . - . - . .-. _- -_._
l Figure 11 15 Trend Analgals For Predicting SLB Leak Rate Vs Bobbin Voltage . ag
. ~
4 4 O 9 eumunia S 9 11 32
Figure 11 16: 2335 psi SLB Leak Rate vs.- Bobbin Amplitude 3/4" Tubes, Model Boller & Field Data a,g M i I e I 4 memen eums e 1
-11 33
.. _ , , . - _. .u. . _. a _ . _ a . - - . - .: . . . -
I 1 1 i 1 I a
. l Figure 11 17: 2650 psi SLB Leak Rate vs. Bobbin Amplitude l 3/4" Tubes, Model Boiler & Field Data a,g l l
r
=
S e S h I manus 4. 11-34
1 t Figure 11 18: 2335 psi SLB Leak Rate vs. Bobbin Amplitude 3/4" Tubes, Model Boiler & Field Data ' a,o G i 4 e O
. buns I
4 11 M
- , --m -
, e-+ ---
.,4 , 7, ,-, - -, ,- , - - ---
i + ~ Figure 11-19: Examination of Regression Residuals 3/4" Tubes, leak Rate @ SLB vs. Bobbin Amplitude .,g 1 I 4 4 9 h g 4 11 38
Section 12 . V.C. SUMMER IPC EVALUATION 12.1 Introductken
. This section provides the tube integrity evaluation performed for the V. C. Summer IPC to demonstrate margins against R.G.1.121 criteria. The V. C. Summer IPC are given in Section 12.2 including the tube repair basis, inspection requirements and operating leak rate limit. An equivalent V. C. Summer voltage repair limit for full implementation of an APC is developed in Section 12.3. This limit is applied to the IPC as the upper limit for leaving bobbin coll flaws in service even if not confirmed by RPC inspection. The Monte Carlo analysis methods used in this section are generally described in Section 12.4. The maximum EOC 7 bobbin voltage is projec'.ed in Section 12.5. Sections 12.6 and 12.7 provide the tube burst margin assessment and the SLB leakage evaluation. Development of the operating leak rate limit is given in Section 12.8 and the principal conclusions of the V. C. Summer IPC evaluation are summarlzed in Section 12.9.
12.2 V. C. Summer Interim Plugging Criteria (IPC) The V. C. Summer interim Plugging Criteria (IPC) follow the precedent approved by the NRC for application to J. M. Farley, D. C. Cook 1 and Catawba-1 steam generators. The IPC include the tube repair basis, inspection requirements and operating leak rate limit as described below: Tube Repair Basis o Bobbin coli indications having flaw voltages greater than 1.0 voit and confirmed as flaws by RPC Inspection shall be repaired. o Tubes with bobbin coilIndications >1.0 volt may be repaired as an attemative to RPC inspection. Bobbin collindications having flaw voltages greater than 2.2 volts shall be repaired independent of RPC confirmation of a flaw. o Projected leakage for a postulated steam line break (SLB) event at end of cycle (EOC) conditions shall De less than 1.0 gpm for the most limiting S/G.' Bobbin coil flaw Indications inspected by RPC and found to have no RPC indication do not need to be included in the leakage analyses. o Tubes identified as subject to significant deformation at a TSP elevation under a postulated LOCA + SSE event shall be excluded from application of the IPC at that TSP location. Inspection Requirements o The inspection shall include 100% bobbin coil inspection of all hot leg intersections and cold leg Intersections down to the lowest TSP for which the IPC is to be applied. o Bobbin coil flaw indications above 1.0 volt and below 2.2 volts shall be inspected by RPC to evaluate for detectable RPC Indications and to support ODSCC as the degradation mechanism, unless the tube is to be plugged or sleeved. 12 1
o Eddy current analysis guidelines shall be consistent with guidelinesgiven Appendix A. o An RPC sampling program of at least 100 TSP Intersections will be performed emphasizing intersections with greater than 5 volt (bobbin coll) dents and including ' some intersections with artifact bobbin indications or indications with unusual phase angles. Operating Leak Rate Limit o The normal operating leak rate requiring plant shutdown shall be limited to 150 gpd per S/G. The remainder of Section 12 demonstrates that the IPC provide margins against the tube integrity criteria of R.G.1.121. The operating leak rate limit is developed in Section 12.8. 12.3 Equivalent V. C. Summer APC Repalt Limit The equivalent APC voltage repair limit for fullimplementation of alternate plugging criteria (APC)ls utillzed in the IPC to establ!sh the maxlmum bobbin coil flaw voltage indication to be left in service even if not confirmed by RPC inspection. The tube repair critoria are developed to preclude freespan tube burst if it is postulated that TSP displacems nt would occur under accident conditions. No analyses for the Model D3 S/Gs in V. C. Summer have been performed for SLB conditions to determine if significant TSP displacement would occur to uncover the tube degradation occurring within the TSPs under normal operating conditions, if significant TSP displacement does not occur, the constraint provided by the TSPs would prevent tube burst and the principal repair criteria would be based on limiting leakage rather than free span burst. The equivalent APC repalt limit is developed to provide R.G.1.121 tube burst margins. The lower IPC repair limits incorporate additional margins against R.G.1.121, For the equlvalent voltage repair limit, the voltage structural requirement for burst capability at three times normal operating pressure differential (3APNO) is reduced by allowances for NDE uncertaintles in the voltage measurements and by voltage growth between inspections as described below. EOC Voltaae Limit for Structurn! Reautrement The recommended correlation between burst pressure and bobbin voltage, as adjusted for temperature and minimum material properties,is developed in Section 10.3. At the lower 95% prediction interval, a bobbin voltage of [ la,g volts establishes the structural requirement for 3APNO (3996 psi) tube burst capability as shown on Figure 10-3. Allowance for NDE Uncertaintv The V. C. Summer NDE uncertainties for bobbin voltage measurements are developed in Section 5.7.2. NDE uncertainties at +90% cumulative probability are applied to develop the tube repair limits. For the V. C. Summer uncertainty evaluation, the NDE uncertainty of 90% cumulative probability is 14% of the voltage measurement. A conservative value of 20% is applied at the ' tube repair limit voltage to develop the APC repair limit. L 12 2
,y- w g 9 .A- ,,. ,-- -s
Allowance for Crack Growth - Voltage growth rates for the V C. Summer S/Gs are developed in Section 9.3 and summarized in Table 9 2. The voltage growth rate average over the entire 800 voltage range is 44% for the 87 indications in tubes plugged during 1991. When averaged over only BOC Indicationc greater than o 0.75 volts, the growth rate is 16% for these Indications. These results show that the V. C. Summer growth rates, as a percentage of BOC voltage levels, tend to decrease with increasing
, BOC volts. This trend is consistent with other domestic planta shown in Figure 6 2. The data from European plants Indicate that percent growth may be approximately independent of amplitude. It is thus conservative to assume that percentage growth is independent of amplitude and to use everall average growth from V. C. Summer operating experience for the growth rate allowance in the plugging limits. The V. C. Summer average growth of 44% is increased to 45%
to establish the equivalent APC repalt limit. Eculvatont APC Ronalt Umit Table 121 summarizes the development of the equivalent APC repair limit based on reducing the structural voltage limit of [ Ja 9 volts by allowances for growth and NDE uncertaintles. The resulting equivalent APC repair limit is 2.2 volts. For IPC applications, the equivalent APC repair limit is used to define an upper bobbin flaw voltage limit for leaving unconfirmed RPC Indications in service rather than as the tube repalt Ilmit. 12.4 Projected EOC Voltages The maximum EOC voltage upon application of an IPC can result from: (1) indications left in service through IPC application with BOC voltages s1.0, (2) from bobbin indications found to be RPC NDD at the prior inspection, or (3) from new Indications which were bobbin NDD at the prior inspection. The latter two sources of indications are independent of IPC implementation. as they result from current inspection practices based on 40% depth repair limits, in fact, the IPC application is more conservative than standard practice by requiring bobbin flaw indications
>2.2 volts to be repaired even if found to be RPC NDD. This section compares the three above sources of EOC indications to demonstrate that an IPC repair limit of 1.0 volt results in lower EOC voltages than found for sources (2) and (3) above, which are applicable to current inspections and IPC applications. Use of V. C. Summer inspection results from the last cycle and data for Plant L at an inspection following IPC implementation are used to demonstrate this point.
Il can be noted, however, that sources (2) and (3) which are not detected by both bobbin and RPC probes would not contribute to EOC SLB leakage. This results as the indications not detected by both probes have a very low liklihood of progressing to a throughwall Indication in one operating cycle. It should also be emphasized that bobbin indications s1.0 volt, even if confirmed by RPC as assumed for SLB leakage analyses, contribute negligible SLB leakage as shown in Section 12.6. This section develops maximum EOC bobbin voltages using both deterministic and Monte Carlo ,, analyses, based on a BOC 1.0 volt Indication left in service by IPC applications. These voltages are compared to the bobbin voltages found at the Cycle 6 inspection which are typical EOC voltages for an outage following application of 40% depth repalt limits. As a demonstration for
. projecting BOC voltage distributions to EOC distributions for use in SLB leakage analyses, example results are presented assuming an APC repair limit of 2.2 volts had been implemented for Cycle 7 operation. Upon implementation of the IPC for V. C. Summer, SLB leak rates are expected to be based on deterministic methods (Section 12.6), unless an unexpectedly large -
12 3
m e number of indications ar found at EOC 7. In the latter case, Monte Carlo methods as demonstrated in this section would be used to develop the EOC voltage distribution fo, i sakage calculation. Determinktic and Monte Carlo Protectiors of Maximum EOC Voltacu To estimate the EOC voltage for an Indication left inservice at the IPC repair limit of 1.0 volt, deterministic and Monte Carlo analyses were performed. Allowances for NDE uncertainty and , growth based on the NDE uncertainty distribution (standard deviation of 11% with 25% maximum) of Section 5.7.2 and voltage growth distribution (Figure 9 8) are added to the BOC voltage to project the EOC voltage. For deterministle analyses, the distribution values at a specified cumulative probability are used to obtain a specific EOC voltage. For Monte Carlo analyses, the distributions are repeatedly (106 samples in present analyses) sampled to develop a cumulative probability for EOC voltage. For an assumed number of BOC indications, an EOC maximum voltage from Monte Carlo analyses can be defined as the voltage value for which the tall of the distribution integrates to one tube (i.e., cumulative probability of (N 1)/N for N Indications). To define a maximum EOC voltage for the Monte Carlo analyses, an assumed 200 indications leads to a 99.5% cumulative probability for the maximum EOC voltage. Table 12 2 summarizes the projected EOC voltages resulting from the BOC IPC limit of 1.0 volt. The deterministic assessments include allowances at +90% cumulative probability for comparisons with 3AP NO tube burst capability and at +99% for comparisons with SLB burst requirements. The deterministic assessn,ents indicate EOC voltages of 1.85 and 3.05 volts at
+90% and +99% cumulative probabilities. The Monte Carlo results at 99.5% cumulative probability for an assumed 200 indications left inservice at 1.0 volt yleid an estimated maxirnum EOC voltage of 2.97 volts (within 0.08% of the +99% deterministic estimate). It can be noted that the sum of +99% probabilities in the deterministic estimate has a net confidence >99.5% for the V. C. Summer application. Thus, the maximum EOC voltage for indications left in service by IPC implementation of a repair limit at >1.0 volt is expected to be between about 1.85 '
and 3.05 volts and statistically is dependent on the number of indications left in service below one volt. Table 12 3 summarizes the Monte Carlo estimated EOC volts as a function of cumulative probability and assumed BOC voltage amplitude. It can be noted that the [ ja.g volts corresponding to 3AP NO burst capability at 95% prediction intervalis: not reached in 106 samples for 800 = 1.0 volt, reached at 99.9% probability for BOC = 1.5 volts, reached at 98.8% probability for BOC = 2.0 volts, and reached at 96.5% probability for BOC - 2.5 volts. Thus, based on the historical growth ratios for V. C. Summer, the IPC repair limit of 1.0 volt precludes reaching the 34P NO urst b pressure at EOC conditions. EOC Voltaaes Resultina from Current 40% Deoth Reoair Umit Current inspection practice for application of the 40% depth repair limit is to identify potential indications by bobbin inspection, RPC inspect the potentialindications and repalt RPC confirmed ,, indications. This practice results as bobbin coil depth indications are often distorted such that tellable depth estimates are not feasible. Since both bobbin and RPC detectability are high for indications >40% depth, this practice is consistent with the 40% repair basis. The RPC confirmation permits conservative bobbin calling criteria for potentialindications and helps to eliminate falso bobbin calls, in some cases, a subsequent inspection leads to RPC confirmation of an indication that was RPC NDD at the prior inspection. The bobbin voltage at the time of RPC 12 4 j
l confirmation can be used as indicative of maximum EOC voltages based on 40% depth criteria for comparison with the above projected maximum voltages based on IPC implementation. The bobbin data re evaluation for growth rates was conducted to Appendix A analysis guidelines, such that voltages are typleal of an inspection implementing an IPC. There were 131 bobbin potentialIndications (Pis) identified in the 1991 inspection. Of these indications,50 were RPC confirmed in 1991. Only 3 of the 1991 bobbin Pls were NDD in 1990, although 39 were not Inspected in 1990 such that the number of new indications (based on reevaluatlon) could be greater than 3. Two of the field Pts were judged NDD by reevaluation in both years and 87 had bobbin Pts in both years. Table 12 4 summarizes the largest bobbin Indications grouped as 1991 RPC confirmed,1991 RPC NDD and new Indications (Bobbin NDD in 1990). The 5 largest
- of 50 RPC confirmed indications ranged from 1.72 to 2.80 volts. These amplitudes are typical of j current SG Inspections which frequently show Indications above 3 volts. The three new bobbin indications have a maximum of 1.39 volts, which may be low due to the small number in this group for the bobbin data reevaluation. The 5 largest bobbin Pls not confirmed by RPC show amplitudes ranging from 1.68 to 2.64 volts. The RPC NDD Indications are left in service. A fraction of this group of Indications can be expected to become RPC Indications at the next inspection and result is some of the largest amplitudes at the next inspection, The voltage range up to 2.80 volts found in the 1991 inspection followed application of 40%
depth repalt criteria at the prior outage. These voltages are essentially the same as the range for 1.85 to 3.05 projected at EOC following implementation of a 1.0 voll IPC repair limit. Thus, the implementation of a 1.0 volt IPC results in approximately the same rnaximum EOC voltage as the 40% depth criteria. Only Plant L (7/8 inch diameter tubing) has had a SG inspection following implementation of a 1.0 volt IPC. Table 12 5 summarizes the largest bobbin voltages (mid-cycle inspection) as grouped for Indications left in service by IPC implementation, new Indications and Indicat!ons
, RPC NDD at the prior cycle, it is clear from these results that the indications left in service based on IPC limits have smaller voltages than either the new indications or the previously RPC NDD Indications which would exist for a 40% depth criteria. These results support the conclusion that the IPC repair limit could be increased significantly above 1.0 volt without impacting tube Integrity, based on the largest voltage indications.
