ML20002A227

Forwards Proprietary & Nonproprietary Info in Response to NRC Questions Re Amend to License DPR-53,fifth Cycle License Application.Affidavit Encl.Proprietary Info Withheld (Ref 10CFR2.790)

ML20002A227
Person / Time
Site:	Calvert Cliffs
Issue date:	11/02/1980
From:	Lundvall A BALTIMORE GAS & ELECTRIC CO.
To:	Clark R Office of Nuclear Reactor Regulation
Shared Package
ML19260G378	List: ML20002A227 ML20002A228
References
NUDOCS 8011050328
Download: ML20002A227 (36)

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November 2, 1980 l

Of fice of Nuclear Reactor Regulation l

U. S. Nuclear Regulatory Commission Washington, D. C. 20555 ATTENTION:

Mr. R. A. Clark, Chief Operating Reactors Branch #3 j

Division of Licensing l

l

SUBJECT:

Calvert Clifis Nuclear Power Piant i

Unit No.1, Docket No. 50-317 Amendment to Operating Licen',e DPR-53 Fif th Cycle License Application Responses to NRC Staff Questions Gentlemen:

Enclosed are our reponses to ques. ions posed by NRC staf f on the subject application.

Very truly yours i

BALTIMORE Gi RIC COA ANY

~[j.

It rfu, -

A. E. Lun vall, Jr.

Vice Pres dent - Supply AEL/W3L/mit Copy To:

3. A. Biddison, Esquire (w/out Encl.)

G. F. Trowbridge, Esquire (w/out Encl.)

Messrs.

E. L. Conner, Jr., NRC P. W. Kruse, CE

' (40 Copies)

Enclosure (Calvert Cliffs Unit 1, Cycle 5, NRC Reload Questions Response -

Answers on CESEC Model Used in Si.B Analysis) - Proprietary Copies

//000001:- 000040, 20 Non-Proprietary Copies 1

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AFFIDAVIT PURSUANT T010 CFR 2.790 Combustion Engineering, Inc.

)

State of Connecticut

)

County of Hartford

)

SS.:

I, A. E. Scherer depose and say that I am the Director, Nuclear Licensing i

of Combustion Engineering, Inc., duly authorized to make this affidavit, and have reviewed or caused to have reviewed the information which is identified as proprietary and referenced in the paragraph immediately 4

below.

I am submitting this affidavit in conformance with the provisions of 10 CFR.2.790 of the Commission's regulations and in conjunction with the application of Baltimore Gas and Electric Company, for withholding this information.

l The information for which proprietary treatment is sought is contained in the following document:

Calvert Cliffs Unit 1, Cycle 5, NRC Reload Question Responses (Answers I

on CESEC Model Used in SLB Analysis)

This document has been appropriately designated as proprietary.

I have personal knowledge of the criteria and procedures utilized by Combustion Engineering in designating information as a trade secret, privileged or as confidential commercial or financial information.

Pursuant to the provisions of paragraph (b) (4) of Section 2.790 of the Commission's regulations, the following is furnished for consideration by the Commission in determining whether the information sought to be withheld from public disclosure, included in the above referenced document, should be withheld.

,. 4 1.

.The information sought to be withheld from public disclosure are selected input data and results from the analysis of a Steam Line Break event, which is owned and has been held in confidence by Combustion Engineering.

i 2.

' The information consists of test data or other similar data concerning a process, method or component, the application of which results l

in a substantial competitive advantage to Combustion Engineering.

3.

The information is of a type customarily held in confidence by Combustion Engineering and not customarily disclosed to the public.

Combustion Engineering has a rational basis for determining the types of information customarily held in confidence by it and, in that connection, I

utilizes a system to determine when and whether to hold certain types of information in confidence. The details.of the aforementioned system were

]

provided to the Nuclear Regulatory Commission via letter DP-537 from j

F.M. Stern to Frank Schroeder dated December 2,1974.

This system was i

applied in determining that the subject documents herein are proprietary.

4.

The information is being transmitted to the Commission in confidence under the provisions of 10 CFR 2.790 with the understanding that it is to be received in confidence by the Conmission.

5.

The information, to the best of my knowledge and belief, is not available in public sources, and any disclosure to third parties has been made pursuant to regulatory provisions or proprietary agreements which provide for maintenance of the information in confidence.

6.

Public disclosure of *he information is likely to cause substantial harm to the competitive position of Combustion Engineering because:

a.

A similar product is manufactured and sold by major pressurized l.

water reactors competitors of Combustion Engineering.

I i

. b.

Development of this information by C-E required thousands of man-hours of effort and tens of thousands of dollars.

To the best of my knowledge and belief a competitor would have to undergo similar expense in generating equivalent information.

c.

In order to acquire such information, a competitor would also require considerable time and inconvenience related to development of methodologies and determination of input parameters for a Steam Line Break

analysis, d.

The information required significant effort and expense to obtain the licensing approvals necessary for application of the information.

Avoidance of this expense would decrease a competitor's cost in applying the information and marketing the product to which the information is applicable.

e.

The information consists of selected input data and results from analyses of Calvert Cliffs, Unit 1 Cycle 5 5 team Line Break event, the application of which provides a competitive economic advantage.

The availability of such information to competitors would enable them to modify their product to better compete with Combustion Engineering, take marketing or other actions to improve their product's position or impair the position of Combustion Engineering's product, and avoid developing similar da.ta and analyses in support of their processes, methods or apparatus.

f.

