ML20028G618

Rev 6 to Analysis of Hydrogen Control Measures at McGuire Nuclear Station.

ML20028G618
Person / Time
Site:	McGuire, Mcguire
Issue date:	02/15/1983
From:	DUKE POWER CO.
To:
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ML20028G612	List: ML20028G611 ML20028G618
References
NUDOCS 8302170144
Download: ML20028G618 (35)

Text

McGuire Nuclear Station An Analysis of Hydrogen Control Measures at McGuire Nuclear Station Revision 6 February 15, 1983 CHANGES AND CORRECTIONS:

Remove and insert pages in accordance with the following tabulations:

Remove these pages _ Insert these pages Volume 3 - Chapter 3 -

3.4-2 3.4-2 3.4-3 Figure 3.4-5 Figure 3.4-5 Figure (s) 3.4-7 thru 3.4-9 Volume 3 - Chapter 4 Table (s) 4.4.3 thru 4.4.8 Volum9 3 - Chapter 7 7.0-98 thru 7.0-116 I

i l

8302170144 330215 PDR ADOCK 05000369 P PDR

-( N the ice condenser upper plenum. The analysis showed that six igniters spaced approximately equally will prevent hydrogen concentrations greater than 10%

by volume in the ice condenser at any location if ignition takes place at 8.5% hydrogen by volume and propagation is at one foot /second, for the maximum hydrogen release rate predicted for the McGuire small break LOCA.

Because 10% is considerably below any plausible detonation limit, this design criterion is conservative.

Power to the igniter system is supplied by 10 circuit breakers, five 'each on Emergency Lighting Panelboards ELA1 and ELBl. These circuit breakers, panel-boards, and associated connections to Class lE power sources are shown on Figures 3.4-7 and 3.4-8 for ELA1 and ELB1 respectively. The connection diagram for the igniters themselves, which indicates the number and location p of !gniters connected to e'ch circuit, is shown in Figure 3.4-9. Periodic LJ measurements to establish the operability of the igniters are made at Emergency 6

, Lighting Panelboards ELA1 and EL81.

l Because each region of containment is supplied by at least one redundant pair .

of igniters, a failure which renders a single igniter inoperable, or which causes ,

the failure of all igniters associated with a single circuit or panelboard, will not affect the ability of the hydrogen ignition system to perform its intended function.

l The function of the Hydrogen Mitigation System is to ignite mixtures of hydrogen and oxygen in the various areas of the containment when the local concentration of hydrogen has reached 8.5%. The early ignition of hydrogen l in the containment has several benefits.

C)

< 3.4-2 Revision 6 l

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_/ a. Hydrogen-oxygen recombination reactions occur at hydrogen concentrations

(

substantially lower than those shown to produce uncontrolled reactions.

Thus detonation of hydrogen is precluded.

b. Reaction at low hydrogen concentration produces lower temperatures and smaller energy release rates. Because more time is available for the containment pressure mitigation systems (ice condenser and containment spray) to dissipate energy, the total containment pressure rise is lower. Lower temperatures also have less adverse effects on other equipment in contai?-ant.

c. The product of deflagration (pure water) does not represent a threat to containment or system integrity or to personnel entering containment following termination of the accident.

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, Table 4.4.3 MCGUIRE CLASIX RESULTS

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CASE 1 - REFERENCE CASE Flame Speed = 2 ft/sec Number of burns LC 5 UP 16

Total H2burned (1bm) 957 H2 remaining (lbm) 581 Peak Temperature (F) LC 1138

LP 228 UP 1517 UC 162 DE 238 Peak Pressure (psig) LC 11.8 LP 11.7 l

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l Ice remaining (1bm) 4.27 x 10 Revision 6

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Table 4.4.4 O MCGUIRE CLASIX RESULTS

SUMMARY

CASE 2 - MINIMUM SAFEGUARDS Flame Speed = 2 ft/sec Number of burns LC 7 UP 10 i

Total H2 burned (1ba) 956

. H2 remaining (1bm) 582 4

Peak Temperature (F)' LC 1101 LP 218 UP 1520 UC 175 DE 245 l Peak Pressure (psig) LC 12.7 LP 12.7  :

