ML20135D366
| ML20135D366 | |
| Person / Time | |
|---|---|
| Site: | Byron  |
| Issue date: | 12/02/1996 |
| From: | Hosmer J COMMONWEALTH EDISON CO. |
| To: | NRC OFFICE OF INFORMATION RESOURCES MANAGEMENT (IRM) |
| References | |
| NUDOCS 9612090352 | |
| Download: ML20135D366 (28) | |
Text
'
commonwealth rAlison Company
.D 1400 Opus Place Downers Grove. IL 60515-5701 December 2,1996 U.S. Nuclear Regulatory Commission Washington, D.C. 20555 Attn: Document Control Desk
Subject:
Response to Request for Additional Information Pertaining to the Byron Cycle Length Byron Station Unit I l
NRC Docket Number 50-454
References:
1.
M. D. Lynch letter to I. Johnson dated October 25,1996 transmitting Third Request for Additional Information Regarding the August 2,1996, Proposal to Delete the Braidwood Unit i Mid-cycle Steam Generator Tube Inspection 2.
M.D. Lynch letter to I. Johnson dated October 30,1996, transmitting Request for Additional Information Regarding an Extension of the Operating Cycle for Byron Unit 1 In Reference 1, the Nuclear Regulatory Commission (NRC) transmitted a Request for Additional Information (RAI) to the Commonwealth Edison Company (Comed) pertaining to the proposal to delete the Braidwood Unit 1 mid-cycle steam generator tube inspection. In previous correspondence, Comed proposed to use the same methodology and database as technicaljustification for the proposal to allow the operation of Byron Unit I for 600 days without having to shut down for steam generator tube inspection.
Because the proposed methodology and accompanying database continue to be relevant to determining the acceptable length of the operating cycle for Byron Unit 1, the RAls transmitted via Reference 1 remains applicable. Additionally, the Staff also issued an additional RAI which was transmitted via Reference 2.
Attached is Comed's response to the RAIs. The response incorporates data from the most current Braidwood steam generator tube insitu pressure tests and industry operational leak rate data. ' Utilization of this data has allowed Comed to establish a correlation between leak rate and 0.080" RPC maximum voltage. The development of this correlation is consistent with Generic Letter 95-05. The Byron cycle length determination calculation is in the process of being formalized. Results to date validate Comed's application of this correlation along with a probability of detection of 0.6 9612090352 961202 f
ADOCK 0500 4
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k:nla bybwd.1I2796. doc:1 A Unicom Company
NRC Document Control Desk December 2,1996 demonstrates that Byron Unit 1 meets the necessmy tube integrity requirement for leakage and burst for 600 days of operation. The finalized calculation will be completed by December 9,1996.
Upon receipt of NRC concurrence with the approach, Comed plans to submit a fechnical Specification Amendment to reduce reactor coolant iodine for the remainder of Cycle 8.
As stated during our November 14,1996 meeting, Comed welcomes the opportunity to discuss Byron's cycle length assessment with the Staffin mid-December. Ifyou have any immediate questions concerning this correspondence, please contact Denise Saccomando (630) 663-7283.
Sincerely l
/
.,6d Ml(%-
JU John B. Hosmer Engineering Vice President Attachment cc:
G. Dick, Byron Project Manager-NRR M.D. Lynch, Senior Project Manager - NRR S. Burgess, Senior Resident Inspector-Byron A.B. Beach, Regional Administrator-Rill Office of Nuclear Safety-IDNS k:nla-bybwd.112796. doc:2
1.
Eddy current data coil correction factors determined by the staff are not consistent with those discussed in the licensee's response to Question 20 in its submittal dated September 24,1996. Based on calculations using the data provided in Table 19b, the coil correction factor for converting data normalized to 10-volts on the through-wall hole is 0.98 and 1.03 for maximum and average voltages, respectively. The value discussed in the response to Question 20 was 0.75. The staff believes that the text in the response to this question is incorrect in stating that the factor is for adjusting voltages normalized to 10-volts on the through wall hole. The basis for this beliefis that the staff's calculations determined a coil correction factor of about 0.76 for converting data normalized to 20-volts on the EDM notch. Discuss this apparent discrepancy in the calculated coil correction factor.
Response
The Reference 1 response to RAl Question 20 incorrectly states "The results from statistical evaluation indicate that the coil size correction factor for the field data 0.75 (for normalization to 10 volts on a 100% throughwall hole) as determined from the ratio...".
The text in the parenthesis should read "(for normalization to 20 volts on a 100% throughwall axial EDM notch)". The discrepancy in the text does not affect the results, the coil size correction factor of 0.75 was applied to the industry data normalized to 20 volts on a 100% throughwall axial EDM notch.
2.
