ML20040B249
| ML20040B249 | |
| Person / Time | |
|---|---|
| Site: | Fort Calhoun  |
| Issue date: | 01/18/1982 |
| From: | William Jones OMAHA PUBLIC POWER DISTRICT |
| To: | Novak T Office of Nuclear Reactor Regulation |
| References | |
| LIC-82-029, LIC-82-29, NUDOCS 8201250310 | |
| Download: ML20040B249 (48) | |
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OMAHA, P4 E D R A S M A f, H 10 2 e TL LE P H ordt 536-4000 AHEA COOf 402 January 18, 1982 LIC-82-029 Mr. Thomas M. Novak Assistant Director for Operating Reactors Division of Licensing U. S. fluclear Reculatory Commission Washington, D.C.
205ES
References:
1.
Letter from D. G. Eisenhut to W. C. Jones dated August 21, 1981 2.
Letter from W. C. Jones to R. A. Clark dated October 20, 1981 3.
Letter f rom W. C. Jones to T. M. Novak da ted November 13, 1981 4.
Letter from T. M. Novak to W. C. Jones dated December 18, 1981 5.
Letter from W. C. Jones to R. A. Clark dated December 31, 1981 6.
CEN-189, " Evaluation of Pressurized Thermal Shock Effects Due to Small Break LOCA's With Loss of Feedwater for Combustion Engineerirg I4SSS," Dated December, 1981
Dear Mr. Novak:
In response to References 1 and 4, the District, as a member of the Combustion Engineering Owners Group (CE0G), is conducting a detailed andlysis of the staff's concerns regarding pressurized thermal shocking (PTS) of the Fort Calhoun Station reactor vessel.
The results of this andlysis to date indicate that the Fort Calhoun Station pressure vessel integrity will be maintained throughout the full life of the plant for postulated PTS events.
Although additional analysis is required, this additional analysis is expected to confirm the present findings.
Accordingly, the District believes that, upon completion of the CEOG program and implementation of measures identified herein, the PTS issue should be resolved.
Your letter (Reference 1) requested certain information regarding pressurized thermal shock be provided for the Fort Calhoun Station within 60 days and additional information be provided within 150 days of your letter.
The District provided our response to your 60 day request in References 2 and 3.
In addition, Reference 4 requested certain additional information be provided for the Fort Calhoun Station in the 150 day response.
This letter provides the District's response to your 150 day request.
8201250310 820118 PDR ADOCK 05000285 p
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Mr. Thomas M. Novak LIC-82-029 Page Two The body of this letter contains the basis for continued operation of tne Fort Calhoun Station, discussion of the four possible actions you requested be considered, and the plan for final resolution of this issue.
The analyses supporting continued operation of the Fort Calhoun Station and the data requested in References 1 and 4 are provided in the enclosures to this letter.
Basis for Continued Operation The results reported in Reference 6 for the Fort Calhoun Station which evaluated the most limiting transient for a range of small break 1.0CA sizes and locations with several different cases of assumed loss of all feedwater indicate that the Fort Calhoun Station reactor vessel would not experience crack initiation for the assumed full plant life. The results reported in Enclosure (B) of this letter indicate that crack initiation will not occur for a stuck open atmospheric dump valve, the limiting anticipated operating occurrences (A00) for the Fort Calhoun Station for the assumed full plant life, as identified in Enclosure (A).
Also reported in Enclosure (B) are the results of evaluations of a transient representative of a postulated full main steam line break at hot standby conditions.
This represents the most severe overcooling initiating event for a CE NSSS.
The results of tnese evaluations indi-cate no crack initiation for the assumed full plant life. The favorable results of these evaluations demonstrate that continued operation of the Fort Calhoun Station is justified.
The acceptance criteria applied to these evaluations are expressed in number of effective full power years (EFPY) for which acceptable con-sequences for a given transient can be justified.
Acceptable conse-quences are defined in terms of the fracture mechanics aspects of the given transient for the vessel material properties.
For higher proba-bility events in the A00 category, no-crack-initiation is considered to be the proper acceptance criterion.
For lower probability events such as an MSLB or LOCA due to a postulated pipe break, or an A00 plus an additional single active failure, crack-arrest is considered to be the appropriate acceptance criterion.
In either case, no violation of the primary system pressure boundary in the vessel would occur.
The dif-ference in criterion is primarily a matter of commercial risk consider-ing the relative probability of the different types of events and con-current assumptions.
In some cases, it is found that an event which would be acceptable under the crack-arrest criterion also satisfies the no-crack-ini tiation limi ts.
Where this is found to be the case it is indicated, as in CEN-189 and the enclosure to this letter.
The District believes that the criteria outlined in the previous para-graph are suitable as quantitative criteria which would assure mainte-nance of an acceptable low risk of vessel failure from pressurized thermal shock.
These vessel criteria are very similar to the current reactor fuel failure criteria applied in all safety analyses. A simpli-fied statement of the current fuel failure criteria is that no fuel failure is to be predicted for A00's (higher probability events) while limited fuel failure is allowed for lower probability events.
