ML19319D691

Chapter 14 to Crystal River 3 & 4 PSAR, Safety Analysis. Includes Revisions 1-10

ML19319D691
Person / Time
Site:	Crystal River, 05000303
Issue date:	08/10/1967
From:	FLORIDA POWER CORP.
To:
References
NUDOCS 8003240674
Download: ML19319D691 (126)

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tN V TABLE OF CO?." RENTS Section Page 14 SAFETY ANALYSIS 14-1 14.1 CORE AND COOLANT BOUNDARY FROTECTION ANALYSIS 14-1 14.1.1 ABNORMALITIES 14-1 14.1.2 ANALYSIS OF EFFECTS AND CONSEQ,UENCES 14-3 l

14.1.2.1 Unco =pensated Operating Reactivity 1 Changes 14-3 l 14.1.2.2 Startup Accident 14-4 l

14.1.2 3 Rod Withdrawal Accident From Rated Power Operation 14-6 14.1.2.4 Moderator Dilution Accident lk-8 14.1.2 5 Cold Water Accident lk-10 3

(O 14.1.2.6 Loss-of-Coolant Flov 1h-11 l 14.1.2 7 Stuck-Out, Stuck-In_, or Dropped-In Control Rod 1h-13 14.1.2.8 Loss of Electric Power 1k-lh 14.1.2 9 Steam Line Failure 1h-16 14.1.2.10 Steam Generator Tube Failures ih-19 14.2 STANDBY SAFEGUARDS ANALYSIS 1h-20 14.2.1 SITUATIONS ANALYZED AND CAUSES 1h-20 14.2.2 ACCIDENT ANALYSES 1h-21 14.2.2.1 Fuel Handling Accidents 1h-21 14.2.2.2 Rod Ejection Accident 1h-22 14.2.2 3 Loss-of-Coolant Accident 1h-28 14.2.2.4 Maximum Hypothetical Accident 14-5h fl t;

14 3 REFERENCES 14-57 0003 14-1

4 J

LIST OF TABLE _S__

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Table No. Page ih-l' Abnormalities Affecting Core and Coolant Boundary 1h-1 lh-2 LUncompensated Reactivity Disturbances 1k-3 lh-3 Situations Analyzed and Causes lh-21 lk-b Reactor Building Structural Heat Capacitance Segments 1k-35 I

lk-5 Core Flooding Tank Performance Data lk-38 lb-5-1 Tabulacion of Loss-of-Coolant Accident Characteristics 1 for Spectrum of Hot Leg Rupture Sizes '

lk-k1b i

l lh-5-2 Tabulation of Loss-of-Coolant Accident Characteristics

for Spectrum of Cold Leg Rupture Sizes 1k kle lh-6 Reactor Operating Conditions for Evaluation lh-h3

14-7 Reactor Building Structure Data for Analysis of Time-Dependent Reactor Building Pressure 1k-kh lh-8 Summary of Reactor Building Pressure Analysis for Three Reactor Building Emergency Cooling Units lh-h8 l1 j 1h-9 . Sensitivity Analysis Showing the Effect of Param-eters on the Two-Hour Iodine Dose at the f Exclusion Area Boundary Following an MHA lb-55 i

4 l

r 0004 O '

g 1h-ii- ~(Revised 1-15-68)

g (j LIST OF FIGURES (At rear of Section)

Figure No. Title 14-1 Startup Accident from 10-9 Rated Power Using a 1.2% a k/k Rod Group; High Pressure Reactor Trip Is Actuated 14-2 Startup Accident from 10-9 Rated Power Using All Rods with a Worth of 9 5% ak/k; High Flux Reactor Trip Is Actuated 14-3 Peak Themal Power versus Rod Withdrawal Rate for a Startup Accident from 10-9 Rated Power 14-4 Peak Neutron Power versus Rod Withdrawal Rate for a Startup Accident from 10-9 Rated Power 14-5 Peak Ther=al Power versus Trip Delay Time for a Startup Acci-dent Using a 1.2% 6 k/k Red Group at 5 8 x 10-5 ( ak/k)/see from 10-9 Rated Power 14-6 Peak Ther=al Power versus Doppler Coefficient for a Startup Accident Using a 1.2% Ak/k Rod Group at 5 8 x 10-5 (6 k/k)/

see from 10-9 Pated Power f

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14-7 Peak Thermal Power versus Trip Dela dent Using All Rods at 5 8 x 10-4 (y Time for a Startup Acci-a k/k)/see from 1 Power 14-8 Peak Ther=al Power versus Doppler Coefficient for a Startup Accident Using All Rods at 5 8 x 10-4 (ak/k)/see from 10-9 Rated Power 14-9 Peak Pressure versus Trip Dela Using All Rods at 5 8 x 10-4 (yak/k)/see Time for from a Startup Accident 10-9 Rated Power 14-10 Peak Pressure versus Tripped Rod Worth for a Startup Accident Using All Rods at 5 8 x 10-4 (Ak/k)/see from 10-9 Rated Power 14-11 Peak Pressure versus Doppler Coefficient for a Startup Acci-dent Using All Rods at 5.8 x 10-4 (a k/k)/see from 10-9 Rated Power 14-12 Peak Pressure versus Moderator C dent Using All Rods at 5 8 x 10 gefficient forfrom (a k/k)/see a Startup Acci-10-9 Rated Power 14-13 Rod Withdrawal Accident from Rated Power Using a 1.2% Ak/k 7

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'J-Rod Group at 5 8 x 10-5 (a k/k)/sec; High Pressure Reactor Trip Is Actuated 0005 14-111

(_) FIGURES (Cont'd)

Figure No. Title 14-14 Peak Pressure versus Rod Withdrawal Rate for a Rod Withdrawal Accident from Rated Power 14-15 k_ essure versus Trip Delay Ti=e for a Rod Withdrawal Acci-dent from Rated Power Using a 1.2% a k/k Rod Group; High Pres-sure Reactor Trip is Actuated 14 16 Peak Pressure versus Doppler Coefficient for a Rod Withdrawal Accident from Rated Power Using a 1.2% A k/k Rod Group 14-17 Maximum Neutron and Thermal Power for an All-Rod Withdrawal Accident from Various Initial Power Levels 14-18 Peak Fuel Temperature in Average Rod and Hot Spot for an All-Rod Withdrawal Accident from Various Initial Power Levels 14-19 Per Cent Reactor Coolant Flow as a Function of Time after Loss of Pump Power 14-20 Minimum DNBR Which Occurs during the Coastdown for Various Initial Power Levels m

,I 14-21 Reactor System Cooling Rate for a Steam Line Break of 4 in.E 14-22 Per Cent Core Experiencing DNB as a Function of Ejected Control Rod Worth at Ultimate Power 14-23 Zr-H2 O Reaction as a Function of Ejected Control Rod Worth at Ultimate Power 14-24 Reactor Neutron Power Variation with Ejected Control Rod Worth 14-25 Reactor Themal Power as a Function of Ejected Control Rod Worth 14-26 Enthalpy Increase to Hottest Fuel Rod versus Ejected Control Rod Worth 14-27 The Effect on Reactor Neutron Power of Varying the Doppler Coefficient - Rod Ejection at 10-9 Ultimate Power 14-28 The Effect on Reactor Neutron Power of Varying the Moderator Coefficient - Rod Ejection at 10-9 Ultimate Power 14-29 The Effect on Reactor Themal Power of Varying the Doppler Coefficient - Rod Ejection at 10-9 Ultimate Power 14-30 The Effect on Reactor Ther=al Power of Varying the Moderator Coefficient - Rod Ejection at 10-9 Ultimate Power n]

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FIGURES (Cont'd)

Fleure No. Title 14-31 Reactor Thermal Power versus Trip Delay Time - Rod Ejection at Ultimate Power lk-32 Enthalpy Increase to the Hottest Fuel Rod versus Trip Delay Time - Rod Ejection th-33 LOFT Semiscale Blowdown Test No. Sh6 - Vessel Pressure versus Time 1k-34 Predicted Per Cent Mass Remaining versus Time - LOFT Test No. 546 14-34-a Neutron Power versus Time for a 36-in. ID, Double-Ended, Hot 1 Leg Pipe Rupture at Ultimate Power Without Trip 14-34-b Reactivity versus Time for a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture at Ultimate Power Without Trip 14-35 Core Flow versus Time for a 36-in. ID, Double-Ended Pipe Rupture

/~'^ 1h-36 Hot Channel Clad Surface Heat Transfer Coefficient after DNB versus Time for a 36-in. ID, Double-Ended Pipe Rupture 14-37 Reactor Vessel Water Volume versus Time for 36-in. ID, Double-Ended Pipe Rupture for 600 psig Core Flooding Tank Operating Pressure 14-38 Reactor Vessel Water Volume versus Time for 36-in. ID, Double-Ended Pipe Rupture for 400 psig and 1,000 psig Core Flooding Tank Operating Pressures 14-39 Core Flooding Tank Analysis; Maximum Clad Temperature versus Time to Quench for a 36-in. ID, Double-Ended Pipe Rupture 1k-40 Maximum Hot Spot Clad Temperature versus Maximum Heat Transfer Coefficient after DNB for a 36-in. ID, Double-Ended Pipe Rupture lk-40-a Maximum Hot Spot Clad Temperature as a Function of Time to 1 Reach DNB for a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture 1k-kl Hot Spot Clad Temperature versus Time for 36-in. ID, Double-Ended Pipe Rupture and Variable Quench Coefficient lk-kl-a Hot Spot Clad Temperature as a Function of Full-Power Seconds 1 Resulting from Void Shutdown for a 36-in. ID, Double-Ended

') Hot Leg P.'pe Rupture ik-v (Revised 1-15-68)

FIGURES (Cont'd) '

Figure No. ~ttle 1h-42 . Hot Spot Clad Temperature versus Time for 36-in. ID, Double-Ended Pipe Rupture and Variable Sink Temperature 9

1h-h3 Mass Released to Reactor Building for the Spectrum of Hot Leg Ruptures lb-kh Reactor Coolant Average Pressure for the Spectrum of Hot Leg Ruptures 14-kh-a . Hot Leg Ruptures - Reactor Vessel Water Volume versus Time 1

Including Effects of Boiloff and Injection

!' ih-kh-b Hot Spot Cladding Temperature versus Time for Spectrum of Hot

( Leg Ruptures lk-kh-c Reactor Coolant Average Pressure - Spectrum'of Cold Leg Rupture Sizes t

l 14-kh-d Cold Leg Ruptures - Reactor Vessel Water Volume versus Time

! Including Effects of Boiloff and Injection

o lk-kh-e Hot Spot Cladding Temperature versus Time for Spectrum of Cold C Leg Ruptures l

1h-kh-f Emergency Core Cooling Systems Capability j 14-h5 Reactor Building pressure versus Time in. ID, Double-Ended Rupture

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46- Reactor Building Pressure versus Time for a 36-in. ID, Double-Ended Rupture with and without Cooling of the Recirculated Spray Water a

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0008 1h-v (a) '(Revised 1-15-68)

O FIGURES (Cont'd)

V' Figure No. Title 14-47 Reactor Building Atmosphere and Su=p Coolant Temperatures Following a 36 in. ID, Double-Ended Rupture 14-48 Reactor Building Pressure versus Time after Rupture - 8 5 ft 2 Rupture 14-49 Reactor Building Pressure versus Time after Rupture - 3 0 fte Rupture 14-50 Reactor Building Pressure versus Ti=e after Rupture - 2.0 ft2 Rupture 14-51 Reactor Building Pressure versus Time after Rupture - 1.0 ft2 Rupture 14-52 Reactor Building Pressure versus Time after Rupture - 0.4 ft2 Rupture 14-53 Reactor Mu11 ding Energy Inventory for 36 in. ID, Double-Ended Rupture 14-54 Reactor Building Energy Inventory for 3 0 fte Rupture 14-55 Reactor Building Vapor and Su=p Temperatures for 36 in. ID, Double-Ended Rupture as a Function of Time after the Rupture 14-56 Reactor Building Vapor and Sump Temperatures as a Function of Time after Rupture - 3 0 fta Rupture 14-57 Criterion 17 Case for 36 in. ID, Double-Ended Rupture 14-58 Reactor Building Zr Reaction Capability for 55 psig Design Pressure 14-59 Thyroid Dose from Loss-of-Coolant Accident Hour, 24-Hour, and 30-Day Doses 14-60' Maximum Hypothetical Accident Thyroid Dose Assuming Fission Product Release per TID-14844 14-61 Integrated Direct Dose 911owing MHA with 3-1/2 Foot Reactor Building Wall Thickness h 0009 a

14-vi l

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(, 14 SAFE 1"I ANALYSIS 14.1 CORE AND COOLANT BOUNDARY PROTECTION ANALYSIS 14.1.1 ABNORMALITIES In previous sections of this report both nomal and abnomal operations of the various systems and components have been discussed. This section su=marizes and further explores abnomalities that are either inherently terminated or require the nomal protective systems to operate to maintain integrity of the fuel and/or the reactor coolant system. These abnor=alities have been evalu-ated for rated power of 2,452 Wt. Whenever a fission product release to the environment occurs, the release is based upon the fission product inventory associated with the ultimate reactor core power level of 2,5k4 Wt. Fission product dispersion in the atmosphere is assu=ed to occur as predicted by the dispersion models developed in 2 3 Table 14-1 su=marizes the potential ab-normalities studied.

Table 14-1 Abnormalities Affecting Core and Coolant Boundary Event Cause Effect Uncompensated Oper- Fuel depletion Reduction in reactor system p!

v ating Reactivity or xenon build- average temperature. Automatic Changes up, reactor trip if uncompensated.

No equipment damage or radiolog-ical hazard.

Startup Accident Uncontrolled Power rise teminated by nega-rod (*) with- tive Doppler effect, reactor drawal. trip from short period, high reactor coolant system pres-sure, or overpower. No equip-ment damage or radiological hazard.

Rod Withdrawal Acci- Uhcontrolled Power rise teminated by over-dent at Rated Power rod withdrawal. power trip or high pressure trip. No equipment damage or radiological hazard.

(*)Cono ol rod, rod, and control rod assembly (CRA) are used interchange-ably in this section and elsewhere in the report.

A control rod grctp consists of a symmetrical arrangement of four or more control red assemblies. See 7.2.2.1.1.

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Table 1h-1 (Cont'd)

Event Cause Effect Moderator Dilution Equipment mal- Slow change of power teminat-Accident function or ed by reactor trip on high operator error. temperature or pressure. Dur-ing shutdown a decrease in shutdown margin occurs, but criticality does not occur.

No radiological hazard.

Ioss of Coolant Flow Mechanical or None. Core protected by reac-electrical tor low-flow trip. No radio-failure of re- logical hazani.

actor coolant Pump (s).

Stuck-Out cr Stuck- Mechanical or None. Suberiticality ct.n be In or Dropped-In electrical achieved if one rod is stuck-Control Rod failure, out. If stuck-in or dropped-in, continued operation is permitted if effect on power peaking not cevere. No radio-logical hazard.

Ioss of Electric Miscellaneous Possible power reduction or re-Power faults. actor trip depending on condi-tion. Redundancy provided for safe shutdown. Radiological hazard within limits of 10 CFR 20.

Steam Line Failure Pipe failure. Reactor automatically trips if rupture is large. No damage to reactor system. Integrated dcces at exclusion distance are 0.002 rem whole body and 0 53 rem thyroid. Radiologi-cal hazard is within limits of 10 CFR 20.

Steam Generator Tube Tube failure. Reactor automatically trips if Failures leakage exceeds normal makeup capacity to reactor coolant system. Integrated doses at exclusion distance are 0.69 rem whole body and 1.0 x 10-4 rem thyroid. Radiological hazard is within limits of 10 CFR 20.

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14.1.2 ANALYSIS OF EFFECTS AND CONSEQUENCES 14.1.2.1 Unccmpensated Operatira Reactivity Changes 14.1.2.1.1 Identification of Cause During normal operation of the reactor, the overall reactivity of the core ch u ges because of fuel depletion and changes in fission product poison concentration. These reactivity changes, if left either uncom-pensated or overcompensated, can cause operating limits to be exceeded.

In all cases, however, the reactor protective system prevents safety limits from being exceeded. No damage occurs from these conditions.

14.1.2.1.2 Analysis and Results During nomal operation, the automatic reactor control system senses any reactivity change in the reactor. Depending on the direction of the re-activity change, the reactor power increases or decreases. Correspond-ingly, the reactor coolant system average temperature increases or de-creases, and the automatic reactor control system acts to restore reac-tor power to the power de=and level and to reestablish this temperature at its set point. If manual corrective action is not taken or if the automatic control system malfunctions, the reactor coolant system aver-age temperature changes to compensate for the reactivity disturbance.

Table 14-2 summarizes these disturbances.

Table 14-2 Uncompensated Reactivity Disturbances Maximum Pate of Average Reactivity Rate, Te ,trature Change Cause (ak/k)/sec (Uncorrected), F/sec Fuel Depletion -6 x 10-9 -0.0006 Xenon Buildup -3 x 10-8 -0.003 These results are based on +6 x 10-3 ( ak/k)/F moderator coefficient and

-1.14 x 10-5 ( ak/k)/F Doppler coefficient. The nominal value of +6 x 10-5(ak/k)/Fisrepresentativeofthemoderatorcoefficientatthebe-ginning of core life for an equilibrium cycle. 'Ihis value is also valid at BOL,for the first cycle after 15 days. A higher value [+10 x 10-5 (ak/

k//]" exists at the start of the first core cycle. However, the -ffect of l1 this slightly higher value has been shown to be of minor importaace by the evaluation of the sensitivity of the reactor to moderator coefficient variations. These reactivity changes are extremely slow and allow the operator to detect and compensate for the change.

O 0012 g l 14-3 (Revised 1-15-68) l

1L.1.2.2 Startun Accident Jh.l.2.2.1 Identification of Cause The objective of a normal startup is to bring a suberitical reactor to the critical or slightly supercritical condition, and then to increase power in a controlled manner until the desired power level and system operating temperatures are obtained. During a startup, an uncontrolled reactivity addition could cause a nuclear excursien. This excursion is terminated by the strong negative Doppler effect if no other trotective action operates.

The following design provisions minimize the possibility of inadvertent continuous rod withdrawal and linit the potential power excursion:

a. The control system is designed so that only one control rod group can be withdrawn at a time, except that there is a ?5 per cent overlap in travel between two successive rod groups.

This overlap occurs at the minimum vorth for each group since one group is at the end of travel and the other is at the be-ginning of travel. Tne maximum worth of any single control rod I

group is 1.25 Ak/k when the reactor is critical.

b. Control rod withdrawal rate is limited to 30 in./ min. lI

c. A short-period withdrawal stop and alarm are provided in the source renge.

d. A short-teriod withdrawal stop, alarn, and trir are provided in the intermediate range.

e. A high flux level and a high pressure trip are t rovided in the power range.

The reactor protective systen is designed to limit (a) the reactor ther-mal power to llh per cent of rated power to prevent fuel danace, and (b) the reactor coolant system pressure to 2,515 psia.

lk.l.2.2.2 Methods of Analysis An analog model of the reactor core and coolant sys' .m was used to deter-mine the characteristics of this accidene. This analog model used full reactor coolant flow, but no heat transfer out of the system and no sprays in the precsurizer. The rated-power Doppler coefficient (-1.lh x 10-5 ( Ak/k)/F] was used although the Doppler is much larger than this for the principal part of the transient. The rods were assumed to be mov-ing along the steepest part of the rod-vorth vs rod-travel curve. A re-actor trip on short period was not incorporated in the ana]ysia. The nominal values of the principal parameters used were: 0.3 see trip delay,

+6 x 10-5 (Ak/k)/F moderator coefficient, and -1.lb x 10-5 (Ak/k)/F Doppler coefficient. The total worth of all the control rods inserted 1 into the reactor core following any trip is 8.45 Ak/k without a stuck e . . . W. a .. , . , . - . y . :. .- e. .' s ..., - . .w . : . . . ; . r, p.s .: ..>., ,

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• n'%3 42 " *

]h b (Revised 7-15-69) 0013 -

- Centrol rod, er 5.h5 A2/k (the nominal case in this study) with a stuck 11 rod.

1h.1.2.2.3 Results of Analysis Ficure lb-1 shows the results of withdrawing the naximum worth control rod group from 1 per cent suberitical. This group is vorth a maximum 17 value of 1.2" (Sk/k)/sec. Theak/k and results Donpler in a reactivity effect begins addition to slow the rate neutron of 5(8)x 10-5 power rise, but the heat to the coolant increases the pressure past the trin point, and the transient is terminated by the high r.ressure trip.