In its letter dated October 10,1996, the licensee states that analyst uncertainty is j
included in the end of cycle (EOC) distributions and does not need to be double counted in the burst and leak correlation's by means of an adjustment for scatter in the data from the coil and calibration procedures. Although there is expected to be some level of analysts uncertainty in the data to support the proposed correction factors, the staff cannot confirm whether the scatter is entirely due to analyst variability. Accordingly, assess the degree of scatur in the correction factors due to differences in eddy current i
data acquisition and setup, i
Response
In order to provide consistent industry data, analysis of tube pull and insitu test ECT data was normalized to 10 volts on a 100% throughwall hole since the throughwall hole was available on the calibration standards for all but one tube. The Byron 1 indication look-back studies were perfomied with a normalization of 20 volts on a 100% throughwall axial EDM notch. In order to compare the two sets of data, a factor was determined to adjust the industry data (normalized to 10 volts on a 100% throughwall hole) to the Byron 1 look-back study data (20 volts on a 100% throughwall axial EDM notch). Additionally, since the industry data was acquired with two different size pancake coils
(.080" and 0.115") and the Byron I look-back data was performed on the 0.080" pancake coil data, a factor to adjust the industry 0.115" data to 0.080" was determined.
I
l
{
The coil size and normalization factors were determined from a field study of 50 indications from Byron Unit I steam generator 1996 inspection results. In the Geld study, data was acquired with the two normalization methods (10 volts on a 100% throughwall hole and 20 volts on a 1007c throughwall axial EDM notch) and with the 0.080" and 0.115" pancake coils. A statistical analysis of the field data was performed using a linear regression analysis. The results indicated that there is a a
good linear fit of the data with a small error band around the linear relationship based upon the r-squared value and the standard error (presented in Reference 4). In Reference 4, it was stated that scatter in the field data is expected as a result of the analyst uncertainty (32% and 30% for average and maximum voltages, respectively) associated with analyzing the data.
l The analyst uncertainty was obtained from the results of a blind test performed in Reference 2 in response to Question 25. The blind test analyst variability includes uncertainty in the setup associated I
with voltage sizing. The root cause of blind test calls with large error were investigated. One of the errors identified in the root cause and included in the analyst uncertainty results is incorrect voltage I
setup. Therefore, any scatter in the correction factors due to differences in eddy current setup is i
included in the analyst variability applied to the EOC distribution calculations.
The field data used to determine the normalization and coil size factors is from Byron 1 1996 and is from steam generator C. The data acquisition of each tube used the same essential variables as j
identified in Table 21 of Reference 2. Therefore, the scatter in the normalization and coil size factors 3
due to differences in data acquisition is minimal and does not affect the resulting factors.
i j
The degree of scatter in the correction factors due to differences in eddy current data acquisition and setup is included directly in the analyst variability in the EOC distribution calculations and does not
{
need to be further considered.
i 4
l 3.
As discussed in Question 2 of the staff's letter dated October 3,1996, voltage correction i
errors could significantly alter the leak rate correlation. Accordingly, assess and discuss i
the potential increase in leakage which could occur under postulated main steam line break (MSLB) accident conditions by applying bounding correction factors to the voltages for the leak rate data.
Response
Even though the degree of scatter in the correction factors due to differences in eddy current data acquisition and setup is bounded by the analyst variability in the EOC distribution calculations, 957c/95% lower bound factors have been determined and applied in response to this question.
Lower bound factors were developed to correlate voltages obtained from 0.080 and 0.115 inch coils normalized at 10 volts on a 100% throughwall hole (TWH) with voltages for an 0.080 coil normalized at 20 volts on an axial EDM notch. To develop the bounding factors, linear regression analyses were 2
l perfonned to obtain the mean factors and the statistical parameters needed to compute lower bound factors for the data.
Two regression analyses were perforrned. The first analysis determined the mean regression line and statistical parameters for voltages from an 0.080 inch coil normalized at 20 volts on a 100% TW axial EDM notch as a function of voltages from a 0080 inch coil normalized at 10 volts on a 100% TWH.
The second analysis determined the mean regression lines and statistical parameters for voltages from a 0.080 inch coil nonnalized at 20 volts on an axial EDM notch as a function of a voltages from a 0.115 inch coil normalized at 10 volts on a 100% TWH. The analyses used maximum and average voltages obtained from 50 tubes for a total of 100 data points for each of the two regression analyses; these data are presented in Table 19b of Reference 6.
The lower bound adjustment factors for each coil size were defm' ed using the 95%/95% lower tolerance limit from the two data sets. Figures 3a and 3b present plots of the data used in the analyses, the mean regression line and the 95%/95% lower tolerance limits for the 0.080 and 0.115 inch, coils, respectively. The lower tolerance limit curves in Figures 3a and 3b have been used to adjust the r
industry burst and leak rate data base. Application of the 95%/95% lower tolerance curves for the voltage adjustments will provide substantial additional conservatism to the burst and leak rate evaluations.
The change in the leak rate results from using the mean normalization factors and the lower bound factors can be qualitatively assessed in Figure 4b by noting the shift in the leak rate correlation using i
the Q vs Mean on V. points versus the lower bound Q vs 95/95 V.
4.
