ML20142A526

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4 to Updated Final Safety Analysis Report, Appendix 6A, Negative Pressure Design Evaluation
ML20142A526
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APPENDIX 6A

~ NEGATIVE PRESSURE DESIGN EVALUATION 6A.l Primary Containment Negative Pressure Evaluation Analysis The primary containment has been designed for a negative pressure of 3 psig. The worst case for this consideration results from the inadvertent actuation of the drywell sprays. During such a transient, cold spray water is passed through the drywell atmosphere, resulting in a drop in gas region temperature and a corresponding drop in gas region pressure.

This condition has been analyzed for the Hope Creek Generating Station. A peak negative pressure of -2.83 psig is obtained in the drywell. The maximum drywell to suppression chamber pressure differential is 2.61 psid.

To determine the temporal pressure and temperature of the primary containment, the conservation equation of mass and energy, along with the state equations for steam and nitrogen {air} noncondensable, are written for the drywell and suppression chamber regions. A schematic of these regions is presented in Figure 6A-1. The various terms for the mass and energy transfer mechanisms are also presented in this figure. The system of equations that were solved are presented below.

~ 6a.2 Nomenclature A Flow area, f t2

  • c Flow coefficient (see equation (6A-43)).

Cv Flow coefficient.

E Energy content, BTU.

~ HCGS-UFSAR 6a-1 Revision 7 December 29, 1995

Standard acceleration due to gravity. 32.174 ft/s 2

  • h Specific enthalpy, BTU/lbm.

k Ratio of the principle specific heats.

K Loss coefficient, defined by the equation 144gc il.P M Mass, lbm.

Condensation rate on a water surface, lbm/s.

Mdrop Dropout rate of mist droplets from a vapor region, lbm/s.

Evaporation rate from a water surface, lbmjs.

p Pressure or partial pressure, psia.

R Gas constant, lbf-ft/lbm- 0 R.

Q Transferred energy, BTU.

T Temperature, °F.

Temperature of spray at inlet to drywell.

Tout Temperature of spray water at outlet from drywell.

Absolute temperature, 0 R.

t Time, s.

u Specific internal energy, BTU/lbm.

HCGS-UFSAR 6a-2 Revision 7 December 29, 1995

v Volume, ft:>.

GREEK SYMBOLS Factor in temperature equation (see text}.

Factor in temperature equation (see text).

Heat exchanger effectiveness.

Spray efficiency.

v Specific volume, ft 3/lbm.

p Density, lbm/ft 3

  • SUBSCRIPTS a Air.

AT Atmosphere.

D Drywell region.

f liquid.

g saturated vapor.

PV Purge valve flow (from the reactor building to the suppression pool vapor region).

s Suppression vessel liquid region (or the "suppression pool").

sv Suppression vessel vapor region.

  • HCGS-UFSAR 6a-3 Revision 7 December 29, 1995

sw Service water.

VB Vacuum breaker flow (from the suppression pool vapor region to the drywell region).

6A.3 Analytical Assumptions Dalton's Law of Partial Pressures is assumed to hold, so that the total pressure in a volume is equal to the sum of the partial pressures. The gas regions are assumed to be saturated, so that the partial pressure of the steam is always equal to its saturation pressure.

The spray water is assumed to enter the drywell at temperature Tm., and to immediately exit to the suppression pool liquid region at a temperature of T~*

These temperatures are related by the spray efficiency, which is defined as E ; T;n -Tout (GA-l)

Tin -TD This functional correlation is determined in the work of Reference 1 and is illustrated in Figure 6A-2.

The spray suction is assumed to be aligned to the suppression pool liquid region.

The drywell region is assumed to remain saturated, and any liquid that condenses as a consequence of this (M~) is assumed to mix with the spray flow (and so immediately enter the suppression pool liquid region at a temperature ofT~)*

The vacuum breaker flow of steam and air is determined from the pressures of the suppression vessel vapor region and the drywell region, using a flow coefficient that varies according to the valve position.

HCGS-UFSAR Ga-4 Revision 7 December 29, 1995

The suppression vessel vapor region is assumed to be saturated, and any liquid that condenses as a consequence of this {Mdrop} is assumed to form as a mist in the vapor region (so that the vapor region receives the latent heat) and then to fall into the liquid region at a temperature of Tw.

The interface between the liquid and vapor regions in the suppression vessel is assumed to transfer heat and mass through the simultaneous processes of evaporation (Me~p) and condensation (M~)*

The energy required for evaporation is assumed to come entirely from the liquid region, and the steam is assumed to be formed and transferred to the vapor region at a temperature of Ts*

The energy liberated by condensation is assumed to go entirely to the liquid region, and the water is assumed to be formed and transferred to the liquid region at a temperature of T~.

The purge valve flow of steam and air is determined from the pressures of the atmosphere and the suppression vessel vapor region, using a flow coefficient that varies according to the valve position.

Any heat supplied from the equipment and structures in the drywell region as it is cooled is neglected.

6A.4 Analytical Models 6A.4.1 Drywell Region A mass balance for the air in this region yields:

{6A-2)

  • HCGS-UFSAR 6a-5 Revision 7 December 29, 1995

A mass balance for the water yields:

d(M) 0 = (6A-3) dt Also, the assumption of saturation implies that (6A-4) so that d(M)D = VD dvgdT0 {6A-5) dt v;(TJ dTD dt From (6A-3) and {6A-5) we obtain the rate of condensation:

(6A-6)

A balance of the total energy in the region yields:

(6A-7}

6a-6 HCGS-UFSAR Revision 7 December 29, 1995

Also, (6A-8) so that (6A-9)

+ (M) _ _ + u (T ) d(Mg)D du gdTD g D dTD dt g D dt Hence the right hand side (RHS) of (6A-7) is equal to the RHS of (6A-9).

Substituting from (GA-5} and (6A-6), and re-arranging gives (6A-10) where du dug VD dv

<<0 = ( M )0 - + ( M ) - +

a DdT, g DdT 2 g(hiT )-ug (TIY')

dT f' out (6A-ll}

D D vg(To) D and

{6A-12)

  • HCGS-UFSAR 6a-7 Revision 7 December 29, 1995

6A.4.2 Suppression Vessel Vapor Region Air mass balance:

d(M)sv (6A-l3}

dt steam mass balance:

{6A-l4}

Also (6A-l5}

so that (6A-16)

Combining (6A-14) and (6A-16), and re-arranging gives:

dTsv (6A-17)

= <<dropdt + pdrop HCGS-UFSAR 6a-8 Revision 7 December 29, 1995

where adrop = (6A-18) and (6A-19)

The energy balance for this region is d;;v = (QPV + MevaA<Ts)l.n (6A-20)

- (QVB + (Mcond)s/tg(Tsv) + MdrophfTsv)} our Also (6A-21) so that (6A-22) dug dT5 v + u (T )d(Mg)sy

+(M) - -

g sv dTsv dt g sv dt

  • HCGS-UFSAR 6a-9 Revision 7 December 29, 1995

Combining (6A-20) and {6A-22), substituting from (6A-16) and (6A-17), andre-arranging gives, after some manipulation (6A-23) where (6A-24) and (6A-25) 6A. 4. 3 Suppression Vessel Liquid Regio.n The mass balance is

~S = (<Mcond)sv + (Mco.)D + Mspray + Mdrop}.,. - (Mevop + Mspray)_

(6A-26)