In summary, the V. C. Summer maximum EOC bobbin voltages for indications left in service by an IPC repalt limit of >1.0 volt are expected to be in the range of 1.85 to 3.05 volts. Based on prior V. C. Summer inspection results and Plant L experience, the IPC repair limit of 1.0 volt results in less than or approximately equal EOC voltages and tube integrity margins as the 40% depth criteria. Examole of Monte Carlo EOC Voltaae Distribution Monte Carlo analyses were performed to project the estimated EOC 7 voltage distribution assuming indications found in the 1991 inspection below the APC repair limit of 2.2 volts weie left in service. This analysis is provided as a demonstration of Monte Carlo projection results based on a 800 distribution of indications. The assumed BOC 7 distribution, growth rate distribution and projected EOC 7 distribution are shown in Figure 12-1. The EOC 7 results for the 68 indications <2.2 volts in SG B show an estimated EOC 7 voltage of 2.85 volts. The maximum EOC voltage is based on evaluating the Monte Carlo cumulative probability distribution at 67/68 - 0.985, which is equivalent to integrating the tall of the distribution to one Indica 0on. If the same distributions were applied to a larger number of indications, the 12 S
projected maximum voltage would increase. These results !ndicate that, statistically, a small sample (68 indications) of BOC voltages up to 1.9 volts (max. 800 in Figure 121) can result in a smaller EOC maximum voltage of 2.85 volts than a larger sample (200 indications) of 1.0 volt Indications which project to a maximum EOC voltage of 2.9 volts. 12.5 Tube Burst Margin Assessment Application of the equivalent APC repair limit developed in Section 12.3 would result in meeting R.G.1.121 criteria at EOC conditions. The objective of the IPC repair limit is to estabilsh additional margins beyond that included in the equivalent APC repalt limit. The limiting R.G.1.121 criterion for V. C. Summer is to satisfy the 34PNO tube burst margin requirement. Thus the additional IPC margins can be expressed as burst pressure margin ratios relawe to 3aPNO. That is, the ratios of BOC and EOC burst pressures to 3aPNO. The burst margin ratios are developed adding 490% cumulative probability on growth and NDE uncertainties to the 1.0 volt repair limit and evaluating the resulting voltages at the lower 95% prediction interval of the burst / voltage correlation (Figure 10 3 for tube burst capab!!!!y.) These uncertainty levels provide that only a few, if any, Indications would exceed 3aPNO at EOC conditions. It is necessary to establish higher confidence levels to prevent tube burst at APSLB-The burst margin ratios relative to APSLB are therefore developed applying +99% curnulative probability on growth and NDE uncertainties to the 1.0 volt repa!r limit and the lower 99% prediction interval of the burst / voltage correlation for tube burst capability. Table 12 6 summarizes the tube burst margin assessment relative to 3aPNO for V. C. Summer. - At +90% cumulative probability on allowances for growth and NDE uncertaintles, the projected EOC volage is 1.85 for a BOC Indication of 1.0 volt. This yields EOC burst pressure capability of , 4740 psi. The EOC burst pressure margin ratlo relative to 3aPNO ls 1.19 or significant margin against the R.G.1.121 criterion. At BOC, the burst margin ratio is 1.35. A typleal case for the 1.0 volt IPC applied to 7/8 inch diameter tubing is also shown in Table 12 5. It is seen that the V. C. Summer burst margin ratios are only about 8% lower than the 7/8 inch tubing example. For a 1.0 volt IPC, the BOC burst margin ratios are comparable for the two tubing sizes which shows that the 1.0 volt IPC establishes essentially equivalent BOC margins (within 4%) between 3/4 and 7/8 inch tubing. While the BOC margin ratios are approximately the same for this particular case, it should be recognized that variations will result for plants with different steam pressures. The EOC burst margins should not be compared for assessing equivalency of the 1.0 volt IPC limit between tubing sizes as modest changes in growth distributions can significantly alter the tube size comparison, it is necessary for IPC - margin demonstration that the EOC margin ratio exceed unity to show that R.G.1.121 is satisfied with some margin. For fullimplementation of the equivalsnt APC repair limits, the EOC burst margin ratlo relative to 3aPho would be expacted to be near unity. The tube burst margin ratio assessment relative to AP SLB is given in Table 12 7 for a 1.0 volt BOC Indication and for the deterministically projected EOC voltage of 3.05 volts. The EOC margin - ratio was found to be 1.54 for V. C. Summer. This demonstrates substantial margin against burst at SLB tor EOC 7 conditions. As a supplemental demonstration of burst margins, the probability 12 6
of tube burst at APSLB (2335 psi) was calculated for a 1.0 volt BOC Indication using Monte Carlo methods and found to be about 6.0x10-6 as given in Table 12 3. For a 2.0 volt BOC indication, the burst probability would be about 7.2X10 5. Thus large margins against burst at SLB conditions are provided by the V. C. Summer IPC.11 can be noted from Table 12 7 that both
. 3/4 and 7/8 inch tubing provide large margins against APSLB although the 7/8 inch margin ratloa are somewhat higher than that for 3/4 inch tubing.
The above assessment shows that the V. C. Summer IPC repalt limit of >1.0 bobbin coilvoltage provides large margins against R.G.1.121 structural critoria for tube burst. These assessments utilized the voltage / burst correlation of Figure 10-3. 12.6 SLB Leak Rate Analyses The IPC require that potentialleakage under SLB conditions at EOC be less than 1 gpm. This section provides the results of the leak rate analyses to demonstrate the methodology for satisfying this requirement for the V. C. Summer IPC repalt limit of >1.0 volt. Two methods of analysis are described for evaluating potentla! EOC SLB leakage for V. C. Summer. If, as expected, the number of indications left in service is less than about a thousand indications in the limiting SG, a deterministic analysis can be app!!ed, if a larger number of Indications are found, a Monte Carlo analysis would be applied. As described in Section 11.3.5, a leakage threshold of 1.0 volt is applied for V. C. Summer. All EOC indications above 1.0 volt are included in the leakage analysis. The threshold for significant leakage (210-3 gpm per Indication) is expected to be above about 2.5 volts (Section 11.3). The
- SLB leak rate per indication is obtained as the probability of leakage from Flgure 117 times the leak rate per leaking indication of Figure 11 16. For the low voltage range (< 3 volts) of interest for IPC applications the leak rate correlation of Figure 11 16 was conservatively developed to bound the SLB leak rate. Guidelines for the SLB analyses are developed in Section 7.4. The analyses are performed for a SLB SG tube pressure differential of 2335 psi.
The reference SLB leak rate analyses are deterministic analyses performed at a +95% confidence leval for EOC voltages >1.0 volt. The BOC voltages are increased by allowances for NDE uncertalntles and growth at +95% cumulative probability to obtain corresponding EOC voltages. At the EOC voltage, the probability of leakage (P.O.L) of Figure 117 is evaluated at the +95% level. The SLB leak rate of Figure 11 16 is evaluated at the +95% confidence level. The total SLB leak rate is obtained by summing the P.O.L. times the leak rate per Indication over all EOC Indications. This methodology conservatively assumes all indications grow at the +95% level and is generally more applicable to a single tube estimate rather than a sum over all indications. Analyses are given in this section for a BOC indication at the repair limit of 1.0 volt to bound the number of Indications that can be left in service by IPC implementation using deterministic analyses, if a large number of indications are found at EOC 7, Monte Carlo analyses can be applied by two alternate methods to reduce the conservatism of the deterministle analyses. The first method is
- to use Monte Carlo to develop an EOC voltage distribution for which lehkage is evaluated at each EOC voltage by applying P.O.L. and leak rates at the +95% level. The second and most accurate method is to apply Monte Carlo analyses for the leak rate calculation. With the Monte Carlo methods, many random samples (105 to 10 6) are made for each BOC voltage level. The NDE uncertalnty distribution (standard deviation of 11% with maximum of 25% per Section 5.7.2) 12 7
and growth distribution (Figure 9 B) are sampled to obtain an EOC voltage sample. The P.O.L. and leak rate distributions are randomly sampled for each EOC voltage sample. The Monte Carlo analyses lead to EOC voltage and SLB leak rate cumulative probability distributions. The SLB leak rate distribution is evaluated at 90% cumulative probability to obtain a leak rate for comparison with the 1.0 ppm limit for the V. C. Summer IPC. The EOC 7 voltage distribution of ' Figure 121 was obtained using Monte Carlo methods and the SLB leak rate is obtained in this section for this distribution. Monte Carlo anlyses are also used in this section to estimate the cumulative probability associated with leak rates obtained by determlnictic analyses. With the use of P.O.L. distributions, the Monte Carlo methods for calculating the cumulative probability distribution for leakage differ from prior APC reports such as WCAP.12871. The cumulative probability of no leakage is calculated as 1.0 minus the sum of all leakage probability samples
- divided by the total number of samples. That is,1.0 minus the average probability of leakage.
The cumulative probability for leakage >0.0 starts at the zero leakage probability and increases based on the leak rates obtained by sampling the leak rate distribution. SLB Leak Rate Per Indicatlon Determ!n!: tic analyses for the SLB leak rate per BOC 1.0 volt Indication are given in Table 12 8. As described above, the reference SLB leak rate is based on evaluating the leak rate correlations at the +95% confidence level on the mean regression fit. The resulting leak rate is 0.00093 gpm per 1.0 volt indication left in service. Thus, for a 1.0 ppm leakage limit, about 1075 Indications at 1.0 volt could be left in service per SG applying this conservative analysis procedure. The associated leak rate corresponds to about 99% cumulative probability from Monte Carlo analyses for a BOC 1.0 volt indication such that only about 1% of the 1.0 volt Indications would have this leak rate. The principal difference is that the deterministic analyses assume allIndications simultaneously have 495% P.O.L and leak rate while the Monte Carlo analyses independently sample these data. Even more conservative analyses are given in Table 12 8 to demonstrate that SLB leakage is not
- a concern for IPC implementation at V. C. Summer. Deterministic analyses are given for a BOC 1.0 volt indication at the +95% and +99% prediction intervals. For these analysis, P.O.L. is evaluated at the +95% (as for reference case) and +99% bounds. The SLB leak rates are evaluated at the additionally concervative prediction interval bounds rather than the confidence on the mean regression line used for the reference analyses. These analyses provide upper bound estimates on the leak rate from the most limiting single indication. However, even if it is postulated that allindications leak at the upper bound rate, the per Indication leak rates of
. 0.0031 and 0.044 gpm permit 323 (+95%) and 23 (+99%) indications per gpm of leakage. It - is seen from the Monte Carlo results of Table 12 8 that only about 0.4% ar.d 0.02% of the tubes would leak at the +95% and +99% estimates. While these levels of conservatism are not appropriate for comparing with the 1.0 gpm allowable SLB leakage limit, it is clear that total SLB leakage from allindications would not approach leak rate levels of greater than 350 gpm associated with a single tube rupture (gulliotine break). Conservatively postulating all indications (rather than expected 0.4% and 0.02%) leak at the +95% and +99% prediction intervals, the number of Indications left in service for 350 gpm would be about 113,050 and 8050, respectively. Thus, assuming all Indications leak at the rete expected for only 0.4% of _ the tubes, the number of indications corresponding to leakage equivalent to a tube rupture exceeds the number of TSP intersections in the Model D3 SG outside of the preheater. Even at an assumed APC repalt limit of 2.5 volts and allindications (rather than expected 0.05%) leaking .
- at the +99% prediction interval,1750 indications at 2.5 BOC volts in service would be required to reach leakage associated with a gulliotino rupture, Consequently, cumulative leakage from TSP Indications approaching levels associated with a tube rupture is unlikely.
12 8
- . .- - ._ - - _ . = - - - - - - - - . -
[ The Monte Carlo leak rates per BOC indication are summarized in Table 12-3 at varlous cumulative probability levels. The cumulative probability per Indication of zero leakage is also given in Table 12 3. For the 1.0 voit IPC repair limit, there is a 96% probability of zero leakage c.- only 4% of the indications would have leakage greater than zero. These results represent an average of the P.O.L distribution of Figure 11 1 over the distribution of potential EOC voltages for a civen 800 voltage. Even for an assumed APC repair limit of 2.5 volts, there is only a 26% probab!!ity that an Indication at EOC would leak. If the P.O.L. curve was developed
. for leakage above a minimum level such as about 10 2 to 10 3 gpm, the probability of leakage above this level would be smaller than obtained for the zero leakage threshold.
SLB Leak Rate for Distribution of EOC Voltacos As noted previously, the use of Monte Carlo analyses to develop an EOC voltage distribution can be I used for SLB leakage rather than the bounding deterministic values. Figure 121 shows the 68 Indications below the equivalent APC repair limit of 2.2 volts found in the 1991 Inspection of SG B. The lower figure shows the Monte Carlo EOC distribution for these 68 Indications. The SLB leak rate can be calculated by applying P.O.L and leak rate at the +95% confidence levels to the EOC Indications of Figure 121 above the 1.0 volt leakage threshold. This leads to a cumulative SLB leak rate of 0.020 gpm for the 68 Indications, if the deterministic reference method at +95% is also used to calculate EOC voltages from the upper figures of Figure 121, the estimated SLB leak rate would be 0.075 gpm as compared to 0.02 gpm using Monte Carlo EOC voltage. This comparison demonstrates the conservatism in the reference, completely - deterministic analyses. Applying complete Monte Carlo totalleak analyses to the BOC distribution would yleid even lower leak rates. Overall, it is concluded that the 1.0 volt IPC repalt limit with voltage growth rates comparable
- to that found last cycle at V. C. Summer will not result in SLB leak rates exceeding 1.0 ppm.
Even if >1000 Indications are found, app!! cation of analysis methods less conservative than the reference deterministic methods would result in <1.0 ppm leakage. 12.7 Operating Leakage Limit R.G.1.121 acceptance criteria for establishing operating leakage limits are based on leak before break (LBB) consideration such that plant shutdown is initiated if the leakage associated with the longest permissible crack is exceeded. The longest permissible crack is the length that provides a factor of safety of 3 against bursting at normal operating pressure differential. As noted above, a voltage amplitude of [ ja,g volts for typical ODSCC cracks corresponds to meeting this tube burst requirement at the lower 95% confidence level on the burst correlation. Attemate crack morphologies could correspond to [ ]a,g volts so that a unique crack length is not defined by the burst pressure to voltage correlation. Consequently, typical burst pressure versus through wall crack length correlations are used below to define the
- longest permissible crack" for evaluating operating leakage limits.