In the staff's review of the data used to develop the proposed leakage correlation,it noted that operational leakage data was not incorporated from at least one past forced outage due excessive primary-to-secondary tube leakage. Specifically, operational leakage data from the shutdown of Arkansas Nuclear One, Unit 2,in March of 1992 was apparently not included in Table 5. In addition, the staff noted that other in-situ pressure test leakage data was not included. Provide a list of all tubes for which leakage (operational or in-situ) data are available and discuss the basis for excluding data in the proposed leak rate correlation. Identify the plant, steam generator, tube identification, the year the data was obtained, and the corresponding data on leakage for each of the tubes in the list. Provide the basis for assuming a maximum leak rate of 0.16 gallons per minute (gpm) when in service leak rate events, similar to that cited above, were determined to have leakage rates at normal operating differential pressures higher than 0.16 gpm.
Response
During evaluation of data to develop a steam generator tube leak rate correlation, only insitu pressure test and tube pullleak test data was considered due to the testing being performed under controlled I
conditions with accurate measurement instruments and a known source of leakage. Additionally, leak 3
rate data was only evaluated frem indications which are characterized with the same morphology as that demonstrated through tube pulls at Byron Unit 1. The morphology includes top-of-the-tube-sheet outside diameter circumferential indications in 3/4" tubing. Insitu testing leak rates from tubes with inside diameter degradation were not included in the correlation due to the differences in morphology.
All available tube pull, insitu pressure and operational leak rate data from indications with the morphology discussed above, are included in the leak rate data included in Table 4.
There have been two outside diameter 1TS circumferential indications which have measured inservice leak rates. The leak rates from the steam generator were estimated to be 0.25 and 0.15 gpm at normal operating pressure differential (norninally 1350 psi and 1375 psi respectively). The voltages of these indications were determined to be 6.65 volts and 14.27 volts, maximum 0.080" pancake voltage and 2.3 and 6.98 volts, average 0.080" pancake voltages, respectively for the 0.25 and 0.15 gpm inservice tube leaks. The data were analyzed to a normalization of 10 volts on a 100% throughwall hole and corrected to 20 volts on a 100(7c throughwall EDM axial notch using the lower bound corrections discussed in Question 3.
The data base used to evaluate leakage for Byron, Unit I has been augmented to include the tubes having measured inservice leak rates. These data are presented in the attached Table 4, where the operational leakage has been corrected to Byron 1 main steam line break conditions. Also included in Table 4 are the leak rate data from TTS circumferential indications, as a result of Braidwood 1 insitu testing corrected to main steam line break conditions. Table 4 includes voltage values adjusted for the bounding correction factors developed in response to Question 3 for those tubes where the voltage could not be normalized at 20 volts on a 100% throughwall axial EDM notch.
Figure 4a provides an updated probability of leak (POL) curve as a function of maximum 0.080" RPC voltage for the data in Table 4. The data set in Table 4 also was used to determine if there is a correlation between leak rate and maximum voltage using the additional leak rate data from the plants with operational leakage and Braidwood insitu test data. The results of the evaluation indicated a linear correlation between leak rate and maximum voltage where the p value for the correlation is 0.30% compared to Generic Letter, GL 95-05 criteria of 5%. The correlation obtained from the linear regression of the data in Table 4 and the data points are shown in Figure 4b, the data is plotted on semi-log scales in order to facilitate presentation of the data. The EOC leak rate calculations use the leak rate and maximum voltage correlation in place of previous analyses which used a bounding leak rate number.
The solid points in Figure 4b are voltages that have been adjusted for the bounding correction factors developed in response to Question 3. The open squares just to the right of the solid squares are the voltages adjusted using the mean correction factors. The open squares that have no associated solid square to the left of the data point are the data where the voltage was obtained directly from normalization at 20 volts with a 100% throughwall EDM notch. The leak rate correlation is represented by the line in Figure 4b and was developed using the voltages adjusted for the bounding correction factors and the voltages obtained directly from the 20 volt normalization on a 100%
throughwall EDM notch.
4
1 e
e The POL and correlation in Figures 4a and 4b meet the guidelines in GL 95-05. The POL curve from Figure 4a and the correlation in Figure 4b were used to compute the 95/95% leak rate. Monte Carlo computational techniques described in WCAP-14277 and accepted for application of Generic Letter 4
i 95-05 were used for the leak rate calculation.
Leak rates were computed for two projected distributions of indications at Byron, Unit I after 600 operating days greater than 50(TF. Both distributions were developed using POD = 0.6 to account for 4
indications that may have been undetected during inspection and for growth of new indications.
l Growth rates obtained from different operating mtervals were used to provide a sensitivity study for the effect of cycle length on growth rate and end of cycle projected indications. The first distribution i
was determined using a POD = 0.6 and growth rate obtained from the operating periods between 1994 l
and 1995, and 1994 and 1996. The computed 95/95% leak rate for this distribution was 28.4 gpm.
4 The second distribution was determined using a POD = 0.6 and growth rate obtained from the operating periods between 1994 and 1995,1995 and 1996, and 1994 and 1996. The computed 95/95% leak rate for this distribution was 35.7 gpm. The binned voltage distributions used in the leak rate computations are provided in response to Question 10 as Table 10.
j The results from the leak rate evaluation indicate the leak rate is not very sensitive to the cycle lengths used to compute growth rates at Byron, Unit 1. The magnitude of the leak rate for the distributions l
projected for 600 operating days will require a change in the allowable iodine concentration; this change will follow the guidelines in GL 95-05. The iodine activity reduction will be submitted under a separate Technical Specification submittal.
i 5.