= (Mcond)SV + (Mcond)D + Mdrop - Mevap HCGS-UFSAR 6a-10 Revision 7 December 29, 1995

The volume of the region is

{6A-27) so that (6A-28}

and, because the liquid and vapor regions share a constant volume, dVsv (6A-29)

=

dt The energy balance is

  • {6A-30}

Also (6A-31) so

  • HCGS-UFSAR 6a-ll Revision 7 December 29, 1995

dEs dt du dTs

= M s dTs dt

+

dM8 ufTs)-

I' dt (6A-32)

Combining (6A-30) and (6A-32} gives (6A-33) where (6A-34) and (6A-35) 6A.4.4 Evaporation and Condensation The evaporation and condensation rates in the suppression vessel are modelled using the kinetic theory of condensation from Reference 2. This results in the following expressions HCGS-UFSAR 6a-12 Revision 7 December 29, 1995

  • (Mcond)SV =

144 f'A cond

~ gc (Pg)SV 2 1tR ~

g yTsv (6A-36)

(6A-37) where (6A-38) r -w {i [1+erj{w)] - e -w 2

w = -vg(Ts) {M~vap + (Mcond>sv) (6A-39)

Acond /2 8c Rg r;v and the erf function is defined as J

w erj{w) =2- e _.,.2d

~ z (6A-40)

/io

  • HCGS-UFSAR 6a-13 Revision 7 December 29, 1995

6A.4.5 Vacuum Breaker and Purge valve Flows The flow coefficients of these flow paths depend on the degree to which the valves are open. This, in turn, depends on the differential pressure across them, and their opening characteristics.

Given the flow coefficient for the vacuum breaker, the flow rate from the suppression vessel vapor region to the drywell is calculated using the non-choked adiabatic flow equations for a perfect gas (Ref 3) as follows.

(6A-41)

(Compared to the equation in Ref 3 the denominator, which has a value less than one, has been neglected. This results in conservatively lower flow rates).

The flows are never expected to become choked, because choked flow requires a pressure ratio of less than about 0.57 (Ref 3). The smallest pressure ratios expected are of order (14.7-3}/14.7 = 0.79. Despite this expectation, a check has been included in the coding to make sure that choked conditions are never calculated to exist.

For a small pressure drop, equation (6A-41) can be manipulated into the form (6A-42)

HCGS-UFSAR 6a-14 Revision 7 December 29, 1995

which serves to define the flow coefficient CVB. In terms of the usual loss coefficient K, it can be seen that 1

c VB = {6A-43)

JKYB The mass fractions of air and steam in this flow are assumed to be the same as the mass fractions in the suppression vessel vapor region.

The energy transferred by this flow is (6A-44)

Exactly the same equations are used for the flow from the atmosphere to the suppression vessel through the purge valves. The mass fraction of steam in the atmosphere is required, and this is obtained from the specified value for the relative humidity, which is defined as (6A-45)

For any given volume of the atmosphere, the mass of vapor in that volume is

{6A-46)

  • HCGS-UFSAR 6a-15 Revision 7 December 29, 1995

Hence the mass fraction of steam is (Mg)AT ::: RH pgCTAT) {6A-47)

MAT PAT 6A.4.6 RHR Heat Exchanger The water for the spray is drawn from the suppression vessel pool and passed through a heat exchanger. The spray temperature is modelled as

. CMspray ~ M' Tspray = TS - S~ mm sw~ (T S

- T '

~

(6A-48)

Msproy .

where ~ is a specified efficiency for the heat exchanger.

6A.5 Solution Technique Equations (6A-2), (6A-10), {6A-13}, (6A-23), (6A-26), (6A-28), and (6A-33) specify the derivatives of the mass of air in the drywell, the temperature of the drywell, the mass of air in the suppression vessel, the temperature of the vapor in the suppression vessel, the mass of water in the suppression vessel, and the temperature of the water in the suppression vessel. Together, these quantities completely define the conditions in the containment. Given a suitable set of initial conditions, these differential equations can be integrated to give a transient solution.

A fourth order Runge-Kutta algorithm was used to perform the integration

  • HCGS-UFSAR 6a-16 Revision 7 December 29, 1995

6A.6 Assumptions and Initial Conditions The concern is that the primary containment pressure will become too low ..

Therefore the following assumptions, which tend to reduce the containment pressure, are conservatisms.

  • The transfer of sensible heat from the equipment and structures to the drywell gas region is neglected, thereby reducing the energy (and so pressure) in this region.
  • Re-evaporation of the condensed drywell steam is disallowed, thereby reducing the mass (and so pressure) in this region.
  • A large volume for the drywell region is maintained by transferring condensed steam mass directly to the suppression pool.
  • The cold leg of the residual heat removal (RHR) heat exchanger is from the hot leg of the Safety Auxiliaries Cooling System. Therefore, the minimum coolant temperature of 65 °F can be higher than the suppression pool temperature. Since it is conservative to neglect any possible heating of the containment spray due to the SACS flow, the RHR heat exchanger effectiveness is set to zero for this analysis .

Two sets of initial conditions will be examined. One set will be the nominal operating conditions of the containment. The other will be chosen to maximize the depressurization transient.

The presence of any noncondensables in the drywell tends to hold up the depressurization of this region following spray actuation. Thus, to maximize the depressurization, a condition is postulated wherein a small break occurs within the drywell, serving to pressurize this region and drive all the noncondensables to the suppression chamber gas space. This sets the initial pressure distribution (and, along with the assumptions regarding saturated conditions for

  • HCGS-UFSAR 6a-17 Revision 7 December 29, 1995

the steam phase, the temperature distribution) for all three regions: drywell, suppression chamber gas region, and suppression pool.

These conditions will be referred to as "post-break" conditions.

The following sections present both sets of initial conditions, and of the valve properties used. The results are presented in Table 6A-1.

6A.6.1 Nominal Conditions At nominal conditions the primary containment is assumed to be at 14.2 psia, with a drywell temperature of 150°F and a suppression vessel temperature of 60°F. The 3

drywell volume is 169,000 ft , and the water in the suppression vessel is assumed to occupy 122,000 ft of the total 255,500 ft 3 , leaving 133,500 ft 3 for the steam.

3 6A.6.2 Post-break Initial Conditions It is postulated that a small break has driven all of the nitrogen from the drywell into the suppression vessel, which remains at 60°F.

6A.6.3 Valve Properties Each suppression chamber to Reactor Building vacuum breaker is installed in series with a butterfly purge valve. Since the purge valves have a higher opening pressure, and take longer to open, the flow will be controlled by the purge valves and not the vacuum breakers. The following table compares these valve parameters to illustrate this.

HCGS-UFSAR 6a-18 Revision 7 December 29, 1995

Vacuum breaker Purge valve Opening pressure (psid} 0.2 {assumed) 0.25 Opening time (s} 0.1 (assumed) 15 Once the opening pressure for the purge valve is exceeded, the valve disc angle is assumed to increase linearly to 90°. Once the pressure differential falls below 0.25 psid, the purge valve is assumed take 5 seconds to close. Closing times in excess of 5 seconds will not adversely affect the calculated peak differential pressures.

6A.7 Determination of Worst Case The worst case of inadvertent spray actuation was determined by running a suite of cases based on a single valve failure. The suite included every combination of the following conditions.

  • Failure of a single suppression chamber to drywell vacuum breaker valve or failure of a single suppression chamber to Reactor Building purge valve.

(The failed valves are assumed to remain closed.}

  • Activation of a single spray loop or activation of two spray loops.
  • Nominal initial conditions or post-break initial conditions .