The CRACKFLO leakage model has been developed for single axial cracks and compared with leak rate test results from pulled tube and laboratory specimens. Fatigue crack and SCC leakage data have been used to compare predicted and measured leak rates. Generally good agreement is obtained between calculation and measurement with the spread of the data being somewhat greater for SCC cracks than for fatigue cracks. Figure 12 2 shows normal operation leak rates including uncertainties as a function of crack length. 12 9
The through wall crack lengths resulting in tube burst at 3 times normal operating pressure differentials (3996 psi) and SLB conditions (2650 psi) are about [ }a, respectively, as shown In Figure 12 3. Nominalleakage at normal operating conditions for these crack lengths would range from about [ Ja would cause undue restrictions on plant operation and result in unnecessary plant outages, radiation exposure and cost of repair. In addition,it is not feasible to satisfy LBB for all tubes by reducing the leak rate limit. Crevice deposits, presence of small ligaments and irregular . fracture faces can, in some cases, reduce leak rates such that LBB cannot be satisfied for all tubes by lowering leak rate limits. An operating leak rate of 150 gpd (-0.1 opm) will be implemented in conjunction with application of the tube plugging criteria. As shown in Figure 12 2, this leakage limit provides for detection of [ ja. Thus, the 150 gpd limit provides for plant shutdown prior to reaching critical crack lengths for SLB conditions at leak rates less than a .95% confidence level and for 3 times normal operating pressure differentials at less than nominalleak rates. The tube plugging limits coupled with 100% Inspection at affected TSP locations provide the principal pretsetion against tube rupture. The 150 gpd leakage limit provides further protection against tube rupture. In addition, the 150 gpd limit provides the capabl!!!y for detecting a rogue crack that might grow at much greater than expected rates and thus provides additional protection against exceeding SLB leakage limits. 12.8 Conclusions Based on the above evaluation of the V. C. Summer IPC repalt limit of >1.0 bobbin volt, it is concluded that: o R.G.1.121 criteria for tube integrity are conservatively satisfied at EOC for an IPC repair limit of 1.0 bobbin volt. o At EOC, burst pressure capability (expressed as margin ratios relative to 3APNO and l APSLB) is expected to have ratios of about 1.19 relative to 3APNO at 90% cumulative probability levels and about 1,54 relative to APSLB at 99% cumulative probability levels. A burst pressure margin ratio of 1.35 relative to 3AP NO for V. C. Summer at BOC conditions is comparable to typical values of 1.4 for plants with 7/8 inch diameter tubing with an IPC repair limit of 1.0 volt. Thus the two tubing sizes can be considered to have comparable BOC margins for an IPC repair limit of 1.0 volt, o Potentiat SLB leakage at EOC conditions is expected to be well below the 1.0 ppm allowablo limit as supported by both Monte Carlo and deterministic evaluations including sensitivity analyses. _ o The maximum EOC bobbin voltage indication resulting from indications at 1.0 volt is projected to be about 2.97 volts for 200 BOC indications and up to 3.08 volts for 500 - BOC Indications. The maximum EOC voltage for indications left in service by IPC implementation is expected to be comparable to or smaller than the indications found in inspections following application of 40% depth repair limits. 12 t o
o The operating leak rate limit of 150 ppd implemented with the IPC satisfies R.G.1.121 guidelines for leak before break. This limit provides for plant shutdown prior to reaching critical crack lengths for SLB conditions at a 95% confidence level on leak rates and for 3.iP conditions at less than nominal leak rates. + e 4 4 a 12- 11 i
f 1 l . Table 121 Equivalent APC Repair Limits to Satisfy Structural Requirements Item Vetts Bacin . . Maximum Voltage Limit to [ la 9 Burst Pressure vs. Voltage Satisfy Tube Bwst Correlation at 95% Structural Requirement confidence level (Fig.101) Allowance for NDE -0.45(20%)(1) From Section 5.7.2, the 14% Uncertainty uncertainty at 90% cumulative probability is conservatively increased to 20% at tube repair limit. Allowance for Crack 1.0(45%)(1) Table 9 2 shows average growth / Growth Between cycle of 44%. Allowance Inspections increased to 45% of Tube Repair Umit. 1 Equivalent APC Repair 2.2 Voltage Umit o Acceptable Umit to Meet Structural Requiremont U211t:
- 1. Voltage percentage allows. ices for NDE and growth rate / cycle applied to Equivalent APO Repair Voltage Limit of 2.2 voits.
12 12
- Table 12 2 ;
i Maximum EOC Voltage Sensitivity Assessment I* baterministic Annantments ; ]i; 90% Cum. Prob. 95% Cum. Prob. 99% Cum. Prob. Monte Carlo , ! BOC Volts 1.00 1.00 1.00 1.00 l 1 . l NDE Uncertainty 0.14 0.18 0.25 - ! i Growth 0.71 0.97 1.80 - EOC Maximum Volts 1.85 2.15 3.05 2.970) . i I i
- blata
- 1. Maximum EOC volts based on integrating the tall of the EOC distribution to one Indication for an assumed 200 indications. The EOC maximum voltage would ,
increase to 3.08 for an assumed 500 indications. i i 4 9 i I
- ~ j p
.12 13.
i Table 12-3 Itonte Carlo Results for EOC Volts, SLB Leek Rate and SLB Burst Prettehility per Indication SLB Leak Rata f% - Curn %e Proham,)f2) $lg
' EOC Vahe f% - De " ^ : P.a
- mal Cum. Prob. gpm gpm gpm Burst BOC Volts '205 EiYa ' 29% 99.5% fi) 99.9 % for 0 Laaks at 95% at 99% at 99.9% Prob. ,
- 1.0 .. 1.69 1.96. 2.77 2.97 3.15 96% 0 0.0009 0.011 0.6x10-5 1.5 2.22 2.48- 3.27 3.47. 3.70 92% 0.0004 0.0044 0.031 2.9 x10-5
- 2.0 2.77 3.00 3.77 3.99 4.27 84 % 0.0018 0.0012 0.069 7.2x10-5 i
2.5 3.32 3.55- 4.30 4.51 4.88 74 % 0.0062 0.028 0.13 1.5x10-4 ; i i' Notes: , 4 o . i
.1) 99.5% corresponds to definition of maximum EOC voltage for 200 indcations.
i 2) . stb ieak rases based on samples of proosbinty of leakage sin >as leak rate. i 12-14 e
Table 1244 Summary of Largest Bobbin Vol', age Indications from 1991 Inspection Bobbin Coll M h '91 Lne. '91 Volts '90 Volta Bobbin Pla Confirmed by RPC C R180103 2H 2.80 0.78 B R42047 8H 2.49 N.I. B R270100 2H 1.75 N.I. C R37C40 5H 1,73 N.I. B R10C97 2H 1,72 1.47 Bobbin Pts Not Confirmed by RPC B R36C50 8H 2.64 N.I. B R42044 12H 2.01 0.46 B R32088 8H 1.81 N.I. ' B R41C68 8H 1.74- N.I. O R47C36 5H 1.68 0.76 New Bobbin Pls, NDD in Prior Outage B R33C43 8H 1.39 POD B R40C54 2H 1.09 . POD B R49041 8H 1.03- . POD th b:. 12 15 t
. . , . a -. ---.. - - , . ., ..-..m , , . . . ,
Table 12 5 Plant L Mid Cycle inspection Hesults Following IPC Implementation 1992 Bobbin 1991 Volts . Projected EOC Rank 1992 Volts
- Re evaluated 1992 RPC EOO Volts" 1991' RPC Indications with Bobbin Volts <1.0:
1 0.93 'O.12 N . I . "* 1.36 2 1.01 0.64 N.I. 1.22 10 0.66 0.66 NJ. 0.65 , 1991 Bobbin Pts (RPC NDD): Independent of IPC 1 2.02 1.11_ Confirrned 2.51
- 2 2.44- 12 . 4 4 MX) 2.44 l 7 1,82 '1.76 Confirmed 1.85 l ..
New Indications (1991 Bobbin e RPC NDD): Independent of IPC 1 1.71 0.87 N.I.*" 2 17 2 1.56 1.11 Confirmed 1.80 l 9 1.23 1.10 Confirmed - 1.30 L.' o Notes: . 1992 inspection - measured bobbin voltages ! Voltages projected to EOC~'oased on plant specific growth to 1992 inspection scaled to EOC EFPDs.
* *
- N.I. = not inspected -
A 12 16
=
i: 4-a- Table 12 6 4 l V. C. Summer Tube Burst Margin Assessment for SAPil o : 4 1 q -~ 2, V. C. Summer:3/4" Tubina 7/8" Tutina Example (i) ! BOC Volts 1.0 1.0
' ?
1 e Allowances @ +90% Cum. Prob. ! o Voltage Growth . 0.71 ; 0.60 ~ v l- o NDE Uncertainty 0.14 0.14-F EOC Volts (+9C%) 1.85 1.74-4- Tube Burst' Capability (psi)I -
- o 3aP NO R equirement- 3996- 4380
- , o Capability at -95% Prod. int.
i At BOC = 1.0 volt 5410 6200-
. At Projected EUC Volts 4740 -5660 4 1 j Burst Capability Ratios to 3aPNO o At BOC - 1.0 volt 1.35 - 1.d1 -
o At Projected EOC volts 1.19' 1.29 [ i t j i i-I f
- .. Nola:
1.- Example for 7/8", Tubing for Plant A 2, WCAP 13464.L 2
)
1 12- 17
Table 12 7 V. C. Summer Tube Burst Margin Assessment for SLB Conditions V. C. Summer 3/4" Tubino 7/8" Tubino Examole(1) BOC Votts 1.00 1.00 Allowances @ +99% Cum. Prob. o Voltage Growth 1.80 2.00 o NDE Uncertainty 0.25 0.25 EOC Volts (+99%) 3.05 3.25 Maximum Projected EOC Volts o Monto Carlo 2.97 (2) Tube Burst Capability (psi) o APSLB Requirement 2335 2335 , o Capability at -99% Pred. Int.: At +99% EOC volts 3590 4420 , At Maximum EOC Volts 3620 - Burst Capability Ratios to APSLB o At +90% EOC volts 1.54 1.89 o At Maximum EOC volts 1.55 - Notes:
- 1) Example for 7/8" Tubing for Plant A-2, WCAP-13464.
- 2) Monte Carlo result for 200 BOC indications. Value increases to 3.08 volts '
for 500 indications. 12 18
Table 12 8 - Deterministic EOC SLB Rates and Equivalent Monte Carlo Probability Reference SLB Leak' Rate Bounding Single Indication SLB Leak Rate (gpm)
-. Inomi ner inelbatinn(1) 95% Prediction int. 99% Predclion Int.
BOC volts 1.0 . 1.0 ' . 1.0 - 2.5 EOC volts - 2.15- 2.15' 3.05- -4.93 per Table 12 9 ' Probability of 0.309 0.309 0.618 : 0.915-Leakage (P.O.L) SLB Leak Rate (gpm) -0.0030 0.0099(2); o,oyj(2) ~ 0.22(2).
- Weighted SLB Leak 0.00093 i 0.003'1 < 0.044 - 0.20 Rate (gpm) _ _ . .
- (P.O.L x Leak Rate)- '
l* lI . l -_ Equivalent Monte Carlo -~99% . ~99.6% '- ~99.98% - ~99.95% . ' Cumulative for Weighted Leak Probability Rate (3). 7 Notes: (1) Obtained evaluating probability of leakage and leak rate at +95% confidence level. (2)1 Obtained evaluating leak rate at specified bound on prediction interval. (3) ' Obtained evaluating Monte Carlo cumulative probability.distributio'n for a given BOC indication at the weighted leak rate from the deterministic analysis.- 4& . 4 r 1 9 a p -y -~ k4 e .%_%2 y m - yee-- %
..~ . _ . . . . . .. - .. . _ - - - - - .--
tf.CEMMER iset TSP IN0eCATIONS-SQ B 12 ) 10-t* L' e d ,1 d 3
~
3 -3 3 3 3 i r b
, I k2 ,
, l Y - ..
lk .. 33, @, .5, k h k Eo
'06' 'os' 04 1 ~ 12 ~ - -
'14' '16' 1s' -2 Banne vons l
V.C. SUMMER PROKCTED GROWTH-ALL SATs - 10 1 14-33 12-io- y a :
^
' .1 .
o os os os is- i.s 2.1 d V C. SUMMER EOC-7 BO9WN WLTAGE-SG B - I e-
, 5 )'d
+
h p - - r a l H, 2 2 I i l 06' li o.7 - - oJ lll
'0S' '1' ' 12' '1 A '
1.1 ' 13
'1A' lll1l2
'13' - '2' lliiiii
'22 2.1_
' 2.4 ' ' 2.4 23 - 2.8 - 2J6
~
Figure 12-1. Example Analysis of Volta 0e Growth for a Distrbution of BOC - .--. deations (Steam Generator B)- 12-20
_. .. _ . . _ . - . ._ - . . _ . . _ . . . .. 4._ . . _ _ _ _ . .. .. _ .. . . - . _ . .- . _ _ . . . a -i
, F .
t 4
; Figure 12-2. ~ Leak Rate Under Normal Operating _ Conditions versus Crack ~-
Length br 3/4 inch Tubing 4 12-21
- i. -
?Y s
- - < - ~ . - ,:.- , .. , , , . , _ _ _ , , , _ , _ _ _ ,
a,g Figure 12-3. Burst Pressure versus Crack Length br 3/4 inch Tubing 12 22
Appendix A NDE Data Acquisition and Analysis Guidelines A.1 INTRODUCTION This appendix documents techniques for inspection of the V. C. Summer SG tubes related to the identification of ODSCC or ODIGA/ SCC at the support plate regions. This appendix contains guidelines which provide direction in applying the ODSCC alternate repair limits described in this report. The procedures for eddy current testing using bobbin coil and rotating pancake coil (RPC) techniques are summarized. The procedures given apply to the bobbin coil inspection, except as explicitly noted for RPC inspection. The methods and techniques detailed in this appendix are requisite for implementation of the alternate repair limit and are to be incorporated in the applicable inspections and analysis procedures. The following sections define specific acquisition and analysis parameters and methods to be used , for the inspections of the steam generator tubing,
, A.2 DATAACQUISITION The V. C. Summer steam generators utilize 3/4" OD x 0.043" wall, Alloy 600 mill-annealed tubing. The carbon steel support plates and baffle plates are designed with drilled holes. The -
holes in the flow distribution baffle are nominally 0.828' in diameter, while the holes in the support plates and the remaining baffle plates are nominally 0.766"in diameter. The following guidelines are specified for non-destructive examination of the tubes within the TSPs at V, C. Summer. A.2.1 Probes Bobbin Col! Probes Eddy current equipment used shall be the ERDAU (Echoram Tester), Zetec MlZ-18 or other equipment with similar specifications. To maximize consistency with laboratory data, differential bobbin probes with the following parameters shall be used: 1 0.610" outer diameter (If other probe sizes are used, cross calibration factors must-be determined by comparative testing with the ASME standard and revised
~A1
t uncertatity estimates developed)
. - two bobbin coils, each 60 mits long, with 60 mits between coils (coli centers separated by 120 mils) in addition, the probe design must incorporate centering fectures that provide for minimum ,
probe wobble and offset; the centering features must maintain constant probe center to tube ID offset for nomenal diameter tubing. 4
- Rotatino Pancake Coi1 Probes The pancake coil diameter shall be s0.125". While any multi-coil (i.e.,1,2 or 3-coll) probe can be utilized, it is recommended that if a 3-coil probe is used, any voltage measurements should be made with the probe's pancake coil rather than its circumferential or axlal coll. The maximum probe pulling speed shall be =0.2 in/sec for the 1 coil or
- 3. coll probe, or 0.4 in/sec, for the 2-coll probe. The maximum rotation speed shall be
=300 rpm. This would result in a pitch of =40 mils for the 3-coll probe.