The
3 Mr. Thomas M. Novak LIC-82-029 Page Three validity of these safety analyses is assured through Technical Speci-fications, Operating Procedures, and Emergency Procedures.
It is sug-gested that the use of vessel integrity criteria could be implemented in a similar fashion through the utilization of Technical Specifications, Operating Procedures, and Emergency Procedures.
Evaluation of Possible Actions The District intends to implement a reduced radial leakage fuel manage-ment scheme during the next fuel cycle (Cycle 8) which is presently scheduled for startup during the second quarter of 1983. This scheme is being implemented at this time to reduce the fluence to the reactor vessel longitudinal welds and is consistent with the District's long range fuel management plan to reduce fuel cycle costs.
The Cycle 8 core will have thrice exposed fuel assemblies, which were scheduled to be loaded in the interior of the core, loaded in the center of the " flats" and the peripheral corners of the core.
These locations significantly contribute to the fluence at the longitudinal welds.
This fuel manage-ment scheme reduces the power produced by assemblies in these locations by a factor of two, resulting in a corresponding reduction of fluence at critical weld locations.
1.
In addition to the program discussed above, the District is parti-cipating in the CEOG program for resolution of the PTS issue, as discussed below. This program includes consideration of possible actions to reduce the rate of further irradiation of the vessel.
Scoping studies are being performed to evaluate the trade-of fs between local fast flux reduction at the vessel wall, changes in fuel management strategies which reduce the power density at selected peripheral fuel assembly locations, and plant power capability.
The following different fuel management strategies are being evaluated to cover the range of core configurations and weld patterns representative of the Fort Calhoun Station.
d.
Dummy fuel assemblies in the outside corner locations, and twice burned fuel assemblies in some peripheral locations.
(Here " dummy" fuel assembly means an assembly which is geo-metrically equivalent to a standard fuel assembly, but is made up of stainless steel rods instead of fuel pins.)
b.
Conventional in-in-out low leakage fuel management normally considered for reduced fuel cycle costs.
c.
Thrice burned fuel assemblies at some outer peripheral lo-cations and conventional out-in f uel management for remainder of core.
d.
Thrice burned fuel assemblies at some outer peripheral lo-cations and twice burned fuel assemblies on outside corners.
e.
Dummy assemblies at some peripheral locations and twice burned fuel assemblies on outside corners.
Mr. Thomas M. Novak LIC-82-029 Page Four Vessel fluence determined for each variation will be compared to the fluence levels established on the basis of surveillance capsule results as reported in CEN-189.
Power capability will be evaluated based on thermal margin using resultant power distributions, and considering the effect on the limiting peak linear heat rate.
These scoping evaluations are intended to indicate the potential benefits possible, and would serve as the basis for further plant-specific evaluations, if necessary.
2.
The District has considered reduction of the thermal shock severity by increasing the ECC water temperature.
The analyses reported in CEN-189 indicated significant mixing and heatup of the cold in-jection water with the reactor coolant system water, using a con-servative empirical calculation based on the preliminary results of EPRI sponsored tests.
Since significant mixing is expected, the sensitivity of the temperature of the ECC water is less than may have been originally thought. We therefore do not plan at this time to take action to raise the temperature of the ECC water supply.
In the analyses performed to date, the ECC water supply has been assumed to be at the Technical Specification limit of 400F.
Plant operating experience indicates that this is conser-l vative based on minimum recorded annual water temperatures of 600F for the Fort Calhoun Station. The potential to account for the added margin remains, as well as the possibility to raise the water temperature.
3.
The program for resolution of the PTS issue includes consideration of recovery of vessel material toughness properties by means of an in-place annealing process.
There has been a considerable amount of laboratory work performed by several organizations which evaluate the parameters necessary to achieve recovery of the material pro-perties. We are continuing to review these results in advance of establishing whether or not such a process is necessary for the Fort Calhoun Station.
We have also begun a review of the procedures and equipment necessary to perform an annealing process in the plant, with a view toward identifying potential problem areas which might need additional development work.
Factors being evaluated on a scoping basis include the effects of possible residual stresses and/or deform-ations in the vessel due to oxial thermal gradients during anneal-ing, and possible dimensional changes at critical mating surfaces such as the closure flange, nozzles, and lower internal support features. Analyses being considered for evaluation of these effects include nonlinear finite element analyses due to heating the material to temperatures near the creep range. Analysis of the l
heat treatment equipment will also require fairly sophisticated techniques.
4.
Control system design changes to mitigate the initial thermal shock aspects of possible overcooling transients, and/or to control repressurization following a thermal shock will be considered
I Mr. Thomas M. Novak LIC-82-029 Page Five l
during the third part of the overall program.
It should be noted that the District has installed as part of the TMI Action Plan requirements safety grade system to automatically initiate flow.
This provides a significant advantage in reducing temperature transient for the PTS events.
