ML20210E511

Rev 13 to Pages from Analysis of Hydrogen Control Measures at McGuire Nuclear Station, Responding to Questions Transmitted by Youngblood

ML20210E511
Person / Time
Site:	Mcguire, Catawba, McGuire, 05000000
Issue date:	03/25/1986
From:	DUKE POWER CO.
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ML20210E509	List: ML20210E506 ML20210E511
References
NUDOCS 8603270369
Download: ML20210E511 (9)

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t Response to questions transmitted by letter from NRC (B. J.

Youngblood) to Duke Power Company (H. B. Tucker) dated December 17, 1985

1. Provide details of the fan response calculation to support the statement in Duke's March 29, 1985 letter that burns occurring at hydrogen concentrations of 6.5% or less do not create sufficient pressure differential across the fans to speed 1 them up to synchronous speed. Include in your response a description of the hydrogen combustion assumptions (e.g., flame I speed, burn completion, compartment. venting, containment spray, l heat removal), and the fan and electrical system models and j assumptions.

Response: To determine a best estimate tire-differential pressure loadino curve for the air return fana, a crASIX run was made using restart information from the containment analysis reported in Reference 1. Starti".g from containment conditions which are typical of those found by analysis during the time that hydrogen is being released and burned (containment preocure approximstely 22 psia), an upper compartment global burn was forced to occur by lowering the ignition limits in the upper compartment until ignition occurred. The time history of the pressure difference between the upper compartment and the deadended compartment, which represents the fan differential pressure was noted. The following are the specific characteristics of this analysis:

a. ~ Hydrogen ignition occurred when the. hydrogen concentration was slightly above 6.5% by volume in the upper compartment.

b. Ice condenser top deck doors were greater than 20% open throughout the transient; ice condenser immediate deck doors were closed due to the reverse differential pressure. Venting from the upper plenum into the ice bed region occurred through the 20 square foot bypass opening around the intermediate deck.

c. The heat and mass removal mechanisms operating in the upper compartment - the containment spray, air return fans, and upper compartment-heat sinks - were operable.

d. The burn time for the upper compartment was 13 seconds, corresponding to a flame speed of approximately 8 feet /second.

Under these conditions, a differential pressure transient with a peak of 5.1 psid and a duration of 16 seconds was imposed on the air return fans. A time plot of this transient is shown in Figure 1.

There is significant conservatism built into this analysis of an upper compartment burn transient. As has been shown in the NTS test series, hydrogen burning in a turbulent atmosphere created by spray can'take place at concentrations as low as 5% by volume.

The NTS test series also showed that hydrogen burning in a transient injection mode, such as at the exit of the ice 8603270369 860325 7.0 - 152 Rev 13 PDR ADOCK 05000369 E PDR

c6ndIn2cr, is likaly to ba by diffusion flames if the exit .

velocity is low; velocities which created diffusion flames in NTS are comparable to those at the exit of the ice condenser.

Therefore, even if one accepts the premise that some type of inhibiting mechanism will move the burning regime from the lower.

compartment to the upper plenum and upper compartment, it is very likely that hydrogen burning will be by diffusion flames at the exit of the ice bed and propagating into the region immediately above the top deck doors, and global accumulations with accompanying whole compartment deflagrations will not occur.

Using the fan blade geometry, this differential pressure was converted into an imposed torque on the air return fan blades.

This torque curve has the same shape as Figure 1 and has a peak of 2200 ft-Ibf.

To determine the effect of this torque on the air return fan speed, the fan and motor were modeled using the information in Reference 2. The general differential equation which describes the transient rotation of the induction machine is J dtY+ Dc/ + 0TeI - T" (1) where cd = operating angular speed 7 = rotating moment of inertia D = total windage coefficient (losses)

T' = shaf t mechanical torque Y = number of phases Te4 = per phase torque of electrical origin Slip (S) is defined as ldsyn - b f l

$= (dsyn (2) 4 where C4g= synchronous speed of the~ machine l For small values of slip in the operating range of steady state, the per phase torque of electrical origin is proportional'to the slip _