A.2.2 Calibration Standards . Bobbin Cell Standards , The botbin coil calibration standards shall contain:
- Four 0.028" diameter through wall holes, 90* apart in a single plane around the?
tube circumference; the hole diameter tolerance shall be 10.001".
- One 0.109* diameter flat bottom hole,60% through from OD
- One 0.187" diameter flat bottom hole,40% through from the OD Four 0.187" diameter flat bottom holes,20% through from the OD, spaced 90* - -
apart in a single plane around the tube circumference. The tolerance on hole
- diameter and depth shall be 0.001".
- A simulated support ring,0.75"long, comprised of SA-285 Grade C carbon steel or equivalent.' -
- This calibration standard will need to be calibrated against the reference standard used for the APC laboratory work by direct testing or through the use of a transfer a standard..
- A probe wear standard for monitoring the degradation of probe centering devices A-2
9 1 leading to off. center coil positioning and potential variations in' flaw amplitude responses. This standard shallinclude four 0.06710.001 inch diameter through , wall holes, spaced 90' apart around the tube circumference with an axial spacing
. such that signals can be clearly distinguished from one another (see Section A.2.3).
RPC Standard The RPC standard shall contain:
- Two axial EDM notches, located at the same axial position but 180' apart circumferentially, each 0.006" wide and 0.5" long, one 80% and one 100%
through wall from the OD.
- Two axial EDM notches, located at the same axial position but 180' apart circumferentially, each 0.006" wide and 0.5"long, one 60% and one 40%
through wall from the OD.
- Two circumferential EDM notches, one 50% throughwall from the OD with a 75'
- (0.57") are length, and one 100% throughwall with a 26' (0.20") arc length, with both notches 0.006" wide.
- A simulated support segment,270* in circumferential extent,0.75" thick, comprised of SA-235 Grade C carbon steel or equivalent.
The center to center distance between the support plate simulation and the nearest slot shall be at least 1.25". The center to center axial distance between the EDM notches shall be at least 1.0". The tolerance for the widths and depths of the notches shall be 0.001". The tolerance for the slot lengths shall be 0.010". A.2.3 Application of Bobbin Coil Probe Wear Standard A calibration standard has been designed to monitor bobbin coil probe wear. During steam generator examination, the bobbin coil probe is inserted into the wear monitoring standard; . the initial (new probe) amplitude response from each of the four holes is determined and compared on an individual basis with subsequent measurements. Signal amplitudes or . voltages from the individual holes - compared with their initial amplitudes - must remain within 15% of each other for an acceptable probe wear condition, if this condition is not satisfied, then the probe must be replaced. If any of the last probe wear standard signal A3
amplitudes prior to probe replacement exceeds the 15% limit, say by a variable value, x*/o,
- then Indications measured since the last acceptable probe wear measurement that are within x% of the plugging limit must be re intpected with the new probe.
A.2.4 Acquisition Parameters The following parameters apply to bobbin coil data acquisition and should be incorporated in the applicable inspection procedures to supplement (not necessarily replace) the parameters normally used. Test Frecuencles This technique requires the use of 550 kHz and 130 kHz test frequencies in the differential mode. It is recommended that the absolute mode also be used, at test frequencies of 130 kHz and 35 kHz. The low frequency (35 kHz) channel should be recorded to provide a positive means of verifying tube support plate edge detection for flaw location purposes. The 550/130 kHz mix or the 550 kHz differential channel is used to access changes in signal amplitudes for the probe wear standard as well as for flaw detection. RPC frequencies should include 550 kHZ,300 kHz and 130 kHz. Dialtizino Rate l A minimum bobbin coil digitizing rate of 30 samples per inch should be used. Combinations of probe speeds and instrument sample rates should be chosen such that: Sample Rate (samples /sec.)
- - - - - - - - - - - --- 2 30 (samples /in.}
Probe Speed (inisec.) Scans and Rotations l l Spans and rotations can be set at the discretion of the user and/or in accordance with applicable procedures, but all TSP intersections must be viewed at a span settin0 one-half A-4
or less than that which provides 3/4 full screen amplitude for 4x20% holes with bobbin probes and 1/10 or less the corresponding span for 0.5"Inroughwall slot (EDM notch) with RPC probes. Missts A bobbin coil differential mix is established with 550 kHz as the primary frequency and 130 kHz as the secondary frequency, and suppression of the tube support plate simulation should be performed. A.2.5 Analysis Parameters This section discusses the methodology for establishing bobbin coil data analysis variables such as spans, rotations, mixes, voltage scales, and calibration curves. Although indicated depth measurement may not be required to support an alternative repair limit, the methodology for establishing the calibration curves is presented. Tht use of these curves is recommended for consbtency in reporting and to provide compatibility of results with subsequent inspections of the same steam generator and for comparison with other steam generators and/or plants. 550 kHz DifferentiM Channel Rotation: The signal from the 100% through wall hole should be set to 40' (t1') with the initial signal excursion down and to the right during probe withdrawal. Voltage Scale:
- 1) Bobbin - The peak-to-peak signal amplitude of the signal from the four 20%
OD flaws should be set to prodtice a field voltage equivalent to that obtained for the EPRI lab standard. The EFRI laboratory standard normalization voltages are .. 4.0 volts at 550 kHz for 20% ASME holes and 2.75 volts at 550/130 kHz mix for 20% ASME holes. The transfer / field standard will be calibrated against the
. laboratory standard using a reference laboratory probe to establish voltages for the field standard that are equivaient to the above laboratory standard. These equivalent voltages are then set on the field standard to establish the calibration A-5
voltages for any other standard. Voltage normalization to'the standard calibration voltages at 550 kHz is the preferred normalization to minimize analyst sensitivity in establishing the mix. However,if the bobbin probes used at V. C. Summer result in 550/130 kHz mix to 550 kHz voltage ratios differing . from the laboratory standard ratio of 0.69 by more than 5% (0.66 to 0.72), the 550/130 kHz mix calibration voltage should be used for voltage normalization.
- 2) RPC The RPC amplitude shall be set to 20 volts for the 0.51nch throughwall notch at 550 kHz and 300 kHz. Each channel shall be set individually.
Calibration Curve: Establish a curve using the measured signal phase angles of the nominal 100%,60% and 20% drilled holes on the calibration standard. The "as-read" depth of the drilled holes should be determined from the 550 kHz differential channel.- This should be achieved by setting the phase angle of the 100% drilled hole to 40 degrees and then determining the "as-read" depth of the 60% and 20% drilled holes from the ZOA-4.1 curve. These "as-read" depths should then be employed for the set uo of all phase angle calibration curves. , 550/130 kHz Differential Mix Channel . Rotation: Set the signal from the probe motion to be horizontal with the initial excursion of the 100% through wall hole signal going down and to the right during probe withdrawal. Calibration Curve: Establish a curve using measured signal phase angles in combination with the "as-read" fisw depths for the 100%, 60%, and 20% flaws on the calibration standard. A.2.6 Analysis Methodology Bobbin coil indications at support plates attributable to ODSCC are quantified using the Mix 1 ., (550 kHz/130 kHz) data channel. This is illustrated with the example shown in Figure A-1. The 550/130 kHz mix channel or other channels appropriate for flaw detection (550 kHz, . 300 kHz or 130 kHz) may be used to locate the indication of interest within the support plate signal. The largest amplitude portion of the lissajous signal representing the flaw should then A-6
be measured using the 550/130 kHz Mix 1 channel to establish the peali to-peak voltage as shown in Figure A 1. Initial placement of the dots for identification of the flaw location may be performed as shown in Figu'res A 2 through A 4, but the final peak to peak
, measurements must be performed on the Mix 1 lissajous signal to include the full flaw segment of the signal. As can be seen in Figures A 2 through A 4, failure to do so can reduce the voltage me:surement of Mix 1 by as much as 65% to 70% due to the interference of the support plate signalin the raw frequencies. The vo;tage as measured from Mix 1 is then entered as the analysis of record for comparison with the repair limit voltage.
A.2.7 Reporting Guidelines The reporting requirements identified below are in addition to any other reporting requirements specified by the user. Minimum Aeouirements At a minimum, flaw signals in the 550/130 mix channel at the tube support plate intersections whose peak to-peak signal amplitude exceeds 1.5 volts, requiring RPC inspection, must be reported. Signals, however small, should also be reported for historical purposes and to provide an assessment of the overall condition of the steam generator (s). Additional Reaulrements For each reported indication, the fo!!owing information should be recorded: Tub identification (row, column) Signal amplitude (volts) Signal phase angle (degrees) indicated depth (%)i Test channel (ch#) .. Axial position in tube (location) Extentof test (extent) i it is recommended that an indicated depth be reported as much as possible rather than some letter code. While this measurement is not required to meet A-7
r. d at a later date the attemate repair limit, this information might be require ~ and/or otherwise be used to develop enhanced analysis techniques. RPC reporting requirements should include a minimum of: type of degradation (axial, , circumferential or other), maximum voltage, phase angle, crack lengths, and location of the center of the crack within the TSP. The crack axial center need not coincide with the position of maximum amplitude. A.3 DATA EVALUATION A.3.1 Use of 550/130 Differential Mix for Extracting the Bobbin Flaw Signal in order to identify a discontinuity in the composite signal as an Indication of a flaw in the tube wall, a simple signal processing procedure of mixing the data from the two test frequencies is used which reduces the interference from the support plate signal by about an order of magnitude. The test frequencies most often used for this signal processing are 550 kHz and 130 kHz where 43 mit wall Alloy 600 tubing is involved. The processed data is referred to as 550/130 mix channel data. This procedure also reduces the interference from magnetite accumulated in the crevices. Any of the differential data channels including the mix channel may be used for flaw detection (though the 130 kHz channelis subject to influence . from many different effects), but the final evaluation of the signal detection, amplitude and phase will be made from the 550/130 differential mix channel. Upon detection of a flaw signal in the differential mix channels, confirmation from other raw channels h.no.1 required. The voltage scale for the 550/130 differential channel should be normalized as described in Section A.2.5. The present evaluation procedure requires that there is no minimum voltage for flaw detection purposes and that all flaw signals, however small, be identified. The intersections with flaw signals 21.5 volts will be inspected with RPC in order to confirm the presence of ODSCC. Although the signal voltage is not a measure of the flaw depth,it is an indicator of the tube burst pressure when the flaw is identified as axial ODSCC with or without minor IGA. . The procedure using the 550/130 mix for reducing the influence of support plate and , magnetite does not totally eliminate the interference from copper, a!!oy property change or dents. These are discussed below. A8
A.3.2 Amplitude Variability
, it has been observed
- hat voltage measurements taken from the same data by different analysts may vary, even when using identical guidelines. This is largely due to the analysts' interpretations of where to place the dots on the lissajous figure for the peak to-peak measurement. Figures A 5 and A-6 show the correct placement of the dots on the Mix 1 lissajous figures for the peak to peak voltage amplitude measurement for two tubes from V.C. Summer, in Figure A-5, the placement is quite obvious in Figure A 6, the placement requires slightly more of a judgement call. Figures A 7 and A 8 show these same two tubes with peak to-peak measurements being made, but in both cases the dots have been placed at locations where the normal max-rate dots would be located. The reduction in the voltage amp!!tude measurement is 19.3% in Figure A-7 and 16.3% in Figure A-8. While this is an accepted method of analysis for phase angle measurements, this is not acceptable for the voltage amplitude measurements required, in Figures A-5 and A-6, the locations of the dots for the peak to-peak measurements being performed from Mix 1 show the corresponding dots on the 550 kHz raw frequencies as also being located at the peak or maximum points of the flaw portion of the lissajous figure, in no case should the dots to measure the voltage amplitude be at locations less than the maximum points of the flaw portion of the 550 kHz raw frequency. Figure A 9 is an example of where the dots have been placed on the transition region of the 550 kHz raw frequency data lissajous figure, it is clear from the Mix 1 lissajous figure that this does not correspond to the maximum voltage measure...ent. The correct placement on the Mix 1 lissajous figure is shown in Figure A 10. This placement also corresponds to the maximum voltage measurement on the 550 kHz raw frequency data channel, in some cases,it will be found that little if any definitive help is available from the use of the i raw frequencies. Such an example is shown in Figure A-11, where there are no significantly sharp transitions in any of the raw frequencies. Consequently, the placement of the measurement dots must be made completely on the basis of the Mix 1 channel lissajous figure a as shown in the upper left of the graphic. An even more difficult example is shown in Figure A 12. The logic behind the placement of the dots on Mix 1 is that sharp transitions in the 5 residual support plate signals can be observed at the locations of both dots, in the following graphic, Figure A 13, somewhat the same logic could be applied in determining the flaw-like portion of the signal from the Mix 1 lissajous pattem. However, inasmuch as there is no A9 l
sharp, clearly defined transition, coupled with the fact that the entry lobelnto the support plate is distorted on all of the raw frequencies, the dots should be placed as shown in Figure A 14. This is a conservative approach and should be taken whenever a degree of doubt as to the dot placement exists. . A.3.3 Copper .nterference in situations where significant copper interference in the eddy current data is noted, the eddy current technique basically becomes unreliable. This results from the unpredictability of the amount and morphology of copper deposit on the tubes which may be found in operating steam generators. The above observation is true both for bobbin and RPC or any other eddy current probe. Fortunately, significant copper interference does not occur in the support p;cte crevice regions of V. C. Summer. This is confirmed by destructive examination of the rupport plate intersections on tubes pulled from V. C. Summer. No plated copper was found on the tube OD within the support plate crevice, although some copper was found in the crevice deposits and some minor plated copper patches outside the crevice region were sometimes observed, inspections with RPC and bobbin probes have shown good correlation for flaw amplitudes exceeding 1.0 volt;i.e., more than "W of the bobbin signals identified have been confirmed to exhbit flaws to the RPC probe. This suggests that spurious signals from conductive deposits do not result in excessively high false call rates. Furthermore, signals judged as NDD with the bobbin guidelines have been confirmed to be free of RPC detectable flaws. Copper is a concem for NDE only when deposited directly on the tube surface. Copper particles with the sludge in the crevice do not significantly influence the eddy current response. To Westinghouse's knowledge, no pulled tubes have been identified with copper
- deposits on the tube at the TSP intersections - in contrast with free span tubing Copper deoosits have not been identified at TSP intersections in the Farley S/Gs and all copper alloys have been removed from the secondary system. Thus, it is not expected that copper interference will significantly influence the TSP signals in the V. C. Summer S/Gs.