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CASE 3 - HIGH H, RELEASE RATE Flame Speed = 2 ft/sec Number of burns LC 11 UP 22 Total H2burned (1bm) 1829 H2 remaining (lbm) 532

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CASE 3 - ICE MELT 00T BEFORE BURN Flame Speed = 2 ft/sec i

Number of burns LC 7 UC 1 Total H2 burned (lbm) 1004 H2 remaining (1ba) 533 Peak Temperature (F) LC 1167 LP 261 UP 282 i UC 407

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CASE 4 - INERTED LC Flame Speed = 2 ft/sec l

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DE 2 Total H2burned (1bm) 958 H2 remaining (lbm) 579 Peak Temperature (F) LC 225 or 690 UP 1427 UC 160 DE 796

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l CASE 5 - INERTED UP Flame Speed = 2 ft/sec Number of burns LC 8 UC 1 Total H2 burned (lbm) 950 H2 remaining (1bm) 542 Peak Temperature (F) LC 1250 LP 270 UP 167 UC 388 DE 208

Peak Pressure (psig) LC 14.4 i LP 14.4 i UP 14.4 UC 14.4 DE 14.4 S

Ice remaining (1bm) 4.92 x 10 1

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s Response to Request For Information On Hydrogen Control

, ) Transmitted By Ms. Elinor G. Adensam's Letter Of January 24, 1983

1. Recent discussions with Duke suggest that the upper plenum igniters in McGuire are located alternately between the crane wall and side and the containment shell side of the upper plenum in a staggered fashion, rather than all on the containment shell side of the upper plenum as depicted in Figure 3.4-5 of the McGuire submittal. Verify the "as-installed" locations for these and all other igniters in the permanent system.

Provide revised drawings as necessary.

Response

See revised Figure 3.4-5.

2. Section 3.4 of the McGuire submittal cites results of an analysis performed by Duke to ensure that 6 of the 12 upper plenum igniters will provide adequate coverage of this region. Based on the information ssupplied, this analysis appears to differ somewhat from one performed by TVA to justify the use of 16 upper plenum igniters in the McGuire and Sequoyah plants, provide details and justification of the McGuire analysis including (1) assumptions regarding vertical and horizontal velocities of the rising mixture and propagating flame, and (2) the method for computing the maximum hydrogen concentration for a given ignition concentration. Address in your response the possibility of the gas mixture bypassing the igniters. If the analysis relies on horizontally propagating flames being carried into the upper compartment by rising gases, discuss the effects of turbulent mixing and dilution of the mixture with the upper compartment atmosphere.

i 1

7.0-98

Response

O The design objective for the igniter spacing in the upper plenum is to ensure that the hydrogen concentration will remain below 10% at all points in the upper plenum when hydrogen is being added at the maximum rate. The following are the assumptions used in this analysis:

1) Hydrogen burning begins at 8.5% hydrogen and propagates in all directions.

2) The burning velocity in the upper plenum is one foot /second in all directions.

3) The flow velocity of the hydrogen, steam, air mixture entering the upper plenum from the ice bed is one foot /second.

4) The maximum hydrogen addition rate to the upper plenum is 70 lbm/

minute. This is identical to the hydrogen release rate predicted by MARCH for the S 2 D accident sequence and thus takes no credit for O hydrogen removal by burning in the lower compartment.

t The igniter spacing used at McGuire (assuming only one train of operable l

igniters) is 50 feet. With this spacing, the burn time in the upper plenum is 24 seconds, corresponding to a hydrogen consumption rate of 190 l lbm/ min for hydrogen ignition at 8.5% by volume. Thus, the upper plenum igniter location and spacing presently used at McGuire'results in hydrogen consumption rates approximately three times greater than the maximum addition rate. At the time that hydrogen ignition occurs in the upper plenum, the upper plenum contains 76 lbs of hydrogen (an 8.5%

concentration). During the 24 second burn period, the maximum hydrogen addition, basec' on the 70 lbs/ minute release rate, is 28 lbm. Since hydrogen burning causes a pressure increase in the upper plenum, some unburned hydrogen will be pushed into the upper compartment. Based on

!O 7.0-99

A the 12 foot width of the upper plenum, we have estimated that 14 lbm of hydrogen will be pushed into the upper compartment during the burn period. This is the basis of our conclusion that the hydrogen concentration in the upper plenum remains below 10% by volume.