In pricing Combustion Engineering's products and services, significant research, development, engineering, analytical, manufacturing, licensing, quality assurance and other costs and expenses must be included.

The ability of Combustion Engineering's competitors to utilize such information without similar expenditure of resources may enable them to sell at prices l

reflecting significantly lower costs.

. 9 Use of the information by competitors in the international marketplace would increase their ability to market nuclear steam supply systems by reducing the costs associated with their technology development.

In addition, disclosure would have an adverse economic impact on Combustion j

Engineering's potential for obtaining or maintaining foreign licensees, s

Further the deponent sayeth not.

l 1

-_ _!ff W j

m. [gsterer Director X

4 l

Nuclear Licensing

}

i Sworn to before me

'his.3N day of NCtOdu / I Lula vt-

@). l lt((-

Notary Public I

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ENCLOSURE 1 Question 1 Why was the code CESEC-SLB needed to simulate the Steam Line Break (SLB) event for Cycle 5?

Answer The SLB event analyses for Cycle 5 included the effects of nanually tripping the Reactor Coolant Pumps (RCPs) on Safety Injection Actuation Signal (SIAS) due to low pressurizer pressure and the automatic initiation of Auxiliary Feedwater (AFW) flow on low steam generator water level signal. The inclusion of these effects in the Cycle 5 analyses required the use of the latest version of CESEC (which has been referred to in Appendix C of Reference 1 as CESEC-SLB) for the following reasons:

1.

The manual trio of RCPs results in a drastic reduction in flow.

The reduced flow causes increased temperature tilt at the reactor vessel inlet which, due to incomplete mixing of the conlant in the vessel inlet plenum produces a much more severe radial temperature asymmetry in the core. The temperature asymmetry experienced during an SLB event with RCP trip requires the use of RCS coolant node scheme which is capable of representing incomplete mixing in the reactor vessel.

2.

The trip of the RCPs also affects the Reactor Ecolant System (RCS) pressure.

The RCS pressure determines the magnitude of Safety Injection flow via the High Pressure Safety Injection (HPSI) pumps, and, thus the total negative reactivity added due to boron injected.

Due to reduced flow through the reactor vessel closure head, a model which explicitly represents the roactor vessel closure head was required to more accurately predict the pressure variation during the event.

3.

The RCP trip also required a more accurate crediction of the way boron injected via the HPSI pumps is distributed in the RC S loop to provide negative reactivity.

Hence, an inproved modelling of the boron transport in the primary coolant and of the safety injection system was required.

4.

The RCP trip rcouired a flow model which is able to exolicitly calculate the time dependent reactor coolant mass flow rate.

5.

The automatic initiation of AFW required modification of the primary to secondary heat transfer model in orde to calculate the RCS cooldown af ter AFW initiation to a poteni.ially dry stean generator.

Hence, the version of CESEC which includes the above mentioned model improvements was used to analyze the SLB event for Cyc!e 5.

References to Question #1:

Unit 1, Cycle 5 License Submittal

I Question 2 Provide a description of the overall conservatism inherent in the CESEC-SLB code.

List all conservative assumptions in the codes and inputs for each parameter.

Answer The conservatism in the Steam Line Break analyses exists in nainly the input data rather than the CESEC code.

The only inherent j

conservatism in this CESEC version is that the heat transfer area I

is calculated assuming that all tubes are covered until the mass i

in the steam generator is equal to 5000 lbn.

This assumption is i

conservative for assc ; sing the potential for a return to power, since it increases the heat transfer rate between the prinary and secondary and thus, produces the maximun cooldown of the RCS.

The conservatisms in the key input data are given below.

Parameter Value Justification j

Power level 27E4 Mwt This is the maximum allowed thermal power including uncertainty. The maximum power level results in maximum coolant average temperature.

The maximum coolant average temperature increases the moderator reactivity inserted during the cooldown by increasing the total change in the coolant temperatures. Also, the maximum power level increases the decay heat.

Core Inlet SE0' F This is the maximum core coolant inlet temperature in-Temperature cluding uncertainty.

The maxinum inlet temperature results in a higher initial steam generator pressure, which increases the blowdown rate from both steam generators.

2 Break Area 6.35 ft The analysis assumes the largest break area of 6.35 ft,

This results in the fastest blowdown and thus the most rapid cooldown of the RCS and the greatest rate of temperature reduction in the reactor core region.

This leads to a maxi."in positive reactivity insertion and the greatest potential for a return to power, l

MTCof-2.2x10goldowncurvecorrespondstoaneffective

!oderator Ccol-See Figure 1 & 2 The moderator c ao/ F and is calcu' lated asswaing that down Curve a Control Element Assembly is stuck in tne fully withdrawn position during the reactor scram.

l

---~ ~ ~... -

i

5

. Pa r no te r Value Jus t i fication I ppler Coeffi-1:0C The EOC lbppler coe f ficient in conhination with the cient D]ppler 1.15 decreasinh fuel tenveratures causes the greatest

ultiplier positive reactivity addition, due to fuel temperature change, during the event. The Doppler uncertainty of 15*, asstmed in the analyses is in a sense that enhances the Doppler feedback.

l Scram L' orth-lif?

7.157, Ap lhis is the minimum available scram worth at EOC and ilZP 4.3' Ap was calculated allowing for the stuck rod which produced the limiting moderator cooldown curve.