A.3.4 Alloy Property Changes . This signal manifests itself as part of the support plate " mix residual" in both the differential . and absolute mix channels. It has often been confused with copper deposit as the cause. Such signals are often found at support plate intersections of operating plants, as well as in the A 1o
model boiler test samples, and are not necessarily indicative of tube walf degradation. Six support plate intersections of Plant A, judged as free of tube wall degradation on the basis of the mixed differential channel using the guidelines given in Section A.P.5 and A.2.6 of this . document, were pulled in 1989. Examples of the bobbin coil field data are shown in Figures A 15 through A 17 (inspection data from a plant with 7/8 inch diameter tubing). The mix residuals for these examples ara between 2 and 3 volts in the differential mix channel and no discontinuity suggestive of a flaw can be found in this channel. All of them do have an offset in the absolute mix channel which could be confused as a possible indication. These signals persisted without any significant change even after chemically cleaning the O.D. and the 1.D. of the tubes. The destructive examination of these intersections showed very minor or no tube wall degradation. Thus, the overall ' residuals" of both the differential and absolute mix channels were not indications of tube wall degradation. One needs to examine the detailed structure of the " mix residuar (as outlined in Sections A.2.5 and A.2.0) in order to assess the possibility that a flaw signalis present in the residual composite. Similar offsets in the absolute channels have been observed at the top of the tubesheet in plants with partial length roll expansions;in such cases, destructive examination of sections pulled from operating plants have shown no indication of tube wall degradation. A 3.5 Dent interference For the population of intersections analyzed that contained flaw-like indications, no denting has been observed in the V. C. Summer steam generators. However, inasmuch as denting could occur at some future time, a brief discussion of the effects of denting interference is being included. The 550/130 kHz (differential) support plate suppression mix reduces or eliminates the support plate and the magnetite which may be present within the support plate, but the resulting processed signal will still be a composite of the flaw, other artifacts and a dent,if present. These composite signals represent vectorial combinations of the constituent effects, and as such they may not conform to the behavior expected from simple flaw simulations as a function of test frequencies. . The effect of the dent on the detection and evaluation of a flaw signal depends on both the relative amplitudes of the flaw and dent signals and the relative spatial relationship between - them. If the flaw is located near the center of the dent signal, interference with flaw detection may become insignificant, even for relatively large dent to flaw signal amplitude ratios. The flaw signal in a typical support plate dent in this event occurs at mid-plane, away from the A 11
[ support plate edges where the dent signal exhibits maximum voltage; tiius the flaw in the - middle section of the support plate appears as a discontinuity in the middle of the composite signal. It can be observed in Figures A 18 through A 23, from Plant A, that one can readily - extract a flaw signal even when the flaw signal to-noise (dent) ratio (S/N) is less than , unity. The question of S/N ratio requirements necessary for flaw detection is not a number that can be readily determined; but as can be seen from these figures, even with ratios as low as 0.184/1.0, the flaw signal can be detected and evaluated. The greatest challenge to flaw detection due to dent interference occurs when the flaw occurs at the peak of the dont signal. Detection of flaw signals of amplitudes equal to or greater than 1.5 volts - the criterion associated with confirmatory RPC testing in the presence of peak dent voltages can be understood by vectorial combination of a 1.5 voit flaw signal across the range of phase angles associated with 40% (110 degrees) to 100% (40 degrees) through wall penetrations with dent signals of various amplitudes. It is easily shown that 1.5 voit flaw signals combined with dent signals up to approximately 8 volts peak to-peak will yield resultant signals with phase angles that fall within the flaw reporting range, and in all cases will exceed 1.5 volts. All such dont signals with a flaw indication signal will bo subjected to RPC testing. To demonstrate this, one-half the dent peak to-peak voltage (entrance of exit lobe) can be combined with the 1.5 volt flaw signal at the desired phase angle. The Plant A inspection data is shown in Figures A 15 through A 20 to permit flaw detection and ! evaluation for flaws situated away from the peak dent voltages. The vector combination analysis shows that for moderate dar.t voltages where flaws occur coincident with dent I entrance or exit locations, flaw detection at the 1.5 volts amplitude levelis successful via
- phase discrimination of combined flaw! dent signals from dent only signals.
The vector addition model for axial cracks coincident with denting at the TSP edge is illustrated as follows: s D e w p A - - - - - - where R - Resultant Signal Amplitude - A - Flaw Signal Amplitude A-12
I D = TSP Dent Amplitude one edge (Peak to Peak- 2D) ' 0 - Flaw Signal Phase Angle (100 40';40%-110') .
$R= Phase Angle of Resultant Signal and 2 R - (D+Acose )2 + (As!nB )2
*R - arctan'1 (Asin0/D+Acose ))
For dents without flaws, a nominal phase angle of 180' is expected. The presence of a flaw results in rotation of the phase angle to <180' and into the flaw plane. A phase angle of 170' (10' away from nominal dent signal) provides a sufficient change to identify a flaw. For dents with peak to-peak amplitudes of 5 volts, D-2.5v and the minimum phase angle rotation 4 ($R) for a 1.5v ODSCC flaw signal greater than 40% throughwallis predicted to be at least 15', sufficiently distinguishable from the 180' (0*) phase angle associated with a simple dent. Supplemental information to reinforce this phase discrimination bas;s for flaw identification can be obtained by examination of a 300/130 kHz mix channel; dent response would be lessened while the O.D. originating flaw response is increased relative to the 550/130 kHz
- mix. RPC testing of indications identified in this fashion will confirm the dependability of flaw signal detection. Intersections with dont voltages exceeding 8 volts, for which flaws -1.5 volts may not be detectable, are candicates for RPC sampling of dented TSP intersections.
A.3.6 RPC Flaw Characterization The RPC inspection of the intersections with bobbin coil flaw indications >1.5 volts is recommended in order to verify the applicability of the attemate repair limit. This is based on estab?!shing the presence of OD9CC with minor IGA as the cause of the bobbin indications. The signal voltage for the RPC data evaluation will be based on 20 volts for the 100% throughwall. 0.5" long EDM notch et all frequencies. The nature of the degradation and its orientation (axial or circumferential) will be determined from careful examination of the isometric plots of the RPC data. The presence of axial ODSCC at the support plate intersections has been well documented, but the presence of circumferential indications related to ODSCC at the support plate intersections has also been established by tube pull in two plants. Figures A-24 to A 26 show examples of single and multiple axial ODSCC from V. C. Summer. Figure A-27 is an example of a circumferentialindication related to ODSCC at a A-13
tube support plate location from another plant. If circumferentialinvolvement results from circumferential cracks as opposed to multiple axial cracks, discrimination between axial and ,
~
circumferential!y oriented cracking can be generally established for affected are lengths of about 45 degrees to 60 degrees. RPC resolution is considered adequate for separation between circumferential and axial
- cracks. This can be supplemented by careful interpretation of 3-coil results.
Circumferential cracking is not expected in the V. C. Summer steam generators due to the absence of deriting, in the event that a circumferential indication is found at a support plate location, ultrasonic exarnination could be employed for verification. If a well defined circumferential indication is identified at a tube support plate location in the V. C. Summer ' steam generators (>60 degrees circumferential extent), guidelines for RPC interpretation will be reviewed and consideration given to a supplemental ultrasonic inspection for resolution of the degradation mode. The isometric graphics which are produced to illustrate the distribution of signals in a TSP may sometimes exhibit distributed extents of flaw content not readily identified with the discrete axial indications associated with cracks; this may occur with or without the presence of crack signals. The underlying tubing condition represented by volumetric flaw indications Is interpreted in the context of the relative sensitivity of various flaw types (pits, wastage / wear, IGA, distributed cracks) potentia!!y present. The response from pits of significant depth is expected to produce geometric features readily identifiable with small area to amplitude characteristics. When multiple pits become so numerous as to overlap in the isometric display, the practical effect is to mimic the response from wastage or wear at comparable depths. In these circumstances the area affected is generally large relative to the peak amplitudes observed. The presence of IGA as a local effect directly adjacent to crack faces is expected to be indistinguishable from the crack responses and as such of no structural consequence. When IGA exists as a general phenomenon, the EC response is proportional to the volume of material affected, with phase angle corresponding to depth of penetration and amplitude relatively larger than that expected for small cracks. The presence of distributed cracking, e.g. cellular SCC, may produce responses from microcracks of sufficient individual dimensions to be - detected but not resolved by the RPC, resulting in apparent volumetric responses similar to A-14
wastage and IGA. . For hot leg TSP locations, there is little industry experience on the basis of tube pulls that true volumetric degradation, i.e. actual wall loss or generalized IGA, actually occurs. All cases reviewed for the APC present morphologies representative of ODSCC with varying density of cracks and penetrations but virtually no loss of materialin the volumetric sense. For cold leg TSP locations, considerable experience has accrued that volumetric degradation in the form of wastage has occurred on peripheral tubes, favoring the lower TSP elevations. Therefore hot leg RPC volumetric flaw indications within the TSP Intersections will be presumed to represent ODSCO, while only peripheral tube, lower TSP locations on the cold leg with RPC volumetric flaw indications will not be so characterized. A.3.7 Confinement of ODSCC/lGA Within The Support Plate Region in order to establish that a bobbin indication is within the support plate, the displacement of each end of the signalis measured relative to the support plate center. The field measurement is then corrected for field spread (look-abead) to determine the true distance from the TSP center to the crack tip. If this distance exceeds one half the support plate axiallength (0.375"), the crack will be considered to have progressed outside the support plate. Per the
~
repair criteria (Section 10), Indications extending outside the support plate require tube repair or removal from service. The measurement of axlal crack lengths from RPC isometrics is presently a standard portion of the V. C. Summer EC interpretation practices. For the location of interest, the low , frequency channel (e.g.10 kHz) is used to set a local scale for measurement. By establishing the midpoint of the support plate response and storing this position in the 300 kHz and 400 kHz channels, a reference point for crack location is established. Calibration of the distance scale is accomplished by setting the displacement between the 10 kHz absolute, upper and lower support plate transitions equal to 0.75 inch. At the analysis frequency, either 300 kHz or 400 kHz, the ends of the crack indication are - located using the slope intercept method;l.e., the leading and trailing edges of the signal pattem are extrapolated to cross the null baseline (see Figure A-28). The difference between j these two positions is the crack length estimate. Altemately, the number of scan lines A-15
r
- indicating the presence of flaw behavior times the pitch of the RPC provides an estimate of the crack length which must be corrected for EC field spread.
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L I 4 % > % i= i . R I I \ I \ r f Figure A-2, Bobbin Coll Amplitude Analysis of ODSCC at TSP - Improper Identifcation of Full Flaw Segment Resulting in Reduced Voltage Measurement When Compared to Figure A-1 A-18 ..
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L I 7 % *> % <= I I I \ l \ Figure A 3. Bobbin Coil Amplitude Analysis of ODSCC at TSP - Improper identifcation of Foll Flaw Segment Resulting in Reduced Voltage Measurement When Compared to Figure A-1 A 19
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1 ( l 1 l e Figure A 13. Incorrect Placement of Dots Marking Mix 1 Ussajous Figure for R30074 A 29
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Figure A 14. Correct Placement of Dots to Effect Max. Voltage Amplitude R30074 l I j A 30
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..1. < p' E
( b c - i __ 400/100 Differential ,,
. < , , . . . . ,n
!( l v., ,.,. x is. .. v., ..., m. ., ..
-. . .,* . : ~.
j (o
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(416 i 1
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e i 1 9 1 ..% % i P .m. '
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Figure A 16. Example of Bobbin Coil Field Data Absolute Mix with No ODSCC A-32
m c. i- ,, . nii a v .... . n, a i 4.. en a . nii a .a , i .L X "" 5' 400/100 Absolute .n wn.c [ s
<c i i i .. . . ,.,, ,,e
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/%
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. i ... 1 4
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- -- . .c.,. .ui,
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= *
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,n. i .,,!.
l
;' w !i,.! '.. , i cs i w y _% - i ii .
) j 3 .} } i. i , i i ..! .
Figure A 17. Example of Bobbin Coil Field Data - Absolute Mix with No ODSCC 1.. A.33
.e # .in i , , :s v i is .se := cm i
- 5.54 i'l als t u ,
OttL ses temast _JL := l es se 1 1 3a j (Ittuti 'It I ttM g SPttp eeeee4 n/d toain
'[_ _
.mm== emme==
n i Out vpp a 5) M LM Omf e _ves e 44 W 144 8:5 #1M 1 298 l it ett us CM i 281 4.8% 1 I ! l .
)
I / J/ '" l VP M
.m i
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. .2 mm in -
.= .., i.n ma -
un _.
' r=.
o I a , m, e ws ii>isi l - 1 i 5 8 og T --y --JE" . , t i i.i .
~
- Figure A 18. Example of Bobbin Coil Field Data - Flaw Signals
- for ODSCO at Dented TSP Intersection from Plant A l
l l A 34
- m. _ _ _ . . _ . _ .
I 1 4 l
- 3. ! . . . c. , . u i. ... c. . ., ,i, .. .i. . >r
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" P
[(.fEMil ,,, ] ff. N [
*1 > (bp(ID ,eeoeei 6m/seg tpa1.