It is expected that the third part of the program will evaluate single active failures and loss of l
offsite power in conjunction with postulated initial failures which would initiate an overcooling transient. This evaluation may identify plant procedure or design changes which will lessen the severity of pressurized thennal shock transients, resulting in additional years of plant operation for which acceptable vessel response could be justified.
Other Issues Your December 18, 1981 letter also requested an assessment of the sensitivity of our analyses to uncertainties in the input values. The evaluations reported in CEN-189 and in the enclosures to this letter assume a range of assumed initial crack sizes from very shallow to very deep.
The approach used to develop the copper content, fluence and initial RTNDT was conservative in order to eliminate the possible con-cerns over uncertainties.
The specific operator actions assumed in the analyses for the Fort Calhoun Station are listed in Enclosure (A) to this letter. The primary action is the inherent assumption that the operator takes control of the plant at 30 minutes for MSLB and 90 minutes for A00 and brings the plant to a safe condition without increasing the severity of the situation.
We believe these are reasonable times at which to assume correct oper-ator action is taken. Our experience indicates that manual control of the plant following a transient is achieved well before 30 minutes.
The results of a review of the Fort Calhoun Station emergency procedures were reported in Reference 2.
The results of this review were that the procedures were adequate to minimize the potential for repressurization following a thermal shock event.
The Fort Calhoun Station operations staff have received refresher training in these procedures within the last year.
In addition, operations personnel have received training on the pressurized thermal shock issue as a part of the Cycle 7 startup training.
Briefings have also been held for the District's Senior Management and the Production Operations Division Management. These briefings have ensured that all levels of management within the District are aware of the pressurized thermal shock issue, its safety significance, and the need for an ultimate resolution of this issue.
The evaluations performed for the Fort Calhoup Station take credit for the warm prestress phenomenom where applicable.
Warm prestress occurs for a particular transient when the crack tip stress intensity Kg is decreasing from a previous condition of higher stress intensity at a
i Mr. Thomas M. Novak LIC-82-029 l
Page Six more ductile (warmer) condition of the material.
If Kj exceeds the critical crack initiation stress intensity K g while Kg is decreasing, i
then the warm prestress aspect of this unloading sequence will prevent crack initiation.
For a situation where crack initiation is prevented due to warm prestress for K1 greater than KIC, then the plant parameters would be above the minimum pressure-temperature (MPT) limits which would be recognized by the operator.
In such a casc, the operators would take actions to bring the plant conditions within the MPT limits, which would preserve the warm prestress assumptions.
The Fort Calhoun Station Emergency Procedures have been reviewed to j
l assure that proper instructions are provided to minimize vessel stress during a pressurized thermal shock event.
Two potential areas of improvement have been identified:
1.
Provide specific criteria for termination of HPI and charging pump l
flow. Possible criteria are restoration of primary system inventory control and an indication of adequate core cooling.
2.
Improved precautions to assure operator compliance with the cool-down curves (i.e., MPT limit).
j 1
The District is continuing to evaluate these possible improvements and j
their incorporation into the Fort Calhoun Station Emergency Procedures, i
Additional training will also be provided by June 1,1982 to the oper-ations staff based on the information given in this submittal.
If the evaluation indicates the need for procedural changes, these procedures will be revised by June 1, 1982.
Program Plan The District is participating in a program being conducted by Combustion Engineering, Inc. for the CE Owners Group (CE0G).
A general discussion of the program being developed was provided by a letter from the CEOG to D. G. Eisenhut dated May 15, 1981.
The program was described to the NRC staff at meetings on July 30, 1981 and October 7, 1981.
The program includes both analytical evaluations of PTS effects and scoping studies of potential actions to lessen the consequences of PTS events.
The.
analytical evaluations as applicable to the Fort Calhoun Station consist of three parts:
(1) the evaluations reported in CEN-189 in response to Item II.K.2.13, Reference 5, (2) the scoping analyses reported in the enclosures to this letter, and (3) a program reported here.
The program also includes evaluations of possible plant changes to lessen the severity of the most challenging events.
The scenarios reported in CEN-189 considered situations of recovery from small break loss of coolant accidents with extended loss of all feed-water.
Scoping evaluations for two types of scenarios are reported in the enclosures to this letter:
(1) main steam line breaks (MSLB) and (2) anticipated operating occurrences (A00). The situations to be evaluated in the third portion of the overall program include more precise analyses of MSLB and A00 events, including assumptions of single active failures and loss of offsite power, where appropriate.
These events will be analyzed to verify that the criteria discussed
1 Mr. Thor as M. flovak LIC-82-029 Page Seven above are met for the lifetime of the plant.
The analyses may identify improvements in Emergency Operating Procedures.
These improvements will j
be factored into the analysis to verify the criteria are met for the full plant lifetime.
The program may also consider changes to the plant j
if the refined analyses combined with cperator actions do not show the l
acceptance criteria being met for the full plant lifetime.
It is cur-rently estimated that the proposed program will be completed by September 15, i
1982.
j The District's complete program will evaluate the pressurized thermal shock aspects of (1) A00, (2) lower probability events such as an MSLB or LOCA or an A00 with an independent single active failure, and (3) extremely low probability events such as the II.K.2.13 scenarios evalu-a ted in CEN-189.