Figure lh-2 shows the results of withdrawing all 61 control red assemblies (with a total worth of 10.0", Ak/k) at 1 pe" cent suberitical. This results 7 in a reactivity addition rate of 5.8 x 10-k ( Ak/k)/sec. About 15.3 see after paasing through criticality, the neutron power peaks at lh7 per cent , where the power rise is stopped by the negative Doppler effect. The high neutron flux trip takes effect 0.25 see after the peak power is reached and terninates the transient. The peak thermal heat flux is only 16 per cent of the rated power heat flux.

A sensitivity analysis was performed on both of these startup accidents to determine the effect of varying several key parameters. Ficures lk-3 through lb-6 show tynical results for the single group, 1.2" Ak/k start-up accident.

g Figures lb-3 and 1h-h shov the effect of varying the reactivity addition rate on the peack thernal power and peak neutron power. This reactivity rate was varied from one order of maenitudy below the single rod group case (1.25 Ak/k) to nore than an order of magnitude above the rate that l7 reeresents all rods (10.05 AN/k) being withdrawn at once. The slower rates - up to about 0.5 x 10-3 (ak/k)/sec - vill result in the pressure l1 trip being actuated, whereas only the very fast rates actuate the high neutron flux level trip.

Figures lk-5 and lb-6 show the peak thermal power variation as a function of a vide range of trin delay times and Doppler coefficients for the 1.2%

ak/k rod group. Only a small change in power is noted. Figures lb-T and 1h-8 are the corresponding results fron the withdrawal of all r _h (10.0', Ak/k). Since this transient inserts reactivity an order of magni-tude faster than does the single control rod group case, there is con- l1 siderably nore variation in the peak tnermal power over these vide ranges.

At high values of the Doppler coefficient, the neutron power rise if vir-tually stopped before reaching the high flux trip level. Reactor power generation continues until sufficient energy is transferred to the reac-tor coolant to initiate a high pressure trip. This results in a higher peak thermal power.

Figures lb-0 through lb-12 show the peak pressure response to variations in several key parameters for the case where all rods are withdrawn. It is seen that the safety valve is ocened when these paraneters are changed considerably from the nominal values, excent in the case of the moderator

(' eutren power is defined as the total sensible enerrv release from fission.

0014 1h-5 (Revised 7-15-69) ,,,

coefficietit which has little effect because of the snort duration of the transient. Again for a high Doppler coefficient, ths high pressure trip g is 2elied upon. W Hone of these postulated startup accidents, except for reactivity addi-tion rates greater than 2 x 10-3 ( Ak/k)/see, which is three tim %

Breater than for withdrawal of all rods at once, causes a thera.1 power peak in excess of 40 per cent rated power or a nominal fuel rod average temperature greater than 1,715 F. The nominal 1.2% Ak/k r >d group with-drawal causes a peak pressure of 2,515 psia, the safety valve set point.

The capacity of the safety valves is adequate to handle the maximum rate of coolant expansion resulting from this startup accident. The 10.0% l1 A k/k withdrawal - using all 69 rods - causes a peak pressure of only 2,465 psia because the flux trip is actuated prior to the pressure trip.

It is concluded that the reactor is completely protected against any startup accident involving the withdrawal of any or all control rods, since in no case does the themal power approach 114 per cent, and the peak pressure never exceec.s 2,515 psia.

14.1.2 3 Rod Withdrat.al Accident From Rated Power Operation 14.1.2 3 1 Identification of cause A rod withdrawal accident presupposes an operator error or equipment failure which results in accidental withdrawal of a control rod group while the reactor is at rated power. As a result of this assu=ed accident, the power level increases; the reactor coolant and fuel rod temperatures increase; and if the withdrawal is not terminated by the operator or pro-tection system, core damage would eventually occur.

The following provisions are made in the design to indicate and terminate this accident:

a. High reactor outlet coolant temperature alams.

b. High reactor coolant system pressure alams.

c. H16h pressurizer level alams. l1 1

d. High reactor outlet coolant temperature trip.

1

e. High reactor coplant system pressure trip.

f. HiS h power level trip. I l

14.1.2 3 2 Methods of Analysis )

An analog computer model was used to detemine the characteristics of this accident. A complete kinetics =odel, pressure model, average fuel rod model, steam demand model with turbine coastdown to 15 per cent of rated load, coolant transport model, and a si=ulation of the instrumen-

's tation for pressure :and ; flux' trip were included.> .Thesinitial conditions; , -. , <

. . . u.J vere normal rated power operation without automatic control. Only the I

lb-6 (Revised 1-15_ss; 00ll g n

moderator and Doppler coefficient of reactivity were used as feedback. The nominal values used for the nain parameters were 0.3 see trip delav tine,

-1.1h x 10-3 (Ak/k)/F Doppler coefficient, +6 x 10-5 (Ak/k)/F moderator co-e fficient , 5.8 x 10-5 (Ak/k)/sec reactivity insertion race, and 1.25 Ak/k l7 control rod group worth. The total vorth in all the centrol rods inserted into the reactor core following any trip is 8.h5 Ak/k without a stuck control l1 rod, or 5.h% Ak/k (the nominal value used) with a stuck rod.

(Sentence deleted.) l7 The reactor protection system is designed to limit (a) the reactor power to llh per cent of rated power to prevent fuel damage, and (b) the coolant system pressure to 2,515 psia to prevent reactor coolant system damage.

14.1.2.3.3 Results of Analysis Figure lh-13 shows the results of the nominal rod withdrawal from rated cover using the 1.25 Ak/k rod group at 5.8 x 10-5 ( Ak/k)/sec. "'he transient is terminated by a high pressure trip, and reactor power is limited to 109 per cent, much less than the design overpower of llh per cent of rated power. The changes in the parameters are all quite small, e.g., 5 F average reactor cool-ant temrerature rise and 200 psi system pressure change.

A sensitivity analysis of inportant parameters was nerformed around this nomi-nal case, and the resultant reactor coolant system cressure responses are shown in Figures lh-lh through lh-16.

73 Figure lk-lh shows the pressure variation for a very wide range of rod with-drawal rates - nore than an order of magnitude smaller and greater than the nominal case. For the very rarid rates, the neutron flux level trip is actu-ated. This is the primary protective device for the reactor core; it also pro-tects the system against high pressure during fast rod withdrawal accidents.

The high pressure trip is relied upon for the slower transients. In no case does t e thernal power exceed 109 per cent rated power.

An analysis has been perforned extending the evalustion of the rod withdrawal accident for various fractional initial power levels up to rated power. This evaluation has been performed assuming simulated withdrawal of all 61 control rods 7 with a reactivity addition rate of 5.8 x 10-4 (Ak/k)/sec. "his rate is a factor of ten higher than used in the cases evaluated at rated power. The results of this analysis are shown in Fip. ires lk-17 and 14-18.

As seen in Figure 1b-17 the neak thermal power occurs for the rated pcwer case and is well below the maximum design power of llh per cent. The teak neu-tron power for all cases is approxinately 117 per cent of rated power and rep-resents a slight overshoot above the trip level of llh ner cent. Figure lh-18 shows that the maximum fuel temperature reached in the average rod and the hot spot are well below melting. Even in the most severe case at rated power, the average fuel te=rerature only increases bv 2A F. It is therefore readily concluded that no fuel dar. age vould result from sirultanecus all-rod withdrawal from any initial power level.

D 0016 lk-7 (Revised 7-15-69) rmy --

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Figures 14-15 and 14-16 show the pressure response to variations in the trip delay time and Doppler coefficient. For the higher vaJues of the Doppler coef-ficient, the pressure trip is always actuated, and, therefore, the pressure levels off. g This analysis shows that the high pressure trip and the high flux level trip adequately protect the reactor against any rod withdrawal accident from rated power.

14.1.2.4 Moderator Dilution Accident 1h.1.2.4.1 Identification of Cause The reactor utilizes boric acid in the reactor coolant to control excess reac-tivity. The boron content of the reactor coolant is periodically reduced to compensate for fuel burnup. The dilution water is supplied to the reactor cool-ant system by the makeup and purification system. This system is designed with ceveral interlocks and alarms to prevent i= proper operation. These are as fol-lows:

a. Flov of dilution water to the makeup tank must be initiated by the operator. The dilution water addition valve can be opened only when the control rods have been withdrawn to the preset position (95 per cent) and the ti=ing device to limit the integrated flow has been set. Dilution water is added at flow rates up to 70 gpm.

b. Flov of dilution water is automatically stopped when either the flow has integrated to a preset value or when the rods have been inserted to a preset position (at about 75 per cent full stroke).

ll

c. A varning light is on whenever dilution is in progress.

The makeup and purification system normally has one pump in operation which 2 supplies up to 70 gpm to the reactor coolant system and the required flow to the reactor coolant pump seals. Thus, the total makeup flow available is limited to 70 gpm unless the operator takes action to increase the a=ount of makeup flow to the reactor coolant system. When the makeup rate is greater than the maximu= letdown rate of 70 gpm, the net water makeup will cause the pressurizer level control to close the makeup valves.

The no=inal moderator dilution event considered is the pu= ping of water with zero boron concentration from the makeup tank to the reactor coolant system by the makeup pump.

It is also possible, however, to have a slightly higher flow rate during tran-cients when the system pressure is lover than the accinal value and the pres-curizer level is below normal. This flov =ight be as high as 100 gpm.

In addition, with a combination of multiple valve failures or maloperations, plus more than one makeup pump cperating and reduced reacter ecolant system pressure, the resulting inflow rate can be as high as 500 gpm. This consti-tutes the maximum dilution accident. A reacter trip would ter=inate unborated water addition to the makeup tank, and total flov into the coolant system would be teroinated by c high pressurizer level.

...s. m . .. . ,, , .m -'  : -: -n . ~~i ~*.~~' .' 'n'.=" P N pg ., 14-8 (Fevised 2-7-68) 0017 g t _

b'O The criteria of reactor protection for this accident are

a. The reactor power vill be limited to less than the design over-power of 114 per cent rated power to prevent fuel damage,

b. The reactor protection system vill limit the reactor coolant sys-tem pressure to less than the system design pressure of 2,500 psig.

c. The reactor minimum suberiticality margin of 1% ak/k vill be main-tained.

d. Administrative procedures vill be imposed to monitor and control 1 the relationship of control rod regulating group patterns and boron concentrations in the reactor coolant over the operating life of the core.

[/

N._

t O 0018 1h-9 (Revised 1-15-68)

~=.4 e e .--i.mw--- . .

IL.1.-.t.2 Analysis and Pecults The reactor is assumed to be operating at rated power with an initial boron l1 concentration (1,800 ppm) in the reactor coolant system. The dilution water is uniformly distributed throughout the reactor coolant volume. Uniform di-lution resc.its from a discharge rate of 70-500 gpm into a reactor coolant f1ci of 88,000 gpm. A change _ concentration of 100 ppm produces a 1% Ak/k l1 reactivity change. The effects o. these three dilution rates on the reactor are as follows:

Average Reactor Dilution Water Reactivity Rate, Coolant Syste=

Flow, 6 a (ak/k)/sec Temp. Change, F /sec 70 - 2.5 x 10-6 0,3 1 100 4 3.6 x 10-6 03 500 + 1. 8 x 10-3 0.h The fastest rate of dilution can be handled by the automatic control sys-tem, which would insert rods to maintain the power level and reactor cool-ant system te=perature. If an interlock failure occurred while the reac-tor was under manual control, these reactivity additions vould cause a high reactor coolant temperature trip or a high pressure trip. In any event the thermal power vill not exceed llh per cent rated power, and the syste= pressure vill not exceed the design pressure of 2,500 psig. There-fore moderator dilution accidents will not cause any damage to the reac-tor system.

h During refueling or maintenance operations when the reactor closure head has been rc=oved, the sources of dilution water makeup to the makeup tank--

and therefore to the reactor coolant syste:--are Iceked closed, and the makeup pumps are not operating. At the beginning of core life when the boror. concentration is highest, the reactor is about 9 5% ak/k suberiti-cal with the maximum vorth rod stuck out. To demonstrate the ability of the reactor to accept moderator dilution during shutdevn, the consequences of accidentally filling the makeup tank with dilutien water and starting the =akeup pumps have been evaluated. 'Ihe entire water volume fro = the makeup tank could be pu= ped into the reactor coolant system (assu=ing only the coolant in the reactor vessel is diluted), and the reactor would still be 6.5% ok/k suberitical, 1k.1.2 5 Cold Water Accident The absence of individual loop isolation valves eliminates the potential scurce of cold water in the reactor coolant system. Therefore, this ac-cident is not credible in this reactor.

9

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q;gg

• 1h-10 (Fevised 1-15-8

• EN rt c ,g;

_ - r 0019

'd 14.1.2.6 Icss-of-coolant Flov 14.1.2.6.1 Identification of Cause A reduction in the reactor coolant flow rate occurs if one or more of the reactor coolant pu=ps should fail. A pumping failure can occur from me-chanical failures or from a loss of electrical power. With four indepen-dent pumps available, a mechanical failure in one pu=p vill not affect op-eration of the others.

Each reactor coolant pu=p receives electrical power from one of the two electrically separate busses of the 6,900 volt system discussed in 8.2.2 3 Loss of a unit auxiliary transfor=er to which the 6,900 volt busses are nomally connected vill initiate a rapid transfer to the startup transformer source without loss of coolant flow. Faults in an individual pu=p motor or its power supply could causy a reduction in flow, but a complete loss of flow is extremely unlikely.

In spite of the low probability of a complete loss of power to all reac-tor coolant pumps, the nuclear unit has been designed so that such a fail-ure would not lead to core damage.

The reactor protection criterion for lon-of-coolant flow conditions start-ing at rated power is that the reactor e 'e vill not reach a Departure from Nucleate Boiling Ratio (DNER) smaller than the DNBR in the hot channel at

,, the steady state design overpower. This corresponds to a DNER of 1 38 at 114 per cent rated power (Table 3 '-) .

(v) 14.1.2.6.2 Methods of Analysis The loss-of-coolant-flow accident is analyzed by a combination of analog and digital computer programs. Analog simulation is used to detemine the reactor flow rate following loss of pu= ping power. Reactor power, coolant flow, and inlet temperature are input data to the digital program which determines the core themal characteristics during the flow coastdova.

The analog model used to detemine the neutron power following reactor trip includes six delayed neutron groups, control rod worth and rod insertion characteristics, and trip delay time. The analog model used to detemine flow coastdown characteristics includes description of flow-pressure drop relations in the reactor coolant loop. Pu=p flow characteristics are de-temined from manufacturers' zone maps. Flow-speed, flow-torque, and flow-head relationships are solved by affinity laws.

A transient, thermal-hydraulic, B&W digital computer program is ured to compute channel DNER continually during the coastdown transient. System flow, neutron power, fission product decay heat, and core entering enthal-py are varied as a function of time. The program maintains a transient inventory of stored heat which is determined from fuel and clad tempera-tures beginning with the initial steady state conditions. The transient cgre pressure drop is detemined for average channel conditions. The representative hot channel flows and corresponding DNBR are obtained by l ) using the average core pressure drop. The hot channel DNER as a function of time is compared with the design DNBR at maximum overpower to deter-mine the degree of heat transfer margin. I Z '

14-11 .

The loss-of-coolant-flow analysis has been carried out in the power range be-tween 102 and 114 per cent rated power. Conditions utilized in the analysis are as follows:

a. Initial core inlet temperature for given power level is assumed to be plus 2 F in error,

b. Initial system pressure is assumed to be minus 65 psi in error.

c. Trip delay time, i.e., time for sensor detection for low flow condi-tion until initial downward movement of control rod, is 300 milli-seconds.

d. The per cent of initial reactor power as a function of time after loss of pumps is as shown in Figure 3-6.

e. The pump inertia is 70,000 lb-ft 2.

14.1.2.6 3 Results of Analysis The results of this analysis show that the reactor can sustain a loss-of-coolant-flow accident without damage to the fuel. The results of the evalua-tion are presented in Figures 14-19 and 14-20. Figure 14-19 shows the per cent reactor flov as a function of time after loss of all pump power. Figure 14-20 shows the minimum DNER's which occur during the coastdown for various initial power levels. The degree of core protection during coastdown is indi-cated by co= paring the DNBR for the coastdown with the design value of 1 38 at 114 per cent rated power. This DNBR (138) in the representative hot channel $

corresponds to a 99 per cent confidence that 99 5 per cent of the core vill not experience a departure from nucleate boiling under steady state conditions at the design overpower (3 2 31).

Under normal conditions, the maximum indicated reactor power level from which a loss-of-coolant-flow accident could occur is,102 per cent rated power (as indi-cated by reacter instrumentation). This power level represents an allowance of plus 2 per cent rated power for transient overshoot. This power level also rep-resents the maximum power demand that vill be permitted to the reactor control system. The 102 per cent rated power is an instrument-indicated value and is subject to the following maximum errors: (a) 2 per cent heat balance and (b) ik per cent nuclear instrumentation. The true power level could be as high as 108 per cent at 102 per cent indicated power. As shown in Figure 1h-20, how-ever, the DNBR at 108 per cent is 1.hh, which is significantly larger than the design DNBR.

The reactor coolant system is capable of providing natural circulation flow after the pu.ps have stopped. The natural circulation characteristics of the reactor coolant system have been calculated using conservative values for all resistance and form loss factors. No voids are asoumed to exist in the core or reactor outlet piping. The following tabulation and Figure 9-10 show the natural circulation flow capability as a function of the decay heat generation.

This material is presented in greater detail in 14.1.2.8 3

,m.e .< . . . . ::., .

. .su-

. ?- .. a , . . ,

.sc 14-12 -

p

Time After Decay Heat Natural Circulation Flev Required for Loss of Core Power, Ccre Flow Available, Heat Pe::! oval, Power, sec  %  % Full Flow To Full Flow 0 36 x 102 5 4.1 23 2.2 x 102 3 '.3 1.2 1.2 x lb 1 1.8 0 36 1.3 x 10 5 1/2 1.2 0.20 w

The flows above provide adequate heat transfer for core cooling and de-cay heat removal by the reactor coolant system.

The reactor is protected against reactor coolant pump failure (s) by the protective system and the ir.tegrated centrol system. The integrated con-trol system initiates a power reduction on pump failure to prevent reac-tor power from exceeding that permissible for the available flow. The reactor is tripped if insufficient reacter coolant flow exists for the power level. The operating limits for less than four pumps in operation have been presented in 4 3 7 14.1.2 7 Stuck-Out, Stuck-In, or Drcpped-In Control Rod 14.1.2 7 1 Identificatier. of Cause

(~) The control rod drives have been described in 3 2.4.3 The results of v continuous control rod withdrawal have been analyzed in 14.1.2.2 and 14.1.2 3 'In the event that a control rod cannot be moved because of electrical faults or mechanical seizure, localized power peaking and sub-critical margin must be considered.

14.1.2 7.2 Analysis and Resulte Adequate hot suberitical margin is provided by requiring a suberiticality of 1% a k/k suberitical witn the control rod of greatest worth fully with-drawn from the core. The nuclear analysis reported in 3.2.2 demonstrates that this criterion can be satisfied.

In the event that an unmovable control rod is partially or fully inserted in the core or a single rod is dropped during cperation, its location and effect on lo:al power distribution determine whether continued power op-eration is permissible. The location of a stuck rod in the core will be studied further to define permissible conditions of operation. The cri-teria for these studies are (a) operation with a stuck rod will not in-crease the DNB probabi?ity above the probability specified for design conditions, and (b) a ht t suberitical margin of 1% A k/k will be main-tained with the stuck rod in its inoperative position and the operating rod of greatest reactivity vorth in the fully withdrawn position.

If a control rod is dropped into the core during power operation,. the same consideration of localized power peaking as for a stuck rod will

^ apply.

w.m

'N.' 14-13

- 0022  %-

1k.1.2.8 Loss of Electric Power 1k.1.2.8.1 Identification of Cause The Crystal River Plant Units 3 and k are designed to withstand the effects of loss of electric load or electric power. Two types of power losses are con-sidered:

a. A " blackout" condition, caused by severe interconnected grid upset.

b. A hypothetical condition resulting in a ec=plete loss of all Plant power.

The reactor protection criteria for these conditions are that fuel da= age vill not occur from an excessive power-to-flow ratio and that the reactor coolant syste= pressure vill not exceed design pressure.

1k.1.2.8.2 Results of " Blackout" Conditions Analysis The net effect of a " blackout" condition on the nuclear units would be opening of all 230 and 500 kv breakers, thus disconnecting the Plant from the entire transmission system. When this occurs on the r.uclear units, a runback signal on the integrated = aster controller car'es an automatic power reduction to 15 per cent reactor power. Other actions 2 at occur are as follows:

a. All vital electrical loads, incluaing reactor coolant pu=ps, con-denser circulating water pumps, condensate and condensate booster pu=ps, and other auxiliary equipment, will continue to obtain power from each unit's generator. Feedvater is supplied to the steam gen-erators by stea=-driven feed pumps.

b. As the electrical load is dropped, the turbine generator accelerates and closes the governor valves, and the reheat stop and interceptor valves. The unit's frequency will peak at less than the overspeed trip point and decay back to set frequency in ko-50 sec.

c. Following closure of the turbine governor valves and the reheat stop and interceptor valves, steem pressure increases to the turbine bypass valve set point and may increase to the stea= system safety valve set point. Steam is relieved to the condenser and to the atmosphere.