As a check that the time integration was not introducing errors, one transient was repeated with a timestep of 0.001 sec instead of 0.01 sec. The results were identical to the precision printed.

A tabulation of the main cases analyzed, and their maximum pressure differences, is presented in Table 6A-2 *

  • HCGS-UFSAR 6a-19 Revision 7 December 29, 1995

The maximum negative drywell pressure achieved is predicted to be -2.83 psig.

This is achieved in the case where two sprays are activated from nominal conditions, and one purge valve fails shut. Some graphical results from this case are presented in Figure 6A-3 and Figure 6A-4.

The maximum pressure differential between the drywell and the suppression vessel is predicted to be 2.61 psid. This is achieved in the case where two sprays are activated from post-break conditions, and one vacuum breaker fails shut. Some graphical results from this case are presented in Figure 6A-5 and Figure 6A-6.

6A.8 Conclusions The conservative calculation methods presented here predict that the drywell pressure will not reach its design limit of -3 psig.

They also indicate that the drywell to suppression vessel pressure differential will not reach the design limit of 3 psid.

6A.9 References 6A-1 Takashi Tagami, "Interim Report on Safety Assessment and Facility Establishment {SAFE) Project,' February 28, 1966, Hitachi Ltd., Tokyo, Japan.

6A-2 Donal J. Wilhelm, "Condensation of Metal vapor-Mercury and the Kinetic Theory of the Condensation", ANL-6984, October 1964.

6A-3 "Marks' Standard Handbook for Mechanical Engineers .. , Ninth Edition, Eugene A. Avallone and Theodore Baumeister I I I .

HCGS-UFSAR 6a-20 Revision 7 December 29, 1995

TABLE 6A-1 Initial and Boundary Conditions Nominal Post-break Drywell Volume (ft3) 169000 Temperature (F) 150.000 244.257 Pressure (psia) 14.200 26.952 Steam partial pressure (psia} 3.719 26.952 Nitrogen partial pressure (psia) 10.481 0.000 Nitrogen mass (lb) 7578.835 0.000 Spray rate (gpmjtrain) 10000 Spray rate (lbjs-train) 1391 suppression chamber Vapor volume (ft3} 133500 Liquid volume (ft3} 122000 Pressure (psia) 14.200 25.509 Temperature (F) 60.000 60.000 Steam partial pressure (psia) 0.258 0.258 Nitrogen partial pressure (psia) 13.942 25.251 Nitrogen mass (lb} 9343.573 16922.408 Liquid specific volume {ft3/lb) 0.016 Depth of spargers (ft} 3.330 Pressure to force water out of spargers (psi} 1.443 Suppression pool free surface area (ft2) 10710 Page 1 of 2 HCGS-UFSAR Revision 7 December 29, 1995

TABLE 6A-1 {Cont)

Suppression Chamber to Drywell Vacuum Breakers Number of valve assemblies 7 or 8 Flow area per assembly (ft2) 1.85 Flow coefficient 0.63 Assumed lifting pressure (psid) 0.2 Assumed opening and closing time {s} 0.1 RHR System - Drywell Spray Mode Service water flow rate (gpm/train) 9000 Service water flow rate (lb/s-train) 1251 Heat exchanger effectiveness 0 Number of trains 1 or 2 Suppression chamber to Reactor Building Purge Valves Flow area per assembly (ft2} 1.85 Number of valve assemblies 1 or 2 Lifting pressure (psid) 0.25 Opening time (s} 15 Closing time (s) 5 Reactor Building Atmosphere Pressure (psia) 14.7 Temperature ( °F} 60 Relative humidity (%} 50 Page 2 of 2 HCGS-UFSAR Revision 7 December 29, 1995

TABLE 6A-2 Maximum Negative Pressure Inside Containment Largest Largest Pressure Depressurization in Difference between Drywell (psig} Suppression Chamber and Drywell (psid) 1 Hx train 7 Vacuum breakers -1.72 (post-break} 0.70 (post-break) 2 Purge valves -1.80 (from nominal} 0.28 (from nominal) 2 Hx train 7 vacuum breakers -2.59 (post-break} 2.61 (post-break) 2 Purge valves -2.60 ( from nomina 1 ) 0.58 {from nominal) 2 Hx train 8 vacuum breakers -2.79 (post-break) 2.05 {post-break) 1 Purge valves -2.83 (from nominal) 0 .. 42 ( from nominal) 1 Hx train 8 Vacuum breakers -2.27 {post-break) 0 .. 54 (post-break) 1 Purge valves -2 .. 34 (from nominal) 0 .. 21 (from nominal)

  • HCGS-UFSAR Page 1 of 1 Revision 7 December 29, 1995

TABLE 6A*3 COMPARISON OF SPRAY ACTUATION FOR SBA AND NORMAL OPERATION Normal Parameter SBA Operation Pressure, psia 14.696 14.696 Temperature, OF 163.3 150 SteamfNoncondensab1e Ratio 0.44 0.21

  • HCGS*UFSAR 1 of 1 Revision 0 April 11, 1988

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APPENDIX 6B

  • SUBCOMPARTMENT DIFFERENTIAL PREBSURE CONSIDERATIONS 6B.l INTRODUCTION Di,fferential pressure analyses were performed for the .reactor pressure ves*sel-(RPV) shield annulus and the drywell head region. The shield annulus analysis is';*done in-- two 'Parts. * 'trhe first- part is a mechanistic ana*l*ysis *to- determine the loads on *the R:PV and its supports, internals, and attached piping~ The second part; is a more conservat'i ve, nonmechanistic, analysis that is used- only to' check the adequacy of the structural design o*f the shield' owall and i 'bs supports. The RPV- shield annulus *'is 48 .-*95..;:fie*eit-high and 1.8'33**;.-*feet. wide at the top.'

68~1.1 Vesse1-Arinulus Pressurization The estimation of mass and energy release is 'based on the guide1ines set- 'forth in GE's Generic Annulus Pressurization Mass Energy Release Methodology (MFN178-78, transmitted as a letter from E. D.. Fuller, GE, to D. -F. Ross, NRC, on May 2, 1978), and Technical Description Annulus Pressurization Load Adequacy Evaluation-- (NED0-24548/78 NED 302). Table 6B-l presents the recircUlation out.let line mass and ene-rgy release data -estimated by applying the'- NED0-2*4'5'48 method of combining 'bldwdown data calculated from finite and instantaneous break opening time approaches ..

The mass flux and area, as £unctions o'f time, that are used for each side of the break are tabulated in Table 6B-2.. In addition, pertinent physical parameters used in the blowdown estimation are also listed in -the table. (Note that the nomenclature used here is the same as that in the previously mentioned GE documents). The break is postulated to occur at the nozzle safe end to pipe weld. Because the break is located within the shield wall

-* See Appendix 3C for* descriptions on the reactor asymmetric load evaluation

£or the MELLLA.conditions.

6B-1 HCGS-UFSAR Revision 15 October 27, 2006

penetration 50 percent of the blowdown is assumed to be released into the annulus with the remaining 50 percent to be vented to the containment atmosphere. Subsequent to this analysis, a flow diverter was added to the shield wall penetration, thus reducing the portion of the blowdown entering the annulus to 25 percent.

Figure 6B-1 shows the flow obstruction and its location relative to the break. The subcompartment pressures calculated with the flow diverter in place are substantially less than the originally calculated pressures presented here.

