ML20235F544
| ML20235F544 | |
| Person / Time | |
|---|---|
| Site: | Fort Saint Vrain  |
| Issue date: | 07/02/1987 |
| From: | Office of Nuclear Reactor Regulation |
| To: | |
| Shared Package | |
| ML20234C612 | List: |
| References | |
| TAC-63576, NUDOCS 8707130433 | |
| Download: ML20235F544 (5) | |
Text
,.
so m
/
'o UNITED STATES NUCLEAR REGULATORY COMMISSION
~
n E
WASHINGTON, D, C. 20555 i
i o.,...../
SAFETY EVALUATION BY THE OFFICE OF NUCLEAR REACTOR REGULATION RELATING TO THE EFFECT OF FIREWATER C00LDOWN ON STEAM l
GENERATOR STRUCTURAL INTEGRITY PUBLIC SERVICE COMPANY OF COLORADO FORT ST. VRAIN NUCLEAR GENERATING STATION DOCKET NO. 50-267,
1.0 INTRODUCTION
On December 30, 1986, Public Service Company of Colorado (PSC) submitted a letter (Ref.1) with an attached analysis to justify operation of Fort St. Vrain, at 82% power.
Safe operation at this level is based on a proposal to circulate available plant firewater through the Evaporator and Economizer /Superheater (EES) tube bundles of the steam generator modules in an event requiring safe shutdown of the plant.
This would permit safe shutdown cooling following a ninety (90) minute interruption of forced helium circulation. (Ref. 2) to the letter represents the I
structural evaluation of the most critical regions of the steam generators during a single cycle cooldown from 82% power.
The objective of this eval-uation was to show that the primary pressure boundary of the steam gener-ator will remain intact even when firewater is used in the EES tube I
bundles to cool down the reactor from a more conservative 100% power condition.
The primary analysis for structural integrity of the steam generator is based on creep collapse.
Creep collapse, is considered to be the only conceivable mode of failure for the hottest steam generator tubes under the conditions analyzed.
2.0 EVALUATION Three regions within the steam generator modules were determined to be of concern during this event:
- 1. the reheater tubes and their supports;
- 2. the EES tubes and'their supports.
The superheater tubes consists of two sets of tubes, labeled Superheater I and Superheater II. Of these, the latter are the most affected.
- 3. The Superheater II downcomer and its support.
The critical region comprises the reheater tubes.
During EES cooldown I
l conditions these tubes have uniform temperatures and low internal pressures, so that the interaction loads between tubes and the tube support plates are low.
However, there is a time interval of about 90 minutes during which the helium flow is stopped.
In this interval the reheater and the EES tubes are purged of water / steam inventory, and are I
B707130433 870702 PDR ADDCK 05000267 P
PDR j
i l
+
. in fact vented to the atmosphere, so that the internal pressure in these tubes is essentially atmospheric.
The helium pressure peaks at 700 psig and the metal temperature drops from 1000 F to 780"F.
These tubes are therefore under net compressive radial loading.
This external pressure, I
combined with the hottest helium temperature impinging on the tubes indicates that a likely mode of failure for the reheater tubes is by creep buckling and collapse.* General Atomics (GA) has performed an evaluation of the reheater tubes subject to these conditions using the GA BUCKLE computer program (Ref. 3).
The basis for this program (Ref. 4) was reviewed and found to be reasonably acceptable.
The program was verified by comparison with the solutions for two test problems involving creep collapse, obtained by using the widely available program MARC-General l
Purpose Finite Element Program (Ref. 8).
GA provided this verification (Ref. 5) and has shown that there is reasonable agreement between the two sets of solutions.
We find this acceptable.
The evaluation consisted of an analysis of a straight tube of nominal diameter and wall thickness, and the maximum allowable small-radius-bend i
manufacturing ovality, subjected to external pressure at high temperature, j
l For evaluation of the creep buckling / collapse of a cross-section, GA l
stated that in this case the analysis of a straight tube is conservative as compared to the analysis of a curved tube since the tubes do not experience restraint of thermal expansion and any bending moments in the i
bends are therefore minimal.
(The tube bundle and its support structure j
were stated to be at the same temperature.) Likewise bending moments due to safe shutdown earthquake (SSE) and low level vibration were determined also to be minimal.
The additional ovalization due to such moments is therefore also insignificant, and a tube bend is therefore more resistant to external pressure than a straight tube.
However, if the bending loads J
had been significant, then the effect of the additional tube ovality caused by bending of an already curved tube would have required evaluation.
We find these arguments acceptable.
The program determines the creep buckling time for a straight cylindrical l
tube with initial ovality, subjected to constant external pressure and constant material temperature.
The criterion for determining that buckling has occurred is that the highest stressed point in the tube cross-section has reached a stress level equal to the yield stress of the material at temperature.
This is a conservative criterion for thicker tubes such as the reheater tubes, which have a diameter-to-wall-thickness l
ratio of eight.
Such tubes were shown not to buckle at stress levels below the yield stress.
We find this procedure acceptable.
Since the program uses constant values of external pressure and tempera-ture, a parametric study was performed to determine the time for achieving the yield stress level for various combinations of pressure and tempera-ture.
It was thus determined that the creep collapse time at 1660 F and 350 psig exceeds 20 hrs.
This is a conservative result since the time
- The term " buckling and collapse" does not imply sudden, catastrophic deformation under creep conditions.
Rather, it describes time dependent deformation.
It is thus characterized by a time limit instead of a load limit.
interval at which the actual tube experiences this temperature during the cooldown event is considerably shorter.
This value of temperature repre-sents the maximum temperature experienced by the tube after helium circu-lation is restarted.