Steam venting to the atmosphere occurs for about 2 min. following blackout frc= 100 per cent rated power until the turbine bypass can handle all excess steam generated. The capacity of the modulating turbine bypass valve is 15 per cent of the valves vide open (WO) steam flev, and that of the safety valves is 100 per cent of WO stea=

flow. The first safety valve banks are set at 1,050 psig with addi-tional banks set at pressures up to 1,10k psis (5 per cent above design pressure as allowed by code). Steam venting pemits energy removal from the reactor coolant system to prevent a high pressure reactor trip.

The initial power runback is to 15 per cent power which is greater than the unit's auxiliary load. This allows sufficient stea= flow for regulating turbine speed control. Excess power above the unit's

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... , - v. ,

u hk 1k-1k 0023 g

i turbine speed control. Excess power above the unit auxiliary load v

is rejected by the turbine bypass valve to the condenser.

d. During the short interval while the turbine speed is high, the vital electrical loads connected to the unit generator vill undergo speed increase in proportien to the generator frequency increase. All motors and electrical gear so connected are designed for the in-creased frequency.

c. After the turbine generator has been stabilized at auxiliary load and set frequency, the Plant operator may reduce reactor power to the auxiliary load as desired.

The blackout accident does not produce any fuel da= age or excessive pressures on the reactor coolant system. There is no resultant ra-diological hazard to Plant operating personnel or to the public from this accident, since only secondary system steam is discharged to the atmosphere.

Unit operation with failed i'uc1 and steam generator tube leakage is shown to be safe by the analysis presented in 11,1.2 5 2 and 14.1.2.10. For the same conditions, the steam relief accompanying a blackout accident would not change the whole body dose. The whole body dose is pri=arily due to the release of Xe and Kr. Release of these gases is not increased by the steam relief be-cause even without relief, all of these gases are released to the atmosphere

- through the condenser vacuum pump exhaust. The rate of reltase of iodine dur-

) in 10gtheapproximately2minofreliervouldbeincreasedby ilmost a factor of

, because the iodir.e is released directly to the atmosphere rather than through the condenser and Plant vents. However, the quantity released during this short time is small, and it would be less than 0.03 MPC at the h,400 ft cxclusion distance.

l1 1L.l.2.8 3 Analysis Fesults of Cc=plete Loss of All Plant Power The second power loss considered is the hypothetical es>e where all Plant power except the Plant batteries is lost. The sequence of eventc and the evaluation of consequences relative to this accident are given below:

a. A loss of power results in gravity insertion of the control rods.

b. The steam generator safety valves actuate after the turbine trips and prevent excessive temperatures and pressures in the reactor coolant system,

c. The reactor coolant system flow decays without fuel damage occurring.

Decay heat removal after coastdown of the reactor coolant pumps is provided by the natural circulation characteristics of the system.

This capability is discussed in the loss-of-coolant-flow evaluation (14.1.2.6).

d. A turbine-driven emergency feedvater pump is provided to supply feedvater any time the main feed pumps cannot operate. The e=er-gency feed pump takes suction from the condenser hotvell and the ensate storage. The emergency pump supplies feedvater to the

'~

JM7 14-15 (Revised 1-15-68)

steem generators. The e=ergency feed pu=p is drive- by steam fro =

either or both stea= generators.

The controls and auxiliary syste=s for the emergency feed pump oper-ate on d-c power frc= the Plant battery.

A recirculation line from the emergency pump discharge back ts the condenser is provided to permit periodic testing

e. The condenser hotvv.1 and the condensate storage tank provide cool-ing water in the unlikely event that all power is lost. The =ini-mu= can'iensate inventory is 200,000 gal. This inventory provides sufficient water for decay heat cooling (assuming infinite irradia-tion at 2,5hk !Gt) for a period of approxi=ately one day.

The features described above permit decay heat cooling of the nuclear unit for an extended period of time following a complete loss of electric power.

The foregoing evaluation demonstrates the design features incorporates in the design to sustain loss of power conditions with just tRe Pirnt batteries to operate syste= controls. I==ediate operation of the emergency ieedvater pump is not of critical nature. The reactor can sustain a complete electric power loss without e=ergency cooling for about 25 min before the steam volume in the pressurizer is filled with reactor coolant. These 25 =in are derived ; fol-lows:

a. Steam generators evaporate to dryness 10 =in

b. Pressurizer safety valves open 5

c. Trassurizer fills with water (due to 10 reactor coolant syste= expansion) 25 min Beyond thir time reactor coolant will boil off, and an additional 90 min vill have elapsed before the boiloff vill start to uncover the core. The emergency feedvater pump can be actuated within this period of time. Accordingly, core protection is insured for the unlikely condition of total loss of Plant elec-tric power.

1k.1.2.9 Steam Line Failure 1L.1.2.9.1 Identification of Cause Analyses have been performed to determine the effect's and consequences of loss of secondary coolant due to failures in the steam lines between the steam g'.nerators and the turbine.

The criteria for Plant protection and the release of fission products to the environment are as follows:

a. The reacter shall trip and remain suberitical without boron addition v.s : 4W ~ until 'a' dentro11ed rat'e"er.'syste= cooldown cah be effe'eted. '

• O pn 2'- 6

. M _

0025

( ) b. The potential environmental consequences from rad 2 activity in the

'd secondary coolant system shall not exceed those specified by 10 CFR 20.

14.1.2 9 2 Analysis and Results The rate of reactor system i:ooling following a steam line break accident is a function of the area of the failure and the steam generator water inventory available for cooling. The steam generator inventory increases with power icvel. The inventory at rated power is 46,000 lb and decreases linearly to 20,000 lb at 15 per cent power. The steam line break accident analysis is per-fomed at . ultimate power in order to determine maximum cooling and inventory l1 release effects.

The 1:Tediate effect of any steam line break accident is a reduction in steam pressure and a reduction in steam flow to the turbine. These effects initially cause the reactor control system to act to restore steam pressure and load generation.

A steam line rupture of a small area causes a relatively slow decrease in steam pressure. This places a demand on the control system for increased feed-water flow. In addition, the turbine control valves will open to maintain power generation. Increased feedvater flow causes the average reactor coolant temperature to decrease, and the resulting te=perature error calls for control rod withdrawal. The limiting action in this condition is the 102 per cent

,e s limit en power demand to the rod drive control system. If the moderator tem-(,) perature coefficient of reactivity is small or slightly positive, the reactor power vill decrease when the control system reaches the power demand limit be-cause of continuing temperature decrease. The reactor will then trip on low reactor coolant system pressure. A reactor trip will initiate a reduction in the feedvater flow to the steam generators.

When the moderator temperature coefficient is negative, the reactor power vill tend to increase with decreasing average coolant temperature. This vill cause control rod insertion to limit reactor power to 102 per cent. With power lim-ited at 102 per cent, additional cooling causes a reduction in reactor coolant pressure, and the reactor trips on low reactor coolant pressure. Turbine trip occurs when the reactor trips. Upon turbine trip the unaffected steam line is isolated by the turbine stop valves as shown in Figure 10-1. The unit with the ruptured steam line continues to blev down to the atmosphere. l1 The maximum cooldown of the reactor coolant system would be that resulting from tne blevdown from one steam generator. A typical cooling rate following reac-tor trip for a steam line rupture of 4 sq. in. is shown in Figure 14-21.

The tabulation below lists the approximate time required to blow down the con-tents of the steam generator with a ruptured steam main.

Leak Area, in.2 Blovdown Time, see I

4 860 ^

c 32 002/a

() 128 27 14-17 (Revised 1-15-68) l J

w 1

A steam line failure of large area results in high steam flow with resulting rapid pressure decrease in the reactor coolant system e.nd steam system. The reactor trips on low reactor coolant system pressure or high flux. Reactor trip causes turbine trip and reduction in feedvater flow to decay heat level.

The turbine trip closes the turbine stop valves which isolate the steem lines l1 and prevent blowdown of the steam generator whose secondary side does not have a pipe rup.ture. The steam generators are designed to maintain reactor system integrity upon loss-of-secondary-side pressure. Therefore, this ac-cident vill not lead to a reactor coolant system failure.

Assuming the blowdown from one steam generator results from a secondary steam system rupture, the maximum cooling rate durin6 this accident occurs during the first 10 see after the break. The maximum cooling rate is approximately 3F/secandalowpressureorhighfluxtripoccurs. The net cooldown of the reactor coolant sptem, assuming total blowdown of one steam generator and ac-counting for transfer of core stored heat and decay heat, is less than 50 F.

Th.4.s results in an average coolant temperature of 530 F vhich is about 10 F lower than the normal ::ero power average coolant temperature.

The minimum shutdown margin at 540 F vith the most reactive rod stuck out is 2 9% a k/k. The reduction in reactivity shutdown mergin associated with cool-ing the moderator temperature 10 F below its normal sautdown temperature of 540 F vould be 0 30% a k/k. Using the maximum value for the moderator temper-ature coefficient (-3 0 x 10-4 Ak/k/F), the shutdown margin at 530 F vould be 2.6% ok/k, which is adequate to prevent return to criticality.

In addition, high pressure injection can be actuated during the cooldown period following a large area steam line failure. This system supplies borated water to the reactor coolant system to increase the shutdown margin further. Boron addition to the reactor coolant during the controlled cooling of the system to atmospheric pressure vill prevent criticality at lower temperatures.

The effect of a steam line rupture inside the reactor building has been evalu-ated by conservativel'y assuming an instantaneous release to the reactor build-ing of the energy associated with this accident. The mass and energy releases per steam generator in this analysis are Mass, lb Energy, Btu x 10-6 Steam Generator 46,000 28.0 Feedvater Flow (6 see full flow plus coastdown to 7 5%

flow @ 16 see) 12,800 5.6 Reactor Coolant System Energy Transferred 17.6 Total 58,800 51.2

- a .c c.

} {-

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14-18 (Revised 1-15-68)

-n -

1. "y C) Based upon the above, a single steam generator release vould result in approx-imately 10 psig pressure rise in the reactor building. This is well celow the reactor building design pressure of 55 psig.

The environ = ental consequences from this accident are calculated by assuming that the nuclear unit has been operating with steam generator tube leakage.

The reactor coolant activity assu=es prior operation with 1 per cent failed fuel rods. With these assumptions, the steam generators contain a total of 0.09 equivalent curies of iodine-131. It is further assu=ed that steam gen-erator leakage continues for three hours before the nuclear unit can be cooled down and the leakage terminated. This additional leakage corresponds to 3 4 equivalent curies of iodine-131. The iodine is assu=ed to be released directly to the atmosphere where it mixes in the vake of the reactor building. With these assu=ptions an integrated dose to the thyroid at the exclusion distance of 0 53 rem is obtained. The corresponding dose to the whole body during this sa=e time period is 0.002 rem. The total release of all activity when aver-aged over a year is 33 per cent of the allowable limits of 10 CFR 20.

14.1.2.10 Steam Generator Tube Failures 14.1.2.10.1 Identificatica of Accident In the event of a reactor coolant leak to the secondary system, such as a com-plete severance of a steam generator tube, the activity contained in the cool-ant would be released to the secondary system. Radioactive gases and some of

(~] the radioactive iodine vould be released to the atmosphere through the con-k/ denser air removal system.

14.1.2.10.2 Analysis and Results In analyzing the consequences of this fail re, the following s>quence of events is assumed to occur:

a. A double-ended rupture of one stesm generator tube occurs with unre-stricted discharge from each end.

b. The initial leak rate, approximately 435 spm, exceeds the norri.

makeup of 70 gpm to the reactor coolant system, and system prescure decreases. No operator action is assumed, and a low reactor coo. ant system pressure trip will occur in about 8 min.

c. Following reactor trip, the reactor coolant system pressure continues to decrease until high pressure injection is actuated at a pressure of 1,800 psig. The capacity of the high pressure injection is suf-ficient to compensate for the leakage and maintains both pressure and volume control of the reactor coolant system. Thereafter, r,ne reactor is conservatively assumed to be cooled down and depressurized at the normal rate of 100 F per hour,

s. d. Following reactor trip, the turbine stop valves vill close. Since a reactor coolant to secondary system leak has occurred, steam line

{v pressure vill increase, opening the steam bypass valves to the con-denser. Each bypass valve actuates at a lover pressure than do the safety valves. The reactor coolant that leaks sa a result of the t _

14-19 00 R .. S -

tube failure is condensed in the condenser. Only the fission prod-ucts that escape from the condensate are released to the atmosphere.

g,

e. The affected steam generator can be isolated by the steam line iso-lation valve when the reactor coolant system pressure falls below the setpoint of the secondary system safety valves, i.e., 1,050 psig.

Cooldown continues with the unaffected steam generator until the tem-perature is reduced to 250 F. Thereafter, cooldown to ambient condi-tions is continued using the decay heat removal system.

f. Atthedesigncoolingrateforthepressurizerof100F/hr,depres-surization to 1,050 psig requires approximately 1 7 hr. Durin8 this time period 1.6 x 100 cc (5,650 f t3) of reactor coolant leaks to the secondary system. This leakage corresponds to approximately 45,800 curies of xenon-133 if the reactor has been operating with 1 per cent failed fuel.

The radioactivity released during this accident is discharged through the tur-bine bypass to the condenser and then out the Plant vents. A partition factor of 104 is assumed for iodine in the condenser.(1,2) Noble gases are assumed to be released directly to the Plant vents. The total dose to the whole body from all the xenon and krypton released is only 0.69 rem at the 4,400 ft exclusion distance. The corresponding dose to the thyroid at the same distance is only 1.0 x 10-4 rem. This calculation conservatively assumes that the Plant vent discharge mdxes in the wake of the building structures rather than remaining at its elevated release height.

14.2 STANDBY SAFEUARDS ANALYSIS O

14.2.1 SITUATIONS ANALYZED AND CAUSES In this section accidents are analyzed in which one or more of the protective barriers are not effective and standby safeguards are required. All accidents evaluated are based on the ultimate power level of 2,544 MJt rather than the rated power level of 2,452 K4t. Table 14-3 summarizes the potential accidents studied.

.- -? w , .

~

14-20 aar y

~

c U Table 14-3 Situations Analyzed and Causes Event Cause Effect Fuel Handling Mechanical de= age Integrated dose at exclu-Accidents durira transfer. sion distance is 0.hk rem thyroid and 0.4k rem whole body.

Rod Ejection Failure of control Some clad failure. Thirty-Accident rod drive pressure day dose e exclusion dis-housing. tance is 1.3h rem thyroid.

Loss-of-Coolant Rupture of reactor No clad melting. Thirty-Accident coolant system. day dose at exclusion dis-tance is 9.9 rem thyroid.

Maximum Release of 100% rare Two-hour dose at exclusion Hypothetical gases, 50% iodine, distance is 65 rem thyroid.

Accident and 1% solid fission Thirty-day dose at low pop-products. ulation distance is 3 4 rem thyroid.

(, 14.2.2 ACCIDENT AW1YSES U

14.2.2.1 Fuel Handling Accidents 14.2.2.1.1 Identification of Accident Spent fuel assemblies are handled entirely under water. Before refueling, the reactor coolant and the fuel transfer canal water above the reactor are in-creased in boron concentration so that, with all control rods removed, the keff of a core is no greater than 0 98. In the spent fuel storage pool, the fuel assemblies are stored under water in storage racks having an eversafe geometric array. Under these conditions, a criticality accident during refueling is not considered credible. Mechanical damage to the fuel assemblies during transfer operations is possible but improbable. This type of accident is considered the maximum potential source of activity release during refueling operations.

14.2.2.1.2 Analysis and Results The fuel assembly is conservatively assu=ed to have operated at 29 MWt, twice the power level of an average fuel assembly. The reactor is assumed to have been shut down for 24 hr, which is the minimum time for reactor cooldown, re-actor closure head removal, and removal of the first fuel assembly. It is further assumed that the entire outer row of fuel rods, 56 of 208, suffers l1 damage to the cladding. Since the fuel pellets are cold, only the gap activity is released. The fuel red gap activity is calculated using the escape rate coefficients and calculational methods discussed in 11.1.1 3 The gases released from the fuel asse=bly pass through the spent fuel storage pool water prior to reaching the auxiliary building at=osphere. As a minimum, 14-21 (Revised 1-15-68) -

the gases pass through 10 ft of water. Although there is experimental evidence that a portion of the noble gases will re=ain in the water, no retention of noble gases is assumed. Based on the data in References 3 and 4, 99 per cent g

of the iodine released from the fuel assembly is assumed to re=ain in the water.

The total activity released to the building at=osphere is therefore Iodine 28.4curigs Noble gases 2 79 x 10 curies The auxiliary building is ventilated and discharges through 90 per cent effi-cient charcoal filters to the Plant vents. The discharge fro = the Plant vents is cssu=ed to =ix in the wake of the building structures rather than rc=ain at its elevated release point. This assumption produces less favorable dilution and, therefore, higher ground concentrations at the exclusion distance.

The activity is assumed to be released as a puff fro = the Plant vents. Atmo-spheric dilution is calculated using the 2-hour dispersion factor of 3 x 10-4 developed in 2 3 The total integrated dose to the whole body at the 4,400 ft exclusion distance is 0.k4 rem, and the thyroid dose at the same distance is 0.k4 re=. In evaluating the sensitivity of this analysis, the thyroid dose at the site boundary is directly proportional to the quantity of iodine released.

For example, if only 90 per cent retention of iodine is assumed by the spent fuel storage pool water, the dose at the exclusion distance is increased by a factor of 10. The dose from this increased iodine release is still a factor of 70 below the 10 CFR 100 guidelines.

14.2.2.2 Rod Ejection Accident 14.2.2.2.1 Identification of Accident Reactivity excursions initiated by uncontrolled rod withdrawal (14.1) were shown to be safely teminated without damage to the reactor core or reactor coolant system integrity. In order for reactivity to be added to the core at a = ore rapid rate, physical failure of the control rod drive housing or con-trol rod drive nozzle must occur. Failure in the drive upper pressure housing can cause a pressure differential to act on a control rod assembly and rapidly eject the asse=bly from the core regicn. The power excursion due to the rapid increase in reactivity is limited by the Doppler effect and teminated by re-actor protection syste= trips.

The criterion for reactor protection, should this condition occur, is that the reactor will be operated in such a =anner that a control rod ejection accident vill not further da= age the reactor coolant syste=.

a. Accident Bases The bases for the rod ejection accident are as follows:

Worth of ejected rod 03% A k/k Rod ejection time 0.150 see Ulti= ate power level 2,544 MWt Reactor trip delay 0 3 see O

1h-22 _

l The severity of the rod eject .. accident is dependent upon the worth

' - of the ejected rt.' and the re.' tor power level. The control rod group of greatest worth is the first of ehe entire rod pattern to be with-drawn from the core. The worth of this rod can be as high as 30 per cent of the total patter- rth of 10.0% Ak/k, i.e., 3% Ak/k. However, 1 the 30% Ak/k value ex1 ., only when the reactor is suberitical. The details of control rod worth calculaticns and the methods of selecting the number of control rods in each group are presented in 3.2.2 and 7.2.2.1.2.

When the reactor is suberitical, the boron concentration is maintained at a level whereby the reactor is at least 1 per cent suberitical with the control rod of greatest worth fully withdrawn from the core. There-fore, rod ejection, when the reactor is subcritical and all other rods are in the core, does not cause a nuclear excursion As criticality is approached, the worth of the remaining control rods decreases. At criticality, rod ejection would result in a maximum reactivity addition of 0.56% Ak/k. l1 At rated power, but before equilibrium xencn is established, the total rod pattern worth remaining in the core is 2.6% Ak/k. At equilibrium xenon the pattern worth is 1.6% Ak/k. Before establishing equilibrium xenon, the greatest single control rod worth is 0.k6% Ak/k. A single rod worth of up to 0.7% Ak/k has been used in the analysis of this ac- 1 cident.