The indication listed in Table 5 with a measured leak rate of 0.29 gpm at a test pressure of 4200 psi appears to be data obtained from a 1996 in-situ pressure test at the Calvert l
Cliffs Nuclear Power Plant, Unit 1. The inspection coil listed for the eddy current data for this indication is stated to be a 0.080 inch pancake coil. However, the information i
provided in Table 21 of the licensee's submittal dated September 24,1996, indicates the data from Calvert Cliffs, was obtained using a 0.115-inch coil. If the data from the tube in question was obtained from inspections at Calvert Cliffs, discuss how the eddy current voltages were obtained. Include in the discussion, a description of the probe and i
coils used and the voltage calibration procedure for each coil. Indicate the voltage measurements listed in Table 5 that were the actual measured values from the data.
l
Response
j Further evaluation of the data included with Reference 2 Table 5 confimis the infom1ation in this question. The data included in Table 5 with Max. Volts of 7.56 and 2.16 normalized to 10 Volts on a 100% TW hole was acquired with the 0.115" RPC and not 0.080" RPC. Previous results for the burst and leak correlation's are not affected by this change since normalization and coil size factors applied to industry burst and leak indication voltage data result in a total factor of 0.51 applied to both 0.080" j
and 0.115" RPC data The correct coil size is included in Table 4 of this submittal. The voltages in 5
Table 4 are corrected using the appropriate bounding normalization correction for 0.115" RPC. The data in the column titled 10 Volts on 1007c TW hole is actual measured voltages prior to application of any factors.
6.
The licensee's response to Question 14 in its submittal dated September 24,1996, states l
that leakage data were adjusted to MSLH differential pressures and temperatures using I
the PICEP computer program. Provide a description of how this computer program was benchmarked to assess its accuracy in applications involving steam generator tube circumferential flaws. Include test data to support your conclusions.
Response
Leak rate calculations using the PICEP code were performed as part of an industry-wide EPRI program on circumferential cracking. Regression analysis of PICEP led to leak rate equations as a function of crack length and crack opening area. Crack opening areas are computed using a formula from the Ductile Fracture Handbook. A single plastic zone correction is applied.
Crack opening area calculations for circumferential cracking are benchmarked against experimental measurements. It is essential to consider the restriction of lateral motion by tube support structures.
j The appropriate solution from the Ductile Fracture Handbook is an internally pressurized cylinder with circumferential crack. This solution essentially ignores bending caused by the end cap pressure axial force, hence lateral motion, in effect, restricted. Figures 6a,6b and 6c illustrate a comparison of calculated and measured crack opening areas as a function of pressure. The agreement is excellent.
The displacements and associated crack opening areas for the two smaller cracks are quite low. In this range, the calculated areas are actually more reliable and the measurements simply demonstrate that the internally pressurized cylinder is the correct solution to apply. Large displacements are observed for the 180 degree crack. The agreement of measured and calculated crack opening areas is excellent.
Leak rate calculations are benchmarked by comparison to other benchmarked calculations and by comparison of measured and calculated leak rates. The EPRI program on axial cracking in roll transitions developed a benchmarked leak rate calculation procedure. An orifice type equation was used. Flow discharge coefficients were determined using a combination of theory and experiment.
The present PICEP regression equations were compared to this earlier methodology. In general, the results agree well but with the PICEP based equations yielding somewhat higher leak rates.
The best illustration of benchmarking of PICEP based calculations is a comparison of measured and calculated leak rates using data for axial cracks. In Figure 6d, calculated and measured leak rates at normal operating conditions are plotted for comparison. The dotted lines are the leak rate calculations l
for axial cracks in 0.750 inch and 0.875 inch diameter tubing. The heavy solid line represents data for fatigue cracks in both sizes of tubing. Other data from fatigue cracked specimens are shown as solid l
diamonds. Finally, data on laboratory stress corrosion specimens is shown as open square symbols.
I 6
~. _ _ _
_ ~- _
As expected, smooth fatigue cracks exhibit higher leak rates than stress corrosion cracks of the same axial through wall length. Both the fatigue cracked and stress corrosion cracked specimens exhibit minimal as cracked, crack opening areas. Laboratory produced cracked specimens, either fatigue cracked or stress corrosion cracked, can easily be blunted to large, as cracked, crack opening areas which are unrealistic compared to cracks produced in service.
Calculated leak rates are shown to be comparable to measured leak rates for specimens with sharp fatigue cracks. The calculated leak rates aie somewhat lower than the measured leak rates for fatigue cracked specimens, yet calculated leak rates are a good upper bound to data from the laboratory stress corrosion cracked specimens.
Figure 6d is a satisfactory illustration of appropriate leak rate calculations. If only stress corrosion samples are used for benchmarking, three issues are likely to develop. The first is matching the as-cracked openings oflaboratory specimens to service produced cracks. The as-cracked opening is very sensitive to the method of specimen preparation. The second issue is a tendency to try to match calculated leak rates to the average measured leak rates using adjustable analysis parameters. This only serves to accentuate the scatter. The third issue is similar to the second, if the crack morphology is known exactly from destructive examination after leak testing, it is very difficult to predict actual measured leak rates. Crack tortuosity is difficult to treat, and, more importantly, mechanical rupture of small ligamer.ts during testing makes it difficult in some cases to identify the correct leaking through wall crack length. Outliers in calculated leak rates for stress corrosion cracks are often due to selecting an inconect leaking crack length. The addition of leak rate data from sharp fatigue cracked specimens provides a firm basis to evaluate leak rate calculations under well defm' ed conditions and construct reasonable upper bound leak rates.