In addition to the analysis for the recirculation outlet line break in the annulus, similar analyses using the same methodology for blowdown rate estimation are performed for a postulated feedwater line break in the annulus. Table 6B-3 presents the mass and energy release rates generated by applying only the very conservative instantaneous break opening time method. No credit is taken for the feedwater pipe whip restraint to limit the break opening area. Also, it is conservatively assumed in the analyses l;

/ that the full blowdown is comple~ly released into the annulus. Mass flux, as a function of time and areas used for each side of the break, is presented in Table 6B-4. Some pertinent physical parameters are also noted in the table.

6B.l.2 Shield Wall Structural Adequacy For the shield wall analysis, a nonmechanistic break equal in area to a double ended rupture of a recirculation outlet line (4.1 square feet) is postulated. As an additional conservatism, the break is postulated to occur at three different locations in the annulus, one near the top (case A) , one near the middle (case B) , and one near the bottom (case C). The most severe loads on the shield wall and its supports resulted from cases B and C.

The mass and energy blowdown rates are calculated using the Moody Steady Slip Flow Model. For case B, this results in a constant mass release rate of 36,300 lbm/s at 514.24 Btu/lbm. For case C, the constant mass release rate is 37,000 lbm/s at 514.21 Btuflbm.

6B-2 HCGS-UFSAR Revision 0 April 11, 1988

6B.1.3 RHR Head Spray Line Break The head spray pipe has been removed. However/ the analysis of a postulated head spray line break discussed in Appendix 6B bounds the effects of an RPV head vent line break. Therefore, the discussion regarding a postulated head spray line break is still valid.

In considering the drywall head region, the maximum blowdown rate results from a break in the residual heat removal (RHR) head spray line. The blowdown mass and energy release rates for this line are calculated using Moody critical flow 2

of 2060 lbm/s-ft and an enthalpy of 1191.8 Btu/lbm. Table 6B-5 shows the blowdown rate for a 6-inch schedule 80S line break with an effective break area 2

of 0.181 ft . Since this is the largest line that could pressurize the drywell

  • head region/ this postulated break is chosen for analysis.

The RPV shield annulus differential pressure calculation was performed by the computer code COPDA, whose computational procedure and analytical techniques are described in Reference 6B-1. These adjusted pressures are combined with the other appropriate loads I e.g., seismic and jet impingement to develop design loads for the affected structures and components . Subcompartment.

venting is used to ensure that the differential pressures developed will remain within the structural capability of component walls.

6B.2 BIOLOGICAL SHIELD ANNULUS SUBCOMPARTMENT MODELING PROCEDURES AND ANALYSIS Two analyses of the pressure distribution around the reactor pressure vessel (RPV) after a feedwater line break and a recirculation outlet line break were performed. The general layout of the shield annulus is shown on Figure 6B-2.

Figures 6B-3a and 6B-3b are schematics of the RPV shield annulus model. The model consists of six major levels. Each level was subdivided into 12 30° segments to form a total of 72 nodes inside the annulus plus an additional node for the remainder of the drywell.

In general, the arrangement of the pipes in the annulus determines the most representative level division, since the pipes constitute the only significant flow restrictions . This 73-pressure node 6B-3 HCGS-UFSAR Revision 14 July 26, 2005

model is considered sufficiently detailed to conservatively predict the maximum pressures on the compartment structure. Therefore, a nodalization sensitivity study is not needed.

For the purpose of determining peak pressure in the reactor vessel shield annulus, all insulation was assumed to move flush against the biological shield wall while still maintaining its original thickness. The volume of the insulation was excluded from the net volume of each subcompartment, and the projected area of the insulation that blocks the venting path was also excluded from the free venting area used in the analysis.

The major vent path to the drywall atmosphere is through the top of the biological shield annulus. Venting through the shield wall is allowed only through the ventilation duct openings at the lower section of the shield wall.

Initial conditions used in this analysis were 15.45 psia, l35°F, and 30 percent relative humidity.

Tables 6B-6 and 6B-7 give the subcompartment volumes, flow areas, L/A ratio, and flow coefficients (including origins} used in this analysis.

The resultant pressure distributions are shown on Figure 6B-4 for the recirculation outlet line break, Figure 6B-5 for the feedwater line break, Figure 6B-6 for shield wall structure case B, and Figure 6B-7 for shield wall structure case c. Additionally, the load forcing functions, which include both peak and transient loadings on the RPV and the reactor shield wall, are presented on Figures 6B-8 and 6B-9 for the recirculation outlet line break, and on Figures 6B-lO and 6B-ll for the feedwater line break. This forcing function represents the time dependent resultant force on the structure, and originates from the vector sum of the product of compartment pressure and area for each of the 72 geometry nodes used to represent the surface.

6B-4 HCGS-UFSAR Revision o April ll, 1988

The locations of the center of each node are given in Table_ 6B-8.

The components of these nodal areas are calculated in the following manner:

(Ax) -RH (Sint1 1 - Sin8 2 ) .

(Ay) -RH (Coss - Coss ).

2 1 where:

R radius of the ith geometry node, in.

H height of the ith geometry node, in.

'1 starting angle (degrees) for ith geometry node

'2 ending angle (degrees) for ith geometry node Therefore, the force generated by a pressure, P, acting on a nodal area A has the following components:

(Fx) - p (Ax)

(Fy) - p (Ay)

The compartment pressure transients resulting from a break in the reactor shield annulus generate a nodal force distribution over exposed surfaces. The resultant of this nodal force distribution is presented on Figures 6B-8 through 6B-ll. There are no external moments generated by this pressure response. Any moments would result from the application of the external force distribution to a structural model. This would generate shear stresses (leading to internal moments) due to bending of the elements used to represent the structure as a result of the nonuniform load distribution. Further discussion of this result is contained in Section 3, where the application of these annulus pressurization results is described in detail .

  • HCGS-UFSAR 6B-5 Revision 0 April 11, 1988

6B.3 DRYWELL HEAD REGION SUBCOMPARTMENT ANALYSIS NOTE: The RHR head spray line has been removed from the RPV. However, this analysis is still valid as the postulated head spray line break bounds the effects of an RPV head vent line break in the drywell head region.

A pressure analysis of the drywell head region for a postulated head spray line break was performed. The effects of a 6-inch residual heat removal (RHR) head spray line break bound those of a 2-inch RPV head vent line, which is. the only other line that runs through the drywell head region.

Figure 6B -12 illustrates the basic arrangement of the head region. Venting from the head region is accomplished through ventilation openings as shown on Figur~. 6B-12 and a manhole opening.

To determine the peak pressure in the drywell head, all insulation was assumed to remain in place. Initial conditions of 15.45 psia, 135°F, and 30 percent relative humidity were used in this analysis.

The maximum pressure in the drywell head region is 26.7 psia and occurs approximately 1. 0 second after the head spray line break. Considering the containment pressure to be atmospheri9 (no drywell air displaced into the suppression chamber), a drywell head to containment pressure differential of 11.3 psid occurs at approximately 1.0 second after the break.

The manufacturer's design differential pressure {between the drywell head and containment region) for the water seal plate is conservatively defined as 20.0 paid. For the refueling bellows, the manufacturer's design differential pressure is defined as 15.0 psid.

6B-6 HCGS-UFSAR Revision 14 July 26, 2005

6B.4 REFERENCES 6B-l "Subcompartment Pressure Analyses," BN-TOP-4, Revision 1, November 1977, Bechtel Power Corporation, San Francisco, California.