Concerning the effect of igniter spacing on hydrogen bypassing the upper plenum, it should be noted that whenever the hydrogen concentration in the upper plenum is less than about 6%, r.o ignition will take place in the upper plenum, and therefore all hydrogen entering the upper plenum will pass into the upper compartment. For hydrogen concentrations greater than 6%, ignition will occur at the upper plenum igniters, and there will be some flame propagation upward into the upper compartment.

Although until the upper compartment itself reaches approximately 6%

hydrogen, these flame: will extinguish quickly. Once the hydrogen con-centration in the upper plenum reaches 8.5%, propagation will be down-ward in the upper plenum, and the quantity of hydrogen able to reach the upper compartment wil' be reduced. Note that in the CLASIX analysis for McGuire, it is assumed that no hydrogen is consumed in the upper plenum until the upper plenum is at 8.5%, thus all hydrogen which escapes the lower compartment reaches the upper compartment. This amount of hydrogen is substantially greater than the amount of hydrogen which could be pushed into the upper compartment during a burn in the upper plenum. It is concluded that, cased on this analysis, the amount of hydrogen which bypasses the igniters in the upper plenum during the time that the con-centration is below 8.5% is much greater than the amount of hydrogen which could be pushed unburned into the upper compartment during a burn in the upper plenum. It appears, therefore, that once a high rate of O

7.0-100

hydrogen consumption has been established by the spacing of hydrogen igniters approximately 50 feet apart, hydrogen bypassing is no longer a consideration in the spacing of igniters.

3. Figure 3.4-2 of the McGuire submittal shows a typical igniter circuit but does not include details of the overall power distribution. In this regard provide a more comprehensive description of the power distribution to the igniters for each train. Specifically, describe the circuit branches, starting from the Class IE power supply and proceeding through I the control switches, fuses and emergency lighting panel board, to the igniters. Identify in your response: the number of control switches and the circuits which they control, the location of the fuses in the system, i

the number of igniters supplied through each fuse, the means by which system status can be monitored during an accident, the location in the

system at which voltage and current readings are taken for surveillance purposes, and the number of igniters on each tested circuit. Assess the potential of a short circuit in a single igniter to inhibit operation of the deliberate ignition system over a critical region of containment.

Response

,

See revised Section 3.4.

l l

4. Verify that the GM glow plug will reliably initiate combustion in a spray environment typical of the ice condenser upper compartment. Acceptable l tests for demonstrating igniter operability should be characterized by a l

O

, 7.0-101

spray droplet density equivalent to that in the upper compartment, droplets at terminal velocity with induced turbulence to simulate upper compartment mixing, and a sufficiently uniform mixture such that hollow cone nozzle effects are eliminated.

4 Res' mise:

The McGuire Hydrogen Mitigation System design, described in Chapter 3, does not directly expose any ignitors to a water spray. However, as part of the hydrogen research program jointly funded by AEP, Duke, and TVA, ignitor assemblies either similar or identical to those installed at McGuire were exposed to a water spray environment. In all cases, the presence of a water spray did not prevent combustion from occurring.

In order to compare the test conditions with the upper compartment environment under post LOCA conditions, spray densities were calculated.

Assuming both spray trains are operating in the upper compartment and that the spray is evenly distributed throughout the volume, a spray density of 0.015 lbm/ft3 was calculated. This value was obtained by I using a mean droplet diameter of 700p and a flow of 15.2 gpm per nozzle.

the assumed atmosphere temperature was 125*F. A description of the McGuire containment spray system is found in Section 6.5 of the FSAR.