Inverse Boron Korth flFP 105 PP:!/ top These corresoond to the r:tininun boron reactivity worths for liZP 100 PP:,l/% as the borop iniected via the Hic;h Pressure Safety Injection j

Pu"ps and minimize the negative activity added by l

Safety lujection.

i IlirJi Pressure The analyses conservatively asstmied that only one Safety Injection

!! PSI pump is operable.

In addition, the maximum a) Ntmber of Pumps 1

Technical Specification tire delay to start the b) Time Lblay to ptmps is also assuned. The maxiatmi volume to be Start Pteps 30 sec swept out prior to when boron injection enters the core c) Volume to be is also asstred. These assumptions are conservative, j

Swept Out Prior since they delay the time at which boron injected

^

to Boron In-via the llPSI pu::ps enters Rcs cold legs.

jocted Enters ICS Cold Ings 76 ft

' lain Feedwater 5', Full laintaining the main feedwater flow to the ruptured Flow Power steam generator increases the mass released dur:ng the Flowrate blowdown, lengthens the blowdown, and aggrevates the cooldown.

rhese are maximwn value allowed by fine to Rampdown 20 sec Technical Specifications.

' lain Feed After frip Feedvater Isola-80 see tion after ! ISIS Auxiliary Feed-350 lbm/sec This value is conservatively calculated assuming that dater Flow l

both auxiliary feedwater ptmps are functional. 'ihe value corresponds to the punp run-out value due to reduced l

back pressure.

In addition, the auxiliary feeduater flow is fed only to the damaged steam generator

'these

! conservative assumptions produce the maximten couldown I of the RCS and thus enhance the potential for Return-To-Power after initiation of auxiliary feedwater flow.

i l'raction

.0060 The maximum EOC S fraction is used in the analyses.

This causes the fastest approach to Return-To-Power due to subtritical multiplicacion.

Initial Steam 853 psia

. The value is the maximtun initial steam generator pressure

' enerator for the initial power, the initial core coolant temperature l 1;rensure and mass flow rate used. This value is conservative because it increases the rate of blowdown of the steam

'Oh~

[g g 0 s

generator.

l 1

' l'a t. dter Value Justification i

'nitial RCS 2300 psia This is the naximum initial pressure allowed.

The use

! : ressure of the naximun pressure delays the time of Safety injection Actuation Signal and thus the amount of negative reactivity contributed by Safety In.jection.

l l

tiin Steam f>.9 seconds This is the maximum time to close the f tSIV's.

The

,olation Valves maximum time prolongs the blowdown from the unaffected j

losure Time af ter Steam Generator.

'iin Steam l

solation Signal 9

4 i

n: ox

i i

i fi 2 LOOP-FULL P0tlER

+6 l

+5 1

1 I

CL

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+4 i

N 5

p

+3 o

Fu 4

e m

f2<

+2 f5 ao

E

+1

.0

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40 45 50 55 60 65 DENSITY, LBM/ CUBIC FT l

BALTIMORE STEAM LINE BREAK EVENT MODERATOR REACTIVITY FEEDBACK vs 1

As' c

co.

c, Nuclear Pcwer Plan, MODERATOR DENSITY l

5.0 2 LOOP-il0 LOAD l

, 4.0 l

i 3.0 p

h$u

+

i M 2'0 l2

&u O

E 1.0 ft 0.0

=

45.0 50.0 55.0 60.0 65.0 MODERATOR DENSITY, LBM/FT3 BALT W ORE STEAM LINE RUPTURE EVENT rigurc GAS & ELECTRIC CO' Coiveri Cirrrs MODERATOR REACTIVITY vs MODERATOR DENSITY 2

lloclear Powcr Plant

~ ~m3

Question 3 Provide a more detailed description of the code CESEC-SLB, including pertinent equations.

Answer The following provides a general description and equations used in the Steam 1.ine Break analysis for Cycle 5.

tiodels pertinent to the Steam Line Break analysis are specifically provided.

Sone models, such as flow from the Safety Injection Tanks are described for completeness but were not credited for in the Cycle 5 analysis.

I l

l I

il 1

SUittARY DESCRIPTIOll 0F TifE 1.RS10fl 0F THE CESEC CODE USED FOR THE CALVERT CLIFFS Uillf 1, t,YCLE 5, STEA!1 LlflE BREAK AllALYSIS In t ro_duc ti_on The CESEC digital computer program (1 to 8) provides for the simulation of c Nnbustion Engineering fluclear Steam Supply System (flSSS). The program

't the plant response for non-LOCA (loss of coolant accident) ini-nts for a wide range of operating conditions.

ogram, which nuncrically integrates the one-dimensional conser-va t.s

. ions, assumes a node flow-path network to nodel the flSSS.

The primary system components considered in the code include the reactor ves-sel, the reactor core, the primary coolant loops, the pressurizer, the stean generators, and the reactor coolant pumps (see Figure 1). The secondary systen components, shown on Figure 2, include the secondary

}

side of the steam generators, the Main Steam System, the Feedwater System, and the variout steam control valves.

In addition, the program models some of the control and plant protection systems.

The code self-initializes for any given, but consistent, set of reactor 1

j power level, reactor coolant flow rate, and stean generator power sharing.

During the transient calculation, the time rate of change in system pres-sure and entha'py are obtained from the solution of the conservation equa-l tions. These oerivatives are then numerically integrated in time, under l

the assumption of thermal equilibriun, to give the system pressure and nodal enthalpies.

The fluid states recognized by the code are subcooled i d saturated;superheating is allowed in the pressurizer. The fluid in the Reactor Coolant System is assumed to be homogeneous.