The program will not evaluate a large spectrum of multiple failure events, but will include events with a single active failure and concurrent loss of offsite power.
The program does include qualitative but not numerical determinations of pressurized thermal shock event probabilities.
Because of the large number of years before the consequences of PTS events approach the acceptance criteria for the Fort Calhoun Station as indicated by CEN-189 and the results of the scoping studies reported in the enclosures to this letter, we believe this systematic approach to the pressurized thermal shock questions is l
prudent. There is not a near-tenn problem requiring insnediate changes prior to the period of time needed to complete a thorough evaluation of further aspects of the issue.
The District Lelieves that following the staff's review of this letter and its enclosures, a meeting to discuss these reports and the Dis-trict's continuing program would be beneficial towards final resolution of the PTS issue.
The District recemmends that this meeting be con-ducted during the week of February 1,1982.
If you concur, we will interface with your staff by telephone to arrange this meeting.
Sincer ly,
[
W. C. Jones DivisiynManager Production Operations 1
i l
WCJ/KJM/JKG:jmm Enclosures cc:
LeBoeuf. Lamb, Leiby & l'acRae 1333 flew Hampshire Avenue, fl.W.
Washington, D.C.
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n Enclosure A Thermal-Hydraulic Evaluation 1.
Introduction l
This enclosure presents the results of a thennal-hydraulic evaluation of the main steam line break (SLB) accident and of overcooling anticipated operational occurrences (A00s) performed for input to the pressurized thermal shock (PTS) stress analyses for the Ft. Calhoun plant.
In order to illustrate how the PTS concern arises, a qualitative discussion follows of a representative SLB transient. Assume that a large break occurs in the main steam piping upstream of the main steam isolation valve (MSIV) associated with one steam generator, to be referred to as the "affected" steam generator. The break increases steam flow from the steam generators, steam generator pressures and temperatures decrease, and heat removal from the reactor coolant system (RCS) increases. Low steam generator pressure causes both a reactor trip signal and a main steam isolation signal (MSIS). Reactor trip terminates, or prevents, fission power generation; MSIS terminates blowdown of the unaffected steam generator by closing the MSIVs and terminates feedwater flow to both steam generators.
A low steam generator water level signal in the affected steam generator will actuate auxiliary feedwater (AFW).
Since the Ft. Calhoun AFW system includes automatic AFW isolation based on pressure difference between the two steam generators, AFW flow will not be initiated to the affected (depressurized) steam generator. Following AFW isolation, the affected steam generator will dry out and RCS cooldown will terminate.
During the RCS cooldown transient, pressurizer pressure decreases to the safety injection actuation signal (SIAS) setpoint.
SI AS starts two high pressure safety injection (HPSI) pumps and three charging pumps.
In addition, following SIAS on low pressurizer pressure the operator will trip all four reactor coolant pumps (RCPs).
The HPSI and charging pumps will repressurize the RCS to the HPSI pump shutoff head, and the charging pumps will further repressurize the RCS at a lower rate.
Conditions identified in the emergency procedures for termination of emergency core cooling flow will be reached and charging and HPSI pump flow will be reduced in order to terminate RCS repressurization.
The PTS concern arises due to the rapid decrease of reactor coolant temperature in the reactor vessel downcomer. The temperature decrease will be largest in the half of the downcomer associated with the affected steam generator. PTS effects are increased by the repressurization of the RCS by the charging and HPSI pumps.
2.
SLB Evaluation Results of a thermal-hydraulic evaluation of the SLB accident are presented in this section.
In order to bound the PTS effects of SLB, the evaluation is performed for an SLB occurring during hot zero power (HZP) operation. This mode of operation maximizes RCS cooldown because steam generator water inventory is largest at HZP and because core decay heat is lowest at HZP.
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To fur *.
- bcund PTS effects, a guillotine rupture of the maia s*.e m line is postulaud, with the assumption of no moisture carryover during the blowdown transient. The assumption of no moisture carryover maximizes total energy removal f rom the af fected steam generator and, theref ore, maximizes integral RCS heat removal. The assumption of no moisture carryover during a guillotine rupture maximizes, in addition, the rate of RCS cooldown.
A complete list of assumptions and plant parameters used for the SLB thermal-hydraulic evaluation is provided in Table 1.
These assumptions and parameters can be expected to give results which provide an upper bound on the rate and magnitude of RCS cooldown which can occur during SLB, primarily due to the following combination of assumptions:
- 1) HZP operating mode, 2) guillotine break with no moisture carryover, and 3) zero decay heat.
Results of the SLB thermal-hydraulic evaluation are provided in Figures 1 and 2.
Figure 1 provides the water temperature versus time in the half of the reactor vessel downcomer associated with the af fected steam generator. Water temperatures in the affected steam generator loop are substantially lower than in the unaffected steam generator loop.
The downcomer water temperature was obtained assuming complete mixing of the cold leg flow with HPSI and charging pump flow.