The test environment selected for comparison was from the combustion tests performed by Acurex corporation. See Section 2.6 for a detailed description of the test program and facility. There were two reasons for selecting this project for comparison: 1) the ignition source used was identical to the McGuire ignitor assemblies and 2) the ignitor was l

O 7.0-102 I _ __

exposed to three different spray environments. Three tests were conducted with a Spraco 1713 nozzle. Operated at a flowrate of 15 gpm, the mean droplet diameter is 700p. Thus, by assuming an even spray distribution throughout the test vessel, a spray density of 0.004 lbm/ft3 was calculated. All three tests were conducted with a continuous injection of a hydrogen / team mixture.

Ten tests were conducted with a spray header consisting of nine micro-fog nozzles in the test vessel. Four of these tests consisted of igniting a l

premixed atmosphere (quiescent tests) and will not be discussed further.

The remaining six tests were conducted with a continuous injection of either hydrogen or hydrogen / steam. Two tests were conducted with a nozzle AP of 20 psi. This yielded a flow per nozzle of 0.12 gpm and a mean droplet diameter of 35p. The calculated spray density was 0.039 lbm/ft2 Four tests were conducted with a nozzle AP of 30 psi. This resulted in a flow per nozzle of 0.16 gpm and a mean droplet diameter of 21p. One of these tests was conducted with a mixing fan operating. The calculated spray density was 0.14 lbm/fts, The following table summarizes the calculations discussed ir, the preceding paragraphs.

Droplet Diameter Spray Density (microns) (lbm/f t3 )

Upper Compartment 700 .015 1

( SPRACO 1713 700 .004 1

Micro fcg (20 psi AP) 35 .039 Micro fog (30 psi AP) 21 .140 7.0-103 l

The data presented above shows that an ignitor assembly identical to that installed at McGuire has been exposed to and successfully performed in combustion tests with calculated spray densities greater than the spray density calculated fer McGuire. The fact that no ignitors at McGuire are directly exposed to spray further ensures that operation of the containment spray system will not adversely affect the operation of the hydrogen ignitor assemblies.

5. Provide the results of the AECL-Whiteshell combustion test with top ignition, 8.5% H 2, and 30% steam, as committed to by response to question 1C of the September 17, 1982, NRC Request for Information.

Response

Five additional low hydrogen concentration tests have been conducted by AECL-Whiteshell. All tests were quiescent with the ignition source located at the top of the test vessel. A description of the test facility is presented in the test reports provided in Appendix 2J. The test results are summarized in the following table:

Hydrogen Concentration Steam Concentration AP (v/o) (v/o) (psi) 8.5 15 0 9.0 15 0.5 9.0 30 0 9.5 30 19.5 8.5* 15 23 The last test was conducted with fans operating. This test is more representative of the dynamic post-accident conditions than the four quiescent tests. The Acurex dynamic test with top ignition yielded AP's of 13-20 psi. Thus, it has been experimentally demonstrated that ignitor lociations near the top of a compartment perform successfully in conditions typical of postuated degraded core accidents.

7.0-104

6. Tables summarizing CLASIX results are presented in the McGuire submittal for only the base case and the flame speed sensitivity case. Provide similar tables for all other sensitivity cases analyzed.

Response

Refer to new tables 4.4-3 thr.ough 4.4-8.

7. With regard to the structural capability of the McGuire containment, provide the following information:

(a) the ASME Eoiler and Pressure Vessel Code, Service Level C Pressure Limit for the McGuire steel containment shell.

(b) a brief description of the calculation method and mate. rial properties used to determine the ASME Boiler and Pressure Vessel Code Service Level C Pressure Limit for the McGuire steel containment shell.

(c) the pressure retention capabilities of the penetrations through the McGuire steel containment shell.

i (d) the pressure capacities of the operating concrete floor for i

! resisting pressures above and below the floow.

Response

(a) The ASME Service Level C pressure limit for the McGuire Containment Vessel is 45 psig based on the analysis described below.

O 7.0-105

) .

L

(b) The analysis to determine the McGuire Steel containment structural i capacity has been described in the Duke Power Company report entitled "An Analysis of Hydrogen Control Measures at McGuire Nuclear Station Volume 4", dated February 13, 1981, which has been previously submitted. Material properties used are those established by a statistical analysis of mill test reports for containment plate. The referenced report provides the detailed analysis methods and material properties. The results of this analysis are used to establish the ASME Service Level C Pressure Limit for the McGuire Steel Containment.