In the subsections which follow, a description of the major models which I

comprise the version of the CESEC code which was used for the Ste'am Line Break analysis for the Calvert Cliffs Unit 1, Cycle 5 safety analysis is given.

Primary Coolant Thermal-Hydraulic Model The CESEC code uses a node / flow-path type network to model the Reactor Coolant System.

The conservation of mass and energy equations are solvec for control volumes or nodes. Uniform pressure is assumed around the reactor coolant loops for the thermal-hydraulic solution.

The conserva-tion of momentum equation is solved independent cf the conservation of mass and energy equations to obtain the pump eiows for the thermal-

)

hydraulic model (see section on Flow Model).

The Reactor Coolant System, consisting of two reactor coolant loops, the reactor vessel, the reactor vessel closure head and the pressurizer was divided into 27 nodes of constant volume for this analysis. The nodal scheme given in Figures 3A ar.d 3B was choseri~ to appropriately sinulate the RCS

~

~

component volumes and, thus, provide an adequate description of the spatial i

variation of the coolant pronerties.

As seen fron Figures 3A and 38, nodes are sr ecified which represent one-half the reactor vessel inlet downcome section, the w w powe=*eWye e*W-w

lower plenun, the core region, the bypass flow, and the upper plenun.

These two synmetrical loops are linked by the cross flow at the reactor vessel inlet and outlet sections and by the flow mixing within the reactor vessel lower and upper plenums. The mixing factors are specified based on test data, fio cross flow is assumed between the parallel regions in the core.

j I

CESEC solves the conservation of nass and energy equations (see Figure 4) to obtain the time derivative of the pressurizer pressure, the internodal flows, the rate of vaporization or water enthalpy tine derivative of the pressurizer water reginns, and the rate of condensation or steam enthalpy time derivative of the pressurizer steam region.

Computation of these parameters allows for the calculation of the RCS pressure time derivative, the time derivatives of the nodal enthalpies, the nodal specific volumes, and the nodal masses.

Closure Head flode During the rapid contraction of the primary ;oolant which takes place as a result of a steam line break, the press irizer empties and voids begin to forn in the RCS.

Since flow throug1 the closure head is only a small fraction of the RCS flow, the temperatures in the closure head remain high and voiding first occurs there.

To some extent, the closure head itself then begins to perform tqe function of a pressurizer.

i Therefore, the reactor vessel closure head r?gion is explicitly modeled in this CESEC version to more accurately predict the RCS pressure.

The coolant flow f^om the core outlet nodes to the vessel head node is specified by a user input fraction.

It is assumed that the vessel head fluid returning into the outlet nodes is evenly distributed between the two loops.

Pressurizer The CESEC. pressurizer model assumes steam and liquid regions to exist in one of the eight thermal-hydraulic states shown in Figure 5.

The model considers such conponents as sprays, heaters, and relief / safety valves. The Pressurizer Level Control Systen which controls cht.rging flow cnd letdown flow by means of pressurizer level setpoints, is also modeled.

The nass and energy transport between the two fluid regions is assumed to occur as a result of liquid vaporization and/or steam condensation, be spray flow which enters the pressurizer is assumed to condense the stean if it is in the saturation state. That is, when the steam region is at saturation, the spray droplets are assumed to reach saturation tempera-ture and will result in bulk condensation of the stean.

However, when the stean region it superheated, the spray droplets are assumed to evaporate into the steam region.

The code nodels' two. spray operating modes, continuous and proportional. The continuous mode spray is a user input constant flow which is added con-tinuously to' the pressurizer.

The proportional mode spray flow originates at the pump discharge in the RCS loop and is linked to the pressurizer as shown on Figure 3.

The spray flow fcr the proportional mode is con-trolled automatically by two pressure setpoints which turn the spray on and off, respectively. Within these two setpoints, the spray flow increases linearly with the pressurizer pressure.

-The code also models two types of heaters located near the bottom of the l-pressurizer:

(1) the proportional heaters which are controlled by the

(

' Pressurizer Pressure Control System to generate heat at -a rate which -de-creases linearly with increasing press ~ure between two pressure setpoints J

and (2) the backup heaters which turn on and off at two pressure setpoints.

4

-In addition, the backup heaters are also controlled by the measured devia-tion of the pressurizer liquid level from the programmed level.

The addi-tion of heat from heaters to the fluid is acco"nted for in the conservation of energy equation.

b Flow ifodel The flow nodel in this CESEC version calculates the mass flow rate (lbm/se:) at the J

pump outlet for each_ reactor coolant system 3 team generator loop. The model in-cludes explicit simulations of the. reactor coolant pumps and of the effects of natural circulation flow. The calculation is based ~on a solution of the one-dimensional momentum equation for each RCS loop. The loops are divided into a number of nodes whose densities, temperatures, and flows are obtained from the CESEC thermohydraulics model.* The flow model utilizes this nodali-zation of the loop to calculate the sum of the various forces around the loop. The forces acting on the fluid volume consist of (1) gravitational forces due to density and elevation changes around the loop, (2) forces due to wall friction and geometric changes in the flow path, and e

(3) forces due to the RCS pumps. The one-dimensional momentum equation for each loop, is written as follows:

t-h n

n i i fric.i + Rggg,$

pump S

+ aP 9E Dii~f O

i=1 1=1 Pin -

(1) dw Pis dt n

?

(t. 5/A )'

j 1=1 oThe average ' f the properties and the flcus frca par-lloi nodes are used for nodes o

representing the reactor vessel.