Figure 2 provides the downcomer pressure versus time.
A rapid repressurization to the HPSI pump shutoff head is seen, with subsequent repressurization at a lcwer rate by the charging pumps.
At 30 minutes the operator reduces charging and HPSI pump flow to terminate RCS repressurization.
3.
Overcooling A00 Evaluation Results of a thermal-hydraulic evaluation of overcooling A00s are presented in tM2 secticn.
The overcooling A00s consist of events which cause increased heat removal via one or more steam generators. Potential causes of increased heat removal include:
- heater, 2)
Increase in main feedwater flow due to a main feedwater control valve malfunction or due to a main feedwater control system malfunction, 3)
Increase in main steam flow due to a turbine control valve malfunction or due to a turbine control system malfunction, 4)
Increase in main steam flow due to a turbine bypass control valve malfunction or due to a turbine bypass control system malfunction, and 5)
Increase in main steam flow due to an atmospheric dump valve malfunction.
If uncontrolled RCS cooldown results from any of the above malfunctions, the cooicoon wili be accompanied by a decrease in steam generator pressure, hwn RCS cold leg temperatures decrease to about 465 F, MSIS will be caused by low
l 1
steam generator pressure (500 psia).
Following MSIS, MSIV closure and main feedwater isolar. inn will occur, thus terminating RCS cooldown for the first f our v i functions '. i sted above.
Since the atmospheric dump valves (ADVs) are upstream of the MSIVs, malfunction (5) can cause RCS cooldown to continue af ter MSIS. Consequently, a thermal-hydraulic evaluation of the atmospheric dump valve malfunction is provided.
The ADV's upstream of the MSIV's at Fort Calhoun Station consist of two manually
~
operated steam safety valves which are located on the two steam lines. The total capacity of each of these valves is less than 5 percent of the full power main steam flow rate at 900 psia steam pressure.
The analysis for Fort Calhoun station assumed a valve capacity of 5 percent of full power main steam flow.
The thennal-hydraulic evaluation assumes that the valve opens and remains open, resulting in uncontrolled RCS cooldown for 90 minutes.
In order to maximize RCS cooldown, the malfunction is assumed to occur in the HZP operating made with no decay heat.
Operator actions assumed for the A00 are the same as listed in Tabl'e 1 for SLB, however, the time frame will differ due to the less rapid cooldown for the A00 transient.
The resulting downcomer water temperature and pressure transients are provided in Figure 3.
l l
l
Table 1.
Assumptions and Plant Parameters Used for SLB Thermal-Hydraulic Evaluation 2
Parameter Value Steam Flow Area 2
a) Affected Steam Generator 3.7 ft b) Unaf fected Steam Generator 2
i) Before MSIV Closure 1.8 ft ii) Af ter MSIV Closure 0.0 Blowdown Quality 1.0 Initial Power Level 0.0 i
i Decay Heat 0.0 MSIS Setpoint 500 psia SIAS Setpoint 1600 psia HPSI Flow Shutoff Pressure 1400 psia AFW Flow 0.0
)
Operator Actions SLB A00 a)
Trip RCPs af ter SI AS on Low Pressurizer Pressure 30 sec 10 min b)
Reduce Charging and HPSI Pump Flow to Terminate RCS Repressurization 1800 sec 90 min i
l i
FIGURE 1 SLB SCOPIfiG CALCULATION, LARGE BREAK AT ZERO PONER, MAlfi FEED ISOLATI0ft AFTER MSIS, AUTOMATIC ISOLATI0fl 0F AUXILIARY FEED Oil STEAM GEf4ERATOR tP l
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Enclosure B i
1.
Results of Fracture Mechanics Analysis for Ft. Calhoun Steam Line Break Transient l
The stress analysis and fracture mechanics analysis were performed j
using the methods outlined in CEN-189.
Plant specific material properties for the controlling longitudinal weld in the Ft. Calhoun i
vessel were used in the analysis as follows:
PCT Ni
.99
=
.35 PCT Cu
=
.012 PCT P
=
j l
l This corresponds to the intennediate shell axial weld at an azimuthal angle of 0 degrees.
The fluence factor at the mid-core level at this location is 92% of the peak fluence in the vessel. At the 12/31/81 level of 5.4 EFPY and peak fluence of.704 x 10 n/cm2 (E >l MeV), this corresponds 19 19 2
j to a point fluence of.648 x 10 n/cm and an adjusted surface RT value NDT j
of 209 F.
j The plot of K vs time for this case is shown in Figure 1 for various assumed y
crack depths. These stress intensities result from the stresses due to the pressure and temperature transient given in Figures 1 A 2 of Enclosure A.
4 The applied stress intensity values were used in detennining the critical l
l crack depth diagram as shown in Figure 2 for the Ft. Calhoun vessel at an l
additional 26.6 EFPY of operation. A lurge region is indicated where K y exceeds the arrest toughness, however, no crack initiation region is present.
Taese results indicate that no crack initiation would occur throughout the Ft. Calhoun plant life for this particular Steam Line Break transient.
j 2.