In crder for any one rod to have this =uch worth, it would necessarily

! be fully inserted in the core. Assu=ing that a pressure housing fail-ure occurs in such a canner that it no longer offers any restrictica for rod ejection, the time and therefore the rate of reactivity addi-tion can be calculated. Purther assu=in6 that there is no viscous drag force it=iting the rate of ejection, control rod travel time to the top of the active region of the core is calculated to be 0.176 sec. To account for the S-shaped reactivity worth versus position of the rod, an ejection ti=c of 0.15u see (75 per cent of active core height) is used in the analysis.

b, Puel Rod Damage Criteria Power excursionc caused by reactivity disturbances of the order cf

=agnitude occurring in rod ejection accidents could lead to three potential modes of fuel rod failure. First, for very rapid and large transients in which there in insufficient ti=c for heat transfer from fuel to cladding, fuel = citing fallowed by vaporization can generate destructive internal press 2-es without incressing cladding tempera-tures significantly. The second acJe occurs when the internal vapor pressure is not sufficient to cause cladding rupture, but subsequent heat transfer raises the temperatare of the cladding and weakens it until failure occurn. The third m.te occurs when the naciear excur-sion has insufficient energy to cauce significant =elting of the fuel, but subsequent heat transfer to clad from fuel may cause excessive cladding te=peratures. In all three cases there is a possible occur-rence of =etal-vater reactions. However, only very rapid and large transients will genern'e a rapid pressure buildup in the reactor coolant syster. l 1

14-23 (Revised 1-15-68s 0032 g

The energy required to initiate UO2 fuel mei 1,n gm, based on an initial te=perature of 68 F. 2)g Theisheat 220oftofusion 225 cal /

requiresanadditional60 cal /g=. Any further energy addition va- 3 porizes the fuel and produces a buildup of vapor pressure within the V fuel rod. The effect of the vapor pressure is dependent upon the te=perature and ulti= ate strength of the cladding. Energy additions of up to 420 cal /gm have been calculated to be necessary before the bursting pressure of cladding is exceeded. The lover limit for pro-ducing significant fuel vapor pressure (14.7 psi) is 325 cal /gm.(6)

The potential cladding failure is a function not only of the fuel vapor pressure, but also of fission product gas pressure, cladding and fuel irradiation exposure, and zirconium hydriding. A:. s lover limit, the potential for bursting of cladding and release or molten fuel to the reactor coolant is conservatively set at a fuel enthalpy 1

of 280 cal /gm in this evaluation.

For power excursions with energy bursts below 280 cal /gs, zirconium-water r yetions are possible. A correlation of the TREAT experiments present's a raethod'of correlating the potential zirconium-water reac-tion as a function of fission energy input.l7) These data are based on initially cold (room te=perature) fuel rods, but are also corre-lated as a function of peak adiabatic core temperature. This corre-lation can be used either by computing the core te=perature or by adding the initial steady state fuel enthalpy to the nuclear energy burst and obtaining an equivalent final fuel enthalpy. Accordingly, a zirconium-vater reaction requires a cinimum fuel enthalpy of 125 cal /gn. Increasing fuel enthalpies cause a linear increase in the percentage of the reaction, which may be approxi=ated by

%Zr-H2 O Reaction = 0.125 (Final Fuel Enthalpy - 125).

O It is assu=ed that DNB will take place when the clad reaches a heat 2

flux of 6.36 x 103 Btu /hr-ft . At this heat flux the hot fuel rod enthalpy would be approximately 140 cal /gm at EOL and 130 cal /g= at BOL. Applying the peaking factors described in 3 2 3 to the results of these analyses, the per cent of the core having an enthalpy greater than the values above can be calculated. Any fuel rod exceeding the enthalpy values above is assu=ed to fail from overheating and releases the gap activity of that fuel rod.

14.2.2.2.2 Method of Analysis The hypothetical control rod ejection acciden was investigated using the exact 1-dimensional WIG 12 digital co=puter program. O) It was found that the point kinetics analog model results agreed with the WIGL2 results to within 10 per cent for rod worths up to 0 75% A h/k. The point kinetics model assumes an initial flux distribution which is undisturbed by local control rod assemblies.

The space-dependent model, however, has significant flux depressions in the vicinity of control rods. Although the flux throughout the core begins to in-crease shortly after the start of the rod ejection, the flux increase in this depressed region rises more quickly so that by the time the average power has reached a level just a few per cent above the initial power level, the flux l shape has almost no perturbation in the region previously occupied by the ,

ejected rod. The entire reactor flux then rises unifom'y until the Doppler l O l 14-24 (Revised 1-15-68) l

' ylAg" l

0033

effect terminates the excursion. Thus by applying the peak-to-average flux factors of 2 92 for EOL and 3.2h for BOL to the point kinetics results, the peak and integrated flux at any point in the reactor can be accurately as-sessed.

14.2.2.2.3 Analysis and Results

a. Source Power A sensitivity study at source level has been done around a single 1 rod worth of 0 5% ak/k. This analysis was performed with the core 0 5% ak/k suberitical so that a total rod worth of 1% ak/k was withdrawn in 0.150 sec. The reactor power was initially at 10-9 of the ultimate power level. The low pressure trip occurs at 1.7 sec after the ejection starts, and the reactor power is terminated at a peak value of 39 per cent ultimate power. This peak neutron power value is not reached until about 15 see after the rod is ejected because Doppler feedback controls the rate of rise and mag-nitude of the neutron power. Therefore, a low pressure trip will terminate the accident before significant power is generated owing to the loss of coolant through the rupture.

An analysis was performed for the accident above without a low pressure trip to demonstrate the capability of the reactor to ac-cept the accident.

In this case the neutron power reaches 1,000 MWt (39 per cent ul-

^

timate power), and the peak fuel temperature is 990 F. This is far below the melting temperature of UO2, and the resultant ther-s' mal power is only 16 per cent of ultimate power. Hence, no fuel damage would result from the rod ejection accident at source power level.

b. Ultimate Power A sensitivity study of ultimate power level has been done around 1 an assumed single rod worth of 0.3% ak/k. The analysis includes rod worths from 0.1 to 0.7% ak/k, however. For the ultimate power case at beginning-of-life (BOL), the ejection of a single control rod worth 0 3% ak/k would result in virtually no Zr-H 2O reaction and approximately 1% of the core experiencing DNB (see Figures lk-22 and 14-23). The hot fuel rod would reach a peak enthalpy of about 166 cal /gm.

For the end-of-life case (EOL), the reactor neutron power peeks at 6,190 MWt, 200 milliseconds after the start of ejection of a 0.3% ak/k control rod. The prompt negative Doppler effect termi-nates the power rice, and control rod insertion from high flux signal terminates the excursion. The total neutron energy burst during the transient is approximately 3,200 MW-sec. The final en-thalpy of the nominal rod is 113 cal /gm, i.e. , the enthalpy of the hot rod is 163 cal /gm. This enthalpy is considerably below the minimum range (220 to 225 cal /sm) for central fuel melting. As a result of the excursion, approximately 13 5 per cent of the core would have DNB (see Figure 14-22).

The power distribution at the beginning of core life, with the ,

higher power peaking factors shown in 3.2 3, was used to determine '

the distribution of the energy of the excursion. With this 14-25 (Revised 1-15-68) ._

, j

zirconium cladding may react (see Figure IL-23) to contribute an addi- g tional 677 !G-see of energy. The resultant te=perature increase is w spread over a relatively long period of time. Consequently, the metal-water reaction energy is liberated over a long period of time, and no damaging pressure pulses are produced in the syste=.

As a result of the postulated pressure housing failure, which produces a rupture size of 0.0k sq ft, reactor coolant is lost from the system.

The rate of mass and energy input to the reactor building is consider-ably lower than that for the 3 sq ft rupture discussed in 14.2.2.3.

This lower rate of energy input results in a lover reactor building pressure than that obtained for the 3 sq ft rupture.

The environmental consequences from this accident are calculated by conservatively assuming that all fuel rods that undergo a DNB result in clad failure and subsequent release of the gap activity. Actually, most of the fuel rods will recover from the DNB, and no fission prod-uct release vill occur. For the case of a 0.3% Ak/k rod ejection from ultimate power at the end of life,13.5 per cent of the fuel rods are assumed to fail, releasing 177,000 equivalent curies of I-131 to the reactor building. Fission product activities for this accident are calculated using the methods discussed in 11.1.1.3. Using the environmental models and dose rate calculations discussed under the loss-of-coolant accident, the total integrated dose to the thyroid at the exclusion distance from this accident is only 2.68 rem in 30 days, which is more than a factor of 100 below the guideline values of 10 CFR 100.

c. Sensitivity Analysis The results of a sensitivity analysis performed on the control rod ejection accident are shown in Figures 1k-2h through 1L-32. Figure 14-2k shows the variation in the peak neutron power as a function of the worth of the ejected control rod. For the nominal 0.3% A k/k case from ultimate power, the peak neutron power is less than 300 per cent, again assuming that a low pressure trip does not occur. The rod ejection from source level results in a Doppler turn-around before the flux trip is reached. Figure IL-25 shows the variation in the corresponding thermal power with control rod worth.

Figure 1k-26 shows the corresponding enthalpy increase of ihe hot fuel rod versus control rod verth. Note the very small spread in values for the BOL and EOL ultimate power conditions. As expected, the enthalpy increases with rod worth.

Figures ik-27 through 1k-30 show the peak reactor neutron and thermal powers as a function of changes in the positive moderator temperature coefficient and negative Doppler coefficient for the nominal 0.5%

A k/k control rod ejection from source level. There was insignifi-cant variation of the peak neutron and thermal power with changes in the two reactivity feedback coefficients.

Figure 1k-31 shows the change in nominal thermal power with varia-i tions in the trip delay time fox the nominal 0.3% A k/k rod ejection h m vss

O from ultimate power (the variation from zero power is negligible).

The trip delay time does not affect the peak neutron power because the Doppler effect controls the power transient. Figure 14-32 shows the corresponding change in the total enthalpy increase of the hot fuel rod versus the trip delay.

The themal power never exceeds 114 per cent ultimate power for any of the variations studied using the no=inal rods (0 3% A k/k for ultimate power and 0 3% Ak/k for source level). The hot fuel rod average temperature never increases by more than 310 F above the ultimate power peak value (4,090 F). It is therefore concluded that each of these larameter variations has relatively little effect on the nominal results.

O O

M 14-27 . _,

00bb

lb.2.2 3 Loss-of-Coolant Accident 14.2.2 3 1 Identification of Accident O '

Failure of the reactor coolant system would allow partial or co=plete release of reactor coolant into the reactor building, thereby interrupting the normal mechanism for removing heat from the reactor core. If all the coolant were not released immediately, the remaining amount would be boiled off owing to residual heat, fission product decay heat, and possible heat from che=ical re-actions unless an alternate means of cooling vere available. In order to pre-vent significant chemical reactions and destructive core heatup, emergency core cocling equipment rapidly recovers the core and provides makeup for decay heat removal.

14.2.2 3 2 Accident Bases All components of the reactor coolant system have been designed and fabricated to insure high integrity and thereby minimize the possibility of their rupture.

The reactor coolant system, the safety factors used in its design, and the special provisions taken in its fabrication to insure quality are described in Section 4.

In addition to the high-integrity system to minimize the possibility of a loss of coolant, emergency core cooling is provided to insure that the core does not celt even if the reactor coolant system should fail and release the coolant.

This emergency core cooling is provided by the core flooding system, the makeup and purification system (high pressure injection), and the decay heat removal g system (low pressure injection). These systems are described in detail in See- W tion 6, and their characteristics are su=marized in the following paragraphs.

The performance criterion for the emergency core cooling equipment is to limit l1 the temperature transient below the clad melting pcint so that fuel geometry is maintained to provide core cooling capability. This equipment has been con-servatively sized to limit the clad temperature transient to 2,300 F or less as [l temperatures in excess of this value promote a faster zirconium-water reaction rate, and the termination of the transient near the melting point would be dif-ficult to demonstrate.

The fuel rods may experience cladding failure during the heatup in the loss-of-coolant accident. This could be due to fission gas internal pressure and weak-ening of the clad due to the increase in clad te=perature. The mechanical stren6th of the Zircaloy cladding is reduced as the te=perature exceeds 1,000 F such that the highly irradiated fuel rods, with high fission gas internal pres-sure, may fail locally and relieve the gas pressure when the te=perature ex-ceeds 1,200 F. Some local ballooning of rods is likely to occur. However, cooling would still be effective since the fuel rods are sub=erged, and cross-channel flow around the ballooned area vill cool the rod. At vorst a local hot spot may occur.

It is calculated that a s=all number of fuel rods operating at peak power vill experience a cladding te=perature transient to 1,950 F in about 18 sec. The injection of emergency coolant, at a time when the cladding is at a temperature of about 1,950 F, may also cause distortion or bowing between supports. As a result'sh5e'er'the fuel rods may crack and allow relief of internal pressure.

lk-28 (Revised 1-15-68) 0037

However, the cladding is expected to remain sufficiently intact to retain the solid fuel material and to prevent gross fuel shifting. The transient would be limited to regions of the core which operate at peak power. The major portion of the core vill not experience as severe a transient.

Heating of the fuel can and the fuel rod spacer grid requires heat flow from the clad to the structure by conduction and radiation; therefore, the structure te=peratures vill lag the cladding temperature transient. As the fuel rod tem-perature rises, the fuel rods are expected to experience some bowing between supports due to the temperature differential existing between the fuel rod and the can. The cans and spacer grids are made from stainless steel and have sub-stantial mechanical strength, even at the maximum expected temperatures. The supporting stainless steel structure vill therefore retain sufficient strength to assure spacing between fuel rods to allow emergency coolant to reach them, j and will keep the fuel rods in the same location in the core to prevent gross fuel shifting.

The core flooding system has two independent core flooding tanka, each of which is connected to a different reactor vessel injection nozzle by a line containing two check valves and a normally open, remotely operated isolation valve. Since these tanks and associated piping are missile-protected and are connected di-rectly to the reactor vessel,st rupture of reactor coolant system piping vill not affect their performance. These tanks provide for automatic flooding when the reactor coolant system pressure decreases belov 600 psi. The flooding water is injected into the reactor vessel and directed to the bottom of the reactor vessel by the thermal shield. The core is flooded from the bottom upward. The combined contents of the two tanks (1,880 ft3 of borated water) rapidly reflood the core immediately after the blowdown to provide cooling until coolant flow can be established by low pressure injection.

High pressure injection, actuated by low reactor coolant system pressure, sup-plies coolant at pressures up to the design pressure of the reactor coolant sys-tem and at a rate up to 1,000 gpm. Low pressure injection actuated by low reac-4 tor coolant system pressure supplies coolant at pressures below 100 psig and at l2 a rate up to 6,000 gpm. Both of these systems can operate at full capacity from l2 the on-site emergency electrical power supply and can be in operation within 25 see after the accident. In the reactor vessel, decay heat is transferred to the injection water.

Injection water is supplied from the borated water storage tank. When this tank empties, water is circulated from the reactor building sump through heat exchangers and returned to the reactor vessel.

Engineered safeguards are also provided to cool the reactor building environ-ment following a loss-of-coolant accident and thereby limit and reduce pressure in the building. Reactor building sprays, actuated on a high building pressure signal of 10 psig, deliver 3,000 gpm to the reactor building atmosphere. This i spray water reaches thermal equilibrium within the building atmosphere during l

its passage from the nozzles to the sump. Spray water is supplied from the borated water storage tank until it is emptied. Thereafter, water collected in the sump is ~ricirculated to the sprays. Cooling is also provided by the re-actor building emergency cooling system in which recirculating fans direct the steam-and-air mixture through emergency co6 ers, where steam is condensed.

Heat absorbed in the emergency coolers is rejected to the nuclear services O

1h-29 (Revised 2-7-68) 0038

cooling water system. The heat removal capacity of either of these two reactor building cooling systems is adequate to preven,t overpressurization of the build-ing during a loss-of-coolant accident. ~

This analysis demonstrates that in the unlikely event of a failure of the reac-tor coolant system, both of the other two boundaries that prevent fission pro-duct release to the atmosphere, i.e. , the reactor core and the reactor building, are protected from failure. Accordingly, the public would be protected against potential radiation hazards.

In order to evaluate this accident, a range of rupture sizes from small leaks up to the complete severance of a 36 in. ID reactor coolant system line has been evaluated. A core cooling analysis is presented for the complete severance of the 36 in. ID reactor coolant piping.,

Since the large rupture removes the least amount of stored energy from the core, this represents the minimum temperature margin to core damage and, therefore, places the most stringent requirements on the core flooding system.

The reactor building pressures have been evaluated for a complete spectrum of rupture sizes without the action of core flooding tanks and, therefore, are conservative. The peak pressure occurs for a 3 ft2 rupture rather than for a 36 in. ID (lb.1 ft2) rupture. Rupture sizes smaller than the 36 in. ID leak result in lenger blevdown times, and the amount of energy transferred to the reactor building atmosphere is increased. As a result the intermediate leak size results in a reactor building presstre greater than that produced by the 36 in. ID rupture.

lb.2.2 3.' Accident S;malation h

a. Hydraulic Model Blovdown of the reactor coolant system following an assumed ru has been simulated by usin6 a , modified version of the FLASH (9)pture

~

code.

This code calculates transient flows, coolant mass and energy inven-tories, pressures, and temperatures during a loss-of-coolant accident.

The code calculates inflow from the emergency cooling and calculates heat transferred from the core to the coolant.

Modifications were made to FLASH to make the model more applicable to this system. The changes are as follows:

(1) The calculation of recetor coolant pump cavitation was based on the vapor preisure of the cold leg instead of the hot leg water.

(2) Core flooding taaks have been added. Water flow from the core flooding tanks is calculated on the basis of the pressure dif-ference between the core flooding tanks and the point of dis-charge into the reactor coolant system. The line resistance and the inertial effects of the water in the pipe are included. The pressures in the tanks are calculated by assuming an adiabatic expansion of the gas above the water level in the tank. Pres-sure, flow rate, and mass inventories are calculated and printed out in the computer output.

O lk-30

. 0039

o

(_) . (3) Additions to the water physical property tables (mainly in the subcooled region) have also bec.n made to improve the accuracy of the calculations.

(h) A change in the steam bubble rire velocity has been made from the constant value in FLASH to a variable velocity as a function of pressure. The bubble velocity term determines the amount of water remaining in the system after depressurization is complete.

For large ruptures, this change in velocity shows no appreciable change in water remaining from that predicted by the constant value in the FLASH code. For smaller ruptures, an appreciable difference exists. The variable bubble velocity is based on data in Reference 10 and adjusted to correspond to data from the LOFT seniscale blowdown tests.

Test No. SL6 from the LOFT semiscale blowdown tests is a typical case for the blowdown through a small rupture area. A comparison of the predicted and experimentally observed pressures is shown in Figure 1k-33. Figure lh-3h shows the per cent mass remaining in the trnk versus time as predicted by the code. At the end of blowdown, the predicted mass remaining is 13 per cent. The mea-sured mass remaining is approximately 22 per ;ent.

The FLASH code describes the reactor coolant system by the use of two

' volumes plus the pressurizer. The system was grouped into two volumes on the basis of the temperature distribution in the system as follows:

O Volume 1 includes half of the core water volume, the reactor out-let plenum, the reactor outlet piping, and approximately 55 per cent of the steam generators.

Volume 2 includes half of the core water volume, the reactor in-let plenum and downcomer section, the reactor inlet piping, pumps, and h5 per cent of the steam generators.

r Volume 3 represents the pressurizer.

The resistances to flow were calculated by breaking the reactor cool-ant system into 2h regions and calculating the volume-veighted resis- l tance to flow for a given rupture location based on normal flow resis-tances. For the double-ended ruptures, all of the leak was assumed to occur in the volume in which that pipe appeared.

The reactor core power was input as a function of time as determined by the CHIC-KIN code in conjunction with the FLASH output. Steam generator heat removal was assumed to cease when the rupture occurred.

While the modified FLASH code now has the capability of simulating

-injection flow from the core flooding tanks, the calculations shown in this report were made prior to the time that the core flooding

~

simulation was added to FLASH. Core flooding was calculated by '

/ taking the reactor vessel pressure .as predicted by FLASH without

> 1 O'. .

core flooding and using this: pressure as input to a separate program .

to get the . flow from the core flooding tanks. Reactor. vessel filling 0040 l 14-31

_U .

w-

,.y.,._ . . , . . .--- -

i

vas calculated by adding the mass remaining in the vessel as predicted g by FLASH to the mass injected from the core flooding tanks. This ce- w thod of calculation is conservative in that condensation of steam by the cold injection water is not taken into account. A more recent analysis using the FLASH code confirms that conservatism used in this analysis.

Pressure, temperature, mass and energy inventories, and hydraulic characteristics as determined by FLASH are input into the core thermal code (SLUMP) and the reactor building pressure buildup code (CONTEMPT).

b. Core Thermal Model The core heat generation and heat transfer to the fluid are dependent upon the blevdown process. The FLASH program includes a core thermal model and the feedbacks of heat transfer and flow on each other.

While the FLASH thermal model is acceptable for determining the ef-feet of core heat transfer on the blowdown process, a more extensive simulation is necessary for evaluation of the core te=perature tran-sient.

Additional analytical models and digital computer program (SLUMP) were developed to simulate the core thermal transient for the period beginning with the initiation of the leak and ending after the core temperature excursion had terminated.