6B-7 HCGS-UFSAR Revision 13 November 14, 2003

TABLE 6B-1 MASS AND ENERGY BLOWDOWN RATE .. RECIRCUlATION LINE BREAK Time, Mass Flow, Enthalpy, s lbm/s Btu/lbm 0.0 0.0 544 .. 5 0.00255 1201.0 544.5 0.00390 2402.0 544.5 0.00496 3603.0 544.5 0.00586 4804.0 544.5 0.00804 8407 . 0 544.5 0.00868 9608.0 544.5 0 .. 00924 10809.0 544.5 0.00980 12010.0 544.5 0.01180 16393.0 544.5 0.01380 20549.0 544.5 0.01580 24500.0 544.5 0.01780 28223.0 544.5 0.01880 30193.0 544.5 0.01910 30781.0 544.5 0.01911 11661.0 544.5 0.01980 12144.0 544 .. 5 0.02180 13545.0 544.5 0.02580 16344.0 544.5 0.02980 19128.0 544.5 0.03380 21863.0 544.5 0.03780 24474.0 544.5 0.04180 26877.0 544.5 0.04680 29366.0 544.5 0.05480 31295.0 544.5 0.05890 32060.0 544.5 5.0 32060.0 544.5 1 of 1 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6B-2 RECIRCUlATION OUTLET LINE BREAK BLOWOWN MASS FLUX TIME HISTORY Mass Flux, Effective Area, 2

Time, s lbm/s/ft ft 2 Vessel Side 0.0000 21260 0 0.00255 21260 0.02825 0.00390 21260 0.05649 0.00496 21260 0.08474 0.00586 21260 0.11298 0.00804 21260 0.19772 0.00868 21260 0.22596 0.00924 21260 0.25421 0.00980 21260 0.28246 0.01180 21260 0.38554 0.01380 21260 0.48328 0.01580 21260 0.57620 0.01780 21260 0.66376 0.01880 21260 0.71009 0.01910 21260 0.72392 0.01911 8055 0.72392 o.-o198o 8055 0.75382 0.02180 8055 0.84078 1.02580 8055 1.01453 1.02980 8055 1.18734 1.03380 8055 1.35711 1.03780 8055 1.51918 1.04180 8055 1.66834 1.04680 8055 1.-82284

  • HCGS-UFSAR 1 of 3 Revision 0 April 11, 1988

TABLE 6B-2 (Cont)

Mass Flux, Effective Area, Time, s lbm/s/ft2 ft 2

1.05480 8055 1.94258 0.05890 8055 1.99007 1.0 8055 3.538 Pipe Side 0.0000 21260 0 0.00255 21260 0.02825 0.00390 21260 0.05649 0.00496 21260 0.08474 0.00586 21260 0.11298 0.00804 21260 0.19772 0.00868 21260 0.22596 0.00924 21260 0.25421 0.00980 21260 0.28246 0.01180 21260 0.38554 0.01380 21260 0.48328 0.01580 21260 0.57620 0.01780 21260 0.66376 0.01880 21260 0.71009 0.01910 21260 0.72392 0.01911 8055 0.72392 0.01980 8055 0.75382 0.02180 8055 0.84078 0.02580 8055 1.01453 0.02980 8055 1.18734 0.03380 8055 1.35711 0.03780 8055 1.51918 0.04180 8055 1.66834 0.04680 8055 1.82284 2 of 3 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6B-2 (Cont)

Mass Flux, Effective Area, ft 2 2

Time 1 s lbm/s/ft 0.05480 8055 1.99007 0.05890 8055 0.442 1.0 8055 0.442 PERTINENT PHYSICAL PARAMETERS USED IN BLOWDOWN ESTIMATION Minimum cross-sectional area between vessel and break - 3.538 ft 2 Minimum cross-sectional area between pipe and break - 0.442 ft 2 Pipe inside diameter at break location- 2.158 ft Break area- 3.658 ft 2 Vessel pressure - 1053 psia Vessel temperature - 546°F Vessel enthalpy- 544.5 Btu/lbm Saturation pressure- 1011.75 psia Inventory length on vessel side - 3 ft Inventory length on pipe side - 51 ft 3

Inventory volume on vessel side - 11 ft 3

Inventory volume on pipe side - 187 ft

  • HCGS-UFSAR 3 of 3 Revision 0 April 11, 1988

TABLE 6B-3 REACTOR PRIMARY SYSTEM BLOWDOWN FLOW RATES AND FLUID ENTHALPY - FEEDWATER LINE BREAIK Mass Flow, Enthalpy, Time. s lbm/s Btu/lbm 0.0 0.0 403.5 0.0001 21180.0 403.5 0.0209 21180.0 403.5 0.0210 19480.0 403.5 5.0 19480.0 403 .. 5

  • HCGS-UFSAR 1 of 1 Revision 0 April 11, 1988

TABLE 6B-4 FEEDWATER LINE BREAK BLOWDOWN(l) MASS FLUX TIME HISTORY Mass Flux, Effective Flow Area, Time, s lbm/s/ft 2 ft 2

Vessel side 0.0 0.0 0.00 20000 0.3529 20000 0.3529 20000 0.2679 1.0 20000 0.2679 Supply pipe side 0.0 0.0 0.00 20000 0.7058 20000 0.7058 20000 0.7058 1.0 20000 0.7058 (1) Pertinent physical parameters used in blowdown estimation:

2 Maximum flow area - 1.0587 ft Vessel pressure - 1155.3 psia Vessel temperature - 425°F Saturation pressure - 327 psia Vessel enthalpy - 403.5 Btuflbm Feedwater specific volume- 0.01891 ft 3/lbm .

  • HCGS-UFSAR.

1 of 1 Revision 0 April 11, 1988

TABLE 6B-5 .

. MASS AND ENERGY RELEASE RATE HEAD SPRAY LINE BREAK(l} ( }

2 Steam Flow, Steam Enthalpy, Time, s lbm/s Btu/lbm 0.0 746 1191.8 0.0030 746 1191.8 0.0031 373 1191.8 10.0 373 1191.8

  • ( 1)

( 2)

Head spray line break is based on a 6-inch schedule 80S pipe with Moody 2

blowdown corresponding to 2060 lbrn/s-ft ; overall containment response is that of a 11 small break accident".

Head spray line has been removed; however, head spray line break is still I the bounding analysis for the drywell head region. *

  • HCGS-UFSAR 1 of 1 Revision 14 July 26 1 2005

TABLE 6B-6

~* HCGS - COMPARTMENT VOLUMES USED IN REACTOR VESSEL SHIELD ANNULUS SUBCOMPARTMENT ANALYSIS Compartment Number Designation Volume, ft 3 1 Vl 72 .. 0 2 V2 72 .. 0 3 V3 72 .. 0 4 V4 72 .. 0 5 V5 72 . 0 6 V6 72.0 7 V7 72.0 8 VB 72.0 9 V9 72.0 10 VlO 72.0 11 V11 72.0 12 V12 72.0 13 Vl3 95.0 14 V14 102.0 15 V15 101 .. 0 16 V16 102 .. 0

.17 V17 102 .. 0 18 V18 95 .. 0 19 V19 95.0 20 V20 102 .. 0 21 V21 101 .. 0 22 V22 102 .. 0 23 V23 102 .. 0 24 V24 95 .. 0 25 V25 79 .. 0 26 V26 77 .. 0 27 V27 77.0 28 V28 77.0 1 of 3 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6B-6 (Cont)