Results of Fracture Mechanics Analysis for Ft. Calhoun A00 Transient The pressure and temperature transient for this A00 case tls presented in Figure 3 of Enclosure A.
The resulting plot of K vs time is given in y
Figure 3 for various assumed crack depths. The projected end-of-life critical crack depth diagram is given in Figure 4 which corresponds to an additional 26.6 EFPY of operation. A small region exists where the arrest toughness is exceeded, however, no crack initiation region is evident. These results indicate that no crack initiation would occur throughout the Ft. Calhoun plant life for this A00 transient.
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Enclosure C FORT CAlliOUf1 REACTOR PRESSURE VESSEL MATERIALS If1FORMATI0ft 150 DAY RESP 0flSE TO DARREL G. EISETHlUT LETTER OF AUGUST 2 C 981, El!EU$URE 1 Request for Additional Information Item #6: Vessel Welds Information Requested:
Axial and azimuthal locations of vessel weld-seams with respect to the core.
Overlay of current fluence map with weld locations.
Identify the critical welds, vertical and circumferential, and give the weld wire heat numbers.
Give weld chemistry for the critical wel ds.
For each weld wire heat number, report the estimated mean copper content, the range and the standard deviation, based on all the reported measurements for that weld wire heat.
The welds may be surveillance weldments for your vessel or others, nozzle dropouts that contain a weld, weld metal qualification data, or archive material.
In the absence of any information, assume that copper content is at its upper limit (0.35 percent when using R.G.1.99, Rev.1) and that the nickel content is high.
Response
a) Axial and Azimuthal Locations of Vessel Weld Seams With Respect the Core --
The vessel weld seams are shown in Figure 1 together with the initial Reference Temperature, RTilDT, for each weld.
The core midplane and the extent of the active core are also indicated.
This information is as submitted in Appendix A, Section A.6 to CEft-189.
I b) Overlay of Current Fluence Map With Weld Locations --
The neutron flux profile is shown relative to the vessel weld map in Figure 2.
Values are given in the Figure as normalized iso-flux values.
Peak flux values (normalized flux equal to 1.0) occur at the 45 azi-muthal locations (eg., 45'), 135", 225, a nd 315" ).
Secondary peaks (normalized flux equal to 0.861) occur at the 900 azimuthal locations (eg,900, 180), 2 7 0", a nd 3 60" ).
c) Chemistry of Critical Welds --
The nickel, copper, and phosphorus content of the critical welds and j
the surveillance weld are given in Table 1.
The basis for these values is discussed in CEft-189.* The wire heat and flux lot numbers are in-dicated in the Table.
The chemistry for weld seams 2-410 (A, B, & C),
3-410 (A, B, & C) and 9-410 is not available; copper and phosphorus content,, therefore, reflec t the Regulatory Guide 1.99, Rev. 1, upper limits, and the nickel content was assumed to be "hiqh" (greater than 0.30 w/o per CEff-lG9*) based on the type of wire and weld processes used.
- see Section 6.3 and Appendix A, Section A.6 Furthermore, no appropriate data are available to perform a statistical analysis of the copper content for these welds.
The phosphorus and i
nickel contents for weld seams 1-410 (A, B, & C) and 8-410 were obtained from records of the coated electrodes used to fabricate those seams, but no copper content was reported.
The 0.07 w/o Cu value reported in Table 1 representr, the upper bound value from 44 lots of similar coated elec-trodes as described in CEff-139 Section 6.3; the mean value for those i
44 lots is 0.0275 w/o Cu with a range of 0.02 to 0.07 w/o and a standard deviation of 0.0236%.
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REQUEST FOR ADDITIONAL INFORMATION BY THOMAS M. NOVAK, LETTER DATED DECEMBER 18, 1981, ENCLOSURE 1 Information Requested Item #1, RTUDT Values:
Initial RTNDT for the welds was the subject of a letter dated November 13, 1981, which supplemented your "60 day" response of October 20, 1981. We still have some questions that lead us to not accept your new value of
-50F for the initial RT f the longitudinal beltline welds.
The NDT archival material that was tested recently was the surveillance weldment.
According to your letter of September 8,1977, the weld wire and weld flux for the surveillance weld were not the same as for the longitudinal welds, as shown in Table 1.
Also from that letter, the drop weight RTNDT of the surveillance welds was reported to be OcF.
It is not clear why the recent test of the archival material should give a value 50 degrees lower.
We are inclined to use the generic upper 2a value of -20 F described in your October 20 and November 13 letters.
This is partly because of the discrepancies described above and partly because -20cF is approximately the Charpy 30 ft-lb level shown for the surveillance weld.
If the peak ID fluences as provided in your letter Sated October 20, 1981, are not the fluence values for the critical longitudinal welds, provide the peak fluence values at the critical longitudinal welds. When the above l
is provided, we will then be able to determine current RTNDT values which we will use in our independent assessments.
Response
The "60 day" response of October 20, 1981 provided an estimated initial RTNDT of -20CF for the longitudinal seam welds.