The model includes the effects of heat generation from neutrons be- g fore reactor trip, neutron decay heat, and fission and activation w product decay heat; the exothermic zirconium-water reaction based on the parabolic rate law; heat transfer within the fuel rods, limited heat convection from the fuel clad surface to any fluid within the core region, heat transfer from reactor vessel valls and internals to the coolant, and heat transfer from fuel rods to the steam neces-sary to sustain a metal-water reaction; and emergency injection flov and boiloff.

The basic model structure provides 50 equal-volume core regions with input provisions to allow any choice of power distribution. The model may be used to simulate the entire core or any subdivision of the core. Therefore, the core geometry may be detailed to the de-gree consistent with the results desired.

The following parabolic law for the zirconium-vater reaction equation (11) with the following constants is simulated for each of the re-gions:

~

" (r - r) **P ~

O 0041 14-32

f~

V) where r = radius of unreacted metal in fuel rod r = original radius of fuel rod t = time K = rate law ccnstant (0.3937 cm27,,,)

AE = activation energy (45,500 cal / mole)

R = gas constant (1.987 cal / mole K)

T = temperature, K The zirconium-water reaction heat is assumed to be generated complete-ly within the clad node. The heat necessary to increase the steam temperature from the bulk temperature to the reaction temperature is transferred from the clad at the point of reaction. The above equa-tion implies no steam limiting. However, the program Joes have pro-vision for steam rate-limiting to any degree desired, but no steam-limiting of the reactions has been assumed. All heat from neutron, beta, and gamma sources is assumed to be generated within the fuel according to the preaccident power distribution and infinite irradia-tion.

Within each of the regions there is e single fuel node and a single clad node with simulation of thermal resistance according to the normal fuel rod geometry. Provision is made to simulate four dif-ferent modes of heat transfer from the cind node to the fluid sink node by specifying'the time-dependent surface coefficient.

The surface heat transfer coefficient input data are determined from calculations which are based on flow and water inventory as furnished from the blowdown and the core flooding tank performan's analysis.

In the event that insufficient cooling is provided, the program will allow clad heating to progress to the melting point. At this point the latent heat of zirconium must be added before the clad melts.

Provisions are also incorporated to allow the clad to be heated to temperatures above the melting point before slump occurs.

~As each region slumps it may be assumed to surrender heat to a water pool or to some'available metal heat sink. If water is available an additional 10 per cent reaction is assumed to_ occur.

The program output includes the following (as a function of time un-less otherwise specified):

> Average fuel temperature of each region.

f-w Average clad temperature of each -region.

d- Per cent metal-water reaction in each region.

0042 ih-33'

-m -____ _ -._.;__.. ._ _ _. _ -

Time for the clad of each region to reach the metal-water thresh-old, the beginning and end of melting, and the slump temperature.

gg Heat transferred to the reactor building from the core.

Heat generation by hydrogen and oxygen reco=bination.

Total zirconium-water reaction.

Total heat stored in metal sinks.

c. Reactor Building Pressure Model The reactor building pressure-temperature analysis is performed using the digital computer code " CONTEMPT" developed by Phillips Petroleum Company in conjunction with the LOFT project. This program and its capabilities are described in Reference 12. With minor modifications this program was adapted for use on the B&W Philco-2000 computer.

In this model, the reactor building is divided into two regions: the atmosphere (vater vapor and air mixture) and the su=p region (liquid water). Each region is considered to be well mixed and in thermal equilibrium, but the temperature of each region may be different.

The reactor building and its internal structures are subdivided into five segmente, as shown in Table 1b-h, and treated as slabs with 1-dimensional heat transfer. Each segment, divided into several heat conducting subregions, may act as a heat source or sink. The pro-gram includes the capability of cooling the reactor building atmo-sphere by air coolers (reactor building emergency cooling units) and

= pray coolers (reactor building spray system), and cooling the liquid region by su=p coolers (decay heat removal coolers).

During blowdown, mass and energy are added directly to the atmosphere where the liquid water present is assumed to fall to the liquid re-gion. After blevdown is over and emergency injection has been ini-tiated, mass and energy are also added directly to the vapor region as steam. When the water level in the reactor vessel reaches the nozzle height, all mass and energy are added directly to the liquid region since no boiling of the injection water occurs after the core has been covered. When the supply of injection water is depleted, recirculation and cooling of sump water is simulated.

The reactor building calculations are begun by computing steady-atate results using initial atmospheric conditions as the input. Following the rupture, the mass and energy addition is determined from the energy input rates for each time step. Heat losses or gains due to the heat-conducting slabs are calculated. Then the pressure and temperature of the liquid and vapor regions are calculated from the mass, volume, and energy balance equations.

O 1h-34 DM3

()

Table 14 h Reactor Building Structural Heat Capacitance Segments Segment Description 1 Reactor Building Walls and Dcme 2 Refueling Cavity (Type 30h SS Liner - One Side) 3 Reactor Building Floor k Internal Concrete 5 Internal Steel The model has been developed so that the effectiveness of the natural heat sinks and the engineered safeguards can be clearly demonstrated.

The model can readily produce the reactor building pressure history for any assumed combination of operable safeguards. Therefore, the effectiveness of any given arrangement can be analyzed.

Ih.2.2.3.h Accident Analysis

a. Core Flooding Tank Design Base Accident The 36 in. ID, double-ended pipe rupture produces the fastest blow-

, down and lovest heat removal from the fuel. This case therefore represents the most stringent emergency core cooling requirements.

Results from the mddified version of FLASH indicate that the core f^/)

A- flooding tank simulation provides for the retention of all injection plus a portion of the original reactor coolant that would otherwise have been released. Thus, the cool injection water provides a cool-ing and condensing effect which reduces overall leakage. For the present analysis, no credit has been taken for the extra accumula-tion of water due to the condensing effect.

(P"ETED) l1 The SLUMP digital computer program, as described in 14. 2. 2. 3. 3.b above, is used to evaluate cere flooding tank performance in terms of core cooling capability. In the analysis, the hottest 5 per cent of the core was simulated in segments of 1/h of one per cent each.

The hottest segment was assigned a peaking factor of 3.1 times the average of the total core power density.

A detailed analysis of the void shutdown and core response was made 1 with the digital computer program CHIC-KIN. This program accounts for changes in flow, pressure, enthalpy, and void fraction. It also computes axially veighted Doppler and moderator coefficients of re-activity for the kinetics calculation. The Doopler coefficient is input as a nonlinear function of fuel temperat tre, and the moderator void coefficient is input as a function of . d fraction. The param-eters describing the coolant were obtained from the digital computer

(] ,

program FLASH, which in turn used the neutron power output from CHIC-V

~

KIN. The core is assumed to be initially at the ultimate power level of 2,5hh MWt.

0044 1h-35 (Revised 1-15-68) -- -

(DELE 2D) l1 h Figures 1k-3ha and 1h .bb show the results of the hot leg, lb.1-ft2 1 rupture simulation without trip action. Figpre lh-3ha is the neutron p wer trace, and Figure 1L-3kb shows the various comoonents of the re-activity feedback.

The transient core flow frcm the FLASH analysis of the 36 in. ID, double-ended rupture was used to determine the core cooling mechanism used in SLUMF. The very high flow Iates during the initial blevdown period provide nucleate boiling conditions. However, the time for Departure from Nucleate Eciling (DNB), especially for the hot regicns, is extremely difficult to determine. Therefore, a conservative ap-proach was adopted by assuming DNS at 0.25 sec. Nucleate boiling surface coefficients at high flow rates may exceed 50,000 Etu/hr-ft 2 -F. A nucleate boilin; surface coefficient of 25,000 Etu/hr-ft 2-F was used in the analysis. .However, the series heat transfer from the clad node to the fluid sink is limited to 6,500 Etu/hr-ft -F2 by the relatively low condu:tance of the clad.

After DNB the surface r' eat transfer was calculated using the flow provided by FLASH results and Quinn's =cdified versicn of the Sieder-Tate (13) correlation:

1+1-X ID I 1E 0.8 bB 0.lk h h_F Ar = 0.023 Dk (I.,ge)0.8(L,pr)1/3 X i

h (pF( ( "k' where h = tv -phase film heat trantfer coefficient, TPF Etu/hr-ft 2_7 k = fluid conductivity, Etu/hr-ft -F D. = hydraulic diameter, ft Up = Reynolds number Np = Prandt1 number x = quality o = density p = viscosity subscript B = " Bulk" subscript F = " Film"  ;

subscript W = " Wall"

... 0045 i

lk-36 (Nevised 1-15-68) l l

l l

_ - .m -

With.this correlation, bulk steam properties are used in the basic f-form, and the last two bracketed terms are modifiers which correct for quality and different conditions at the vall.

Figure 1h-35'shows the core flow vs time for the 14.1 ft2 leak as calculated by FLASH.

Figure ik-36 shows the clad surface heat transfer coefficient versus time based on the flow of Figure lh-35 and the modified Sieder-Tate

equation. The straight line in Figure lk-36 indicates the surface heat transfer values which were used in SLU'G, and which are conser-vative as compared to the results obtained from the Sieder-Tate equa-tion.

In applying the Sieder-Tate equation constant values of bulk steam quality and temperature corresponding to the most conservative as-sumptions were used.

A sensitivity analysis was made for maximum coefficients in SLU G ranging from h00 to 2,000 Etu/hr-ft2 -F initially and decreasing to zero at the end of blowdown. Results are discussed below.

After blowdown no core cooling is assumed until after core recover-ing. starts. khen the water level reaches the core bottom and starts to rise up on the core, the submerged pcrtion vill be cooled by pool boiling, and any steam thur produced vill provide some cooling for that portion of the core above the water line. However, in deter-mining peak clad temperatures no cooling is assumed for that portion of the core which is above the water line.

1

. e the point of initial contact of cool water against hot cladding the heat flux and temperature differences vill be such that film boil'ing is the probable mode of heat transfer. This mode provides the lowest surface coefficients which would be in the range of 100 to 300 Btu /hr-ft 2 -F. However, in evaluating the core flooding tank design a conservative approach was used by assuming a value of 20 Btu /hr-ft 2 -F. This value is adequate for terminating the temperature excursion in the clad.

The core flooding tank analysis incorporated the study of performance sensitivity to three significant core flooding tank parameters: (a) gas pressure (h00 to 1,000 psig), (b) ratio of nitrogen gas volume to total volume (1/3 and 1/2). (c) and size of piping between the I

core flooding tanks and the reactor vessel (12 in. and lh in. ID).

Figure 1h-37 shows the reactor vessel vater level versus time for core flooding tanks operating at 600 psig with different combinations of volume ratio and line size. This figure includes an allowance for boiloff and also shows the effect of the flow provided by high pressure and Icv pressure injection beginning at 25 see when emer-gency power is available. Similar curves for h00 psig and 1,000 4

psig core flooding tanks are shown in Figure lh-38. Figure lh-39 shows_the maximum clad temperature reached by the hot spot and by.

~

the 1, 2, 3, h and 5 percentiles of the core as a function of quench

! Itime, o

d lu-3r og46

Tha qu:nch tima for o giv;n p;rc;ntil2 is tak;n as th t ti=2 wh;n the water level reaches the highest point in the core at which the peaking factor corresponding to that percentile exists. The fact that the same peaking factor may exist at sc=e lower point in the .

core provides an inherent conservatis in the data as plotted. The axial peaking factor profile for the beginning of core life was used as it inposes the most stringent requirements en the core flooding tank design.

Peak clad temperatures for the core flooding systems described above 2 are also shown on Figure 1k-39 These curves demonstrate that all of the systems presented are capable of keeping the peak te=perature at the hot spot more than 1,000 F below the melting temperature of the clad. The amount of zirconium-vater reaction which occurs for each of these core flooding syste=s is shown in Table 1k-5. While this preliminary analysis indicates sc=e difference in the performance of the systems, it is not considered to be a significant difference since the analysis was performed vitncut considering the effects of condensation by the core flooding coolant or of possible bypass to the leak of part of the coolant.

The preliminary core f1 coding tank design selected is for a 600 psi charge pressure, 9k0 ft3 of water, L70 ft3 cf nitrogen, and a 1k in.

supply line. The performance of this system in limiting core tempera-tures is approximately in the center of the range for the syste=s described. The parameters selected for the final syste design vill be based on the results of core =elting analyses to be ccnducted as part of the final design of the reactor. For this 600 psi charge pressure, Figure 1k-39 indicates that the hot spot clad temperature would reach 1.950 F at 17.5 see and that less than 5 per cent of the core vould exceed 1,690 F. For this same case calculations indicate less than 0.005 per cent total circonium-water reaction for the whole core.

Table IL-5 Core Flooding Tank Performance Data Line Nitrogen Total Metal Size, Volume, Water Reaction.

Pressure in.  % of Total  %

LOO lb 33 .022 LOO 1L 50 .009 600 14 33 .005 600 1h 50 .002 600 12 33 .022 600 12 50 .010 1,000 12 33 .003 1,000 12 50 0 Additional analysis was performed to evaluate the sensitivity of the maximum clad te=perature to three important thermal parameters. All cases discussed below have in ec==cn the following parameters:

lk-38 (Eevised 2-7-68) 0047

k

\ Leak size: 14.1 ft2 Time of DNB: 0.25 see Time at ultimate pover: 2 see Time that blevdown cooling ends: 9 5 see Core region: Hot spot Time to initiate quenching: 18 see Dependent variable examined: Clad temperature for hottest 5 per cent of core.

Figure 14-h0 shows the clad maximum temperature sensitivity to the initial surface heat transfer coefficient after the 0.25 see nucleate boiling period. The coefficient is linearly de-creased to zero at 9 5 sec. Zero cooling is maintained until quenching is initiated with a clad surface coefficient of 20 2

Btu /hr-ft-F. Previous discussion indicated Justification for assuming 1,000 Btu /hr-ft -F for the clad surface at 0.25 sec.

2 Figure lh h0 shows that a value of 1,000 is not on the most sensitive part of the curve and a 20 per cent decrease in h would only result in increasing the peak clad te=perature 120 F.

The assumption that DNB occurs at 0.25 see is quite conservative. 1 The duration of the nucleate boiling period has been evaluated to show the sensitivity of the maximum fuel temperature to this param-eter. Figure lk-hk shows the effect of variation of time to reach a DNB. It should be noted.that if DNB occurred at the time of rupture, the peak. temperature would only increase about 30 F above 1,950 F.

Figure 14-kl shows hot spot clad temperature transients for a range of injection cooling coefficients. All cases have a clad surface coefficient of 1,000 Btu /hr-f t2-F at 0.25 sec, decreas-ing to zero at 9 5 sec. Heat removal is then zero until the effect of injection cooling is simulated. Figure 14 h1 shows that without any cooling the temperature reaches the melting point in approximately 50 sec.

The analysis of core cooling has been based upon 2.1 full-power sec- 1 onds resulting from a void shutdown using the maximum positive mod-erator temperature coefficient of +1.0 x 10-h (Ak/k)/F. The effect of variation of the integrated power on hot spot clad temperature is shown in Figure ik h6. The resultant integrated power before a void shutdown occurs could increase to 3.h full-power seconds before the p hot spot clad temperature would reach 2,300 F, the temperature a j'j

() which 1.0 per cent Zr-veter reaction occurs. V lh-39 (Revised 1-15-68)

An h value of 15 stcpa the fast te=perature excursson and al-lovs only a low rate of increase thereafter. Since the contin-uously increasirg cepth of coverage provided by the flooding 'g tanks and the p up:d flow injection syste=s provide additional coolirg capability ith ti=e, an initial cooling v:uue as lov as 15 is probably aaequate.

An h value of 20 provides i==ediate quenchinc, cetion and a slow coolin6 rate the: .after An h value of 100 provides very fast cooling. Even though the 100 is a realistic value fer film boilin6 in a pool - the prob-able = ode for the sub=erged portion of the core - a more con-servative value of 20 has been used as the reference for eval-unting core floodirs tank perfor=ance.

Figure IL-42 shows hot spot clad te=perature transients for a range of poo) fluid sink temperatures. Parc=eters for heat O

O ll- 39 m. (Revised 1-15-68) 0049

transfer prior to 18 see are the sa=e as discussed in the preceding paragraph. At18seeasurfacecoefficientof20 Btu /hr-ft-Fvas 2 applied with sink temperatt.res as indicated. All results reported h

herein previously have had a sink te=perature of 280 F during the quenching period. Prior to quenching the sink te=perature in all cases is based on the transient fluid pressure which results from the FLASH analysis. Figure 14-42 shows that any sink te=perature below approximately 500 F is adequate for holding or reducin6 the chd tem-perature which existed at 18 sec. The core flooding tanks vill pro-vide a h'.gh linv of cool water. Although some heating vill occur from contact with hot metal before the injection water reaches the core, the temperature rise could not be over 50 F assuming that the water came in centact vii.h all reactor coolant system metal below the nozzle level before it contacted the core. Using a reference value of 280 F provides an added conservatism to the analysis.

In conclusion, the analysis has shown tnat the preliminary design of the core flooding system vill provide for covering approxi=ately 80 per cent of tne core at 25 see after the double-ended rupture of the 36 in. ID pipe first occurs. Beyond this time high pressure and low pressure injection vill provide a continuous increase in the water level.

The clad hot spot temperature excursion is ter=inated at 1,950 F and less than 5 per cent of the total cladding exceeds 1,690 F. Only a minute amount (0.005 per cent) of zirconium-water reaction occurs, and the maximum temperature is at least 1,400 F below the clad melt-ing point.

The temperature transient in the core can produce significantly higher than normal temperatures in components other than fuel rods. There-fore a possibility of eutectic fomation between dissimilar core ma-terials exists. Considering the general area of eutectic formation in the entire core and reactor veseel internals, the following dis-similar metals are present, with rajor elements being in the approxi-mate proportions shown.

Type 30h Stainless Steel 19 per cent Chromium 10 per cent Nickel Balance Iron C6ntrol Rod 80 per cent Silver 15 per cent Indium 5 per cent Cadmium Zircaloy-h 98 per cent zirconium 1-3/L per cent Tin m2 0050 O

1h-40

i 1

All these elements have relatively high melting points, i.e., greater v than 2,700 F, except those for silver, cadmium, and indium which, in the case of indium, is as low as approximat ly 300 F.

The binary phase diagram indicates that zirconium in the proportion of 75 to 80 per cent has a eutectic point with either iron, nickel, or chromium at the temperatures of approximately 1,710, 1,760, and 2,370 F, respectively. If these dissimilar metals are in contact and if these eutectic points are reached, the materials could theoreti-cally melt even though the temperature is below the melting point of either material taken singularly.

One point of such dissimilar metal contact is between Zircaloy-clad fuel rods and stainless steel spacers. The analysis of the perfor-mance of the core flooding tanks during a loss-of-coolant accident indicated that only 4 per cent of the cladding would ever exceed the zirconium-iron eutectic point. Since the ~ C ers arc located at 21 in. intervals along the assembly and each has a very small con-tact area, only a fraction of the 4 per cen. ,ould be in contact with stainless steel. The approximate time period that the 4 per cent of the cladding is above the eutectic point is 30 sec. Because of the relatively small area of contact, the condition could not progress very far and fuel geometry would be maintained. Unless the proper ratio of metals is available, the melting point is higher than the eutectic point.

Another area of dissimilar metal c7ntact is that of a zirconium guide tube with the stainless steel cladding of the control rod. Fnllowing blowdown, heat can be generated in the control rods by absorption of gamma rays. Beta ray decay heat vill be deposited in the fuel rods where generated. Since ga==a decay heat is only about one-half the total decay heat, and the control rod is shielded from the fuel by a guide tube, heat generation rates in control rods will be less than l one-half the rates in the fuel. As a result, the peak heat generation rate in control rods adjacent to hot spot fuel vould not exceed an estimated one-half times the rate in these fuel rods which have a 3 1 power ratio. The contribution frcm radiant heat transfer from higher

]

powered fuel rods vould be relatively small. The analysis of core melting shows that, with core flooding tanks, fuel rods with a 15 power ratio vill not exceed 1,500 F. This is well below the eutectic melting point.

The reactor core will re=ain suberitical after flooding without con-trol rods in the core because the injection water contains sufficient boron U,270 ppm) to hold the reactor suberitical et reduced tempera-tures. Tne meat stringent boron requirement for shutdown without any control rods is at the beginning of core life when the reactor is in a cold, clean condition and 1,820 ppm boron are required to maintain l1 keft of 0 99 (see Table 3-6, soluble Boron Levels and Worth.) The concentration existing in the reactor building sump after a loss-of-coolant accident from operating power at the beginning of core life i is.2,17h ppm boron. This concentration represents a boron margin of 1 l O 35' >>= * **e ="cr1*1c 11*F ae 81 = v 1"e =er8 1 = r 1 > r c =*- 1 lk kl (Revised 1-15-68) 0051

b. Core Cooling Analysis for Spectrum of Leak Sizes 1 An analysis of the loss-of-coolant accident has been made for a spec-trum of leak sizes and locations. This information has been analyzed g j and is reported according to the following grouping: (1) hot leg ruptur,es, (2) cold leg ruptures (3) injection line failures, and (h) injection system capability.