Y, Y "" {
/;*~
at I - I
. ,, . ,,, -, . . . , m. ,
i . Q+
~
s V
.. .. m a na e. mm we == i e na m .!,: , ~
Figure A-19. Example o' Bobbin Call Field Data Flaw Signals for ODSCC at Dented TSP Intersection from Plant A A.35
9 al *e4 the Cm I 5 14 1% als .g og i els 4 e > CM 1* 63 MI
. to n at
' ~
i
! ce l et 04 l rl tw ENYtufl f(C l f ta 5
; p @ intra e esiiu tesim vse 1 ll M 108 WT vpp 3 El M IM *t
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" 4 L
t %__ 3 i
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- unmei e setti em
- ande., e ettMr%7IMI
- tait m-
_ vpp t.30 M 188 698 vpp l.13 M 73 795 N5 e 4 e 813!*ll!e iie! 8
;- ( l ) . , ,e
- ! s' % b e 5 t i is !
j - 4 g . , i i. Figure A 20. Example of Bobbin Coil Field Data - Flaw Signals , for ODSCC at Donted TSP Intersection from Plant A A.36
.e l a m i e ct, si n .= c= i in in ermi en wt.
1/5 i t.agg in I .. m I,i tittati ttc I stu en ( _. m ia .e n ,s.i. tui,
- m. 4 'h \
_ 6 t
% /
t I
; . , is u a iii av .. . n m i.e we
~~ il el .e. tas CM i 3M $ es iat als i gn i
I
.=.
- O I
. l I
-- t m l w........ (Wik j
_. v. e.m m im in v. i.ie an m r= o,.ini. .iini
. i i o 1 i
1 b k ,b ' '! l "J ]lll p _ y 4 4- -- Figure A-21. Example of Bobbin Coll Field Data Flaw Signals l for ODSCO at Danted TSP Intersection from Plant A 1 l A 37 l l
.e i . . . c. i . , is ... ta, c., i o i ). i i aii i . ,,
\ .flL
_l __
=
ses seems
- . - 1 . = 1. s i-
~~~
uitwil etc l ttu i 7 g _. v v wern i s =1., tui. v % ,.
' ~~
.n eo a sn ~53 ., t ., m s.e aus d'
t il .et tas 04 i O l 14 ist als i e4
.J _ .
- __ W
- %) -
- m ..
. e. ....m ...
. m
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via _a en
= = ._i : .
i v.>inv:,
- o I 4 l b 4 b ", '
i y .- , !!'. i4 ,
; 7 i .g 4 ,
Figuro A 22. Example of Bobbin Coll Field Data - Flaw Signals for OCSCC at Dented TSP Intersection from Plant A A 38
e 3 1 .i. i , en > . e *. . = cm i 3 .n i, .i ,, __- em d 1/6 14 I
- i + m i11 'i- l
""' l~
- oftati N 1 tra i 7 5
wun ....... w r i.,
": ?~ f a summmmmme J l e ,
- i. . m iu i l _n. et
- 3. .,M s.t1 m in ot -
6.** *00 sh. CN 1 23 les mis t M e ~ ' 6 sa-,m. e-wi. i = n . . .m c.u
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=
iu i .iiiiiiii-s i p / 2 t' < w " 4 - --
> :=--
,' l .' ;
s - s- s c, .. i, ,
> r i r e
Figuie A 23. Example of Bobbin Coll Field Data . Flaw Signals for ODSCC at Dented TSP Intersection from Plant A A-39
T ex 4 , en e v . .a >w m c= 4 a wM u > g gg ;g 3 l'0.n , ,'l 2D ED 'E
,,...n., .. m .n i
. . = = u n i, n
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. 2H+4.M 4.n Me
# X
~
Aulu famu AXl E EXTs 8.$3 IN i Figuro A 24. Example of RPC Data for Axial UDSCC Indication from V. C. Summer A-40
Je a4y cn a v 4.as aos na cm e e5 GG M E > g g.g Y I'O.,i., f MttflMt M. 9F Stas till Ele E W il 3D S/s 3D it (El i sileAsas til/scun styttlit:4/tlls gM j eg,3} l,l llgg
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l uim. vitu g , ciecertarmis crt 04 us-
..u
[ 4194.42
-0.61 M.
MIE EXis
%/
e.24 ill una taux v ~ !' Figure A 25. RPC Data for Axlal ODSCC Indication at V. C. Summer A-41
.m m o. cx . in a wua, a y g g,g
- 3. cn v os i v ge s.
u 3D 3D QE3 A : wien=,, wien.. .. its en iart es. OF Was ties: 69 gran niunu mex., nuna niune. _ e., 1 ..ii,g
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9r - x 388 _ una vita . CleCWFEMX1141. pti 34 des l 4.43 / _ f -
. , a ,-
uis ext. e..e in [ Figure A-26. RPC Data for Multiple Axial ODSCC Indications (MAI) at V. C. Summ~ l A-42 l l
.. - . ..-. ..- -. -..- -._.- - .-_..__ . - . - - ....- - -.--._ ..-.. . - - .~....~ .-_.. _ _ ... -
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(+ -e.sa l} 88' d e '4u 9
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Figure A 27. RPC Data for Circumferential ODSCC Indication at TSP i l A 43
i. l
.<l 1
1 . I j a l i.cH 2 WRT .,' . cH 2Hom:;l cHee<L Ho - t lro e now 3 ccL s R n.nnce cam - e i i J F'E9 400 kHz 9M*4 1% p DOTATICM - 364 DEe i p ' h LUTSTWIP DWT m ..... k. ......, CH @ed - 24 FWG 440 kHz CPAM 1% 4 ' ROTRTICM - 354 DEO t19tfI.itt_1f.StfEI i cHee .- aH FRE9 400 kHz 1 i SMD4 1% ' 3 5TD +1se.9 ROTRTION - 35< (Et i ECIlllLCQtf M 11IBf IWE - CKF ALA.T -
. a em - <
l 12:23:29 fut
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-lM 11 !
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4 1 I 88 sun sors 2 k 4 4 Figure A 28. Location of ona end of an Indication using an RPC probe . \ . A 44
Appendix B: Flegression Analysis , B.1 Introduction The analysis of the relationship between two variables is generally termed either
, regression analysis or correlation analysis, in addition, one may also find the term confluence analysis. For each, the objective is to establish a mathematical model . describing a predictive relationship between the variables. The use of the term regression is frequently interpreted to imply that some sort of causal relationship exists while correlation has been reserved for non causal relationships. Other differentiations between the two terms involve the nature of the variables, i.e.,
whether or not one or both is stochastic. In addition, the term regression is also frequently used to mean the process by which the parameters of a relationship are determined. For the purposes of the evaluations reported herein the name regression analysis is
- used in the broad sense of covering the aspects of the fitting of a curve, i.e, equa-tion, to observed data points, where concern is with the slope and position of the curve that best fits the data, and to the analysis of how well the data points can be represented oy the curve, i.e., the correlation analysis. The correlation analysis has two aspects, one is a measure of the degree of covariability between two variables, and the second is as a measure of the closeness of fit of a regression line to the distribution of the observations. The statistical analysis is performed for the purpose of establishing a stochastic dependence, and does not, nor does it have to, demon-
- strate the existence of a causal dependence.
For the analyses dealing with the APC it is desired that models be developed relating the burst strength and leak rate of degraded tubes to the morphology of the degrada-tion. Unfortunately, the degradation morphology is only known exactly for tubes which have been destructively examined. However, a third variable, based on the ! non destructive examination of the tubes, is available which is also directly related to the morphology of the degradation. Each degradation state is taken to correspond to a set of quantifiable characteristics or variables, such as the burst strength (mea-sured by a burst pressure test), the leak rate (measured as a function of differential pressure), and a non-destructive examination (NDE) response, e.g., eddy current bobbin coil) signal amplitude in either an absolute or differential mode. Since the field examination of the tubes la based on the NDE response it is appropriate to examine the relationships between the first two variables and the third. The degradation process determines the magnitude of the evolution of each variable, however, the degradation process is complex and the morphology and time history
- will vary even under conditions which would normally be termed identical. Thus, it is expected that the correlation between any pair of the three variables may have significant scatter. This is expected even if each of the variables is measured with ^
perfect accuracy and contains no measurement error, cas.=, , 8.1 m 1
lt is to be noted that the causative factor relative to the magnitude of each variable is the crack morphology, and that none of the three characteristic variables can be considered to be the cause of the other. This rneans that for any pair, either may be treated as the predictor and the remaining variable treated as the response. ' Once a correlating relationship has been established, either variable may be used to predict an expected value for the other. For example, a correlating relationship may be '
' mathematically determined using burst pressure as the response and bobbin ampfl.
tude as the predictor. Once the relatinnship is known the bobbin amplitude associat- ~ ed with a given burst pressure can be calculated. The general, linear, first order analysis model relating two variables is given by y, = a o+ a, x, + c (B.1) where y, is taken here as the response or predicted variable, and x, as the predictor, or regressor. The 4, or error, term accounts for deviations from a perfect prediction. In order to establish confidence and prediction limits on yj, the error is assumed to be normally distributed with a mean value of zero and a variance that is uniform over the range of interest. An analysis it then performed to determine the best values of so and a, to use in equation (B.1). Three methods are commonly used for the analysis, ma>Jmum likelihood octimation, least squares (LS), ; and weighted least squares (WLS). For maximum likelihood analysis the values of soand a, are found that maximize the probability of obtaining the observed re-sponses. The use of maximum likelihood analysis is formally correct, however, if the errors are normally distributed, the maximum likelihood estimators (MLE) will be ! identical the estimators obtained using least squares. If both variables are stochastic l and the errors are normally distributed then the application of least squares still leads to the maximum likelihood estimators of so and a 3. The LS method is based on minimizing the sum of the squares of the errors, also referred to as reciduals, between the observed and predicted values, thus, the best values of so and a are 3 those that make I n { (Yj- 9)2 f (B.2) l=1 a minimum, where the caret indicates the predicted value, 9, = a o+ a 3x, . (B.3) Expression B.2 is differentiated with respect to ab and a3and the resulting expres-slons set equal to zero and solved for the coefficients. For WLS the same expres-
,, _ B.2 .
sion for the errors is established by considering the error term, r, to be weighted non-uniformly, i.e., the error distribution is - t, - N(0.I,o2) (B.4) and the expression to be minimized becomes f w,( y, - Y',)2, (B.5) s.1 where the I andf hence the w are j known. In situations where the variance of the response is not uniform it is possible to find appropriate weights such that the resulting estimators are MLE's. For the unweighted LS analysis the slope of the regression or correlation line is found to be
, , E (Xi -I)(Fi-E) , (B.6)
E (xi-E)* where the summation limits are understood. The intercept is then found as ao = y - a, F (B.7) . If y has beer' regressed on x. If xis regressed on y the slope will be E (Xs - X)(Yi-E) (8.8)
~
[ (Yi-E)* relative to the ordinate, or y, axis. If this is reckoned to the x axis, i.e., the abscissa of the original coordinates, the slope is a, . b (Y'~ E) _
, (B.9)
E (x -x)(yi-y) o If the data used for the analysis contains significant scatter the values found by (B.6) and (B.9) can be quite different. A rough visualization of this can be obtained by picturing the smallest ellipse that can be drawn that envelopes all of the data points. A line connecting the largest and smallest abscissa values of the ellipse will approxi-mate the regression of the y variable on the x variable, while the line connecting the maximum and minimum ordinate values will approximate the regression of the x a variable on the y variable. For the APC analyses the objective is to relate burst pressure and leak rate to bobbin voltage. This means that bobbin amplitude is depicted as the abscissa variable while burst pressure and leak rate are depicted as ordinate variables respectively. Sub.h depiction does not imply the direction of the regression analysis performed. B.3 .=
t i ! 5.2 Consideration of Variable Error ! If there is measurement error present in the predictor variable the slope obtained : i from the regression analysis will be biased. For the predictor, say X, the total , : j variance will be 1 og , y 2gam a i, (B.10) t v,here te subscript m indicates measurement error, it can be shown that in this ! I case the expected value of the calculated slope, a3, will be l i ! a, ,
?
i ag = l om (B.11) j 1 +q j. og i j where ag, is the true value or the value that would result if no measurement error l was present, and - 1 o n , E (4- E , (B.12) . n-1 a It is noted that sowould be found from equation (B.7) as before. A key point to note , j is that the calculated slope under predicts the true slope (without measurement - , e error). - If the measurement error is known, and is uniform, its effect on the analysis slope can be calculated directly and the appropriate slope to be used for prediction , l would be .
' )- '
!. (B.13) a3 = ai 1 . O i g x, l When the error variance is known and can be expressed as's fraction of the variable 1 variance the slope will be affected by a like amount.. i I When the error variance is not known an estimate of the'true slope can be made-i- using the partitioning . technique developed by Wald and 9A:: ;i=itly improved upon - by Bartlett. The technique consists of partitioning the data into three groups based I upon the ordered predictor variable. The line joining the controids of.the upper and lower groups.is an unbiased _and consistent estimator of the true siope, if the slope thus found is close to the slope determined without considering measurement errors ,
; then the measurement errors are not significant;
! It is noted that the consideration of measurement errors does not lead to any criteria- - I- for selection of a regression directionc. J 8
~. .~ _.--...-. _ _..--_.u.. ..._._....._....m ..a_ ._ .
a.. . - ,- . - . . . , ,J. .;..,_ _. . .. C ~ n. _._-
B.3 Detection cf Outilers if the errors are normally distributed the application of LS to det' ermine the coeffi. cients of the regression equation minimizes the variance of these estimators. The coefficients are also the MLE's. A drawback of the LS technique is that it is not very robust. This means that the fitted line may not be the best estimator of the correct
. relationship because it can be significantly influenced by potentially outlying data. In addition, the resulting fit may be such that potential outliers become hidden if exam- .