Subsequent to that letter, drop weight specimens machined from an archive section of the surveillance weld were tested, resulting in a drop weight NDTT of -50 F, I
as shown in Table 2.
The RTNDT for the surveillance weld wi.s then determined to be -500F applying the Charpy and drop weight data in ac-cordance with paragraph NB2331 of the ASME Code,Section III.
The sur-veillance weld was then used as a benchmark to establish improved estimates of the vessel weld seam RTUDT's (as discussed in the November 13, 1981, letter and CEN-189")which were transmitted in the November 13, 1981, letter.
It has been previously acknowledged that the surveillance weld was fab-ricated with a different heat of wire and lot of flux than the vessel wel dmen t s. However, the fact that the weld process and procedures and the types of wcld wire and flux were the same for both the surveillance weld and the vesse' longitudinal seam welds (2-410 and 3-410) makes the surveillance weldment the best available benchmark for estimating the 1
properties of the vessel welds.
It should also be noted that no
- section 6.4 I
discrepancy exists concerning the surveillance weld RT DT; the 00F N
value reported in the September 8,1977, letter was a conservatively er cimated value based on just Charpy impact data; at that time, no drop weight test results were available. The recent drop weight tests provided the first opportunity to establish the RTNDT in accordance with the ASME Code (i.e., use of both drop weight and Charpy impact da ta ). With regard to the " approximate 30 f t-lb level", it should be noted that this is a rule of thumb developed by the Naval Research Laboratory (long before the RTNDT concept was formalized) to approxi-mate the lower " knee" of the transition curve, and it was not intended to define NDTT or RTNDT. As a basis of comparison, the baseline test results for the Calvert Cliffs Unit #1 surveillance weld are reproduced in Table 3.
This material exhibited a drop weight NDTT of -80 F, but.
30 ft-lb impact energy v.as not attained until about -50CF.
Based on this very limited sample (two submerged arc welds), it appears that the 30 ft-lb temperature lies well above the drop weight NOT temperature and RTNDT.
Information requested concerning neutron fluence values for the critical longitudinal seam welds is indicated by the iso-flux contours in Figure 2.
None of the longitudinal seam welds are in peak flux locations.
The maximum relative flux and the fluence as of December 31, 1981, are as follows:
Estimated Maximum Relative Vessel Weld Weld Seam Azimuthal Location Neutron Flux Fluence 1 2/31 /81 2
2-410 00 0.861 0.606x10l9n/cm l9 2
2-41 0 120c& 2400 0.680 0.479x10 n/cm I9 3-410 1800 0.852 0.600x10 n/cm2 2
3-41 0 600& 3000 0.673 0.474x10l9n/cm.
TABLE 1 Fort Calhoun Reactor Vessel Weld 14aterials Weld Sean Wire Heat Flux Lot Chemical Content (%)
Nickel Copper Phosphorus 1-410 A,B,C HBEGa 1.75 0.07b 0.01 0 d
d 2-410 A,B,C 51989 3687 0.99c 0.35 0.01 2 JBFGa 1.29 0.07b 0.008 d
3 al0 A,B,C 13253 3774 0.99c 0.35 0.01 2d 12008 3774 0.99c 0.35d 0.012d d
27204 3774 1.07 0.35 0.012d E0AG 1.04 0.07b 0.008 a
.m' 8-41 0 LODGa 0.93 0.07b 0.009 9 -41 0 20291 3833 0.99c 0.35d d
0.01 2 b
HADHa 0.94 0.07 0.012 Surveillance 30541 3951 0.60 0.35 0.01 3 a
Coated electrode b
Upper bound for coated electrodes (see text and CEN-189) c Estimated nickel content (high nickel type wire or weld process; see text and CEN-189) d Regulatory Guide 1.99 upper bound prediction limit
TABLE 2 Drop Weight NDTT Determination for Fort Calhoun Surveillance Weldment Specimen Test Individual Number Temperature Test Results NDTT OMW 3
-40 F No Break OMW A
-400F No Break 0
OMW l
-50cF Break
-50 F 9
r TABLE 3 Calvert Cliffs Unit #1 i
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Test Impact lateral Specimen Temp.
Energy Expansion Shear Code
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(Mils) 334
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346
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- 80 15.3 15 10 35A
- 40 38.6 35 30 34Y
- 40 44.5 38 35 4
33P
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- 20 91.8 67 50 344
- 20 100.7 71 65 356 0
62 52 60 332 0
85 63 50 361 40 132.8 90 85 31Y 40 133 100 90 36A 80 140 95 100 32L 80 142 97 100 34U 120 153.9 98 100 31J 120 156 100 100 34T 160 158.4 101 100 322 160 160.6 97 100 Drop Ueight UDT = -80"r RT
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Enclosure (D)_
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i Fluence Data for Fort Calhoun 4
j 1.
Geometry i
The geometry applicable to surveillance capsule analysis is shown j
in Figures 1.1 and 1.2, and the applicable engineering drawings, E-232-426 Rev. 09 and E-232-551 Rev. 4.