Compartment Number Designation Volume, ft 3 29 V29 77.0 30 V30 79.0 31 V31 79.0 32 V32 77.0 33 V33 77.0 34 V34 77.0 35 V35 77.0 36 V36 79.0 37 V37 80.0 38 V38 77.0 39 V39 80.0 40 V40 80.0 41 V41 77.0 42 V42 80.0 43 V43 80.0 44 V44 77.0 45 V45 80.0 46 V46 80.0 47 V47 77.0 48 V48 80.0 49 V49 103.0 50 vso 100.0 51 V51 103.0 52 V52 97.0 53 V53 95.0 54 V54 103.0 55 vss 103.0 56 V56 95.0 57 V57 97.0 58 V58 103.0 59 V59 100.0 2 of 3 HCGS*UFSAR Revision 0 April 11, 1988

TABLE 6B-6 (Cont)

Compartment 3

Number Designation Volume, ft 60 V60 103.0 61 V61 46.0 62 V62 46.0 63 V63 46.0 64 V64 46.0 65 V65 46.0 66 V66 46.0 67 V67 46 .. 0 68 V68 46.0 69 V69 46.0 70 V70 46.0 71 V71 46.0 72 V72 46.0 73 V73 169,000

  • HCGS-UFSAR 3 of 3 Revision 0 April 11, 1988

TABLE 6B-7 HCGS - FLOY AREA AND COEFFICIENTS USED IN REACTOR VESSEL SHIELD ANNULUS SUBCOMPARTMENT ANALYSIS Flow K L/A Flow Flow Paths 2 Description(!) ft-l Area, ft Factor Coefficient 1*2 13.19 0.1008 30° turn 0.481 0.953 1-12 13.19 0.1008 30° turn 0.481 0.953 2-3 13.19 0.1008 30° turn 0.481 0.953 3-4 13.19 0.1008 30° turn 0.481 0.953 4-5 13.19 0.1008 30° turn 0.481 0.953 5-6 13.19 0.1008 30° turn 0.481 0.953 6 .. 7 13.19 0.1008 30° turn 0.481 0.953 7-8 13.19 0.1008 30° turn 0.481 0.953 8-9 13.19 0.1008 30° turn 0.481 0.953 9-10 13.19 0.1008 30° turn 0.481 0.953 10*11 13.19 0.1008 30° turn 0.481 0.953 11-12 13.19 0.1008 30° turn 0.481 0.953 1-13 11.47 0.038 Friction 0.767 0.982 2-14 11.47 0.038 Friction 0.767 0.982 1 of 13 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6B-7 (Cont)

Flow K L/A Flow Flow Paths Area, ft 2 Factor Description(!) ft- 1 Coefficient 3*15 11.47 0.038 Friction 0.767 0.982 4-16 11.47 0.038 Friction 0.767 0.982 5-17 11.47 0.038 Friction 0.767 0.982 6*18 11.47 0.038 Friction 0.767 0.982 7*19 11.47 0.038 Friction 0.767 0.982 8-20 11.47 0.038 Friction 0.767 0.982 9-21 11.47 0.038 Friction 0.767 0.982 10-22 11.47 0.038 Friction 0.767 0.982 11-23 11.47 0.038 Friction 0.767 0.982 12-24 11.47 0.038 Friction 0.767 0.982 1-73 0.85 0.41 Contraction 1.274 0.842 3-73 0.85 0.41 Contraction 1.274 0.842 4-73 0.85 0.41 Contraction 1.274 0.842 6-73 0.85 0.41 Contraction 1.274 0.842 7-73 0.85 0.41 Contraction 1.274 0.842 9-73 0.85 0.41 Contraction 1.274 0.842 2 of 13 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6B-7 (Cont)

Flow K L/A Flow Flow Paths Area, ft 2 Factor Description(!) ft-l Coefficient 10-73 0.85 0.41 Contraction 1.274 0.842 12-73 0.85 0.41 Contraction 1.274 0.842 13-14 17.26 0.1008 30° turn 0.333 0.898 0.14 Around pipe 14-15 17.26 0.1008 30° turn 0.333 0.898 0.14 Around pipe 15-16 17.26 0.1008 30° turn 0.333 0.898 0.14 Around pipe 16-17 17.26 0.1008 30° turn 0.333 0.898 0.14 Around pipe 17-18 17.26 0.1008 30° turn 0.333 0.898 0.14 Around pipe 19-20 17.26 0.1008 30° turn 0.333 0.898 0.14 Around pipe 20-21 17.26 0.1008 30° turn 0.333 0.898 0.14 Around pipe 22-23 17.26 0.1008 30° turn 0.333 0.898 0.14 Around pipe 23-24 17.26 0.1008 30° turn 0.333 0.898 0.14 Around pipe 3 of 13 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6B-7 (Cont)

Flow K L/A Flow Flow Paths Area, ft 2 Factor Description(!) ft- 1 Coefficient 21-22 17.26 0.1008 30° turn 0.333 0.898 0.14 Around pipe 13-24 12.67 0.1008 30° turn 0.421 0.671 1.12 Around pipe 18-19 12.67 0 .. 1008 30° turn 0.421 0.671 1.12 Around pipe 13-25 6.434 1:3 Around pipe 1.177 0.624 0.27 Around pipe

  • 18-30 19-31 6.434 6.434 1.3 0.27 1.3 Around pipe Around pipe A~ound pipe 1.177 1.177 0.624 0.624 0.27 Around pipe 24-36 6.434 1.3 Around pipe 1.177 0.624 0.27 Around pipe 14-26 7.81 0.27 Around pipe 0 .. 968 0.806 0.27 Around pipe 16-28 7.81 0.27 Around pipe 0.968 0.806 0.27 Around pipe 17-29 7.81 0.27 Around pipe 0.968 0.806 0.27 Around pipe 4 of 13 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6B-7 (Cont)

Flow K L/A Flow Flow Paths Area, ft 2 Description(!) ft-l Factor Coefficient 20-32 7.81 0.27 Around pipe 0.968 0.806 0.27 Around pipe 22-34 7.81 0.27 Around pipe 0.968 0.806 0.27 Around pipe 23-35 7.81 0.27 Around pipe 0.968 0.806 0.27 Around pipe 15-27 6.43 0.27 Around pipe 0.992 0.737 0.27 Around pipe 0.30 Around pipe 21-33 6.43 0.27 Around pipe 0.992 0.737 0.27 Around pipe 0.30 Around pipe 25-36 14.62 0.105 30° turn 0.429 0.951 30-31 14.62 0.105 Jo* turn 0.429 0.951 25-26 12.77 0.105 3o* turn 0.451 0.889 0.16 Around pipe 26-27 12.77 0.105 30* turn 0.451 0.889 0.16 Around pipe 27-28 12.77 0.105 30* turn 0.451 0.889 0.16 Around pipe 5 of 13 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6B-7 (Cont)

Flow K L/A Flow Flow Paths Area, ft 2 Factor Description(!) ft- 1 Coefficient 28*29 12.77 0.105 30° turn 0.451 0.889 0.16 Around pipe 29-30 12.77 0.105 30° turn 0.451 0.889 0.16 Around pipe 31-32 12.77 0.105 30° turn 0.451 0.889 0.16 Around pipe 32-33 12.77 0.105 30° turn 0.451 0.889 0.16 Around pipe