(1) Hot Leg Ruptures In Ih.2.2.3.ha an analysis of the 36-in. ID, double-ended pipe rupture was presented. This accident produced the fastest blow-down and lowest heat removal from the fuel, therefore producing the highest cladding temperatures of any loss-of-coolant acci-dent. This was therefore the basis for design of the core flood-ing equipment. A decrease in the rupture size assumed results in decreased maximum clad temperature during a loss-of-coolant accident.

Core cooling evaluations have been performed for a spectrum of four additional rupture sizes using the same basic calculatic:c.1 technique and assu=ptions as for the large rupture case. These rupture sizes are 8.5, 3.0, 1.0, and 0.h ft2 The reactor coolant system pressure-time history for these rupture sizes is shown in Figure.14-44.

The reactor vessel water volume as a function of time after the rupture for the various rupture sizes is shown in. Figure 14-hk-a.

These water volume curves were generated utilizing the flov &

available from core flooding tanks, one high pressure injection 5 W pump, and one low pressure injection pump. The pumping systems were assumed to have a combined capacity of at least 3,500 gpm with the high pressure pump running on emergency power within 25 sec after the rupture, and the low pressure pump delivering 3,000 gpm when the pressure has decayed to 100 psi, or at 25 sec, whichever occurs later.

Figure 1k-kh-b shows the hot spot clad temperature as a function of time for the various rupture sizes. As can be seen from this figure, the small-sized ruptures yield maximum clad temperatures which are considerably lower than those resulting from the larger sizes. The results of this study are shown in the follow-ing Table 1k-5-1.

e Table IL-5-1 1

Tabulation of was-of-Coolant Accident Characteristics for Spectrum of Het Les Purture Sites Rupture Min. Wats: Leve, Below Hot Spot Size, Full-Pever Bottom of C:.re, Max. Temp.,

ft2 Seeen24 ft F ik.1 2.1 -6.8 1,950 8.5 3.k -5.2 1,916

.3.0 -2.2 1,235 1.0 1. 5( * ))

1.5(* +4.7 1.075 0.h 1.5 +12.0 1,015

(*) Blowdown forces on control rods are equal to, or less than, normal pressure drop, and control rods will insert with ncrmal velocities. These values are for trip shutdown rather than for a void shutdown, but include void reactiv-ity effects.

(2) Cold Leg Ruptures A similar analysis of a spectrum of rupture sizes has been made i for the cold leg piping. The rupture sizes tabulated are the double-ended, 28-in. ID inlet pipe, which yields 8.5 ft2 of rup-ture area, and the 3.0, 1.0 and 0.4-ft2 sizes.

O w/ The reactor coolant system average pressure for this spectrum of rupture sizes as a function of time is shown in Figure lk-kh-c. The water level as a function of time is shown on Figure lk-kh-d. The water level calculation has been based up- I on uninhibited flooding as the check valves are provided in the core support barrel to equalize pressures and permit the trapped steam above the core to escape out the rupture.

The hot spot temperature as a function of time for the spectrum  ;

of cold leg leak sizes is shown in Figure lk-kh-e. The results ,

of this analysis are shown in the following Table 14-5-2. j i

)

Table lk-5-2 l Tabulation of Less-cf-Ccolant Accident Chsr e c istics j for Erectrr: Of Cold leg Furture Si en 1

Fupture Min. Water Level Below Het Spct Size Full-Power (*) hettom of Core, Max. Temp.,

ft2- Seconds ft F 8.5 0.L(*) -6.7 1,785 30 1.0l") -L.8 1.575 1.0 1.3(*) +3.6 1,250

-0.L 1.3 +7.0 1,090

(*) Blowdown forces on cor. trol rods are equal to, or less tnan, normal pressure drop, and centrol rods will insert with nor-mal velocity. These values are fer trip shutdevn rather thsn l

-.g' void. shutdown, but include reactivity effects. i l

lk-k1b (Revised 4-8-68) 0053 I

(3) Evaluation of Emergency Coolant Injection Line Failure 1 The evaluation of a low pressure injection line failure has been '

made, and the results of the analysis show that the reactor is protected. The rupture of a pipe which connects a core flood-ing tank and the low pressure injection flow to the reactor ves-sel was assumed to fail adjacent to reactor vessel and before the first check valve. (See Figure 6-1.) This pipe has an in-ternal diameter of 115 in. , and the resultant rupture area is 0 72 ft 2, Interpolation of available blowdown calculations has been used to evaluate this rupture size, and the data show that a rupture of this size would result in the core being uncovered several feet below the top of the core. However, the hot spot vill never be uncovered, and peak cladding temperatures will be slightly less than that shown in Figure 14-kh-e for the 1.0 ft2 cold leg rupture.

Since this small rupture size leaves a considerable water in-ventory in the reactor vessel, the remaining core flooding tank inventory is more than adequate to completely reflood the core.

The other low pressure system can supply 3,000 gpm of water to 5 the reactor vessel and provide coolant to keep the core cooled.

The combined capacity of the two high pressure pumps is 1,000 gpm which is in excess of the boiloff rate (680 gpm) due to de-cay heat immediately after blowdown. With a single 500 gpm g high pressure injection pump the excess water above the core is adequate to prevent the core from being uncovered below the three quarter elevation and beyond 300 see the water level vill begin to increase.

The high pressure injection system has two independent chains of flow to supply borated coolant to the system. If a rupture of high pressure injection piping were to occur in one of the four lines between the attachment to the primary pipe and the check valve, the other chain of this system would have adequate capacity to protect the core against this small leak. In the event of a component failure in the second high pressure injec-tion loop, the ruptured flow path can be monitored by the opera-tor and spillage flow can be stopped by isolation of the affect-ed piping. The entire capacity of one pump can then be utilized to handle the small rupture and protect the core.

(h) Evaluation of Emergency Core Injection System 1 Performance for Various Rupture Sizes The loss-of-coolant analysis is based upon the operation of one 5 high pressure injection pump (500 gpm), one low pressure injec-tion pump (3,000 gpm), and the operation of the core flooding tanks. The capability of other combinations of engineered safe-guards to provide core protection he.s been evaluated in a pre- g liminary analysis. This capability is shown on Figure lk-kh-f. W lk-kle (Revised k-8-68) ()(154 u _

In this ev41uation the core is considered protected if the com- l' bination c.-emergency cooling systems considered will prevent ~

core damage which would interfere with further core cooling.

([

The high pressure injectica_ equipment with one pump operating l5 can accommodate leaks up to approximately 3 in. in diameter.

The preliminary analysis upon which this conclusion is based in-dicates that this pump will probably have the capability to pro- l 5 tect the core for leaks somewhat larger.

A combination of one high pressure pump and one low pressure l5 injection pump will protect the core up to a 0.h-ft2 leak.

This is equivalent to the rupture of a pressurizer surge line.

One high pressure injection pump plus two low pressure injec- l5 tion pumps can protect the core up to leak eizes of 3.0 ft2 This is considerably in excess of any of the piping connecting to the reactor coolant system. High pressure injection, plus the ccie flooding tanks and one low pressure injection pump, can protect the core up to 14.1 ft2 which is a double-ended rupture of the 36-in. ID, hot leg piping.

The corethe protect flooding tanks core from and one about low leak a- 3-in. pressure up toinjection pung can the lh.1-ft leak. Figure lk-kh-f demonstrates that high pressure' injection system provides core protection for normal operating leakage and for small leaks in which pressure decay of the system may be s ov. For intermediate leak sizes, either the core flooding tanks or low pressure injection protects the core following the

("]

\m/

lcss-of-coolant accident. For very large leaks in the category of a double-ended rupture of the reactor coolant piping, the core flooding tanks and low pressure injection together protect the core. For these leaks the core flooding tanks provide imme-diate protection and can protect the core for severa1' minutes following the rupture. Due to their limited volume, they must be supplemented by the high flow from the luv pressure injection pumps within several minutes following the leak in order to pre-vent the core from again becoming uncovered as a result of boil-ing off the core flooding tank coolant.

This evaluation of emergency core cooling capability demonstrates.

that the core is protected for the entire spectrum of leak sizes in both hot and cold leg piping, i

D U 0055 lk-kld (Revised h-8-68) .. l l

t l

• 1

. + ~ . - ]

c. Retor Building D: sign Baso Accidnnt l1 A range of leak sizes between 0.h ft2 and 14.1 ft2 has been evaluated.

The 14.1 ft2 ia equivalent to a double-ended rupture of the 36 in. ID re-actor outlet piping. The reactor operating conditions used in this analy-sis are listed in Table 1b-6. g,

/

The basis for this analysis is that only the makeup and purification sys-tem and the decay heat removal system are verking. It was assumed that the makeup and purification system (high pressure injection) had one of the pumps available for operation and that the decay heat removal system 2 (low pressure injection) had both of the two pumps available for operation.

These systems are assumed to operate on emergency power and can be in op-eration to deliver a total injection flow of 6,500 gpm within 2h see after the accident occurs.

This approach is conservative since any combination of two flooding tank operations and minimum flow from the high and low pressure pumps vill pro- 2 vide a lover energy release rate and peak reactor building pressures than those resulting from the 6,500 gpm flow.

During blowdown mass and energy releases to the reactor building are cal-culated by FLASH. Figure lb L3 is a plot of mass released to reactor build-ing and Figure lk kh is a plot of reactor coolant average pressure, each calculated by FLASH for the spectrum of hot leg ruptures. Following blow-down a 20-region SLUMP model vss used to simulate the core thermal trans-ient. This simulation includes fuel heat generation e metal-vater reaction, and quenching when the injection water provided cooling by contact with the core.

As any given segment reached h,800 F it was assumed to drop into water be-lov the core and release all heat down to a datum of 281 F. Also, it was assumed that 10 per cent additional zirconium-vater reaction occurred.

When the water covered approximately 25 per cent of the core, the surface heat transfer coefficient from all the core clad to the water was assumed to be 100 Btu /hr-ft 2 -F. The determination of water level was based on in-jection flow and included the effects of boiloff.

Assuming a pool boiling coefficient of 100 for the whole core when only 1/h was covered was conservative for reactor building pressure analysis because it compressed overall energy transport into the shortest credible period.

Heat was also released from the hot metal of reactor coolant system and the reactor vessel internals. During the blevdown period a surface heat transfer coefficient of 1,000 Btu /hr-ft 2 -F vas used. After blevdown this coefficient was changed to 100 Btu /hr-ft2 -F for the metal below the leak and 5 Btu /hr-ft 2-F above the leak. The coolant sink temperature was nro-vided by FLASH for the blevdown period and assumed to be 281 F theresiter.

The internal heat transfer of the metal was based on a multilayer finite difference model. The whole process of reactor coolant system metal heat transfer was simulated with a digital computer program.

All heat transferred from the core and the reactor coolant system metal was assumed to generate steam without taking credit for the subcooled con-dition of the injection water (except for the portion which was boiled off) until the reactor vessel was filled to the leak 1h-h2 (Revised 2-7-68) O 0056

,~

(') height. Thereafter all energy was removed by low pressure injection flow of subcooled water, and the energy release to the reactor build-ing atmosphere terminated. No delay was assumed in transporting steam to the reactor building. The heat from hydrogen burning was added directly to the reactor building as hydrogen was evolved from the metal-water reaction.

Both reactor inlet (cold) and reactor outlet (hot) line breaks were analyzed with FLASH. However, a complete analysis was made only for the hot line breaks since they provided for the most rapid heat trans-port from the core. This was true because the hot line breaks had longer blowdown and better heat transfer during blowdown than did the cold lir.e breaks.

The results of calculations of fluid and heat transport to the re-actor building as determined by FLASH, SLUMP, and other analytical models were-used as input to the 1_ actor building pressure analysis program, CONTEMPT.

Table 14-6 Reactor Operating Conditions for Evaluation Parameter Value

() Reactor Coolant System Pressure, psig Reactor Coolant Average Temperature, F 2,185 58h Reactor Power Level (ultimate), MWt 2,5hh Reactor Coolant System Mass, lb 519,173 Initial Reactor Building Temperature, F 110 Initial Reactor Building Relative Humidity, % 0 Initial Reactor Building Pressure, psig 0 In calculating the reactor building pressure, it was assumed that the 3 average temperature of the building atmosphere and structural mate-rials was 110 F. Upon release of hot reactor coolant, the steel and concrete act as heat sinks which reduce the reactor building pressure.

The heat sinks considered in this analysis are specified in Table 14-7 O(/

0057, 14-43 (Revised 1-15.68)

'O Table 14-7 Reactor Building Structure Data for Analysis of Time-Dependent Reactor Building Pressure Parameter Value Reactor Building Free Volume, ft3 2,000,000 Exposed Liner Plate Surface, ft 2 87,220 Mass, lb 1,238,000 Dome and Wall Liner Thickness, in. 0 375 Refueling Cavity Liner Thickness, in. 0.250 Reactor Building Concrete Enclosure Consisting of a 3-ft-Thick Dome and 3-ft, 6-in.-Thick Walls and a 2-ft-Thick Flocr Wall and Dome Surface, ft 81,700 Wall and Dome Mass, lb 41,100,000 Dcposed Floor Surface, ft2 11,000 Exposed Floor Mass, lb 3,190,000 Structural and Misceilaneous Steel Exposed to Reactor BuiMing At=osphere Surface, ft2 < 80,000 Mass, lb 500,000 Internal Concrete Surface, ft2 102,280 Mass, lb 23,398,500 Refuelira Cavity Concrete Surface, ft2 5,520 Mass, lb 3,001,500 1h-hk .

Heat transfer from the reactor build ng atmosphere to the steel liner was calculated using a condensing coefficient of 620 Btu /hr-ft2 -F fl until a totcl heat input of 110 Btu /ft2 had been achieved. There-R' after, a condensing coefficient of 40 Btu /hr-ft -F2 was used.

For heat transfer from the reactor building atmosphere to the con-crete, a condensing reefficient of 40 Btu /hr-ft2-F vas used. For heat transfer from the sump water to the concrete floor a coefficient 2

of 20 Btu /hr-ft -F was used. No credit was taken for heat transfer to reinforcing steel in the internal concrete structures.

For structural and miscellaneous steel, one heat transfer section with an equivalent thickness of 0.153 in. vas used. condensing coef-ficients of 620 and 40. Btu /hr-ft 2-F vere used.

Following a loss-of-coolant accident, the reactor building is cooled by three reactor building emergency cooling units and a spray system.

Each cooling arrangement has a heat removal capability of 240 x 106 Btu /hr at a vapor temperature of 281 F. Two cooling units plus 1,500 gpm sprays, or 3,000 gpm sprays, provide cooling that is at least equivalent to the three reactor building emergency cooling units.

Each system is designed so that it alone can protect the reactor building against overpressure. Each system was assu=ed to operate on e=ergency power and was delayed until 35 see after the rupture occurred.

Figure 14-45 shows the reactor building pressure for ccm anceofa36in.IDreactorcoolantsystempipe(14.1ftgletesever- rupture

\-

area) with 6,500 gpm of barated water injection into the reactor cool-ant system beginning 25 sec after the rupture. Reactor building cool-ing is provided by three e=ergency cooling units. The peak pressure r resulting from this accident occurs 181 see after the rupture at a value of 52.1 psig.

An analysis of the reactor building pressure for the 36 in. ID pipe rupture and spray cooling of the building has also been performed to de=onstrate the effectiveness of this system. Initially coolant for the building sprays and for injection to the core is pumped from the borated water storage tank. When water from the borated water stor-age tank is .epleted, the water collected in the reactor building sump is reci culated through the reactor building sprays and through the decay he.t removal coolers to supply the. low pressure injection water. The result is an increased injectionMad spray water tempera-ture. No boiling of the injection water results from this decrease in subcooling. The reactor building spray effectiveness vill decrease.

The net result is a decrease in the energy removal rate from the reac-tor building atmosphere.

-The requirements for cooling the water recirculated from the reactor building sump to the reactor building spray system are set by the de-sign basis of this system. The design basis is to =aintain the post-accident reactor building pressure below the design value. This criterion can be met by spraying the sump water directly into the re-

, actor building atmosphere without additional cooling, other than that p provided by the decay heat removal system.

o 0059 14-45 .

e

_ _ _m. _

The water temperature in the reactor building su=p during the recir-culation phase of a loss-of-coolant accident is maintained below sat-uration temperature by the decay heat removal coolers. These coolers ,

reduce the temperature of water recirculated to the reactor vessel and returned to the reactor building sump. The heat transfer surface of these coolers is set by the normal operating conditions under the decay heat removal operation mode. The cooling capability of this mode of operation vill maintain the reactor coolant at 140 F or less at 20 hours2.314815e-4 days <br />0.00556 hours <br />3.306878e-5 weeks <br />7.61e-6 months <br /> after extended rated power operation and is in excess of that required under accident conditions. The performance of these coolers tt various inlet temperatures is shown in Figure 6-4.

Figure lk h6 shows that the reactor building pressure decays to less than 5 psig in 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />. For co=parison purposes and to show that the effect of spraying cooler water into the reactor building is smil, a second curve is presented on Figure 14 46 which is based upon a spray recirculation cooling rate of 100 x 106 Btu /hr (approximately equivalent to two decay heat re= oval coolers) at a sump temperature of 195 F. . (This is the te=perature of the suup when recirculation to the l23 sprays begins.) Figure 1k-47 shows the te=perature of the reactor building and sump coolant for the two conditions.

These curves demonstrate that cooling of the recirculated spray water has no effect on peak building pressure and only a =inor effect on the rate of pressure decay during the first 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />. Accordingly, it is concluded that no cooling of the reeirculated spray water is required for this accident.

Figures 14-48 through 14-52 show the reactor building pressure for the other rupture sizes analyzed with the same cooling capability as the 14.1 ft2 rupture above. A su==ary of the input parameters snd results for the spectrum analysis are tabulated in Table 14-8.

A 3 0 ft2 rupture area results in the highest postaccident reactor building pressure (see Figure 14 k9).

Figures 14-53 and 14-54 show the reactor building energy inventory as a function of time after rupture for 14.1 and 3 fta rupture areas with three emergency coolers operating. These curves show the effective-ness of the reactor building structures and emergency cooling units.

Figures 14-55 and 14-56 show the reactor building vapor temperatures and sump te=peratures following 14.1 and 3 0 fta ruptures.

The peak reactor building pressure shown in this evaluation for the spectrum of leak sizes results is 52.1 psig and is the result of a 3 0 fta rupture in the reactor outlet piping. The reactor building design pressure is 55 psig and a design margin of about 3 psi exists.

With core flooding tank operation this margin would be increased further.

The above analyses conservatively assu=e that the hydrogen liberated vill burn at the rate for=ed, and that no core flooding tank operation occurs. The core flooding tanks J imit the amount of zirconium water 3 ,

14-46 (Revised 3-1-68)

O

... 0060 - _

I'}

b> reaction to 0.005 per cent for a 36 in. ID pipe rupture, and the po-tential hydrogen energy release is approximately 4,000 Btu. This amount of energy will not significantly affect reactor building pres-sure if ignition is delayed or if the hydrogen burns as formed.

For the case of no core flooding tcnks, as used in the abcve reactor building design pressure evaluation, the amount of metal-water reac-tion is somewhat greater. The zirconium-water reaction begins at 40 see and stops at 130 sec, by which time the 6,500 gpm of injection flow provides sufficient coolant inventory to the reactor vessel to recover the hot spot and quench the reaction. The steam flow during this period is assumed to provide the transport mechanism for the hy-drogen generated. The resultant concentration of hydro 6en (at time of maximum metal-water reaction rate) in the steam leaving the reac-tor vessel is 7 2 volume per cent. This concentration is below the flammability limit. Further dilution vill occur as the steam enters the reactor building, and combustion will not occur, even as the re-actor building is depressurized.

Criterion 17 of the AEC General Design Criteria states that the con-tainment (reactor building) be designed to accommodate the largest credible energy release including the effects of credible metal-water reactions uninhibited by active quenching systems. Although the eval-uation of the emergency injection systems demonstrates that only a small amount of metal-water reaction can occur, the case *of no injec-(7 tion flow has been evaluated in response to the above criterion.

O This case assumed that, after blowdown, the reactor vessel would have water up to the bottom of the core. The core was allowed to heat up by decay heat and metal-vater reaction heat.

Steam flow rate-limiting of the reaction was not considered so long as any water was assumed to be in the vessel. If and when the clad reached the melting temperature, it was assumed that the whole region slumped into the bottom of the vessel with the attendant reaction of 10 per cent more of the remaining zirconium and with the release to the reactor building of all sensible and latent heat above 281 F.