Using the 1980 Edition of the ASME Code,Section III, Subsection NE, Service Level C limits are met when the Primary Membrane Stress Intensity (Tresca stress intensity) reaches the material yield stress Sy. Using a 98% confidence level, this statistical yield point has been established to be 42.1 ksi for McGuire Nuclear Station. Since the stress intensity due to dead load is negligible compared to that due to internal pressure, and tends to negate the effects of internal pressure, the contribution of dead load is conservatively neglected. Inspection of the results of the containment capacity analysis described above shows that for a containment internal pressure of 45 psig, Tresca stress intensities are less than the statistical yield point for containment plate.

For a containment internal pressure of 50 psig, the calculated Tresca stress intensity exceeds the statistical yield point for the containment plate at isolated points, though yielding has not yet begun baset upon the maximua distortion strain energy (Von Mises) yield criterion.

7.0-106

] (c) All penetrations through the containment shell have been reviewed to

) verify their structural capability. In all cases it was determined that these penetrations were capable of withstanding a pressure of at least 67.5 psig.

(d) The calculated differential pressure capacity of the cperating floor is 41 psid.

8. In response to Item 2 of the NRC letter dated February 10, 1982, Duke has stated that the basis for the assumptions that the equipment did indeed reach an equilibriuc temperatura at least equivalent to the MSLB peak te.mperature (qualification temperature) was engineering judgment. Please confirm that the judgment and/or analysis was performed for all the equipment required for the hydrogen burn event. Also, provide the reference to individual summary component evaluation worksheets (SCEWS) together with the qualification profile for all the equipment whose survivability is demonstrated based on the qualification test performed in accordance with NUREG-0588. For any other equipment which is required for the hydrogen burn event but is not in the EQ program, provide the justification for survivability during the hydrogen burn event.

Response

To establish the temperature response of the cables in question to the LOCA qualification test chamger environment, a model of the cable identical to that contained in Section 5.4 was used. This model uses cable parameters obtained from the vendor. The following assumptions were used in this analysis:

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1. The test chamber is heated using a flow of steam. Accordingly, a forced convection flow coefficient based on a velocity of 1 foot /sec. was used. No credit was taken for test chamber preheating

- it was assumed that the cable began the test at ambient temperature.

2. Heat transfer to the cable was considered only as convection.

Radiation from the test chamber walls and from the steam in the chamber to the cable were neglected for conservatism. Note that the radiation contribution to the heat transfer is a significant component during the early part of the thermal transient and its neglect is a particularly conservative assumption when temperature rise of the cable exterior is considered.

O The results of the calculation may be summarized as follows:

1. The Duke supplied RTD Cable (0.810 inches 00, Okonite type FMR) -

Jacket temperature rise was very rapid, reaching thermal equilibrium within a few seconds of the start of the steam flow. The insulation and conductor layers under the cable armor come into approximate thermal equilibrium less than two hours later. The cable remains in this environment an additional four hours. For the core exit thermocouple cable, we have shown by previous analysis that its response to the temperature profile created by hydrogen burning is very similar to that of the duke supplied RTD cable. Accordingly, its response during the LOCA qualification test will also be similar (its total exposure was for eight hours rather tnan six) and

therefore was not analyzed explicity.

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2. The Westinghouse supplied RTD cable - because this cable has lower s.) heat capacity and smaller diameter than the Duke supplied RTD cable, its response was expected to be much more rapid. The cable reached approximate thermal equilibrium through its cross section in approximately one minute. /.dditional exposure to the same environment continued for nineteen additional minutes.