ined after the analysis is performed. There are established methods for identifying influential data that may result in a distortion of the regression line. Such methods fallinto the categories of regression diagnostics and robust regression. Robust regression met, hods are designed to be insensitive to potential outilers, and can be used to identify outliers based on the residual errors from the robust regression line. A rather simple example of improving the robustness of the fit would be simply minimize the sum of the absolute values of the residuals instead of the sum of the squares. This provides significant improvement if the outiler is in the y direction for a y on x regression, but is not resistant to outliers in the x direction. One very robust technique is termed the 'least median of squares,' or LMS. The best regression line (or polynomial) is the one for which the median of the squared residuals is a minimum. The drawbacks to this technique are that there is no closed form solution, and techniques for the determination of inference regions would be difficult to apply. However, the determination of a reasonable solution is quite easy using a computer. The algorithm proceeds by u. awing sub-samples of a given size l from the data set. For each sub sampio, regression line coefficients and the median of the squared residuals are calculated. The coefficients of the minimum median solution are designated as the LMS solution. A median based scale estimate (analogous to the standard deviation) is determined for the identification of outilers at a two-sided 98% confidence level, or o one sided 99% confidence level. The data for th( APC were examined using the LMS robust regrassion program PROGRESS by Rousseouw and Leroy. It is noted that the application of robust regression is not for the automatic deletion of improbable data points, only for the identification of potential outfiers. The rejection of any data is then based on an evaluation of the circumstances surrounding the data collection to search for possible errors. B.4' Selection of a Regression Coordinate System For the analysis of continuous variable data four, alternatives were examined for each correlation. These choices are listed in Table B.1. For each case, the correla. - tion coefficient, r, measuring the ' goodness of fit' of the regression line was calculat-ed. The correlation coefficient is a measure of the variation of the data explained by the regression line, thus the largest value is indicative vi the best fit. The expression l I B.5 n.= L
._7__._
l Table B.1: Fitting Options' Considered for LS Regiession 1 Abscissa Ordinate Relation Linear Unear y, = a , + a , x, y, = a, + a, log (x,) Logarithmic Unear - Unear Logarithmic log (y,) = a, - + a, x, - Logarithmic Logarithmic log (y,) = a, + a, log (x,) l for the square of the correlation coefficient, known as the index of determination, is , r , E (fi-E) . (8.14) E (Fi-E)* The index of determination is the proportion of the total variation about the mean of - the predicted variable that is explained by the regression line ;The scale combination i yielding the largest index of determinadon, and, hence, correlation coefficient, was: riected for the analysis, in the event that the predicted variable for the logression is the Imarithmic transformation of a physical variable, the above calculation is perform ed on the untransformed variable, it is readily apparent,- however, that for data with a range of several orders of magnitude, e.g. bobbin amplitudes ranging from O(0.1 volt) to O(100 volts), the use of a loganthmic scale is appropriate, it is also to be expected that the variation of observed. voltages would be normally distributed about ,. the log of the voltage. The same is true for the leak rat 6 which ranged from O(0.02 _ 1/hr) to O(5001/hr) for specimens for which lerking was obes.wd. E.5' Selection W 4 Regression Diiection As'noted in he inkoduction to this appendix, the bobbiri amplitude does not cause t L the'olmerved burst pressum and vice versa. The same is true for the relation - betweer bobbin voltaga and leak rate. Thus, the regression drection is not specified - by the chdos cf varimes. L l- . The objective (A performing the regression analysis is prec5ction. For all practical.
- purposes the bobbin voltage will be used as a predictor of burst pressure and leak -
rate. However, the intended use does not dictate the designation of predictor and: - response for the regression analysis._ The LS fit simpiy finds'the line such that the variance of the_ responses is minimized relative to the regression line? As'previously " J , noted, once the LS fit has been performed either variable can be predicted from the l {other. in addition, inference regions or bands established for prediction in one-K e._, . B.6 - ,, --
, . ._ . _ _ __ ... m _
7 i i 1 L direction m:y be similarly use in tha revsrso direction (although the tIrminology is ' changed to discrimination), . For a regression of y on x, the average of future values of yo for a given xo.is
)
bounded with a (1-a) 100% level of confidence by , i-(yo - a o - a 4]2 = f*-e.n-2 s 3 3 a 1. .(% E) , (B.15) U [ (xi-x)2 I and an individual future value of yo for a given xois bounded with a (1 a) 100% - I j level of confidence by ( } (yo - ao - ai xo]2. , f,2- e.n-2 8* 1 +1+ . (B.16) - l ,. - E (x1-X)* , I where s 2is the' standard error of regression,' l.e., 4
,2gE(yi-fi)* (B.17) i n-2:
Y and t 1-e.n-2 s found 1 from the Student's t-distribution. _ However, for a given yo the - l bounds on xoare four.d by solving equation (B.16) for the values of xothat' satisfy.
- . tne equality. ,
I If the scatter of the data is small, the regressions of x on y and y on x will yield
- - slopes that are similar. . However, for APC analyses the data exhibit significant i scatter and the two regression lines. have significantty different slopes. - In this case it -
p is appropriate to select the regression line based on non statistical considerations. l - Such considerations may be known end points of the regression lir v,.e.g., burst : pressure for_ non-degraded tubes, or comparison of the slope with theory based
- _ - results. For either regression direction, inference regions can be determined.
B.6 Signifloance;of the Regreeston. u i..
- - The signiilcanoe of the regression is evaluated by calculating the improvement in the l
estimate of the predicted variable based on knowledge of the regressor variable.= For
- the' APC analysee this is the same as determining whether or. not the estimate of the.
burst pressure 'or leak rate for a tube is improved by knowing the bobbin coil voltage amplitude. For a linear,1 8order regression this is the same as testing toLdetermine if:
~
r
. - a zero slope is probable. Elf the confidsnce interval for the slope includes zero then the relationship between the predicted variable and the regressor could be accidental, i.e., due to random errorf The ' actual determination may be made by calculating the :
J l
. B.7 , , . _ ,
- cosm _
- . , . . .--~ ,. , . ;,;...A .- .:., ,-
a confidence interval on the slope, a3, to see if it includes zero, or by testing the null hypothesis that the true slope is zero. In practice this is stated'as 8 Ho: a3 =0 (B.18) , Hi : a*
- 0. .
if the null hypothesis, Ho, is true, then ratio of the mean square due to regression . (SSR) to the mean square due to error (SSE), i.e., the mean square of the residuals, follows an "F" distribution with the regression degrees of freedom (DOF) in the numerator and the residuals degrees of freedom in the denominator. - For a linear analysis with k regressor variables and n data points, then MS Regression = F1 -a,r n-r-1, (B.19) MS Error where 100-(1-a)% is the associated confidence level. (Note that'a is the area in the . _
- . tall of the distribution.) If the true value of the slope is zero ihon both mean square -
(MS) Regression and MS Error are independent estimators of the true value of the error variance.1 Since they are both . estimates they.would not ce expected to be_- exactly equal, however, it would be expected that they would be nearly_ equal so that their ratio would not be too far from t'nity, it is noted that the F ratio and the Index of Determination, ra, are both calculated from the sums of squares of the variables, so - SSR/k- , r 2 , n- k-1: (B 20) -
~ F*'"'"*" k
= SSEl( n- k-1 ) .1 - r*
and a critical value of 2r for a selected critical a can be fourv'ss - k f -a.k.n-k;s 1 p (n- k-1) + k F3 _,,g,,.u.3
- If the value of r2 found from tne regression is greater than the critical value from-(B.20) the null hypothesis, Ho, i.e. that the slope is zero, would be rejected, and the attemate hypothesis, H3 , that the true slope is not zero would belaocepted. For .
. example, consider k=1, n=15, and a=0.01. Then, from equation (B.20) we find a critical value of raof 0.411 and a critical value of r of 30.64_1 . If the' regression value of r2exceeds 0.411 the regression is significant at a level greater than 99%
~
I 4' r _ B.8 ., A
,- - - e ,e m+ , , , .____.i________________.m
For a 1dorder regression, equation (B.19) can be_ rearranged as r i t -a/2,n-2 " 1-r2 /(B.22) 3 n ' i.e , a t distribution with n-2 DOF's. Given a value of r from the regression analysis,' . a value of t can be calculated and a significance level determined.- For the same 2 example as above, we consider r =0.41_1 and n=15. _ From equation (B.22) we find t=3.0 3, and a significance level of 100{1 a)=99.1%,_wh_ich agrees with the _above determination it is to be noted that for a small number.of regressor variables and a large number of data ' points, the square of the correlation coefficient'does not have to be veiy close to one to reject the null hypothesis and accept the attemate hypothesis that the slope is not equal to zero, thus implying that a correlation does exist, e {
^
4 l. I l l
,t,- ,
y_ B.9 ~ .- ! . . . 1 i._ , _
m
REFERENCES:
( . l 1. Fuller, W.A., Measurement Error Models. John Wiley & Sons, New York (1987). Y . L 2. Draper, N.R., and Smith, H., Acolied Reoression Analysis. Second Edition. Second Edition.' John Wiley & Sons, New York (1981), 3 3. Hald, A., Statistical Theory with Enoineerino Acolications John Wiley & Sons, , [ New York (1952). -- l 4. Bardatt, M.S., Fitting a Straight Line When Both Variables are Subject to Error,
- Annsjs of Mathematical Statistics, Vol. 5, pp. 207-212 (1949).
- 5. Wald, A. Fitting a Straight Line When Both Variables are Subject to Error, Annais
- of Mathematical Statistics, Vol.11, pp. 284-300 (1940).
- 6. Rousseeuw, P.J., and Leroy, A.M., Robust Rearession and Outlier Detection.
John Wiley and Sons, New York (1967).
- 7. Yamane, T., Statistics An Introductorv Analggia,2nd Edition, Harper'& Row, j New York (1967).
Hosmer, D.W., and Lemeshow, S., Apolled Loaistic Regression. John Wiley &-- 8. Sons, New York (1989).
- 9. Weisberg, S., Anolied Linear' Rearession. John Wiley & Sons,- New York (1965). , ,
- 10. Lipson, C., and Sheth, N. J.,3.tStistical Design and Analvsis of Enaineerina Exooriments. McGraw HiH, New York (1973).
- 11. Mandel, J. The_StatisScal Analves of Experimental Data. Dover Publications, New York (1984).
- 12. Doming, W. E., Stabstical Adiustment of Data. Dover Publications, New York L (1964).-
~
- 13. CSS: Stabstica Users Manual.LStatsoft (1991).
- 14. M' Revision 6 Fourth Edition, Volume 2, Chapter 27,;7he?
Logistic Procedure, SAS Institute, Inc..--_Cary, North Carolina (19907).
. g-s a
B.10 4 _- i
1 5 1 Appendix C . - Adjustment Procedure for SLB Leak Rates , 'i. C1 Objectives Laboreiec has conducted tests of leak _ rate on pulled tubes from Model E steam generators at Plant ; E.4. Tests were conducted at room temperature with both normal operating (NOP) and steam line - break (SLB) pressure differentials.:These test data broaden the leak rate data base. This - ' appendix evaluates the Belcian test data, and describes an adjustment procedure to scale the room . l temperature data to hot, operating conditions. The adjustment procedure was originally developed
- by Laborelec and slightly modfied as described in this appendix to enhance agreement with more - .
detailed crack leakage analyses obtained using the_CRACKFLO computer code. . l Westinghouse measurement! at SLB conditions (SLB leak rate tests) were obtained at a 2650 pal; pressure differential but with a 350 psia secondary side pressure rather than a prototypic 15 : psia pressure to improve the maintenance of constant test conditions in the measurement facility. I The adjustment procedure developed here is also applied to the Westinghouse measurements to : ' adjust the data to prototypic SLB conditions, in addition, the adjustment procedure is applied to ootain data for a SLB pressure differential of 2335 psi. i
- In summary, the objectives of this appendix are as 'follows. -
- 1. ' Adjust the Belgian room temperature measurements of leak rate to operating temperature,-
and adjust to reference pressure differentials from test pressure differentials.'
~
- 2. Adjust Westinghouse measurements at operating temperature from high primary pressure to the primary pressure expected during steam line break.
- 3. Adjust leak rates at temperature between measured and reference pressure differentials.
This appendix uses the Westinghouse CRACKFLO code to verify the Belgian adjustment procedure. C2. Leak Rate Adjustment Procedure ' Laborelec developed adjustment procedures to scale the B_elgian test data to operating temperature l: and pressures different from the measured ones. The leak rate may be defined as: - L = KA{Ap,T) 10P
'(1)
;p(T)
, where
. L = volumetric leak rate
,, - K= discharge coefficient-A = leakage flow area.~a function of 4p and temperature T
- p = water density, a function of temperature T. ~
lC-1
~
'l r -
, ma ,we- w -t , -, .- -,,,w- h,+E ,i
T- water temperature . ap- pressure differential;py -pp if no f: ashing or py -pyf if flashing at pyf py - upstream pressure, or primary side pressure , pp - downstream pressure, or secondary side pressure p;f a saturction pressure corresponding to upstream temeperature Ty , The pressure differentialis the difference between upstream and downstream pressures of the , cracked passage of the tube. During steam line break, the leak flow may flash on the secondary side. So the effective downstream back pressure is equal to the saturation pressure p;f if it flashes at the saturation point of the upstream temperature Ty. C2.1 Adjustment Under the Same Temperature Considering the measurement conditions of leak rate at c reference temperature To one can write L(dpmo,T,) = KA(4pmo,T, DkT",) (2) L(dPo,To) = KA(4po,T, D# (3) YIch where the subscript o indicates the condition at the reference temporature To. and the subscript m refers to measurement, in addition, -
- apo - desired pressure differential at the reference temperature d#mo - measured pressure differential at the reference temperature Both Eqs (2) and (3) descrbe the leak rate at the refersnce temperature but at different pressure differentials. The leak rate is to be adjusted from the measured pressure differential apm to a desired pressure differential apo under the same temperature To. Eq (2) refers to the test measurements, whereas Eq (3) is for desired conditions, to which the leak rate is to be adjusted. The adjustment is defined by the equation:
L(dPo,T,) = 6L(apmo,T,) @) where the coefficient 6 is a function of pressure differentials derived from Eqs (2) and (3) as: 5 = A(APo,To) spo A(Apmo,T,) APmo . This coefficient Sinvolves two parameters. The first one is the area of crack flow opening. The second one is the pressure differential. Taking into consideration that the crack area A will be . dependent on the total pressure differential ap* acrc'ss the tube, the above expression may be rearranged as follows. C.2
3 6 = A(sp;,To) sp; golsp; (g) A(#Lo,To) sp;o_ &nl#l,,o a
- where the following definition is used.-
a
^ Ap * # ps -p2
- The superscript
- emphasizes the total difference of the upstream and downstream pressures,' ,
regardless of whether water fleshing occurs or not. - When 4p is used without the superscnpt
- it ;
can be py _-pp for, non flashing situation or py - pyffor flashing situation. Explicit function of . Eq (5) will be given later. Eqs (4) and (5) adjust the laboratory test data to any 4po 40 be analyzed at the same temperature. C2.2 Adjustment Under Different Temperature During a steam line break, tube leak occurs at operating temperature 7, say 616*F. The Belgian
- data were measured at the room temperature. One has to scale the room temperature data to the operating temperature.
#o,Tol
- L(4p,T ) = L(4p,T)L(4po,To) (6)
- , - Using Eq (4) Eq (6) becomes L(@,T)
-L(Ap,T) = 6L(4p,,,o To)L(spo,To) (7) .
.- Using Eq (1), Eq (7) appears as follows.