7.
Given the limited available data for the burst pressure correlation and the fact that only two pulled tube specimens burst circumferentially during testing, the staff believes that a j
correlation between burst pressure and circumferential crack voltage may not be feasible at present. Specifically, the staff believes the extrapolating a curve to higher voltages to predict ihe voltage at the structural limis is questionable with the data currently available. Additional data is necessary to support such a correlation. (Refer to Question 6 in the staff's RAI dated September 9,1996.) Discuss the potential for using a lower bound voltage threshold below which the likelihood for circumferential flaw burst is considered to be low based on available burst and in-situ pressure test data.
Response
Deterministic relationships between pressure and average and maximum voltages have been developed to define conditions where tube burst would be unlikely. These relationships were developed using available industry insitu pressure test data. Tubes at normal operating pressure whose measured voltages were higher than the industry insitu and burst data base also were included in the data base.
7
In general, the deterministic bounding non-burst curve was constructed by bounding the available insitu pressure test data in three distinct regions. These regions included a region where burst was likely to occur axially, and two regions where burst was likely to occur circumferentially. The region where burst was likely to occur axially was modeled by a horizontal line from zero volts to a voltage less than the lowest voltage at which a tube burst circumferentially during a burst test. The pressure in this voltage region was a constant 5.3 ksi, which was the highest insitu test pressure from the industry data base. The second region extended from the point identified by 5.3 ksi and the voltage threshold for circumferential cracking to the insitu data point near the three times normal operating pressure differential. The third region extends from the point near the three times normal operating differential pressure to zero pressure through the lowest of the four points obtained from normal operation. The lowest of the four voltage normal operating pressure data points was used (not extrapolated out to higher voltage data points) to define the bounding curves to ensure conservatism. The results are presented in Figures 7a and 7b for maximum and average voltages, respectively.
Burst pressures identified in Table 4 of this submittal used for the bounding curves were corrected by tube geometry differences and material LTL properties (95(7c/95% at 650*F).
The results in Figure 7a for maximum voltage show the data bounded by the non-burst curve. The structural limit associated with the non-burst line is 3.7 volts,0.080" maximum voltage. Similar results are presented in Figure 7b for average voltage. The structurallimit associated with the non-burst line is 1.27 volts,0.080" average voltage.
The non-burst curves in Figures 7a and 7b were evaluated with two criteria described in Reference 7 to ensure adequate margin against tube burst for the two distributions described in response to Question 10. These criteria include: (1) the frequency of tubes greater than the structural limit is less 4
than 2x10, and (2) the conditional failure probability per tube is less than 10". These criteria ensure that there are relatively few tubes in the EOC distribution beyond the structural limit, and that the contribution to probability of burst would come from relatively few tubes or fractions of tubes at voltages beyond the structural limit, where the likelihood of having tubes with voltages this high in service would be relatively low. These criteria were satisfied for both distributions.
8.
The staff has ldentified some differences in the essential variables that were not identified or discussed in the licensee's response to Question 21 provided in its submittal dated September 24,1996. Inspections at three of the plants listed in Table 21 (ANO, Millstone, and Palo Verde) used 4(M) kilz as the prime frequency (PF). The remaining data were obtained in inspections using 300 kHz as the PF. If 300 kHz data were not available to size indications for the three plants using a PF of 400 kHz, discuss, how the data were corrected to compensate for the differences in flaw voltage resp (mse. Provide data that supports your conclusions. Also, the probe extension cable type was r.ot indicated in Table 21. Discuss the influence on flaw voltage response resulting from differences in extension cable length and type,if applicable.
8 l
Response
An evaluation by Comed was performed to determine the flaw voltage response for frequencies of 400 and 300 kHz to substantiate earlier evaluation results. The objective of the evaluation was to determine if the flaw voltage responses were significant enough to apply a correction factor to the 400 kilz voltage measurements. The essential variables used for the evaluation were configured to simulate or closely simulate the essential variables used at all plants. A Zetec Miz-30A was used to acquire the data with a.610" MRPC probe head containing a 0.080" mid-range pancake coil, + Point coil, and 0.115", pancake coil. The eddy current standard used was a Zetec (Z-14627) guide tube standard containing axial and circumferential EDM notches, as well as flat bottom holes.
For the analysis portion of this test, probe motion was set horizontal (flaws going up). The 100 7c axial EDM notch was set to 20.00 volts for both the 400 kHz and 300 kHz. Voltage measurements were made for the 60%,40%, and 20% circumferential OD EDM notches, as well as the 100% drilled hole. The largest response from all scan lines was used for each flaw measurement. Table - 8 shows the voltage responses for the specific flaws.