4

./

L 1

l where w = mss flow rate at the pu:rp, lbn/sec th f

w. = cass flow rate of i node, Ibm /sec i

th 3

= average fluid density of i noda, lbm/ft 3 h p5" = single-phase fluid density of itl' node, Ibm /ft pjg

= the elevation difference c. cross the i'h node, ft h$=hin, i - hout i th ncde Dj = Thom two-phase multiplice for the i th node g = Darcy friction factor for the i f

th node, ft L'4 = effective flow path length for the i th 2

node, ft

$ = effective cross sectional flow area of i A

L /D fric,i=1['d and R 2Aj K

o,1 Ngeo i [1 th

. = effective diameter of i node, ft where D,1 e

g,j = flow loss coefficient for geometric changes in the flow path, dimen:.ionless K

Il The first term en the right hand side of Ecuation 1 represents the net pressure change crtund the loop due to tha gr.vitational force act-ing on each fluid node.

The sccond term represents the total pressure j

change arcund the lccp due to the~ frictional loss and the gsentric changes in the flow path. The Darcy friction factcr, f, which is a function of the Reynold's number, Re, is determined by the following correlations:

Re < 1250 f = 64/Re j

(2) f = (-0.000004)Rc+0.056

'1250 5,Re < 600 f = 0.184/Re.2)

Re 3,6000 0

The third tarm in Equatico 1 represents the cressere difference APpumn, across the RCS nu:7.a, wnich is calculated by the dynar.lic puma trndel described in the folicwing paragrcphs.

The cump model calculates the pressure difference across the numa.The pres'sure difference, for use in the conservaticn of actrentua ecuation.

The speed is calculated or oumo head, is decendent cn the cumo speed and flow.

thus, giving a como head dependent on transient througncut the transient,The purc.o speed is determined frca the following equation:

flow conditions.

h-T,g)y (3) dg=(Tel - T g

m 1

_a._.,___

7-

^ ~~~ ' - ~ ~ - ~

D**D

• D we e

A

=

uhere w = angular velocity of the rotating assembly I

t = time T

= electrically induced torque actiag on the matcr retor g

T = hydraulic torque exerted on the fluid by the.pucp impeller h

= torque exerted on the rotating asscrbiy due to bearing friction T 7'g and windage' losses g = gravitational ccastant I = m.craent of inertic of the rotatinq assembly el, T, and Tf,g) are calculated ts folicws:

Ti.e tarte torques (T h

a) 1he hydraulic torque is calculated from the following equations:

lv/a (S/a )(a )(T )(p/p )

for 1.0 (4) g g

T

,i h

f f

2 (0/v )(y2)(T )(0/p )

for v/a 1.]

where S = Ratio of the hydraulic torque to the rated hydraulic torques, SE T /TR h

= Ratio of the pump speed to the rated pump speed, as g/g a

r V = Ratio of the volumetric flow rate to the rated volumetric flow rate, vs Q/Qr 3

p = Density of coolant, lbm/ft 3

R = Density corresponding to pump rated conditions, lbm/ft L

P The values of B/a and 6/v as a functions of v/a and a/v, respectively, are determined from the single phase honologous pump curves.

b) The friction and windage torque is calculated from the following equation:

T

=al alt (O

fgg fgg; l

where Tg; is the input friction and windage torque at rated speed, c)

Electrical torque is found by interpolating in an input table of speed vs. electrical torque using the pump speed fron the previous time step.

/

T

]D 'd d o

0 9 "D

/

o ju g6 liith all quantitic knc.m in 9;uation 3, the rate of chhnge of pump speed can 'cc calculated.

The pump speed is then changed by the product of this rate with the time step size for the next time step.

The pump head, H, is calculated frcm the follouing equation:

(u )(g )

(af,2) for lv/al 1

1.0 2

(6) g (v )(il )

(h/v?)

forlv/ol 2

1.0 g

where:

(

11

- Puup head, ft of water li

= Rated pump head, ft of water g

h

= Ratio of the pump head to the rated pump head, h E II/Ho Tne value of h/a and h/v as a function of 9/a and a/v, respectively, are i

determined from the single phase homologous pump curves.

The pressure differcnce across the punp is then calculated from:

(7)

AP

• HP9 pump Reactor Kinetics The energy source in the CESEC code is fron fission in the fuel.

The core is represented by a cylindrical fuel rod located in an average coolant channel.

This fission energy consists of two parts, the instantaneous fission power and the decay power released by the fission products. The instantaneous power is determined by solving the standard point kinetics neutron equations with six delayed neutron groups while the decay power is calculated from an 11 fission product group decay heat nodel.

The total reactivity in the point kinetics couation is calculated as the sum of the control rods, noderator, fuel temperature (Doppler), and boron contributions.

The code also has an explicit function of time simulating the control rod reactivity A table of rod reactivity versus time after initiation of scram is user insertion.

input.

The moderator feedback effects considered include the moderator density or The noderator and Doppler reactivity feedback terms are the moderator tenperature.

calculated at each time step by interpolation of user input tables. The boron reactivity effect includes the contribution from the Safety Injection System and the letdown and charging portions of the Chemical and Volume Control System.

The kinetics equation is solved numerically by a fourth order Runge-Kutta/Nerson method for the power generation at each time step.

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m

)

Heat Transfer Within the Core i

The CESEC core heat transfer nodel represents a fuel rod at core average conditions. The cylindrical configuration models the fuel, gap and clad.