Yep = ~ K b where the negative sign indicates the torque direction. ' Equation (1) may be rewritten in terms of' slip rather than angular speed, then split into two equations - one which represents the steady state operating condition and one which represents the' transient deviation from that steady state condition. The transient equation may then be solved to yield:

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AT* -4 AS =

-( l - e 7] I

(if) D + JK where T is the induction machine time constant.

q _~ J g+ n9K (5)'

w n = number of pole pairs in the machine u) = excitation angular frequency In plugging actual numDers into this analysis, the windage coefficient was neglected for conservatism. The other values needed for the analysis were obtained from the_ fan.and motor vendor. The time constant of the fan / motor was found to be very short, so it was assumed that the machine follows the imposed differential pressure transient exactly. The calcuJated peak speed of the fan is 1308 rpm corresponding to a value of slip of

-0.09. Thus the induction motor attached to the fan will become an induction generator.

To find the electrical transient associated with motoring the fan, it was assumed that the running current / torque curve is symmetric about the synchronous speed. This assumption is valid for small values of slip, as in this case, and was verified to be reasonable in discussions with the motor vendor. The current through the motor breaker as a function of time is shown in Figure 2.

There are a number of'significant'conservatisms in.the analysis

, which establish a degree of margin. Aerodynamic and frictional F forces (windage) which oppose an increase in motor speed have been ignored. The imposed differential pressure transient is assumed to be converted to torque on the fan with 100%

j efficiency. Finally, in the calculation of differential pressure across the fan, no allowance'was made ror the fan flow to increase beyond the maximum shown on the fan head / flow curve.

Clearly if the fan is to be speeded up by imposed torque, the flow through the fan must increase and upper compartment venting I take place even faster than in the CLASIX analysis. This final i conservatism would tend to reduce the imposed differential  !

pressure to something less than 5.1 paid. l Given an electrical transient consisting of a current reversal l for several seconds, then a second reversal and return of the l fan / motor to normal operation, the effect on the fan motor j circuit breaker was assessed to determine whether either the l

' direction or magnitude of the transient would cause it to trip.

Current reversal in itself will not cause a trip of the breaker.

It is necessary then to compare the current versus time curve 7.0 - 154 Rev 13

o 4

(Figure 2) with the trip current versus time curve for the thermal overload devices on the specific circuit breaker installed for the air return fan. When this was done, it was found that the peak current shown in Figure 2 would have to exist for at least ten seconds in order to cause the breaker to trip on thermal overload. The actual 10 second trip condition for this breaker corresponds to 10-time locked rotor current, or 449 amps, which is 7% higher than the calculated peak current. We conclude that an upper compartment burn will not cause a trip of the air return fan, particularly when the conservatism of the analysis used to arrive at that conclusion.is considered. It is therefore unnecessary to assess further the operator action required to mitigate the effects of fan trip, however a general discussion of this point'has already been included in Reference 1, Revisions 8 and 10.

To summarize the margin in this analysis:

a. The CLASIX analysis of an upper compartment burn assumed that the burn was global at 6.5%. Data from the NTS test series shows that global burning in the presence of spray will occur at concentrations as low as 5%, therefore a compartmental concentration of 6.5% will never occur. The NTS data also shows that a more likely burning scenario for hyurogen which does not burn in the lower compartment is that it will burn as diffusion flames in the upper plenum of the ice condenser or in the upper compartment in the region immediately above the ice condenser.

The survivability of equipment in the upper plenum of the ice condenser for sustained burning has been shown in reference 1.

There is no vital equipment in the upper compartment which would be affected by diffusion flames.

2. Using an appropriate modelling technique, provide a quantitative assessment of the pressure loading on the ice condenser doors created by hydrogen combustion in a) the upper plenum and b) the upper compartment. Describe and justify the assumed or calculated door positions. Provide an evaluation of the ultimate capability of the ice condenser doors to withstand reverse differential pressures. Discuss the probable failure modes and the consequences of such failures.