Table 8 Flaw Voltage Response Comparison 400 kHz vs. 300 kHz 400 kHz (volts)
FLAW =
100 %
62 %
44r7e 100(7c TWH Run #1 20 6.38 2.65 6.92 Run #2 20 5.65 2.85 7.17 Run #3 20 6.31 2.68 7
Avg. (volts) 20
- 6. I 1 2.73 7.03 300 kHz (volts)
FLAW =
100'7c 62(7c 44%
100r7c TWH Run #1 20 6.15 2.74 6.43 l
Run #2 20 5.59 2.91 6.36 Run #3 20 6.17 2.85 6.88 l
Avg. (volts) 20 5.97 2.83 6.56 Avg.%
Voltage
-2.4(7c 3.5%
-7.2%
Difference The 400 kHz voltage response is larger for the drill hole and 62(7c EDM notch, while the 300 kHz l
voltage responses are larger for the shallower EDM notch,44%. The shallower flaw will show larger voltage responses for the 300 kHz due to the eddy current depth of penetration. The evaluation l
9
IU results indicate that the voltages of indications where data was acquired at 400 kHz are not significantly different than data acquired at 300 kHz.
The most commonly used extension cables for MRPC inspections are the Zetec " Low Loss" and the Westinghouse RPC extensions. The essential variables from Table 21 show that in most cases the Zetec Low Loss cable was used at lengths of 50 or 60 foot. One exception showed that a Westinghouse extension cable was used. Impedance sweeps have been performed for the MRPC 0.080" mid-range coil using the Zetec low loss and Westinghouse RPC extensions. The extension i
type cable characteristics for the Zetec low loss and Westinghouse RPC consist of an impedance of 80 and 50 ohms respectively, with a 17.5 pf/ft (nom.) and 26.0i 2.5 pf/ft respectively. The impedance trace for the 0.080" mid-range pancake coil using a 50' Zetec low loss cable and a 60' Westinghouse cable show equivalent responses, both at the center frequency of the 0.080" coil between 280-290 kHz at 725 Ohms. Refer to Westinghouse calculation note DDM-96-009 entitled " Appendix H i
l Compliance and Equivalency" i
i j
Essential variable differences identified in this question do not significantly affect the results of the data presented in Table 4 of this submittal.
9.
One focus of Question 23 provided in the staff's RAI dated September 9,1996, was to assess the voltage threshold at which the 0.080 inch coil detected potentially more significant tube circumferential flaws with a high degree of assurance. This was the focus of requesting that the licensee provide a relationship between 0.080-inch coil voltages and those measured with a plus point coil. However, the response provided to Question 23 in the submittal dated September 24,1996, did not include an answer to the staff's question on the relationship between these voltages. Explain the relationship between the circumferentialindication eddy current voltages as measured with the 0.080-inch coil to those measured using the plus point coil.
Response
Figures 9a and 9b provide plots of 0.080" coil voltage versus + point voltage for maximum and average voltages, respectively. The information in the figures represent indications detected in Braidwood, Unit 1 Steam Generator C. The results in the figures include 604 total indications detected by + point, with 87% of the indications detected by both + point and the 0.080" coil. The 0.080" coil detected indications with greater than 0.3 average + point volts and greater than 0.7 maximum + point volts with a high degree of reliability. From the data presented in Figures 9a and 9b there does not appear to be a direct relationship between the two coils.
10
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- a R AI Ouestion Dated October 30.1996*
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10.
The submittal dated October 18,1996, proposes the use of a probability of detection (POD) for inspection of 0.8. As stated in Section 4.3 of the submittal, this value is a conservative estimate of POD based on the results of the Byron blind test results. The POD value oro.6 applied in voltage-based tube repair criteria applications following the l
guidance of NRC Generic Letter 95-05 was established considering not only the detection capabilities of the analysts and the inspection technology, but also the initiation and growth of new indications over the next operating cycle. Discuss how the Hyron, Unit 1 operational assessment accounts for the initiation and growth of new 4
i indications over the proposed extended operating cycle.
?
Response
in order to be consistent with the methodoldgy in NRC GL 95-05 to account for not only detection l
capabilities of the analysts and inspection technology, but also the initiation and growth of new indications over the next operating cycle the Byron EOC assessment has been performed applying a l
POD of 0.6. The resulting EOC distributions for maximum and average voltage are included in Table
- 10. Deterministic burst and probabilistic leak rate assessments have been completed using the new EOC distributions applying a POD of 0.6 as described in response to Questions 7 and 4, respectively, i
References:
i 1.
NRC Request for Additional Information Dated September 9,1996 2.
Comed September 24,1996 response to NRC RAI Dated September 9,1996 3.
NRC Request for Additional Information Dated October 3,1996 4.
Comed October 10,1996 response to NRC RAI Dated October 3,1996 5.
NRC Request for Additional Information Dated October 30,1996 6.
Comed Letter Byron 96-0266 Dated October 18,1996, Operating Interval Between Eddy Current Inspections for Circumferential Indications in Byron Unit 1 Steam Generators 7.