The fuel rod is divided into three equal-volume radial nodes (see Figure 6).

The third radial node is assumed to contain the outer portion of the fuel, 3

the gap, and the clad.

The radial energy equation (see Figure 6) is formulated for each node with the nodal properties (e.g., specific heat and thermal conductivity of fuel and clad) determined by temperature depend-ent correlations. The input paraneters required by the nodel include the 3

fractions of power generated witliin the fuel, the clad, and the moderator, i

i respectively, and the gap conducts:nce which is assuned to be a constant.

Ilithin the fuel region, a uniform power distribution is assumed by the code.

The heat transfer at the clad - coolant interface is assumed to be given by the following correlation for all #1uid conditions (Reference 6):

2 0

h = 0.148 (1+0.0lT - 0.00001T ) y.8 (8) f where T = fluid temperature V = fluid velocity D = channel hydraulic diameter Initially, the steady state fuel temperature distribution is determined by a scheme which solves the radial energy equation iteratively based on the initial reactor power output, the gap conductance, and the initial coolant condition. The radial energy equation is solved numerically at each time step by a fourth order Runge-Kutta/tterson method.

Charging and Letdown The CESEC co3e provides a nodel for calculating the charging and letdown flows. The contributions from the charging and letdown flows are included in the conservation of mass and energy equations for the corresponding RCS nodes.

Included in the model is a Pressurizer level Control System which deternines the deviation between the measured pressurizer water level and the programmed level.

The progranmed level is given by an input table as a function of either power or average RCS temperature.

The algorithm by which the neasured level is calculated is described in Reference 5.

The charging flow is.provided by a set of constant speed charging pumpe, wi th the charging flow rate automatically controlled by suitchi g each pump on n

or off at two input level deviation setpoints.

The letdown flow control is provided either by a set of letdown control and backpressure valves, with the flow rate either controlled by the opening or the closing of each set of valves at two level deviation setpoints, or by a linear letdown flow control model.

i

+

1 The charging and letdown fluid temperatures are user input.

In addition, the letdown fluid tenperature can be selected to be that corresponding to the steam generator outlet temperature. The boron concentration from the letdown and charging portion of the Chemical and Volume Control' System (CVCS) is only accounted for in CESEC when the Safety Injection Systen is activated.

However, the user can optionally turn off the letdown and charging systems and take no credit for the boron reactivity contribution from the letdown and charging systems.

The calculation of the boron con-centration in the reactor coolant is described in the Safety Injection Sys.en section.

Reactor Protective System Trips The reactor is shutdown by the insertion of the control element assemblies (CEAs) following the generation of a trip signal.

A trip signal is ini-j tiated when a certain system parameter reaches a value which exceeds the corresponding user input trip setpoints.

The delay time between the ini-tiation of the trin signal and the start of CEA motion is accounted for in CESEC. The CEA motion is represented by an input rod worth versus time table.

The following trips are programmed in the CESEC code:

1.

high power trip, l

2.

high pressurizer pressure trip, 3.

low pressurizer pressure trip, j

4.

low coolant flow trip, 5.

low steam generator pressure trip, 6.

low steam generator level trip, and 7.

nanual trip.

To generate the trip signal on the low steam generator water level, the stean generator water level is determined from a set of steady state input data and the transient inventory in the steam generator.

The set of steady state curves relates steam generator water level to secondary water mass I

and power level. This data is then used in a table look-up routine to obtain the steam generator water level for the purpose of determining the trip signal.

t

• ~~

- & n

..m1-w _

..u_i n

m

.._ -.e w.m t.m --

-+:.

- - -ha-

Safety Injection System The boreted safety injection water from the high and low pressure safety injection pumps is injected into each cold leg node downstream of the reactor coolant pumps. The borated injection flow rates versus pressure are speci-fied by input tables. Once the safety injection flow reaches the cold leg node, it is assumed to mix homogeneously with the reactor coolant in that node. The boron is transported through the RCS by solving at each time step the con-tinuity equation for each coolant node for the boron concentration:

Li M h = 1;lin in - N C

N C

out where l

C is the boron concentration C

is the inlet boron concentration in Win is the inlet flow rate W

is tie outlet flow rate out it is the reass inventory in the node The boron concentration for the reactor core node is used to calculate the reactivity contribution due to boron via an input reciprocal boron worth.

A time delay is input to CESEC to account for the time required to start the diesel generator and/or -to bring the safety injection pumps to full-

' speed. An additional time delay is calculated to account for the time required for the unborated water in the safety injection line (from the outlet of

.the safety injection pumps to the injection nozzles) to be swept out before borated water from the refueling water tanks enters the cold legs.

CESEC alsosolves an orifice equation to determine the rate of safety injection flow from the safety injection ta.cks into the RCS as a function of time. The input parameters :are the initial nitrogen pressure, volume of water, volume of gas _, flow coefficient, flow area, water specific volume, and elevation head of the safety injection tanks.

In addition to the nitrogren pressure within'the safety injection ~ tank, the static head of fluid within the safety injection piping.is considered when calculating the instantaneous pressure difference across the orifice. The nitrogen expansion process is assumed to be isentropic.