Response

The top deck doors of the ice condenser consist of flexible blanket insulation supported on grating. During the initial blowdown phase of a LOCA, these blankets are displaced from.their normal positions and do not again form a seal. Westinghouse estimates that the area of the top deck will be at least 20% open at all times following this initial displacement of the top deck doors. There are no failure modes associated with the top deck doors because this open area assures that a reverse differential pressure loading cannot occur. The displacement of the top deck doors also ensures that the upper plenum burns vent to the upper compartment, so that burning in the upper plenum creates very 7.0 - 155 Rev 13

e I

4 small pressure differentials across the intermediate deck' doors.

The limiting condition for the intermediate deck doors is therefore the global upper compartment burn. For analysis purposes, the same upper' compartment burn which was used in.the discussion of air return fan behavior was used to assess what happens to the intermediate deck doors if a global upper compartment burn occurs.

There is a permanent bypass area associated with the intermediate deck doors of 20 square feet. This bypass tends to limit the imposed differential pressure by allowing venting around the doors. For conservatism, this bypass area was neglected and the full differential pressure, 5.1 psid, was assumed to be imposed on the intermediate deck. An analyJis was performed of the capability of the intermediate deck to withstand differential

. pressure loading. The vendor drawings of this area were l consulted, and nominal material properties were used. The l reverse differential capability of the intermediate deck was l found to be 6 psid which agrees with the capability calculated  ;

independently by TVA and previously reported to NRC. Similarly )

the lower inlet. door capability previously reported by TVA as 7 psid applies to Catawba because the structural components of the lower inlet doors are identical between Sequoyah-and McGuire/ Catawba.

We note in Reference 3 that NRC is concerned about the different numbers reported by the utilities in assessing the structural capability of the ice condenser doors. Some structural changes were made by Westinghouse in these areas between D. C.. Cook and McGuire/Sequoyah which may account for some differences.

Structural analysis methods may also make some difference; although we have great confidence in our independent verification of capabilities-previously reported by TVA. Differences in the loading created by hydrogen burning may be attributed to .

assumptions made by the different utilities in performing CLASIX analysis or in interpreting the results. This most recent work we are reporting is considered to be the best because it uses the latest test data available from NTS to establish a conservative set of hydrogen burning parameters. Previous analyses which assumed global burning at 8% or 8.5% (as we did in the~ analysis reported in Reference 1) would necessarily have resulted in much larger imposed transient differential pressures.

We have also reexamined the Sandia analysis in which they us MARCH and HECTR to predict.that upper compartment global burn 2ng is a likely outcome of a LOCA in which core degradation occurs.

As we discussed in our previous submittal, the inhibition of lower compartment burning due to partial steam inerting and the consequential large number of upper plenum and upper compartment burns as predicted by Sandia is contradicted.by work we have done using the MAAP code. We note.that these Sandia results have also been contradicted by the results reported by AEP and their contractors in Reference 4. When one examines the graphs of hydrogen and steam release rates reported by AEP, one notes a substantially smaller rate of flow of steam into the lower 7.0 - 156 Rev 13 i

- - ~ , - - - , _ . - - , _ - _ - - _ - - _ . _ .- .. _- - .. .-. ,.. . _ _ - . . - _. - -,_ -._ _ ,, ~ _ . _ -,.-

O I

containment than that reported by Sandia in Reference 5 during the time that hydrogen is being released. This apparent' error by-Sandia is particularly pronounced for the S2D sequence as seen in the comparison of Figure 14 of Enclosure 2 of Reference 4.with

~

Figure 3-3 of Reference 5. We are led to the conclusion that the Sandia results have incorrectly predicted steam release rates which accompany hydrogen releases to the extent that their analysis incorrectly predicts hydrogen concentrations below the ignition level in the lower compartment-during the duration of hydrogen release. Thus hydrogen burning has been forced to occur in regions other than the lower compartment, and as a consequence upper compartment burning occurs frequently. This upper compartment burning would be eliminated if Sandia had correctly

predicted the steam release rates in MARCH.and used a realistic ignition level in the lower compartment, such as 7.04.