After about 7 hours8.101852e-5 days <br />0.00194 hours <br />1.157407e-5 weeks <br />2.6635e-6 months <br /> it was shown to achieve a value of approximately 350 F, after which there is a further decay with time.
Likewise, the external pressure was shown to drop to about 300 psi, after which it remained approximately constant.
Thus, the actual creep collapse l
time appears to be much longer than 20 hours2.314815e-4 days <br />0.00556 hours <br />3.306878e-5 weeks <br />7.61e-6 months <br />, and certainly much longer J
than needed to achieve cold shut down.
However, prior creep fatigue damage can affect the creep-collapse failure mode if the prior accumulated creep-fatigue damage is high.
GA has therefore stated that the fatigue l
analysis performed and reported in the original steam generator design i
report (Ref. 6) showed low fatigue damage even in those regions of the steam generator where there were significant cyclic loads.
Since the l
thermal loads in the reheater are low it is expected that there will not be any significant creep-fatigue damage in the reheater tubes, particularly adjacent to the small radius bends where the maximum ovality is expected to occur.
Creep-fatigue damage will therefore not likely affect signifi-cantly the predicted creep-collapse times.
We find this acceptable.
The reheater tubes and their supports are basically at the same temperature during EES firewater cooldown. The interaction loads batween the tubes and the tube support structures due to restrained thermal expansion are therefore minimal.
Likewise,' the stresses due to combined SSE loading, dead weight,. internal pressure and helium flow drag loads were shown in Ref. 6 to satisfy the ASME Code,Section III, requirements for primary stresses at 100% power. During the firewater cooldown, the stresses due to pressure and thermal effects were determined to be lower than the stresses at 100% power, while at the same time the allowable primary stress was higher, since this is a one time, short term plant event.
GA has therefore provided reasonable assurance that an SSE occurring during firewater cooldown will not fail the reheater tubes. We find this accept-able.
A structural investigation of the Superheater II helical bundle under combined sustained and thermal expansion loading was also performed.
The stresses under firewater cooldown were determined by multiplying the full power operation stresses by the ratios of the corresponding thermal and mechanical loads for the two conditions.
The full power operation stresses were obtained from Ref. 6.
The pressures and tube wall thermal gradients are lower under firewater cooldown than those under full power operation.
However, the loads resulting from tube and tube-support plate l
interaction increase due to restrained thermal expansion. The stresses due to this interaction are, however, classified as secondary and are therefore not required to be evaluated for faulted conditions. GA has evaluated the primary stresses in the tubes when subjected to combined l
flow induced vibration, dead weight, SSE and pressure loading and has determined that the ASME Section III limit at temperature as stated in Ref. 7 is satisfied.
Although not required, GA has also evaluated the combined total primary plus-secondary stresses, to verify that even under upset conditions, the required stress limit at 385'F* is also satisfied.
This stress limit was obtained from Ref. 7, which is applicable at
^1his is the peak tube surface temperature.
e L
l 1
1 l
temperatures below 800 F.
The EES tube-support plate stresses were also l
stated to be within design allowables at the firewater cooling tempera-i tures.
We find these evaluations acceptable.
For the Superheater downcomer and its support structure, GA has stated that during full power operation the differences in temperature are very i
small, and will not cause significant differential thermal expansion.
The temperatures, as well as the pressure in the downcomer, during firewater cooldown are considerably lower, while the SSE stress remains the same.
Because of the lower temperature the stress limit also increases, so that the safety margins increase considerably as compared to those at full power operation.
Thus, on the basis of a comparison of the temperatures l
used for analysis at 100% power with those predicted during firewater l
cooldown, GA determined that the stress limits would not be exceeded under this condition.
We find this acceptable.
3.0 CONCLUSION
l The results of the structural evaluation by GA of the most severely loaded l
regions of the steam generator, during a single cycle of cooldown from power using plant firewater, has shown that the most critical region is l
in the reheater.
GA has demonstrated that the reheater tubes will not experience collapse due to creep during such a cooldown cycle, including 1-1/2 hours of interrupted forced circulation, provided that the power l
level is limited so that the maximum helium temperature is less than 1660 F for less than 20 hours2.314815e-4 days <br />0.00556 hours <br />3.306878e-5 weeks <br />7.61e-6 months <br /> and the corresponding maximum helium pres-sure is no more than 350 psig.
GA has also demonstrated that in other regions, such as the Superheater II tubes and the Superheater downcomer, the stresses are below the corresponding ASME Section III design limits.
i We find these demonstrations acceptable.
4.0 REFERENCES
1.
Letter No. P-86683, from D. W. Warenburg, PSC, to H. N. Berkow, NRC, dated December 30, 1986.
2.
Lewis, A. G., "Effect of Firewater Cooldown Using EES Bundle on Steam Generator Structural Integrity." GA Documents 909204/NC and 909204/A, December 4, 1986 and June 5, 1987.
3.
Almajian, A. T., " Buckle Users Manual." GA Document 909410/NC, April 1987.
4.
Pan, Y. S. " Creep Buckling of thin Walled Circular Shells Subject to Radial Pressure and Thermal Gradients," J. of Applied Mech., March 1971, pp. 206-216.
5.
Almajian, A. T., " Verification Report for the BUCKLE Computer Program."
GA Document 909438/NC, May 28, 1987.
6.
GA Report No. GADR-9, " Stress Report for Fort St. Vrain Steam Generator," October 1, 1970.
7.
ASME B & PV Section III Summer 1968 Edition, including Code Case 1342-1, 1
1 8.
" MARC General Purpose Finite Element Program," MARC Analysis Corp.,
Palo Alto, California, March 1983.
Date: July 2,1987 Principal Contributor: Mark Hartzman, MEB l