  • 33-34 34-35 12.77 12.77 0.105 0.16 0.105 30° turn Around pipe 30° turn 0.451 0.451 0.889 0.889 0.16 Around pipe 35 .. 36 12.77 0.105 30° turn 0.451 0.889 0.16 Around pipe 25-37 11.47 0.06 Friction 0.695 0.971 26-38 11.47 0.06 _Friction 0.695 0.971 27-39 11.47 0.06 Friction 0.695 0.971 28-40 11.47 0.06 Friction 0.695 0.971 29-41 11.47 0.06 Friction 0.695 0.971 6 of 13 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6B-7 (Cont)

Flow K L/A Flow 2

Area, ft Factor Description{!) ft -1 Coefficient Flow Paths 30-42 11.47 0.06 Friction 0.695 0.971 31-43 11.47 0.06 Friction 0.695 0.971 32-44 11.47 0.06 Friction 0.695 0.971 33-45 11.47 0.06 Friction 0.695 0.971 34-46 11.47 0.006 Friction 0.695 0.971 35-47 11.47 0.06 Friction 0.695 0.971

  • 36-48 37-38 11.47 13.02 0.06 0.105 0.16 Friction 30° turn Around LPCI 0.695 0.442 0.971 0.886 0.01 Around instrument 41-42 13.02 0.105 30° turn 0.442 0.886

~.16 Around LPCI 0.01 Around instrument 43-44 13.02 0.105 30° turn 0.442 0.886 0.16 Around LPCI 0.01 Around instrument 47 .. 48 13.02 0.105 30° turn 0.442 0.886 0.16 Around LPCI 0.01 Around instrument 7 of 13 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6B-7 (Cont)

Flow K L/A Flow 2 Description(l) ft'"' 1 Flow Paths Area! ft Factor Coefficient 38-39 13.02 0.15 30° turn 0.442 0.951 40-41 13.02 0.15 30° turn 0.442 0.951 44-45 13.02 0.15 30° turn 0 .. 442 0.951 46-47 13.02 o*.l5 30° turn 0.442 0.951 39-40 14.62 0.15 30° turn 0.429 0.951 45-46 14.62 0.15 30° turn 0.429 0.951

  • 37-48 14.31 0.105 30° turn 0.428 0.947 0.01 Around instrument 42-43 14.31 0.105 30° turn 0.428 0.947 0.01 Around instrument 37-49 10.86 0.0788 Around instrument 0.806 0.937 0.06 Friction 42-54 10.86 0.0788 Around instrument 0.806 0.937 0.06 Friction 43-55 10.86 0.0788 Around instrument 0.806 0.937 0.06 Friction 48-60 10.86 0.0788 Around instrument 0.806 0.937 0.06 Friction 39-51 11.47 0.06 Friction 0.805 0.971 8 of 13 HCGS .. UFSAR Revision 0 April 11, 1988

TABLE 6B-7 (Cont)

Flow K L/A Flow Flow Paths Area, ft 2 Factor Description(!) ft"' 1 Coefficient 40-52 11.47 0.06 Friction 0.805 0.971 45-57 11.47 0.06 Friction 0.805 0.971 46-58 11.47 0.06 Friction 0.805 0.971 38*50 8.27 1.075 Around LPCI 0.864 0.694 41-53 8.27 1.075 Around LPCI 0.864 0.694 44-56 8.27 1.075 Around LPCI 0.864 0.694 47-59 8.27 1.075 Around LPCI 0.864 0.694 49-61 9.18 0.468 Around pipe 0.724 0.804 0.0788 Around instrument 50 . . 62 9.18 . 0.468 Around pipe 0.724 0.804 0.0788 Around instrument 51-63 9.18 0.468 Around pipe 0 .. 724 0.804 0.0788 Around instrument 54-66 9.18 0.468 Around pipe 0.724 0.804 0.0788 Around instrument 55-67 9.18 0.468 Around pipe 0.724 0.804 0.0788 Around instrument 58-70 9.18 0.468 Around pipe 0.724 0.804 0.0788 Around instrument 9 of 13 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6B-7 (Cont)

Flow K L/A Flow 2

Flow Paths Area! ft Factor Description(!) ft- 1 Coefficient 60*72 9.18 0.468 Around pipe 0.724 0.804 0.0788 Around instrument 59*71 9.18 0.468 Around pipe 0.724 0.804 0.0788 Around instrument 49-50 15.35 0.006 Around instrument 0.358 0.846 0.1005 30° turn 0.146 Around LPCI 0.146 Around pipe 53-54 15.35 0.1005 30° turn 0.358 0.846 0.146 Around LPCI 0.146 Around pipe 0.006 Around instrument 55*56 15.35 0.1005 30°turn 0.358 0.846 0.146 Around LPCI 0.146 Around pipe 0.006 Around instrument 59*60 15.35 0.1005 30° turn 0.358 0.846 0.146 Around LPCI 0.146 Around pipe 0.006 Around instrument 58*59 16.84 0.1005 30° turn 0.338 0.953 50*51 16.84 0.1005 30° turn 0.338 0.953 10 of 13 HCGS*UFSAR Revision 0 A ril 11~ 1988

TABLE 6B-7 (Cont)

Flow K L/A Flow Flow Paths Area, ft 2 Factor Description(!) ft*l Coefficient 52-53 14.66 0.1005 30° turn 0.365 0.798 0.469 Around pipe 56-57 14.66 0.1005 30° turn 0.365 0.798 0.469 Around pipe 49-60 19.21 0.1005 30° turn 0.326 0.953 54-55 19.21 0.1005 30° turn 0.326 0.953 51-52 16.95 0.1005 30° turn 0.342 0.896 0.146 Around pipe 57-58 16.95 0.1005 30° turn 0.342 0.896 0.146 Around pipe 52-64 6.89 0.468 Around pipe 0.959 0.718 0.468 Around pipe 53-65 6.89 0.468 Around pipe 0.959 0.718 0.468 Around pipe 56-68 6.89 0.468 Around pipe 0.959 0.718 0.468 Around p~pe 57 .. 69 6.89 0.468 Around pipe 0.959 0.718 0.468 Around pipe 61-62 6.03 0.53 Around pipe 0.853 0.761 0.104 Around instrument 9.094 30° turn 11 of 13 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6B-7 (Cont)

Flow K L/A Flow Flow Paths Area, ft 2 Factor Description(!) ft-l Coefficient 67-68 6.03 0.53 Around pipe 0 .. 853 0.761 0.104 Around instrument 0.094 30° turn 63-64 6.64 0.53 Around pipe 0.796 0.785 0.094 30° turn 65-66 6.64 0.53 Around pipe 0.796 0.785 0.094 30° turn 69-70 6.64 0.53 Around pipe 0.796 0.785 0.094 30° turn 71*72 6.64 0.53 Around pipe 0.796 0.785 0.094 30° turn 62-63 8.91 0.094 30° turn 0.703 0.956 66*67 8.91 0.094 30° turn 0.703 0.956 70-71 8.91 0.094 30° turn 0.703 0.956 61*72 8.91 0 .. 094 30° turn 0.703 0.956 64-65 8.91 0.094 30° turn 0.703 0.956 68-69 8.91 0.094 30° turn , 0.703 0.956 61*73 10.33 0.12 Contraction 0.212 0.945 62-73 10.33 0.12 Contraction 0.212 0.945 12 of 13 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6B-7 (Cont)

Flow K L/A Flow Flow Paths Area, ft 2 Factor Description(!) ft- 1 Coefficient 63*73 10.33 0.12 Contraction 0.212 0.945 64-73 10.33 0.12 Contraction 0.212 0.945 65-73 10.33 0.12 Contraction 0.212 0.945 66-73 10.33 0.12 Contraction 0.212 0.945 67-73 10.33 0.12 Contraction 0.212 0.945 68*73 10.33 0.12 Contraction 0.212 0.945 69-73 10.33 0.12 Contraction 0.212 0.945 70-73 10.33 0.12 Contraction 0.212 0.945 71-73 10.33 0.12 Contraction 0.212 0.945 72-73 10.33 0.12 Contraction 0.212 0.945 (1) Factor for final expansion is included in the flow coefficient calculation .