If imposed differential pressures were to cause some blockage of' ice condenser doors, this is not a problem for the following reasons. The ice condenser doors are sized to pass the very large flows of air which result from the initial blowdown following a LOCA. Once most of the air in the lower compartment has been replaced by steam, the flow through the ice condenser is reduced significantly because most of the steam condenses on the ice. Flow through the intermediate deck doors then consists mainly of the air being recirculated by the air return fans (60,000 to 80,000 scfm) which is very small compared to the 1000 to 2000 square feet of flow area available. Even if this area were reduced substantially by a partial failure of the' ice condenser doors late in a transient (the time that hydrogen is being released), there would be sufficient flow area to ensure that no adverse effects would occur because~so little area is needed to ensure operability of the heat removal mechanisms in containment, and the energy being released to containment is considerably less than that of the earlier phases of blowdown.

To summarize our response to the two questions contained in the NRC letter:

1. We have performed an analysis of the response of the containment to an upper compartment global hydrogen burn at 6.5%

hydrogen by volume'. This analysis contains a number of conservatisms such as the hydrogen volume percentage, the assumed method of burning (as global deflagrations), the assumed burnout time for the compartment, limitation of the venting rate associated with the air return fans, and the position of ice condenser doors. The resulting differential pressure of 5.1 psid is considered to be an upper bound.

2. This conservative upper bound case-has been examined in great detail in order to assess quantitatively the possible effects on i containment structure and components. . It has been found that the I containment air return fan motors will undergo motoring on the i 1

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1

r 3

bus end cet na gsnsratore for a chart pariod of time, but the time duration and magnitude of current are not sufficient to cause tripping of_the circuit breakers associated with the fans.

Neither is the fan overspeed sufficient to cause any threat to the structural integrity of the fan assembly. Following the transient, the fans will resume normal operation.

3. An independent assessment of the capability of the ice condenser doors to withstand differential pressure has been performed which has verified numbers previously reported to NRC

.by TVA. This capability is greater than the worst imposed differential pressure loading, therefore the doors will not be adversely affected by hydrogen burning.

4. Further contradiction of previously reported Sandia results which show a preference for upper compartment burning and partial steam inerting of the lower compartment can be found in work reportec by Reference 4.

References:

1. Duke Power Company, An Analysis of Hydrocen Control Measures at McGuire Nuclear Station, complete through Revision 12, March 29, 1985.

2. Meisel, J., Principles of Electromechanical Enerav Conversion, McGraw-Hill, New York, NY, 1966, pp. 536-619.

3. " Safety Evaluation Report related to the operation of Catawba Nuclear Station, Units 1 and 2," NUREG-0954, Supplement No. 5, February, 1986.

4. NRC memo from D. L. Wigginton dated December 16, 1985, reporting the summary of a meeting held between NRC and IMEC on December 5, 1985.

5. Camp, et. al., MARCH-HECTR Analysis of Selected Accidents in an Ice Condenser Containment, NUREG/CR-3912, dated December, 1984.

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5- -

FIGURE 1 AIR RETURN FAN DIFFERENTIAL PRESSURE VS. TIE FOR GLOBAL 4,. HYDROGEN BURN IN UPPER CorARTENT 4

S H2 CONCENTRATION - 6.5%

IC FLAE SPEED -

8 ft/sec BURN COWLETION - 60%

3--

E W

W t:'

y 2--

i..

2 4 6 10 l2 l4 )6 TIE AFTER IGNITION (SECONDS) 7.0-159

500 --

FIGURE 2 ,

AIR RETURN FAN MDTOR CURRENT

~ 400 --

VS. TIE FOR GLOBAL UPPER COWARTENT BURN ASSUWTIONS:

a. D/P ACROSS FAN CONVERTS TO TORQUE 300 -- AT 1001 EFFICIENCY f b. WINDAGE LOSSES NEGLECTED

c. MOTOR CURRENT CURVE SYleETRIC ABOUT h"

• SYNCHRONOUS SPEED FOR SMALL VALVE!,

OF SLIP W

g 200 --

i 8 5

E i

100 1

i l  ;  ;  ;  ;  ! .

2 4 6 8 10 12 14 16 18 TIE AFTER IGNITION (SECONDS) 4 4'

, 7.0-160

-_ _ _ _ _ - _ - _ _ _ _ _ _ _ - _ _ _ . _ _ _ _ _ _ - - _ _ _ -_.