In order to provide an additional check on this calculation, another calculation was performed using the unrealistically conservative assump-tion that the LOCA qualification test took place in a quiescent chamber where the only heat transfer to the cable was by natural convection from still air. The results of this calculation showed that thermal equilib-rium occurred for the Duke supplied RTD cable after five hours, thus allowing an additional hour of exposure with the whole cable at the LOCA qualification temperature. For the core exit thermocouple cable, as previously justified, similar exposure would permit at least three hours at the LOCA qualification temperature. For the Westinghouse supplied RTD cable, the short duration of the test does not permit any conclusion to be drawn - the cable temperature after 20 minutes reached 70% of its total rise indicating a time constant in still air of approximately 16 min ~1 However, the justification for survivability of the Westinghouse supplied RTD cable is based on testing of this cable by TVA in a hydrogen burn environment as described in section 5 4.2.4.

It is concluded that for reasonable assumptions concerning the test environment of the cable during the LOCA qualification test, thermal

, equilibrium between the cable jacket and the environment is reached very l

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quickly. Therefore, the use of the LOCA qualification temperature as measured by a thermocouple close to the cable during the test is a reason-able standard by which to assess cable survivability. It is concluded that all of the necessary cables will survive hydrogen burning.

With regard to the instrumentation transmitters, the method used to assess survivability is described in Section 5.4.3. As noted, hydrogen burning has very little effect on the interior temperature of the transmitter. The total temperature rise, even if the environment of the transaitter were to include hydrogen burning (which is does not), would ,

be well below any relevant test temperature attained during qualification testing. If more massive equipment is considered, e.g., an air return fan, hydrogen burn events ca not transfer sufficient energy to massive equipment to have a significant effect on the temperature of the d equipment. The effect represents only a small fraction of the total temperature rise created by the environmental effects of the LOCA, including the large steam release. It is concluded that adequate consideration has been given to all aspects of equipment survivability

, and that appropriate action has been taken to ensure operability of all l

essential equipment.

Summary component evaluation worksheets (SCEWS) were not submitted as part of the McGuire NUREG-0588 submittal. Environmental conditions to which equipment was qualified are specified in the appropriate attachments in the NUREG-0588.

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9. The lists of equipment provided in Section 5.2 of the McGuire submittal do not include all essential equipment, e.g., isolation and FORV valves.

In this regard, provide a summary table, such as Table 2.2-1 in Att.achment 4 to TVA's October 1, 1981 submittal for Sequoyah, and justification for the survivability treatment of each item. As a minimum, add the following equipment to the list of the equipment required for the hydrogen burn event and provide the analysis to demonstrate the survivability of this equipment during the hydrogen burn event:

(a) Hydrogen Recombiner (b) Reactor Vessel Vent Valves (c) PORV and Block Valves

Response

In establishing the list of essential equipment, it was determined that equipment should be included if the answer to either one of the following questions is yes:

1. Is the equipment essential to achieve and maintain the reactor coolant system in a safe shutdown condition?

2. Is the equipment essential co maintain containment integrity?

The key word in each question is " essential", implying that the function could not be performed if that particular equipment were not available.

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In the case of the hydrogen recombiner, it was concluded that it is not essential to operate this equipment in order to maintain containment integrity. The design of the hydrogen recombiner is based on the requirements for hydrogen control within the design basis as defined in 10CFR50.44. The total hydrogen release involved is less than 10% of that considered in the beyond design basis sequence of events which led to the installation of the hydrogen mitigation system. The hydrogen mitigation system has been shown to be adequate to protect the containment from the effects of hydrogen for the beyond design basis event, by consuming hydrogen at rates far in excess of the capability of the hydrogen recombiners. The hydrogen recombiners do not make a contribution to the control of the beyond design basis hydrogen release event and therefore are not essential for the maintenance of containment integrity.

l c The second consideration is whether the equipment is essential to achieve l

and maintain safe shutdown and adequate core cooling. The essential function of any active component on the reactor coolant pressure boundary in this case is to assume and remain in a closed position so that the i integrity of the reactor coolant pressure boundary is maintained. For that reason, the reactor vessel head vent valves fail closed on loss of i

power. There are no essential functions of these valves which require the valves to be opened in oraer to promote natural circulation cooling.