Upon ecmpletion of boiloff, heat input to the reactor building was assumed to cease. Figure 14-57 shows a reactor building pressure of 53 2 psig at 220 seconds, the time at which the reactor vessel boils dry. This peak pressure is below the 55 psig design pressure of the reactor building.

'f3 U

... 0061 14-47 e )

.. a .- .n_. --- - . - - - - .

Table 14-8 Su==ary of Reactor Building Pressure Analysjs for Reactor Building Fr.ergency Cocling (240 x 106 Btu /hr)

Rupture Size, ft 2 14.1 85 30 2.0 1.0 0.4 Reference Figure No. lk-45 lk-h8 lk-h9 lk-50 lk-51 14-52 Time Blovdown Ends, see 15 20 48 68 141 351 Time lov Pressure In-jection Begins, see 25 25 39 59 121 321 Fraction of Core Zr-react.ed 0.08 0.05 <0.01 zero zero zero Time Zr-reaction Beginc, sec 40 50 130 -- -- --

Time Zr-reaction Ends, see 130 130 131 -- -- --

Time to Reach Peak Pressure, sec 181 181 41 67 181 261 Peak Building Pressure, O psig 52.0 51 3 52.1 50 7 48 9 43 2 Vapor Temperature at Peak Pressure, F 278 277 278 276 274 266 Sump Temperature at Peak Pressure, F 232 230 221 215 210 196 Conditions for All Cases

a. 500 gpm high pressure injection

b. 6,000 gpm low prescure injection

c. Reactor hot leg rupture

d. No core flooding

e. No reactor building sprays

f. Tnree emergenc:/ cooling units start 35 see after the rupture.

O 14 48 00bA

i I/ d. Reactor Building Zirconium Reaction Capability l1 In order to determine the theoretical ultimate circonium reaction capability of the reactor building a series of hypothetical accidents was investigated.

Blowdown was based on the IL.1 ft2 leak case. Heat transfer from the core and all reactor coolant system metal below the leak height was assu=ed to transfer to a 281'F sink based on a surface coefficient of 50,000 Etu/hr-ft 2 -F. For reactor coolant system metal above the leak height 5 Btu /hr-ft2 -F was used.

Available core heat consisted of the initial secred heat, the equiva-lent of two ftl1 power seconds, decay heat, and metal-water reaction heat , whien was added at arbitrary linear rates. The total heat transferred from the core and reactor coolant system metal was as-st=ed to produce steam frc= vater initially at the saturated condi-tion. Hydrogen recombination energy was added to the reactor build-ing as superheat at the rate of hydrogen producticn from the zirco-niu=-vater reaction.

A series of calculations for each of the various cooling capacities was made varying the energy input rate, i.e., Zr-H2 O reaction rate.

For exa=ple, a 1 per cent per second zirconiu=-vater reaction produces g 1.173 x 100 Etu/see of =etal-water energy and 0 902 x 100 Btu /sec

(_ hydrogen reco=bi..ation energy. In all cases the energy was input at a linear rate bejinning 10 see after the rupture. The emergency cool-ing units and spray coolers were started 35 see after the rupture.

The " time to complete reaction" is the time it takes to reach reactor building design pressure (55 psig).

The results of this study are presented in Figure 1L-58. This amount of allovable zirconius reaction at any time after blowdown depends upon the a=ount of reactor building cooling in operation. The capa-bility curves show that at approximately 10 sec, when the blevdown pressure peak occurs, the reactor building could accept an instan-taneous zirconium-vater reaction of k per cent. This capability in-creases greatly after the blevdown pressure peak with reactor build-

.ing cooling equipment in operation.

With three emergency cooling units in operation a 100 per cent reac-tion in 4,200 see vill not exceed the design pressure of 55 psig.

With three emergency cooling units and two sprays operating, a 100 per cent reaction in 1,L20 seconds will not exceed the design pres-sure.

.. 0063 lh-h9 (Revised 1-15-68)

14.2.2 3 5 Environmental Analysis of Loss-of-Coolant Accidents Safety injection is designed to prevent significant clad melting in the event of a loss-of-coolant accident. The analyses in the preceding sections have demonstrated that safety injection vill prevent clad melting for loss-of-coolant accidents resulting from reactor coolant system ruptures ranging in size from small leaks to the complete severance of a 36 in. ID main coolant pipe. Without clad melting, only the radioactive material in the coolant at the time of the accident plus some gap activity is released to the reactor building.

The environmental consequences from a loss-of-reactor-coolant accident are ana-lyzed by assuming that 1 per cent of the fuel rods are defective before the re-lease of reactor coolant to the reactor building. Table 11-3 lists the total activity in the coolant. In addition to the coolant activity, the activity associated with the gap of all fuel rods is also assumed to be released. Cal-culations indicate that 77 per cent of the fuel rods vill have some point along their lengths with temperatures in excess of1,200 F at the time of core flood-ing tank injection. While perforation of fuel cladding vill require some time, it is conservatively assumed that all of the fuel rods release their gap activ-ity during the accident.

Half of the iodine released is assumed to plate out en exposed surfaces in the reactor building. The other half is assumed to remain in the reactor building atmosphere where it is available for leakage. The sodium thiosulfate in the reactor building spray reduces the airborne iodine as described below. Of the iodine available for leakage, 5 per cent has been conservatively assumed to be unavailable for removal by the spray.

g The rate at which the elemental iodine can be removed from the reactor building atmosphere by the reactive spray is calculated using Griffith's methods.(14)

This method is based on the work of Taylor,(15) who shoved that the rate at which elemental iodine can be transferred into reactive solutions is controlled by the gas film resistance, and on the work of Ranz and Marshall,(16) who shoved that the equation below can be used to calculate the mass transfer co-efficient when the rate of transfer is controlled by the gas film resistance:

DpM7 /2 1/3 -

kg = -2 + 0.6 [dvP -

MmdP ,

\F Dp ,

where 2

kG = gas film mass transfer coefficient, gm/cm -see-atmos D = diffusivity of iodine in air, em 2/see p = density of air, gm/cm3 MI = molecular weight of iodine, gm/gm-mole 4 = mean molecular weight of the air-iodine mixture in the boundary layer P = partial pressure of air in the gas film, atmos 14-50 0064 e, mob

~--

O-V d = drop diameter, em v = relative velocity between the drop and the gas phase, or approximately the terminal velocity of the drop, em/sec

= viscosity of the air Since the mass transfer of iodine is gas-film-controlled, kg is approximately equal to Kg (below), and the foregoing equation can be rewritten in tems of the velocity of deposition, V g:

Rr RT b p 1/ 1/3-

,2 + 0.6\Fldyp l\DP y =-K =

g My g My M di P wherc Vg =overallvelocityofdeposition,em/sec R = universal gac constant = 82.057, atmos-cm3/K_gm_ mole T = absolute temperature, K Kg=overallmasstransfercoefficient,gm/cm-sec-atmos 2 Since the maximum possible iodine concentration in the large volume in the re-actorbuildingislessthan10-7gm/ce,thepartialpressureofairinthegas film, P, can be taken as the total pressure, and the mean molecular weight, Mms can be taken as the molecular weight of air,gM . If the gas equation is used, the equation may be simplified somewhat by substituting MgET for p/P, as follows:

,D- 2 + 0.6 [dvpgl/2 [.F_jl/3' g d \V) \ D p)

The surface area of drops available for iodine absorption can be calculated from the next equation, which is based on the assumption that all the drops are spherical and have the same diameter.

2 6F0

• 6FH 8

• nd F0

• d w d3 dv

'd .

1 where S = surface area of drops suspended in the gas phase, em2 F = spray flow rate, em3/sec 0 = drop fall time, sec  !

d = drop diameter, em H = drop fall height, em v = drop fall velocity or teminal velocity, em/sec . 0065 14-51 i

If there is a large excess of chemical reagent to react with the iodine and convert it to a nonvolatile form with little or no tendency to return to the gas phase, then the iodine removal rate can be expressed by d1,,,[,VSh_g,_

,I , , sI dt (V ej where V

c

= free v lume of reactor building, em3 or ft 3 Ag = iodine removal time constant, br-1 The fraction remaining in the reactor building atmosphere is expressed as a function of time by the solution of the equation above as follows:

I ~As t.

[o=e where I- = fraction of initial inventory remaining o

t = spray time, hr When the specific parameters for Unit 3 or 4 of the crystal River Plant are used:

F = 3,000 gPm v = 397 cm/see H = 90 ft Vs = 5 06 cm/see 6

V = 2 x 10 ft3 6V FH E

A 8

=

Ve dy = 25 3 hr-1 d = 1,000 microns These iodine removal calculations have conservatively corrected the iodine deposition velocity (V g) to the peak temperature and pressure in the reactor building. A sensitivity analysis was performed on the iodine removal calcula-tions, and the results are shown in 14.2.2.4 3 in ter=s of the 2-hour iodine dose at the exclusion distance following an MHA.

Although the reactor building leakage rate vill decrease as the pressure de-caya, the leakage is assumed to remain constant at the rate of 0.25 per cent per day for the first 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />. Thereafter, since the reactor building vill i have returned to nearly atmosphere pressure, the rate is assu:: led to be reduced l to 0.125 per cent per day and remain at this value for the duration of the ac-l cident.

l The atmospheric dispersion characteristics of the Plant site are described in 23 The site dispersion factors for the duration of the accident are listed in Table 2-3 A breathing rate of 3 47 x 10-4 m3/see is assumed for the 2-hour g 14-52 0066

exposure. For the 24-hour exposure, a breathing rate of 3 47 x 10-4 m3/seeis assumed for the first 8 hours9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br />, and a rate of 1 74 x 10-4 m3/seeisassumedfor the remaining 16 hours1.851852e-4 days <br />0.00444 hours <br />2.645503e-5 weeks <br />6.088e-6 months <br />. For the 30-day exposure, a breathing rate of 2 32 x 10-4m3/seeisassumed.

The iodine doses to the thyroid per curie inhaled are obtained from the values

-given in TID-14844:

, I-131 1.48 x 106 rem per curie I-132 5 35 x 104 rem per curie I-133 4.0 x 105 rem per curie 25x10 rem4 per curie I-134 I-135 1.24 x 105 rem per curie Figure 1k-59 shows the total integrated dose to the thyroid as a function of distance from the reactor building for 2-hour, 24-hour, and 30-day exposures.

The total thyroid dose at the 4,400 ft exclusion distance is 1.45 rem for a 2-hour exposure, 8.0 rem for a 24-hour exposure, and 9 9 rem for a 30-day ex- l1 posure. These doses are vell below the guideline values of 10 CFR 100. The direct dose from this accident is insignificant since it is several orders of magnitude below 10 CFR 100.

14.2.2 3 6 Effects of Reactor Building Purging At times during the nor=al operation of the reactor, it may be desirable to purge the reactor building while the reactor is operating. In the event a loss-of-coolant accident were to occur during purging operations, activity would be released to the environment. The purge values will be completely closed in 5 sec. During this time, assuming a 36 in. ID, double-ended rupture, essentially all of the reactor coolant will have been blown down. The activity in the reactor building is due to the reactor coolant activity after operation with 1 per cent failed fuel. For this case, 0 53 per cent of the reactor building atmosphere vill escape through the purge valves before they close, ,

corresponding to a release of 3 equivalent curies of iodine-131. This analy- 1 sis assumes unrestricted flow through the purge line for the full 5-second  ;

closing time. No reduction in flow is assumed as the valve closes, and there- i fore the results are conservative. The release of this iodine results in a l total integrated thyroid dose of 0.48 rem at the exclusion distance. This dose, i when added to the thyroid dose for a loss-of-coolant accident without purging, j is well below the 10 CFR 100 guidelines. Therefore, purging operations can be perfor.ned during reactor c. stion.

i i

p L) .

0067

! 14-53 (Revised 1-15-68)

lk.2.2.k Maximum Hypothetical Accident 14.2.2.4.1 Identification of Accident -

/

The analyses in the preceding sections have demonstrated that even in the event of a loss-of-coolant accident, no significant core melting vill occur.

HovcVer, to demonstrate that the ope;ation of a nuclear power plant at the proposed site does not present any undue hasard to the general public, a hypo-thetical accident involving a gross release of fission products is evaluated.

No meenanism whereby such a release occurs is postulated since f t would re-quire a cultitude of failures in the engineered safeguarxis provided to prevent its occurrence. Fission products are assumed to be released from the core as stated in TID-1484h, namely, 100 per cent of the noble gases, 50 per cent of the halogens, and 1 per cent of the solids.

Further, 50 per cent of the iodines released to the reactor building are as-sumed to plate out. Other parameters, such as meteorological conditions, io-dine inventory of the fuel, reactor building leak rate, reactor builaing io-dine removal rate, etc., are the same as those assumed for the loss-of-coolant accident in 14.2.2 3 5 The average iodine inventory, in terms of equivalent curies of iodine-131 available for leakage at different time periods after the accident, is as follows:

0 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> 28.7 x 10 curies O to 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> 22.8 x 10 curies 1 to 30 days 5 1 r 10 curies 14.2.2.4.2 Analysis and Results of Environmental Analysis Figure lh-60 presents the total integrated dose to the thyroid as a function O of distance from the reactor , building for 2-hour, 24-hour, and 30-day expo-sures. It can be seen that the 2-hour thyroid dose of 65 rem at the exclusion distrunce of h,400 ft and the 30-day thyroid dose of 3.h rem at the k1-mile low population zone distance are less than the guideline values of 10 CFR 100.

In the year 2015, the projected population within a 5-mile radius of the Plant vill be less than 1,000. The corresponding 30-day thyroid dose from the EA at the 5-mile zone boundary is 38 rem.

The direct dose to the whole body following the accident is shown in Figure 1L-61. No significant dose exists from this source at the exclusion distance.

The dose to the whole body from the passing cloud has been calculated using the same meteorological conditions used for determining the thyroid dose. The 2-hour whole body dose at the exclusion distance is only 1 9 rem, and the 30-day dose at the kl-mile low population zone dictance is 0.11 rem. The 30-day dose at the 5-mile zone boundary is 1.2 rem.

14.2.2.4 3 Effects of a Sensitivity Analysis of the Reactor Building Sprays for Iodine Removal A sensitivity analysis on the calculation of iodine removal was performed us-ing the reactive chemical sprays in the reactor building. The results are shown in Table 14-9 in terms of the 2-hour iodine dose at the exclusion dis- -

tance following an MdA.

14-54 0068

- - . = - .. - - -. .. . ._ -

.. . . . - . . ~-

T

_Y  %)

Table 11-94 Sensitivity Analysis Showing the Effect of Parameters on the Two-Hour Iodine Dose at the Exclusion Distance Following an !! iia

. Todine(l) Iodine (2)

Reswal Removal Drop Drop Fall. Velocity of Tim Iodine (l) Time Iodine (2)

Case Size, Velacity, Deposition, Temp, Pre ss . , Constant, Dose, Constant, Dose, No. microns en/sec en/see F psig br*l rem hr-1 rem Remar ks

.I 1,000 3c/T 5.c6 281 55 12.65 & 25 5 65 operation of the reactor building spray system at -

marfen building tempera-ture and pressu'v.

2 1,000 3CJr 6.44 212 25 15.05 75 32.1 51 Operation of the reactor g building spray system after g- partial cooling, about 1 1

.vif hour, 3 1,000 397 11 55 100 0 28.8 62 57.6 54 operation of the reactor building spray system after cooling to ambient condi-

, tions.

4 1,000 3,(170 14.85 281 55 5.7 171 7.39 109 Effect of drop falling at 10 times its terminal velocity.

5 2,000 649 4.25 281 55 5.25 192 6.50 116 Effect of large drop site.

6 200 76 7.24 281 55 471 46 942 46 Effect of small drop size.

For at! cases, reactor building free volume = 2 x 10 ft . and drop fall height = 90 ft.

, Notes : (1) Flow rate of sprays = 1,500 spm.

(2) Flow rate of sprays = 3,000 gpm.

O CD Ch

14.2.2.4.4 Effects of Engineered Safeguads Leakage During the Maximum Hypothetical Accident An additional source of fission product leakage during the maximum hypotheti-cal accident can occur from leakage of the engineered safeguads external to the reactor building during the recirculation phase for long-tem core cooling.

A detailed analysis of the potential leakage from these systems is presented in 63 Thatanalysisdemonstratedthatthemaximumleakageisabout5,000cc/hr.

It is assumed that the water being recirculated from the reactor building sump through the external system piping contains 50 per cent of the core saturation iodine inventory. This is the entire amount of iodine release from the reac-tor cooling system. The 50 per cent escaping from the reactor coolant system is consistent with TID-14844. The assumption that all the iodine escaping from the reactor coolant system is absorbed by the water in the reactor build-ing ir conservative since much of the iodine released from the fuel vill be plated out on the building valls. The activity in the recirculation vnter is equal to 0.037 equivalent curies of I-131 per cc of water. The iodine is chem-ically bound to the sodium thiosulfate, and vill not be released to the atmo-sphere. However, it is conservatively assu=ed that iodina release does occur.

Since the temperature of water in the reactor buildin6 cump is less than 200 F vhenrecirculationoccurs,theiodinereleaseiscalculatedusingagas/ liquid partition coefficient of 9 x 10-3 Leakage from the auxiliary building is caused by exfiltration. The most re-strictive case for a ground release occurs during inversion conditions. It is assumed that the building leaks at the rate of 100 per cent per day with atmo-spheric dilution occurring in the wake of the building. For this building leak rate and the ir ersion condition, the iodine vill produce an integrated dose to the thyroid af 0.005 rem in 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> at the 4,400 ft exclusion distance.

u,, m "

0070 . -

14-56 c .

14 3 REFERENCES (1) Watson, L. C., Bancroft, A. R., and Hoelke, C. W., Iodine Containment by Dousing in NPD-11, AECL-ll30.

(2) Styrikovich, M. A., et, al., " Transfer of Iodine from Aqueous Solutions to Saturated Vapor", Soviet Journal of Atomic Energy E , July 1964.

(3) Dispersion of Soluble Radioactive Material in Water, CF-58-3-109 (4) International Symposium on Fission Product Release and Transport Under Accident Conditions, Oak Ridge, Tec essee, April 1965 (5) Liimatainen, R. C., et al., Studies of Metal-Water Reactions at High Temperature, ANL-6250.

(6) Ackerman, R., et al., "High Te=perature Vapor Pressure of UO ", Journal 2

of Chemical Physics, December 1956.

(7) Reactor Development Program Progress Report, ANL-6912, June 1964.

(8) AEC Research and Development Reports, WIGL2 - A Program for the Solution of the One-Dimensional Two-Group, Space-Time Diffusion Equations Ac-counting for Temperature, Xenon and Control Feedback, WAPD-TM-532, October 1%5 O (9) Margolis, S. G. and Redfield, J. A., FLASH: A Pzugram for Digital Simu-lation of the Loss-or-Coolant Accident, WAPD-TM-534, May 1966.

(10) Grenda, R. J. and Patterson, J. F., "The Velocity of Rising Steam in a Bubbling Two-Phase Mixture", Transactions of The ANS 5, No. 1, p 151, June 1962.

(11) Possible Zirconium Water Reactions in Water Reactors, AEC Regulatory Staff Symposium, April 27, 1965 (12) Wagner, R. J. and Finnegan, L. J., "An Analytical Model for Predicting the Pressure-Temperature History Within a Containment Vessel in Response to a Ioss-of-Coolant Accident", Phillips Petroleum Company, Atomic Energy Division, Idaho Falls, Idaho, Presented at ANS Meeting, Washington, D.C.,

, November 1965 (13) Quinn, E. P., Forced-Flow Heat Transfer to High-Pressure Water Beyond the Critical Heat Flux, ASME 66WA/HT-36, November 27, 1966.

(14) Griffiths, V., The Removal of Iodine from the Atmosphere by Sprays, AHSB (S) R 45, 1963 (15) Taylor, R. F., " Absorption of Iodine Vapor by Aqueous Solutions", Chem.

Eng.Sa.,X,Eo.1/2,pp68-80, April 1959 (16) Ranz, W. E. and Marshall, W. R., Chem. Eng. Progress, Q , 141, 173, 1952.

0071 14-57

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j , i 10 ' ' ' '

/ .t  ! i i 0 Xl IM ' i i 20  !

I I

! I I THERMAL 10 j/ N w l

P OWE R , T' 0 ' -

i I I t 200 i FUEL 100 ^

j w ,I TEMPERATURE O

CHANGE,F l

l ,

/^T

(_) io '

I i !

l l' l l AVERAGE i i i i

- m

~ ~ ~~

CORE MODERATOR ' M TEMPERATURE O CHANGE,F i

l i i

"- 5 SEC+

2500 I I I

' L

2350

,  ;.pd- --

REACTOR , j SYSTEM '

2200 -

PRESSURE, PSIA ___

START UP ACCIDENT FROM 10 ' RATED POWER USING A 1.27.OK/K ROD CROUP; (Q

RJ HIGH PRESSURE REACTOR TRIP 15 ACTUATED CRYSTAL RIVER UNITS 3 & 4 0072 _

-- FIGURE 14 1

1 '

_q 1 ,

NEUTRON _

POWER,% 150 100

! I t 0  ! 11 \  !