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2.
Material Description The composition of the materials which are used in typical neutronic i
models for neutron transport calculations from the core to the vessel are shown in Table 1.
The typical arrangement of these materials is shown in Table 2.2.
j 3.
The neutron source data for Fort Calhoun Cycles 1 through 6 were provided in a letter from W. C. Jones to R. A. Clark dated November 12, 1981.
The basis for this present study has been the linear extrapola-i tion of the current fluence level to end of the life.
Fuel management changes anticipated for Cycle 8 should reduce the power in assemblies adjacent to longitudinal welds by a factor of two.
4.
The vessel fluence distribution data reported in CEN-189 for the i
Fort Calhoun reactor as part of the C-E owners group pressurized thermal shock effort is included here for information.
5.
The surveillance capsule analysis data is given in Reference 4-1 of l
Appendix A to CEN-189.
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4 TABLE 2.2 Typical Material Arrangement i
Homogenized Core Water (Adjusted by Appropriate Densit; Factor)
Stainless Steel (Core Support Barrel)
Water (Adjusted by Appropriate Densit3 Factor)
Stainless Steel (Thermal Shield, if present)
Water (Adjusted by Appropriate Density Factor)
Stainless Steel (Vessel Cladding)
Carbon Steel (Reactor Vessel).
4 1
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' e 1
l Omaha Public Power District supplied the reactor power history and detailed l
radial power distributions needed to update the azimuthal fluence distribu-tion from the time at which the surveillance capsule was removed to December 31, 1981.
Through December 31, 1981 a cumulative energy generation of 66,655,037 Megawatt hours was quoted as shown in Table A5-1.
Defining full power as 3
1420 Megawatts-thermal (Mwt) this energy output yields 5.36 Effective Full Power Years (EFPY). According to the surveillance capsule analysis (AS-1) i the peak fast neutron fluence on the reactor vessel was 3.4 x1018 (n/cm2) af ter 2.59 EFPY at 1420 Mwt. This implies a rate of accumulation in the 2
18 (n/cm ) per EFPY at 1420 Mwt. As a result the
]
peak fluence of 1.31 x 10 value of the peak wall fluence as of December 31, 1981 (5.36 EFPY) is 2
7.04 x 1018 (n/cm ).
A comparison of the surveillance capsule dosimetry results with the neutron energy spectrum calculated at the surveillance capsule position using a
)
DOT-R9 model yielded a calculated value within 4% of that quoted in the i
surveillance capsule report. Therefore there is good confidence in the surveillance capsule analysis results.
The full power level was cefined as 1420 f+dt because this was the full power level during the exposure experienced by the surveillance capsule and most of tne time period up to December 31, 1981.
For extrapolation to future times the full power level is referenced to 1500 r.ht and an annrnvimate estimata of the increased fluence accumulation rate is obtained by multiplying the previous rate by the ratio of the power densities (1500/1420 = 1.056).
4 A summary of the results obtained from the surveillance capsule analysis report is shown in Table AS-2.
TABLE AS-2 l
Peak Vessel Effective Full Peak Fluence Full Power Level Power Years AccumglationRate WallF{uence (n/cm )
(Mwt)
3.4 x 1018 1420 2.59 1.31 x 1018 7.04 x 1018 1420 5.36 (12/31/81) 1.31 x 1018 1
i
The azimuthal shape of the fluence distribution was obtained by updating i
the 00T-R9 azimuthal distribution corresponding to the surveillance capsule analysis (2.59 EFPY) to December 31, 1981 (5.36 EFPY). The adjustment factors were calculated using the SHADRAC code as described in Section 5.2.2.
j l
The detailed radial power distribution corresponding to 5.36 EFPY was obtained by combining the power distributions supplied by Omaha Public Power District for cycles 4 through 7 with the distribution used to represent the time up to the end of cycle 3 in the surveillance capsule analysis. The nodalization of the power distribution is shown in Figures A5-1 and AS-2.
The detailed power distributions for cycles 4 through 7 are shown in Figures A5-3 through AS-6.
The resulting azimuthal fluence distribution is shown in Figure A5-7.
The 00 reference point for the azimuthal distribution is shown in Figure A5-8.
One-eighth core symmetry was assumed. The axial and radial fluence distributions in the reactor vessel are obtained from 00T-RZ calculations and are as shown in Figures AS-9 and A5-10, respectively.
References:
4-1.
Omaha Public Power District Fort Calhoun Station Unit No.1, Evaluation of Irradiated Capsule W-225, Combustion Engineering, TR-0-MCM-001 Rev. 1, August 1980.
i s
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Fort Calhoun Power History
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4 TABLE AS-1 i
1 I
i f
i i
1' i
i 1
i Power Level Cycle EFPD (MW,h)
MW-HR i
1 307.56 1420 10,481,698 2
371.58 1420 12,663,429 3
265.08 1420 9,033,881 4
280.65 1420 9,564,550 1
5 360.18 1420 12,274,999 6
77.05 1420 2,625,937 4
6 278.07 1500 10,010,543 i
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