AN 22.! OP L(Ap,T) = L(Ap,,,o,To) S (sp;,To) (8) A P. .spo . The above _ expression describes the procedure to adjust a measured leak rate at dPmoand To to an equivalent leak rate at 4p*o (= 4p*) and T. ' The leak rate L(4p,7) is the volumetric rate at operating temperature T. . But, leak rate detected - in actual plant monitoring is collected at room temperature To. '.aboratory test of leak rate under.- the operating temperature is also generally collected at room temperature. - An equivalent room . , [, : . temperature, volumetric leak rate, say L(Ap,7)Tocan be obtained from ' pL(ApiT) = pol (Ap,T)r, (9) Using Eqs (8) and (9),= Ca3 '
,-ll , - - , - - , -n, , , . , , r
_ _ . . _ . . . _ _ . _ _ _ _ . _ . _ . . . _ . _.. _ _ _ _ . . _ . ~ .. M 1 L(@,7)r, = L(#mo,To) A T6 ^(&;,' Po c) (2- . &o . (10) ! The above equalbn may be written as: i c L(4,T)r, = a yL(@.o,To) -(11) , where l y, A(#;,To) @;
-(12)
A(#Lo,To) 4:,0 F [ p . A(4*,7) g (13) l A(4;,To) Po o ! -y. 4/ME -
- (14) _
4,,ap;,, - The a factor involves a ratio of flow area of the crack o' pening as a function of ap*moandap*o ., i under the same upstream temperature To. The a factor captures a mechanical effect of total -
- pressure differential on crack opening area.' The p factor adjusts temperature effect on the leak -
l - rate traugh flow area change and upstream water density variation under the same total pressure -. j differential (i.e., ap*o - ap'). The yfactor takes care of hydraulic effect of pressure differential ~ on leak flow rate. (- i C2.3 The Mechanical Factor a -
~
4 According to report by Laborelec, significant tearing of crack length took place during tube , pulling. Measured leak rates in Table C-1 represent the behavior of significant tearing of crack length. For each of the eight iratersections tested, the leak rates were plotted as a function c f l pressure differential on a semi-logarithmic paper. There are usualy 4 data points per tube and !- p ; they reasonably align along a straighi line. According to these straight line fitting of leak rate, j- Laborelec developed the following correlation for scaling measured leak rate from ono' pressure , ' differential to another at the same (room) temperature.- a- lo(ap,- ap )725 - ,-_ (15)1 o where the pressure differentials are in psi. This corre_lation of the mechanical adjustment 1actorf .. i fa fits the Belgian pulled tube data of Table C-1. Behavior of the pressure adjustment factor may vary from one pulled tube to another, ' C.4c ' i' i
, . . -- - _. - ...,..-.- - .. . - . . - _ . . n.. +- ,
y-
.- C2.4- The Temperature Factor p.
- The flow area of the crack is approximately (confirmed by CRACKFLO analyses) inversely _..
i proportional to the product of flow stress a f and Youngs modulus Eof the tub 6 metal. Both flow - -
, stress and Youngs modulus are functions of temperature. Tnerefore,it follows that cro p_
f = Eo - {3gy E- of Po
- where p is the water density corresponding to upstream tenyperature Ty and pressure p; and po is water density at room temperature, say 70'F.
1-l C2.5 The Hydraulic Factor y The yfactor describes the hydraulic effect of pressure differential on leakage flow rate. When there is no' water flashing at To and T,4p- ap'o and 4pmo = dP*mo , so y= 1. When there is water flashing, effective pressure differentials should be used for op and 4Pmo. Therefore: y=
' (Pt - Cphsp; for flashing at To and T -
(Pmot - Cpmopmotj)Mpk
, = V(pi - Cppy)/dpf for flashing at T- only :
=1 for non flashing at To amt T ' (17) where Cpis a parameter to correlate the effective flashing pressare. For an isentropic process of
- water choking (or flashing), the flashing takes place at the saturation pressure corresponding toi
! the upstream water temperature Ty (e.g., Ty, Toy ). :So C p - 1 is for an Isentropic process, if l the real process deviates significanti/ from isentropic, the parameter C p will be less than unity. C3 The Mechanical Factor a - C3.1 Belgian Factor a and Leak Rate Adjustment br Plant E-4 Data-- ! ; Table C-1 presents measured leak rates for the Ned tubes from Plant E-4 steam generators.- l - The leak rate tests of the eight pulled _ tube inte' sections were conducted at room temperature -
- (70*F).; These data have been used to estat+sh Eq (13) for the mechanical adjustment factor a. L
,.,_ The CRACKFLO code was developed based on Westinghouse fatigue' crack and pulled _ tube datar_.The L CRACKFLO pressure adjustment factor _ is different from the Belgian. The difference is believed to ' be due to minimal ligarnent tearing used in the CRACKFLO model compared to significant tearing in
' the Belgian puIled tubes. Next subsection wGI discuss the mechanical adjustment factor for
!J . minimalligament tearing. i.
- j. C5 g
~ .
- y -r - - - -re- y s -e+ -* prey - --#--, 4-,v m- e +- .---:---- +
C3.2 General Mechanical Factor a and Leak Rate Adjustment for Minimal Ugament Tearing A mechcanical adjustment factor a for minimalligament tearing is proposed as follows. (dP,- dp.,W25 For the Plant E 4 data, the constant (725 psi) in Eq (15) principally varied from 725 to 940 with one indication at about 1450. The larger constants tended to occur when there was increased , evidence of damage from tube pulling effects which would likely lead to higher than expected leak rates at low pressure differentials such as normal operating conditions. Leak rate at higher pressure differentials such as SLB conditions would be less influenced by ligament tearing in tube pulung operations as some ligaments would likely tear away at the higher pressure differential. For this reason, Laborelee has recommended the constant of 725 psi for all Plant E-4 pulled tube intersections, i Eq (18) is similar to Eq (15), except the base is e (=2.718) rather than 10. Applicability of Eq (18) is verified with the CRACKFLO code. Table C.2 presents the CRACKFLO code results of leak rate at different crack lengths under a variety of pressure differeittials duing power operation. Table C 2 also includes comparison of the CRACKFLO adjustment factor and the a factor oktained from Eq (18). Similar to Table C 2, Table C 3 presents the comparison for a variety of pressura differentials under steam line break. The Belgian a adjustment factor for Plant E, based on Eq (15), decreases faster than the CRACKFLO calculation and th1 exponential (Eq (18)) adjustment factors for decreases in op from the reference point. Use of the exponentialadjustment factor is then conservattve in estimating the leak rate when the adjustment is toward a lower ap. Adjustment to Westinghouse , measurements in this report are to lower pressure differ 6nt'als such that the use of the exponential form (Eq (18)) leads to higher adjusted leak rates than if Eq (15) were applied. C4 Temperature and Flashing Adjustment Factors Temperature adjustment tactor is defined by Eq (16). Flashing adjustment factor yls defined by Eq (17). C4.1 Temperature Factor Only and Leak Rate Adjustment (i.e.,without Flas11ng) For water without flashing at To and T and ap - dPo" dPmo, it fonows that 1.fAp,T)n = } Q4pmo,To[ An equation for factor was provided by Eq (16). This will be compared to the CRACKFLO calculated adjustment factor under non. flashing conditions. Table C-4 presents the comparison - for the case witnout water flashing for a crack length of 0.2 inch. Similar results were obtained
. for crack lengths of 0.4 and 0.6 inch. These results indicate that the approximation of Eq (13) as Eq (16) yields good agreement with the more detailed crack Grec models in CRACKFLO for leak rate ratios between different temperatures.
C6
C4.2 Adjustment Factors p and y from Room to Operating Temperature Under SLB Conditions For water flashing at Tand non flash!ng at To and opp = dpmo. it follows that L(6p,T)7, = 13yL(Apmo,To) where is given in Eq (16) and y by the second expression of Eq (17). Table C 5 presents the comparison of leak rate adjustment between the CRACKFLO code and the proposed expressions with e pressure coefficient Cp = 1.0. For Cp= 1.0, the proposed expressions of and y factors yield an adjustment 18% to 25% lower than the CRACKFLO calculated factor. The deviation comes from the flashing point. A C p - 1.0 implies that flashing takes place at the saturation pressure corresponding to upstream tamperature Ty. For Ty = 616'F, the saturation p essure pyf- 1735 psia. However, because of - heat trarisfer and friction along the leakage path, the water flashing will occur at lower pressure, or a saturation pressure corresponding to a temperature lower than the upstream temperature Ty. In fact, the CRACKFLO code predicts a flow choking at a pressure less than py; To bring the proposed equation to yleid a result equal to the CRACKFLO calculation, we obtain a Cp ' for each case. Table C-6 presents the results Dr Cp. The longer the crack length the larger the pressure coefficient for a given pressure differential. A longer crack length means a larger crack - opening, and thus less friction and smaller heat transfer effect. It approaches the ideal case: an isentropic (i.e., frictionless and adiabatic) processt a situation with less leak rate. Table C-6 suggests use of Cp= 0.85 for py - 2350 psla (ap = 2335 psi), and Op 0.77 for py - 2665 psia (op- 2650 ps9 C4.3 Adjustment Factors a and y for Primary op Changes at Temperatures with Flashing Westinghouse leak rate tests were conducted at a primary pressure higher than the typical pressure during steam line break. The tests were conducted at the operating temperature of 616'F. We will adjust the leak rate to the typical primary pressure. The relevant adjustment factors are the mechanical adjustment factor a and the flow flashing adjustment factor y. Now it fo!!ows that L(Ap,T)7, = ayL{Apmo,To) where the a factor is defined oy Eq (18), and the yfactor by the first expression of Eq (17). Note that T o- T= 616'F. We would like to convert the leak rate to the equivalent volumetric rate at the room temperature (70*F). To do so, the following expression is used. .o L(4p,T)70 = &L(Ap,T}r, - P70 Eq (18) invotves a constant parameter A = 725 psi. Uss cf a = 725 waq ritted from the Belgian 0-7
data. For Westinghouse data, a specific parameter A will be fitted for each tube tested. As an illustration, Table C 7 presents some results of the adjustment factors ior scaling from one set of flash l rig conditions to another. Table C 7 also lists the comparison between the CRACKFLO code prediction and the proposed expressions with appropriate choice of the pressure coefficient Cp There is good agreement between adjustment factors derived from CRACKFLOW and Eqs (17) and (18), This provides confidence in the proposed expressions for the adjustment factors a and 7 . C5 Leak Rate Adjustment to Belgian Plant E 4 Data Table C 1 presents leak rates measured at Laborelec for both normal operating and steam line break conditions. Both conditions simulate the typicai pressure oifferentials across the tube, but tests were conducted at room temperature, it is desired to scale these room temperature data to hot temperature of 616'F. First, the normal operating pressure data la ocated. C5.1 Adjustment for Normal Operatlag Pressure Data All of normal operating conditioris involve no flashing of water. So the adjustment factors are as follows. L(Ap,T }r, = af L(Ap,,w,To) where the factor a ls defined by Eq (15) and the factor p by Eq (16). The dPmo is 1500 psi (see Table C 1), except tube R19C35, which is 1305 psi, andTo = 70*F. These will be scaled to apo - 1300 psi and T= 616*F. When there is no water flashing ap = apo so y= 1. The properties of flow stress, Youngs modulus and water density can be f,und in Table C4. Table C-8 presents results of leak rate at the target conditions. C5.2 Adjustment for Steam Line Break Data f l Similarly, djustments may be made to the Belgian leak rate data at SLB pressure differential. Table C-1 lists the measured leak rates and pressure differentials d#moat room temperature. These are to be scaled to T= 616*F and 4po - 2650 psi and 2335 psi, respectively. Now it j follows that L(6p,T}r, = af)L{Ap,,w,To) l where the a and factors are defined by Eqs (15) and (16), respectively, and the yfactor by the second expression of Eq (17) with Cp = 1.0. Tables C-9 and C-10 list the adjusted leak rate at a l i pressure differential of 2650 psi and 2335 psi, respectively, e. C8 l
C6 Adjustment to Leak Rate Data Base for Alternate op - -- Table C 11 lists the model boiler and pulled tube test results from leak rate tests conductedby -.- Westinghouse at elevated temperature. Leak rate adjustments for these data were made using both mechanical and hydraulic factors as described in Section C4.3. Table C 12 presents the adjusted
,- leak rates for the normal operating (NOP) and SLB mnditions.
O 9 C9
Table C 1 Results of Leak Rate Tests Under Room Temperature , and Typleal Pressure Differentials During Steam Line Break Normal Ooeratino Condhrons Steam Line Break Conditions . Plant E-4 Tube Pressure Measured Pressure Measured Tube Support Differential Leak Rate Differential Leak Rate . Row Col Plate No. Bat pgl lilent ga Bag psi liteth gm _g
*~
e k 9 i C 10
Table C 2 , Comparison of the Pressure Adjustment Factor a Between . CRACKFLO Prodletion and Proposed Correlation at Temperature
- Normal Operating Pressure Differentials -
Crad Upstream Downstream Upstream Leak Adjustment Factor a Length Pressure Pressure _ Temor sture Bata CRACKR.O EQ.ll.B1 Eq (15) (in) (psia) (psla) ('F) (gpm)' a 9 l l 3 ,,,,,, . l e l L- ! C-11 l-
l l Table C-3 l , Comparison of the Pressure Adjustment Factor cr Between CRACKFLO Prediction and Proposed Correlation at Temperature
- Steam Line Break Pressure Differentials -
Crack Upstream Downstream Upstream Leak Adjustment Factor a , L2D21h Pressure Pressure Tomoerature Balg CRACKFLO Eq11]Il Eq.(151 (in) (psia) (psia) ('F) (gpm) _a 9
-i C-12
Table C4 Comparison of the Temperature Adjustment Factor 8 Between CRACKFLO Prediction and Proposed Expression for the Case Without Water Flashing 4 l Pressure Flow. Youngs Leak Temperature . Crack (psia) Temp. Denshy Stress Mcdulus Rate - Adjustment Factor Length Rf R2 IfCB ^ Ibm'ft3 of Ossa E(Eta Inoml fcRacKRo # Expression
- l i
l
- - Calculated using Eq (16)
' i.
l u l 3 l
~.
C 13 1
-y y & V W
Table C-5 Comparison of the Temperature and Flashing Adjustment Factors ,' Between CRACKFl.O Prediction and Proposed Expression
'for the Case with Water Flashing i
Pressure Flow Youngs Laak Temperature Crack (psla) Temp. Density Stress Modulus . Rate - Adjustment Factor - Lansib or as IfL'El sJbrndt3 of (Esu E(EsD 12Dm1 #)CRACKFE) # Expression
- a
-(
'4 -
Calculated using Eq (1}}