Comed Letter Dated August 2,1996, Operating Interval Between Eddy Current Inspections for Circumferential Indications in the Draidwood Unit 1 Steam Generators 11 i
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i Table 10 Byron Unit 1 End-of-Cycle 8 Distribution Maximum and Average Voltage Max Volts Averes e Volts Top of Bin Topof Bin (volts) Tubes (volts) Tubes 0.10 0.00 0.05 0.15 020 3.87 0.10 11.94 0.30 27.21 0.15 51.46 0.40 56.42 0.20 78.06 0.50 76.32 0.25 86.58 0.60 88.19 0.30 89.67 0.70 8920 0.35 8622 i 0.80 83.39 0.40 76.00 0.90 73.97 0.45 62.33 1.00 61.50 0.50 48.22 1.10 48.41 0.55 37.30 1.20 - 36.85 0.60 29.32 1.30 29.28 0.65 21.48 1.40 21.76 0.70 17.82 l 1.50 16.51 0.75 13.54 1.60 11.82 0.80 10.56 1.70 8.80 0.85 8.34 1.80 6.19 0.90 5.68 1.90 5.20 0.95 5.37 2.00 3.98 1.00 420 2.10 2.87 1.05 3.64 i 220 2.21 1.10 3.09 2.30 1.99 -1.15 2.30 2.40 1.91 1.20 2.02 2.50 1.96 1.25 1.64 i 2.60 2.01 1.30 1.76 2.70 1.89 1.35 1.51 2.80 1.38 1.40 1.57 ] 2.90 1.48 1.45 1.33 3.00 127 1.50 1.23 3.10 1.13 1.55 1.01 320 0.70 1.60 0.63 3.30 0.66 1.65 0.65 3.40 0.43 1.70 0.82 3.50 0.36 1.75 0.58 3.60 025 1.80 0.49 3.70 0.24 1.85 0.35 3.80 0.11 1.90 0.32 3.90 0.20 1.95 0.24 4.00 0.06 2.00 0.25 4.10 0.08 2.05 0.29 420 0.05 2.10 0.27 4.30 0.09 2.15 029 4.40 0.11 2.20 0.30 1 4.50 0.06 2.25 027 i 4.60 0.05 2.30 0.19 j 4.70 0.05 2.35 0.12 4.80 0.02 2.40 0.14 4.90 0.08 2.45 0.12 5.00 0.06 2.50 0.11
Figure Sa: Correlation of 80 Mil / 20V to 115 Mil /10V Data Regression Curve and One-Sided Tolerance Bound 1.500 80 Mil / 20V Norm. (Avg & Max) a Regression Curve g
95'7c / 95% Tolerance Bound I
/ / = 1.000 f V g l~ l f f ") MI,.a*' .D] A l aI / E E I. ,.-i' a y ** yy, ~ ^ S f; l a
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^ E a M,.a ^ t. S a am !.- ~- 7,l l l 0.000 O.000 0.500 1.000 1.500 2.000 2.500 115 Mil /10V Normalization Amplitude (Volts) [-ME1B05 XLS] B0 20vs115.10 1 RFK: 11/8S6,5:10 PM
Figure 3b: Correlation of 80 Mil / 20V to 80 Mil /10V Data Regression Curve and 95% / 95% One-Sided Tolerance Bound 1.50 y j l a 80 Mil / 20V Norm. (Avg & Max) I /- Regression Curve /i 3
95% / 95% Tolerance Bound d-../
s / l / P' / l / ~ 1'00 [ ,/ ' a g _,f_.< 7 u
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la },.- l !-}i 0.00 i O.00 0.50 1.00 1.50 2.00 2.50 3.0G 80 Mil /10V Norn;alization Amplitude (Volts) [-ME1B05 XLS] 80 20vs80.10 1 RFK; 11/896. 5.11 PM
~ Figure 4a: Probability of Leak vs. Maximum Voltage 1'0 _- '[ ' ~ V / \\ 0'9 Adjusted Data j / e New LogLogistic / / 0.8 ,/ /
New 95% Conf.
I !/ 0.7 ,I M .I I d / j j 0.6 / l .$0.5 / I = o f c3 / 10.4 d: I ! / l i/ 0.3 j 1 I i / / I 0.2 / 'f:,/ I 0.1 ) ..........-s' - J 0.0 0.1 1 10 0.080" RPC Maximum Voltage, volts [RK11-26.XLS] PoLVmax RFK: 11/27/96,1:51 PM
Figure 4b: Leak Rate vs. Maximum Voltage 1000.0 I e n m ,0 D p 100.0 s l i i l I }' i i i > i u)2 Leak Rate Data [ / Converted from gpm to g j e j ma Liters / Hour using a y 10.0 correction of 227.092 r ~ j j[ l LPH/gpm z g f 1 i i i t-I e Q vs 95/95 Vmax l m0 l 4 f Q on 95/95 Vmax E D 9 ** M*"" V""* 1.0 l l l l 0.1 0.00 2.00 4.00 6.00 8.00 10.00 12.00 14.00 0.080" RPC Maximum Voltage, volts [RK11-26.XLSI QvsVmax RFK: 11/27/96.1 A7 PM .. =
t Figure 6a: Crack Opening Area Versus Pressure,60* Circumferential Crack. l Lateral Motion Restricted by Tube Support Plate P (LOOD45 - f +EXPERA4ENTN. ......caanne 0 0004-0.00035-E f 4 / /% / 1 5% A 8 t A 6 i ) i O D 1000 2000 3000 4000 5000 PRESSINE, PSI f I I i
t Figure 6b: Crack Opening Area Versus Pressure,120* Orcumferential Crack, Lateral Motion Restricted by Tube Support Plate 0.005 l t 0.00 6 - t 0 3PERIMENTAL i ...... CALCULATED t 0.004 - 0.0036-I e k g naas. I f.n l l
- o.ow.