  • HCGS-UFSAR 13 of 13 Revision 0 April 11, 1988

TABLE 6B-8 GEOMETRY NODE LOCATIONS(!)

Elevation Node Numbers 116' - 3.77" 1 - 12 125' - 1.48" 13

  • 24 134' - 3.88" 25 - 36 142' - 3.63" 37 - 48 151' - 6.50" 49 - 60 159' - 2.75" 61 ... 72 Node Angles Node Numbers 345° 1, 13, 25, 37, 49, 61 315° 2, 14, 26, 38~ 50, 62 285° 3, 14, 27, 39, 51, 63 255° 4, 15, 28, 40, 52, 64 225° 57 16, 29, 41, 53, 65 195° 6, 17, 30, 42, 54, 66 165° 7, 18, 31, 43, 55, 67 135° 8, 19, 32, 44, 56, 68 105° 9, 20, 33, 45, 57, 69 75° 10, 21, 34, 46, 58, 70 45° 11, 22, 35, 47, 59, 71 15° 12, 23, 30, 48, 60, 72 (1) Elevations and node angles are for center of geometry nodes .
  • HCGS-UFSAR 1 of 1 Revision 0 April 11, 1988

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REVISION0 APRIL 11. 1988 PUBLIC SERVICE ELECTRIC AND GAS COMPANY HOPE CREEK NUCLEAR GENERATING STATION PRESSURETRANSIENTIN SHIELD ANNULUSFOLLOWINGA RECIRCULATION LINEBREAK AT THE NOZZLESAFE END UPDATEDFSAR Sheet 6 of 6 FIGURE68-4

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REVISION0 APRIL 11, 1988 PUBLIC SERVICE ELECTRIC AND GAS COMPANY HOPE CREEK NUCLEAR GENERATING STATION PRESSURETRANSIENTIN SHIELD ANNULUSFOLLOWINGA FEEDWATER LINEBREAK UPDATED FSAR Sheet1 of 6 FIGURE68-5

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REVISION0 APRIL 11, 1988 PUBLICSERVICEELECTRICAND GAS COMPANY HOPE CREEK NUCLEARGENERATINGSTATION PRESSURETRANSIENTIN SHIELD ANNULUSFOLLOWINGA FEEDWATER LINEBREAK UPDATEDFSAR Sheet 2 of 6 FIGURE68-5

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REVISION0 APRIL 11, 1988 PUBLIC SERVICE ELECTRIC AND GAS COMPANY HOPE CREEK NUCLEAR GENERATING STATION PRESSURETRANSIENTIN SHIELD ANNULUSFOLLOWINGA FEEDWATER LINEBREAK UPDATEDFSAR Sheet 3 of 6 FIGURE68-5

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REVISION0 APRIL 11, 1988 PUBLIC SERVICE ELECTRIC AND GAS COMPANY HOPE CREEK NUCLEAR GENERATING STATION PRESSURETRANSIENTIN SHIELD ANNULUSFOR STRUCTURAL ANALYSIS CASE C UPDATEDFSAR Sheet4 of 8 FIGURE68-7

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REVISION0 APRIL 11, 1988 PUBLIC SERVICE ELECTRIC AND GAS COMPANY HOPE CREEK NUCLEAR GENERATING STATION PRESSURETRANSIENTIN SHIELD ANNULUSFOR STRUCTURAL ANALYSIS CASEC UPDATED FSAR Sheet 8 of 8 FIGURE68-7

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APPENDIX 6C STRUCTURAL DESIGN CRITERIA FOR SEISMIC CATEGORY I HVAC DUCTS AND DUCT SUPPORTS Seismic Category I HVAC ducts are fabricated from sheet metal and/or plate steel. Stiffener plates are provided at 48*1/2 inch maximum spacing. Maximum allowable stiffener plate spacing is based on analysis and tests.

Seismic Category I HVAC duct supports are fabricated from rolled shapes and unistrut members. Transverse bracing is generally provided at every duct support. However, for several duct support types for duct perimeter 150 inches and less, transverse bracing is provided at every alternate duct support. Longitudinal bracing is provided at 32 foot maximum spacing. A minimum of one longitudinal brace is provided in every straight duct run. Duct sul?ports in the control room are attached to a steel liner plate anchored to the concrete ceiling slab (expansion anchor bolts are*

not used) . Expansion anchor bolts are not used to support HVAC equipment with rotating machinery {e.g., fans) attached to walls or ceilings except as follows:

1. Vane axial fans located in service water intake structure.
2. Unit heaters.

Since the above equipment are relatively light in weight, the expansion anchors supporting them provide very large factors of safety relative to the allowable design loads. Vibratory loads on the expansion anchors are insignificant relative to their design capacity .

Seismic analysis of Seismic Category I HVAC ducts and duct supports is performed using the response spectrum method. Damping values used are 2 percent of critical damping for the operating basis earthquake (OBE) and 4 percent of critical damping for the safe shutdown earthquake (SSE) for welded construction. For bolted construction the corresponding damping values are 4 percent and 7 percent in accordance with the NRC Regulatory Guide 1. 61. Load combinations and allowable stresses for Seismic Category I HVAC ducts and duct supports are given in Tables 6C-l and 6C-2, respectively.

6C-2 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6C-l LOAD COMBINATIONS FOR HVAC DUCTS Allowable Sheet Loading Combination(l) Corner Stress D + p 0.6 Fy D + P + E 0.75 Fy D+ P +E 0.9 Fy D+P+W 0.9 Fy

  • (1) Symbo_ls used in load combinations:

D P

Dead weight of duct Maximum operating pressure inside duct E Operating basis earthquake load Es Safe shutdown earthquake load Wtp Tornado differential pressure Fy Minimum specified yield strength of duct material

    • 1 of 1 HCGS-UFSAR Revision 0 April 11, 1988

TABLE 6C-2 LOAD COMBINATIONS FOR HVAC DUCT SUPPORTS Load Loading Allowable Case Combination(!) Stress( 2)(3) 1 D+L F s

2 D+ E Fs 3 D + Es 1.5 F or 0.9 Fy whichever s

is smaller (1) Symbols used in load combinations:

-1 D Dead load including weight of duct, support frame, stiffeners, volume dampers, instrumentation, etc.

L Live load - 250 lbs applied to a one square foot area at points of maximum moment and shear (representative of a man walking or crawling on a duct).

E Operating basis earthquake load E Safe shutdown earthquake load s

F Allowable stress for support material governed by s

AISC or AISI as applicable Fy Minimum specified yield strength of support material.

(2) Where expansion anchor bolts are used to attach duct supports to concrete structures, 33 percent increase in allowable design loads is permitted for load case 3, for expansion anchors .

    • 1 of 1 HCGS-UFSAR Revision 0 April 11, 1988
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