In computing the safety injection flow rate by means of an orifice equation, the code takes into account the effect of piping friction, turning losses, and

~

expansion / contraction losses through the use of a single equivalent loss coefficient which is based on the minimun cross-sectional flow area. The instantaneous liquid discharge rate at time t is given by i

c a

L

7 D**)D *]D'l 6

ow M w M.2UN a

288 AP(t) 1/2 (10) g a(t) i A(=

)

gy g

6P(t) = P (t) + P RCS(t)

(11)~

~

g E

tchere '.lis the tress fica rate in iba/sec 2

A is the flow area of safety injection tank line in ft l

X is the friction loss coefficient for the flcu crea.

y is the specific voluca of liquid in ft /lbm J

i P (t) =_ nitrogen pressure at time t g

PE = clevati n head RCS(t) = RCS pressure at time t' P

If 6P is less than or equal to zero, the code sets this variable equal to zero in order that no, liquid mass be ejected,from the tank for this condition.

The instantaneous liquid volume V in the tank at time t is then i.

V(t)=

V(t-at)

'.l(t) at v (l?)

where at is the tima step interval.

The instantaneous gas volume V in the tank at time t is given by g

V (t) = V (0)tV.(0)- V (t)

(13) g G

where V (0) and Y (0) are'the initial gas volume ar.d li, quid volume, respective'y.

G The instantaneous ga.s pressure,in the tank at time t is given by:

I Y (0)g 4

P (t) = PO (0),f G

(14)-

g i T[y/

I w

_a

e Critical Flow !!odel For steam escaping from the ruptured steam line, the mass flow rate is calculated in CESEC from the following empirical criti. cal flow correlati.on(4):

11 = 1977.6 x P

(15)

A F-T85.0 where:

ll is the mass flow rate, lbm/sec A is the effective flow area, ft P is the steam pressure, Psia h is the steam enthalpy, Btu /lbm Steam Generator flodel The CESEC steam jenerator model performs a detailed computation of the

(

overall heat transfer coefficient for each steam generator.

The heat transfer correlation used in the primary side for calculating the film re-I sistance is the same as for the core. The secondary side heat transfer mechanism is pool boiling. The boiling resistance is calculated using the modified Rohsenow pool boiling correlation (6),

L g

2/3 (16) h,cc = K where q = heat rate A = heat transfer area i

K = a Pressure dependent coefficient given by a C-E proprietary correlation.

R (Refe: ence 6)

In the CESEC version used for this analysis, the heat transfer between the primary and secondary sides of the steam generator is calculated using algorithms based upon a polynomial spatial variation of the primary side temperature. The polynomial algorithm selects the mid-point temperatures of the nodes simulating.

the stean acnerator tubes and the inlit Y-

--m-=

-,.-w.

M-' % n. 6 M 4 -- M 5

Ma'bu

and outlet pleaua to obtain the spatial variation of the primary side tempera-ture along the steam generator tubes.

The difference between this temperature variation and the secondary side temperature is integrated between the node mid-points and divided by the distance between these points to obtain the average primary-to-secondary temperature difference.

This temperature difference is than used to calculate the steam generator heat transfer rate on a node-center-to-node-center basis at each time step.

The overall heat transfer coefficient is determined from the film resistance of the primary and secondary sides, the tube wall resistance, and the tube fouling resistance.

The tube wall / fouling resistance is determined initially by the design full power condition and is assumed to be constant thereafter during the transient. These heat transfer coefficients are calculated on a node-center-to-node-center basis rather than for the stean generator as a whole.

For low heat flux predictions, the secondary-side heat transfer coefficient is limited to a constant input value, rather than being allowed to go to zero with the heat flux as ir. Equation 16.

This minimum value is also used for tFe secondary-side heat transfer coefficient for conditions of reverse (secondary-to-primary) heat transfer.

The~ primary-side coefficients are calculated using the same correlation (Eq;ation 8 ) for forward and reverse heat transfer.

The secondary side of each steam generator is represented by a control vol-ume. The control volume consists of saturated liquid and steam. The fluic properties and nass inventory are determined oy solving the conservation of nass and energy equations shown on Figure 7.

The initial conditions of the secondary s:de of the steam generator are determined by iterati'19 on the secondary pressure,given the initial heat flux and power demand as i

specified by the user.

To avoid singul6rities in the solution of the conservation equations shown t

in Figure 7, the liquid inventory in each steam generator is limited to a nininum of 7500 lbn. Thus, steam is allowed to escape through the ruptured steam line with a critical flow velocity as long as the steam pressure is above atnospheric pressure and the stean generator liquii inventory is greater than 2500 lbm.

Smooth transition fron normal neat transfer con-ditions to a condition which does not reduce the liquid inventory below the minimum of 2500 lbm and also maintains the saturated conditions assumed by the model for the secondary side is achieved by a rampdown of the product of the averall heat transfer coefficient with the heat transfer area, VA. For tteam generator inventories greater than 5000 lbm, the total heat i

transfer area is used together with a primary-to-secondary heat transfer coefficient calculated using the correlations given above.

For liquid inventories less than 2500 lbm the product of heat transfer coefficient and heat transfer area is assumed to be just sufficient to raise the

)

enthalpy of any incoming feedwater to that of the saturated liquid.

For 11 quid inventories between 2500 lbm and 5000 lbm the product UA is scaled l

linearly betwee,1 values calculated for inventories g'reater than 5000 lbm and inventories less than 2500 lbm.

--m

_ _ _ _ _. m

_. - -. _. ~ ~

1 The feedwater flow is optionally determined in CESEC hy the following three methods:

1) matching the stean flow, 2) input table of flow rate ver-sus time, 3) automatic feedwater control on the stean generator water level.

The initial feedwater flow is assumed to match that corresponding to the power level at time zero.