In fact, there are several disadvantages associated with the operation of these valves, including reduction of system pressure, increasing the

release rate of steam and hydrogen into the containment, and interfering I

with reactor vessel level monitoring. In summary, it is concluded that l

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O the reactor vessel vent valves cannot be considered essential, that their use during a period of recovery of core cooling should be avoided, and it is therefore inappropriate to include them in equipment survivability consideration:i.

In previous analyses of equipment survivability, it was established that large, massive objects experienced only small temperature increases due to hydrogen burning. This is ode to the relatively small amount of energy available in a given volume when hydrogen combustion occurs at low concentrations. For example, instrumentation transmitters, with dimensions in inches and weights of a few pounds, experienced temperature increases of 20-30 F per hydrogen flame passing over them. Extending that analysis to the pressurizer PORV or electric hydrogen recombiner, both of which have dimensions measured in feet and weights on the order of tens or hundreds of pounds, it is concluded that the temperature rise due to hydrogen burning would be an insignificant fraction of the total temperature rise due to the loss of coolant accident. The probability of hydrogen combustion at the location of the pressurizer PORV in the top of the pressurizer cavity was examined. The combined action of the lower compartment igniters burning hydrogen closest to the source and the hydrogen skimmer fan drawing hydrogen out of the top of the cavity make it unlikely that hydrogen flames would-initiate in the pressurizer cavity. Hydrogen flames propagating into the cavity from the lower compartment would be burning at concentrations around 6% by volume and therefore would have much less effect than the 8.5% flames previously considered in equipment survivability analysis. It is concluded that O

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neither the pressurizer PORV nor the electric hydrogen recombiner, even V

if considered to be essential equipment, would be adversely affected by the deliberate ignition of hydrogen in cantainment.

10. In response to Question 10 of the February 10, 1982, NRC Request for Information, Duke justified exclusion of the T B scenario from B2 consideration on the basis of offsite and onsite power reliability. A description of the utility grid system was provided to support the Duke position. To further justify the reliability of AC power, provide the probability values for the station blackout (total loss of AC power) at McGuire Nuclear Station lasting longer than 90 minutes, but shorter than 140 minutes.

Response

Duke Power has not developed a quantitative assessment of the probability of loss of all AC power at McGuire. The basis of our previous response was that the compounding of low probability events required to produce the accident sequence specified by NRC resulted in a probability much less than that of the events which lead to the degraded core and production of hydrogen in the first place. Accordingly, additional failures beyond those which lead to the degraded core are not considered in assessing the performance of the hydrogen mitigation system. The quantitative measure of probability in this case is not necessary to justify exclusion from consideration of a transient involving loss of all AC power.

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11. Provide a description of the procedural instructions for turning on and turning off the hydrogen igniters to clarify and supplement information provided in Section 3.7.1 of the McGuire submittal. Your description should include the following items:

(a) the procedures in which operator actions are required to turn on the igniters.

(b) the conditions which would require the actions (e.g., any diagnosed 1 loss of coolant, any safety injection actuation signal, only if inadequate core cooling is diagnosed).

(c) the procedures in which operator actions are required to turn off or verify the igniters are turned off.

(d) the conditions which would require those actions (e.g., upon reaching cold shutdown, return to power operation, diagnosed spurious safety injection).

Response

Operator action to turn on the igniters is contained in Energency Procedure 1/A/5000/02, " Loss of Reactor Coolant". This emergency procedure is initiated by the operator on actuation of Safety injection if the following conditions indicate that a loss of reactor coolant has occurred: rapid decrease in pressurizer level and pressure, increase in containment sump level, change in containment atmosphere thermodynamic state, containment high activity alarms, etc. Based upon analysis, the time of actuation will be approximately one hour prior to the first release of hydrogen to containment under S2D conditions.

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The operator is directed to take the following action with respect to deenergizing the igniters. If any indication of inadequate core cooling exists or has existed, it is required that the hydrogen igniters remain energized for 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> after the establishment of adequate core cooling.

If adequate core cooling has always existed for the duration of the transient, the hydrogen igniters are deenergized when containment pressure drops below 0.25 psig. If at any subsequent time indications of inadequate cere cooling exist, the hydrogen igniters are reenergized.

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