,e***.... l ,,,a Reefs. e** ,e e#*# h Reef. e m te o S me e g e ep s e a ep Ir ep de@e 4 he9 40005 W 4 W S 9eS 99 4 4 Se# O P l 0 10D0 2000 3000 6 5000 m asune.es F I t I
4 Figure 6c: Crack Opening Area Versus Pressure,180* Circumferential Crack, Lateral Motion Resnicted by Tube Support Plate i D.018 - 1 0 EXPERBENTAL' ...... CALCULATED 0.016 - 0.014 - N k D.012 - 9 l' I i (LODS - e* O* ~~y (LD04-I g. s 8,e 9 .==,,....
- ,3 g.
f ,,.O m 0 0 1000 2000 3000 40D0 5000 MINE.PBt i i I l - :- 2
l . e .' i O l 100 l 10 ,L i e.- .I J.- (.*' 2 l 3 r ( .r-s o.. f) o' p / ' '_ fun" g ],3 _f O.1 ~ x ' 'J..- u o . 5 D,. 4.' + n 1' 8 0.01 r .z* u 5* E _ f' L 5,VFAUGUE CRACKED DATA _' ** g ..... 0.876" OD BY 0.02 Wall 0.001 .~ - - 0.7W OD BY 0.043" WALL a CEGB 800 + CEGB FATIGUE CRACKED DATA I I I 0.0001 0: O.1 1 CRACK LENGTH,94CHES i l Figure 6d: Calculated and Measured Leak Rates for Axial Cracks in Alloy 600 Tubing at Normal Operating Conditions l i
Figure 7a. Maximum Voltage vs Burst or insitu Pressure 10 i i i m, n e ins tu (No Bmt) 9 s a a Burst Test-Pulled Tube g e a Normal Operation (No Burst) g a. 7 ei Deterministic Bound $a 6 E OOOO w E oh Structural Limit = 3.7 Volts ( g 4 o-a %* 20 o E 3 o 2 1 A A_A N 0 0.0 2.0 4.0 6.0 8.0 10.0 12.0 14.0 16.0 Maximum Voltage, Volts INDATA9 XLS 11f2fV96 3 54 PM
Figure 7b. Average Voltage vs Burst and insitu Pressure 10 i I I I .g e e o insitu (No Burst) e ~ m Burst Test-Pulled Tube 8 m a Normal Operation (No Burst) s ._E7 I Deterministic Bound 3e 6 e a 2 I5 o 'i5 Structural Limit = 1.2iVolts ^ c E 4 o O ~ O l-3 3 E O 2 i ^ ^ ^ 1 V g 0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 8.0 9.0 10.0 Average Voltage, Volts INDATA9 wts 11/2tW96 4 04 PM m
9 i Figure 9a: Maximum 0.080" Coil Voltage Versus Maximum + Point Voltage Braidwood, Unit 1, Steam Generator C 1.6 o 1.4 .Eg 1.2 7 g o CD O 10 ^ 5 o 0 0 0.8 a ol S o o co go s 82 8 Ao $ ,o l 0.6 E g,- ,,g o y I 0 o To
- 8 o
o Indications Detected by +PT and 0.080 Coil 3 o o y g o Regression Line e 0.4 3-oo 2 il dcp o 0 o i o a Indications Detected by +PT Alone geo 0.2 0 6"' oo t 0.0 '= --: u - O.0 0.5 1.0 1.5 2.0 2.5 Maximum + Point Voltage, Volts PPT-80ml
m 1 l i .o r i G G CU O 2ll: ? O to d E O b Q. ~ + G OO o o Eb 1 O g "O v d <E g o e o O O
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0 00 O O g 0 00 m g =g, y = 3 0 00 00 000000 ,oO O ooo oo O < 3 e o CD O m O e O 3 O 9 O CO O 00 00 00 d *"3 oOO-G - X-x-:-:-:- -o 3 9 D 3 O g g g { O O O 0000 b O 00 O 00 O 3 ) g O g g n. a. O OO Oc K)O 3 i C5 3 b + + 0 000 0 N N O OOO < 3 j O 00 O 3 ( g g 0 00000 00 3 h g } =* e 00 w:.::::. :. ""fb 3 ) g .] E O O OO ~ ^ d ~'""I"" - _ _ i_ _ _ _ _------ m C m 3 L c .9 c 0 3 .9 .9 3 ,G 'y 8 ,y 00 O O 3 LL, O 6 CC O O 3 O <3 0 i o 6 9 9 v. 9 N n 9 o o o o o o o y siloA 'e6elloA llo0. 080'0 e6eJeAV f i}}