The flow during an event is calculated according to the user option selected. The feedwater isolation valves are programmed to close at a specified rate of closure followinq the main steam isolation signal which is actuated on low steam generator pressure.

The feedwater enthalpy can be specified by input tables of enthalpy as a function of either power level or time.

Auxiliary feedwater flow and enthalpy are modeled using the input table option given above.

The path of the steam flow from the secondary side of the steam generator is illustrated on Figure 2 Downstream of the main steam isolation valve, the main steam lines from each steam generator are connected together at a conmon steam header. At the initial steady state, the steam flow in each stean line is detennined consistent with the reactor coolant flow rate in each steam generator loop.

During the transient calculation, the.

steam flow is determined by the turbine power demand, the operating of the secondary valves, and the break flow rate.

The steam flow *.nrough each valve is assumed to be choked. Thus, a critical flow ~correlacion (Equation 15) for steam is useJ to calculate the flow rate.

j The code simulates two main steam isolation valves, one for each main steam line. These valves are normally open and do not affect steam generator operation unless the steam generator pressure drops below a soecified e,et-1 point.

Once this occurs, the MSIVs begin to close after an input delay 0

tine. As the MSIVs close, steam flow to the turbine-and other downstrean components terminates.

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f(ODE PHYSICAL DESCRIPTIC4 1

COLD LEG UPSTRENi HALF OF IitLET PLEitUM (BEFORE FLCW MIXIffG) 2 00WilSTREAM HALF 0F I;iLET PLEt(UM (Ar'TER FLOW MIXIliG) 3 4

- BYPASS FLON 5

CORE 6

UPSTREAM HALF 0F OUTLET PLE:1UM 7

00WriSTREAM HALF 0F OUTLET PLE?iUM 8

HOT LEG 9

STE%i GENERATOR IlLET PLEiUM 10 UPSTREAM HALF OF STEAM GE?iE?ATOR TUBES 11 009tiSTRENi HALF 0F STENi GE?tEPATOR TUSES STEAM GEtiERATOR OUTLET PLE?iUM '

12 13 SMiE AS 'l' Ill OTHER STEAM GE?iERATOR LOOP 14 SAME AS 2 IN OTHER STEAM GE?iE?ATOR LOOP 15 SN4E AS 3 Itt OTHER STEAM GEttERATOR LCOP 16 SAME AS 4 Ifi OTHER STEAM GENERATOR LOOP 17 SAME AS 5 Ill OTHER STEAM GE:iERATOR LOOP 18 SAME AS 6.IN OTHER STEAM GE?iERATOR LOOP 19 SAME AS 7 It! OTHER STEAM GEiERATOR LOOP 20 SAME AS 8 '?i OTHER STEAM GE:tERATOR LOOP 21 SAME AS 9 Iti OTHER STEAM GE?tE?ATOR LOOP 22 SAME AS 10 IN OTHER STEAM GEiERATOR LOOP 23 SAME AS 11 Ill OTHER STEAM GENERATOR LOOP 24

.SMiE.AS 12JN OTHER STEAM GENERATOR LOOP 25 REACTOR VESSEL CLOSURE HEAD 26-SURGE LINE 27 PRESSURIZER F

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~ ~ ~ - ~ ~ ~ ~ - - -

=.-.-

1 PRbSURIZER TWO VOLUT*ES (STEAM AND WATER)

SAFETY VALVE CONSTANT VOLUME CONSTRAINT L

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TIME

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SHELL5:DE C3 BALTIMCRC GAS AND ELECTRIC b

CALVERT CLIFFS ::UCLEAR PC'.iER PL.M;T CC!;SERVATION EQUATIONS FOR 'IHE e

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FIGURE 7 t

~

f

i-

{

.1:1.1 l'.1:l;I;Cl:S F0lt QUESTIOtt 3

),

cFfll'Is-J O /, Cl4LC-lill;i t i;) Siruil.iLi on of a C<, abunt i rni 1.ngint e r i ng i

1;uc l eis t St reia Supply Sys t en," C-!: Pr oprietary 1;eporf ( April, 1974).

7.

C1.1:1'18-10 7, Supplement 1, "Al';S !!odel :tod i.f i cat i ons to Cl:SEC," C-E Proprietary Iteport (Se; tenber, 1974).

3.

Cl.NpD-1(17, Suppleinent 2, "ATUS 1:odels for Reactivity Feedback and 1,ffect of Pressure or, Fuel," C-E Proprietary Report, ( S e p t eiabe r, 1974).

/..

CENPD-107, Supplement 3, "ATUS !;odel !!odi ficat ions t o Cl:SSC, C-E Non-Proprietary keport (August, 1975).

5.

Cf.1:PD-107 Supple e nt

], /c.c nd:ae nt 1-P.

"A1W3 !!ade1 !!cdiIicaLicus to CI:SEC," C-E l'roprietary Keport ( :ovenber, 1975).

6.

CI:llPD-107, Supplement 4, "Alus 1:odel !!odifications to CESEC," C-E Proprietary 1:eport (December, 1975).

7.

CE".PD-107, Supplet;e nt 5, "ATWS !!odel !:odi f ications to CESEC," C-E Proprietary 1:eport (J:.:ne, 1976).

8.

Cl:11PD-107, Supplemant 6, "CESEC - Digital Simulation of a Combustion 1:ngineering iluclear Stea:a Supply System," C-E Non-Proprietary Report

( Aut.us t, 1978).

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