I 20 l THERMAL 10 / 's POWER,fo ' -

0 1 '

t I i i 200  !

FUEL

/ '

TEMPERATURE N CHANGE,F 0 O

10  ! I i '

AVERAGE 5 , , ,_

CORE MODERATOR ,_ , , , ,

TEMPERATURE O /

CHANGE,F

+ 5 S E C-+-

2500 REACTOR SYSTEM 2350 -

[_q_-

PRESSURE, #

2200 PSIA l

START-UP ACCIDENT FROM 10 9 RATED POWER U5tNG ALL RODS WITH A WORTH OF 9,5% OK/K; HIGH FLUX REACTOR TRIP 15 ACTUATED CRYSTAL RIVER UNITS 3 & 4

~

0073

- FIGURE 14-2

O V.

1 e i i i ll 4 i i 6 ieit i i i i s 6ai i i e i i i il 80 _ . _ _

e

}. 60 E

g E

i

]g 40 -

High Pressure c = High Flux level Trip Level Trip A.

s.-< n e One control h

\

/

eo as croup 1

/

6 Nominal All Rods bo o i i i! !, , i e i,,,! i . . ie ,,. , e i ! t i ei 4 6 -5 2 19 4 6 -4 2 L 6 -3 2 o 4 6 10 1o-2 Rod Withdrawal Rate, ( \k/k)/sec PEAK THERMAL POWER VERSUS ROD

/

WITHDRAWAL RATE FOR A START-UP ACCIDENT FROM 10-9 RATED POWER CRYSTAL RIVER UNITS 3 & 4

. Em FIGUR E 14-3

O i 1 6 1 6ill '

i i i iliii e i i iiill i

~

I i 4 4ili 6 .

4 -

6 g .

I E

71 or .

g 6 -

• k _

J ~ All Rods m _

I E e -

8 2

go Single m Control Rod -

Group -

E 6 '~

~

k -

/ -

P .

1 10 t I t t ,,t, t , , ,,t!i t i . ,.iee i ,t

!  !!tt 2

10 4 6 10*5 6 10 4 2 4 6 10*3 2 4 6 0-2 Rod Withdrawal Rate, ( Ak/k)/see PEAK NEUTRON POWER VERSUS ROD CITHDRAWAL RATE FOR A START UP ACCIDENT FROM 10-9 RATED POWER CRYSTAL RIVER UNITS 3 & 4 3.- FIGURE 14 4 0075

O 30 25 e

20 2 \

\

)b 15  %

^

N _

~

4 Nominal g 10 2

5 0

0 -0.h -0.8 -1.6

-1.2 Doooler Coef ficient. ( Ak/k)/F x 10

-2.0 0

1 PEAK THERMAL POWER VER135 DOPPLER COEFFICIENT FOR A START-UP ACCIDENT U5lHG A 1.2% 6K/K ROD GROUP AT 5.8 x 10-5 (OK/K)/SEC. FROM 104 RATED POWER .

CRYSTAL RIVER UNITS 3 & 4 O

"~"

FIGURE 14-6 l

O P'

[

'O

, /

i 19 e o.

I 16 5

2

] 17 >

14 [

/\ Nominal 15 .

! ( I O C. 4 0.5 1.0 1.6 O Trip Delay Time. Sec PEAK THERMAL POWER VERSUS TRIP DELAY TIME FOR A START UP ACCIDENT USING A 1.2% OK/K ROD GROUP AT 5.8 x 10 5 (OK/K)/SEC.

FROM 10 RATED POWER CRYSTAL RIVER UNITS 3 & 4 0076 _

-- F iGUR E 14-5  ;

O ho r

/ ,

• 30 #

, /

! /

=

x 20 /

8 a.

( I Nominal 10

//

0 0.4 0.8 1.2 1.6 Trip Delay Time. sec l

l PEAK THERMAL POWER VERSUS TRIP DELAY TIME l FOR A START UP ACCIDENT USING ALL RODS AT  !

5.8 x 10-d (OK/K)/SEC. FROM 10-9 RATED POWER l CRYSTAL RIVER UNITS 3 & 4

= - = = FIGUR E 14-7 l

l l

O 40 I I I

( High Fitz IAvel Trip -

High Pressure

' Imvel Trip w.

\

b M \  ;

1 m I

E i E I a .

1 \ '

i Nomina.1 lo W 0 -0.4 -0.8 -1.2 -1.6 -2.0 Doppler Coef ficient. ( \k/k)/F x 10 PEAK THERMAL POWER VERSUS DOPPLER COEFFICIENT FOR A START UP ACCIDENT USING ALL RODS AT 5.8 x 10 d (OK/K)/SEC. FROM 10-9 RATED POWER -'

CRYSTAL RIVER UNITS 3 & 4

== FIGURE 14-8

O 26 .

2550 a

$ 2500

Safety Valve

a. 2450

/ Set Point

% [ e Nominal 2400

) ,

2550 0 0.4 0.8 1.2 1.6 P.0 Trip Delay Time. sec O

PEAK PRESSURE VERSUS TRIP DELAY TIME FOR A START UP ACCIDENT USING ALL RODS AT 5.8 x 10*d (OK/K)/SEC. FROM 10*' RATED POWER QQg] CRYSTAL RIVER UNITS 3 & 4

$~ FIGURE 14-9

O 2550 8*#'tY V"l

2525 Set Point 2500 \ T m

21 95

! s \

f 2450 N w g 2425 2400 4 6 8 yo Tripped Rod Worth, $ g /c PEAK PRESSURE VERSUS TRIPPED ROD WORTH FOR A START.UP ACCIDENT USING ALL RODS AT 5.8 x 10'd (OK/K)/SEC. FROM 10 RATED POWER CRYSTAL RIVER UNITS 3 & 4 g hE--. PicuRe u. i0

I l

O 2550 j g High Flux _ High Pressure 2525 Ievel Trip - Invel Trip

=

2500 k I S fety 3ve set Po nt I

I \

h 2475 i

nominal N

s i 2h50 1 1 2

2k25 2400 l 0 -0.4 -0.5 -1.2 -1.6 - 2. 0 Doppler Coefficient, (ak/k)/F x 10 5 l

1

~

PEAK PRESSURE VERSUS DOPPLER COEFFICIENT FOR A START UP ACCIDENT U5lHG ALL RODS AT 5.8 x 10 4 (OK/K)/SEC. FROM 10 RATED POWER CRYSTAL RIVER UNITS 3 & 4 no 1

== FIGUR E 14- 11 l 1

'O 2500 Pk90 E Pk80 s p #

e Pk70 ~

b #

2460 p

.g "

2 s Pk50 0 +8 41p .

Moderator Coefficient. (t.k/k)/' x M 5 6 l

l PEAK PRE 55UR.9 VER5US MODERATOR COEFFICIENT FOR A START UP ACCIDENT USING ALL ROD 5 AT 5.8 x 10 4 (OK/K)/5EC. FROM 10 RATED POWER ,

CRYSTAL RIVER UNITS 3 & 4 Od3 4 hE. FIGURE 14- 12

i i! i t '

i l i i i i I

(] NEUTRON i i I  ! I I

fu 120 ll'!! ,

POWER, % , j, 14 0 l l l k!

O ' ' I iI I I 1 ! i l l i 1 THERMAL 120 ll l POWER,5 80

\

l' l -q 40 l, I

O

- 4_. _

FUEL 200 TEMPERATURE '

O CHANGE,'F i \

-200 y

-400 , l ,

n v ,,

5 - - g AVERAGE ,

CORE MODERATOR TEMPERATURE 0 ,

I

! \ -

-5 CHANGE, F

-10 , , l  ;

=- 5 S E C-+

2500 2350 y

REACTOR - -- " \

SYSTEM 2200

(\

PRESSURE, 2050 -

PSIA -A 1900 ROD WITHDRAWAL ACCIDENT FROM RATED POWER USING A 1.2P. AK/K ROD GROUP AT r3 5.8 x 10 5 (OK/K)/SEC; HIGH PRESSURE s_) hQ@4 REACTOR TRIP 15 ACTUATED CRYSTAL RIVER UNITS 3 & 4 E. FIGURE 14 13

7

' ' _ S

_ _ e

_ _ a lA# _ a h

Mb 4 u

{ e a

m

_ 1 _ e

_ o _. g

- _ @ a

- N _ 1 a

4 g

Y w g a =

j 1 3 L

me - ~

33 O.

W m

i

_ _ S

_ _ a t f I l t R 8 R 8 R Q

$ $ $ $ N m sisd 8unseMd Mead PEAK PRESSURE VER$US ROD WITHDRAWAL RA?v FOR A ROD WITHDRAWAL ACCIDENT '

FROM RATED POWER CRYSTAL RIVER UNITS 3 & 4

3. FIGURE 14 14 0085

O l 2500 2k75 m

2450 >

c. f h' 2h25 A /

w f

/

2kOO ,

Nominal 2375 2350 Q 0 0.4 0.8 L2 L6 Trip Delay Time, see i

, PEAK PRESSURE VERSUS TRIP DELAY TIME FOR A 8.;J ROD WITHDRAWAL ACCIDENT FROM RATED POWER USING A 1.2P.6K/K ROD GROUP; HIGH PRESSURE REACTOR TRIP 15 ACTUATED CRYSTAL RIVER UNITS 3 & 4 0086 _

w. - FIGURE 1415

. .,..._..,...___2_.- . _ - - . -

l O\ l l

1 l

2425 , , ,

y High Flux = 1 High Pressure Trip g 2400 Trip a

{ 2375

/ Nomi nal i 2350 8

' /,  :

i 2325 e

CL 2300 0 -0.4 -0.8 -1.2 -1.6 -2.0 Doppler Coefficienr. (3k/'. )/ F x 10 l

l l

l I

l l

PEAK PRESSURE VERSUS DOPPLER COEFFICIENT FOR A RCD WITHDRAWAL ACCIDENT FROM RATED POWER USING A 1.2r.dK/K ROD GROUP CRYSTAL RIVER UNITS 3 & 4 dhs_ PioURe i4.ie

/ 120 s

Maximum Neutron Power (1177,),

100 7

we.

J Maximum Thermal y

so Power

\,

/

/

80 -

o /

70

[

60'

'O- 20 H) 60 80 100 initial Power, % of Rated

~-

MAXIMUM NEUTRON AND THERMAL POWER

• b FOR AN ALL ROD WITHDRAWAL ACCIDENT FROM YARIOUS INITIAL POWER LEVELS CRYSTAL RIVER UNITS 3 & 4 5o FIGURE 14-17

l I

i i i i i i

_ e l

~

Maximum Temp. in Hottest Fuel

~

] Rod (Hot Spot) y _ _

o x 30 u /

g -

= _

e

g. _ _

,o _

= _

. 20 x _

Max. Temp. in Average Fuel Rod _

p IX _

10 ' ' ' ' '

O 20 18 0 60 80 100 Initial Power, %

f PE AK FUEL TEMPER ATURE IN AVERAGE ROD AND HOT SPOT FOR AN ALL ROD WITHDRAWAL ACCIDENT FROM VARIOUS INITIAL POWER LEVELS O'089 CRYSTAL RIVER UNITS 3 & 4 E. FIGURE 14-18

O 100 80 2

I - 70,000 lb-ft N

60 E N N 3

N 3

e 40 d!

O 20 0

0 h 8 y y Time, see rX PERCENT REACTOR COOLANT FLOW AS A FUNCTION Q OF TIME AFTER LOSS OF PUMP POWER CRYSTAL RIVER UNITS 3 & 4

$~ FIGURE 14- 19

17 Trip Set Point Maximu;n (10'(.5%) l Overpower -

l N 1.6 ,

g Minit:u:n DNE Ratio x

in Hot Channel at li4T j \ l Power Steady State (1.38) a h de N l

\

q 15 -

\ g S@O N  : I = 70,000 lb-ft N I S ".,

o rE Fated -er -ith  :

Y e : ~ t,. t in; ..c a d%'. J l  %

g3 1.4 ' i u r, . '

as I N

]

E O

' N' l

f t I

• a 13 l l

l 1.2 !l 100 102 104 106 108 110 112 lik Overpower at 'n'hich Coastdown Begins, %

MINIMUM DNBR WHICH OCCURS DURING THE .

COASTDOWN FOR VARIOUS INITIAL POWER LEVELS CRYST AL RIVER UNITS 3 & 4 0091 hE_ eiouRe i4 2o

O 0.c6 0.05

\

N\

\

0.04 -

u \

2 g 0.03 \g I

o 0.0p N '

m

) 0.01 -

0 0 100 200 300 koo 500 600 Time after Break, sec REACTOR SYSTEM COOLING RATE FOR O 4 IN.2 STEAM LINE BREAK *--

CRYSTAL RIVER UNITS 3 & 4

{g 7 bd o FIGURE 14 21

.m. . _ _ , - - . ..

40 I

BOL Parameters g = -1.14 x 10'E (& /k)/F

" *

• I#I")/I 35 "M T

= 0 3 see Delay I = $.L7 x 10 see 50 '

EOL Parameters SL -

a = -1 56 x 10-5 (& /k)/F aM

• A' *""" 2##

25 t = 0 3 see Delay _

k t

= 2 75 x 10-5 sec [

r E

~; 20 E

a 3

/ O

/ ,_ a at Case 5

O o.1 0.2 03 0.4 05 0.6 07 Cont ol Rod Worth. hk/k t

PCCENT CORE ':PERIENCl"G DNB AS A FUNCTION CF EJECTED CONTROL ROD WO (H AT ULTIMATE POWER -

CRYSTAL RIVER UNITS 3 & 4 l - FIGURE 14 22

0 25 BOL Parameters EOL cr D * ~1*14 x to-5 (3/k)/F 2.0 - g = 6.0 x 10-5 (&/k)r T . o,3 see Delay g ta = 5.kT x 10 sec g EOL Parameters

• 15 -a 3

= -1 56 x 10-5 (3 73)jy a

o M " ^** "

9 T = 0 5 see Delay N t' = 2 75 x 10~ sec 1.0 0* $

f'/ -

~ Nominal Case I

BOL 0

0.1 0.2 0.3 0.4 0.5 0.6 0.7 ControlRodWorth,%ok/k q

ZR H2O REACTION AS A FUNCTION OF EJECTED V CONTROL ROD ki;RTH AT ULTIMATE POWER CRYSTAL RIVER l NITS 3 & 4 0094 , rA.-

= j us c"o."'. FIGURE 14- 23 o

I

_ E0L Parameters /

4 D

. -l . Jo x 10 ' (ak/k)/F /

- a . Assue.e Zero 0.3 see De'ay 2

• -5 g 2.75 x 10 see bitimate

/

BOL Parameters l 3 ECL 10 . -1.14 x 10 .$(sk/k)/F f ,

-a

_ g 6.0 x 10'5 ($k/h)/F # '

f j

- r 0.3 see Delay / /

4 g* 5.47 x 10 see m

' 2

/

/[/ j g Ultimate Nominal # '*

10 Case 3 ,

x /

w

' /

$ /

- 4

/

,/

Mominal 2 Case 10 Ultimate Power.

10 00'

/

/

6 7

4 ,

2 -

10 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 Control Rod korth. Tak/k REACTOR NEUTRON POWER VARIATION WITH EJECTED CONTROL RCD WORTH CRYST AL RIVER UNITS 3 & 4 m FIGURE 14- 24 0095 L

u40 J Ultimate Power.

EOL L20

/

A

/ Ultimate Power.

BOL 100 Nominal Case e w*

, BOL Parameters

{ ,0 . -l.14 x 0 5 ( \k/k /F 80 .,

. 6.0 x 10' (sk/k)/F e

, 0.3 see Delay 2' . S e

17 x 10" sec 60 EOL Parame_ters

• ., , -1.36 x 10 -5 ( sk/k)/F

- Assume Zero

. 0.3 see Delay 40

-5 -3

, 2.75 x 10 sec 10 Ultimate Power.

PCL 20 s

Nominal l Case 0.0 0.i 0.2 0.3 0.4 0.5 0.6 0.7 Cont;al God Wo'th. 'ak/k l

i e

,s .

I

,g REACTOR THERMAL POWER AS A FUNCTION

)

tv; 0F EJECTED CONTROL ROD WORTH I l

CRYSTAL RIVER UNITS 3 & 4 1 t

0096 -

FIGUR E 14 25

O 60 , , , , ,

BOL Parameters g = - 114 x 10-5 (a/k)/r g = 6.0 x 10-5 (g jg)jy ( g3 gi,,,, p,,,,,

50 - t = 0 5 see Delay t' = 5.41 x 10-5 ,,e O EOL Parameters /

a _ g = -1.% x 10-5 gaf,)fy ,

/ ,

} g - Assume Zero

- _T = 0.5 sec *

• • - 2 75 x 10gmysec

/

i'

~

1. s#

U1 imate Case s 4 20 Power, BOL -

s E

/

/

- r O

• Nominal Case

~

/

/

0p #

0.1 0.2 05 0.k O.4 05 0.6 Control Rod Worth, M/k l

l l

i ENTHALPY INCREASE TO THE IlOTTEST FUEL ROD VERSUS EJECTED CONTROL ROD 50RTH 0

CRYST Al RIVER UNITS 3 & 4

• ~**

-- FIGURE 14 26

0 N

BOL Parameters sq = 6.0 x 10~5 @A)/F C 60 N &/k = 0 54 t . 0 5 see Delay N f* = 5.k7 x 10*3 see

\ l I

40 N A Naminal Case  %

3 N  %

N 20

-0 5 -0 7 -0 9 -1.1 -1. 5 -1 5 -1 7 O Doppler Coef ficient. (t.k/k)/F x 10 r

P THE EFFECT ON REACTOR NEUTRON POWER OF l VARYING THE DOPPLER COEFFICIENT ROD EJECTION AT 10 ULTIMATE POWER l ...

0098 CRYSTAL RIVER UNITS 3 & 4

) s Em FIGURE 14-27 j

'O 44 i i i i i i BOL Parameters fi g = -1.14 x 10-5 (&/k)/F w 3 /k= 0 5%

C 40 -t

=03seenegay~

1. = 5 47 x 10 sec

{

o g Nominal Case o m 3

e 32 o

i 3

i 6

e 9

i 12 I

15 g

Moderator Coefficicnt, (Sk/k)/F x 10 THE EFFECT ON REACTOR NEUTRON POWER OF VARYING THE MODERATOR COEFFICIENT ROD EJECTION AT 10-' ULTIMATE POWER CRYSTAL RIVER UNITS 3 & 4 . 0099 6

3. FIGURE 14-28

O 2h BOL Parameters

\ q = 6.0 x 10-5 (t)c/k)/F

% t1/k = 0 5f,

,3 A T = 0.3 sec Delay u' - 5.47 x 10 sec j s to i

l 16 Nominal Case W - l k  %

.9 Nm 2

S 12

-05 -0 7 -09 -1.1 -1 3 -1 3 -1.T Doppler Coefficient, (.1k/k)/F x 10 5 O

THE EFFECT ON REACTOR THERMAL POWER OF VARYING THE DOPPLER COEFFICIENT ROD EJECTION AT 10 9 ULTIMATE POWER CRYSTAL RIVER UNITS 3 & 4 0100 _

lllllll.- FIGURE 14-29

O i i . 1 i  ! i i i BOL Par luneters w

_g= 1.14 x 10'5(ak/k)/r _

l .

C ak/k = 0 5% '

[

f _t = 0 3 see Delay z* = 5.A7 x 10 ~5 see 1 /

$ l7 f gNcuninal Case 11

-d E

15 i i e i , ,

0 3 6 9 12 15 18 Moderator Coefficient, (ak/k)/F x 10 O

l l

l l

THE EFFECT ON REACTOR THERMAL POWER OF VARYING THE MODERATOR COEFFICIENT ROD EJECTION AT 104 ULTIMATE POWER l CRYST AL RIVER UNITS 3 & 4 _

E FIGURE 14 30

112 i  : i i 80L Parameters ill .-

5 "D

110 -

u H

- 6.0 x 10" (Sk/k)/F /

Ak/k . 0.37c 109 -

-5 z' 5.u7 x 10 sec EOL Parameters 108 we-a . -1.36 x 10 -5(ok/k)/F [

J O

E 107

- a Assume Zero l

_ ak/k . 0.3"e /

y I6 - ~

2.75 x 10 sec I

105 >

Ultimate Power. 'r a: 104 -

O / >

103 102 '

Ultimate Power. -

101

\ EOL

' ' I

/[ '

- Nominal Case 100 0 0.l 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 Trip De.ay, sec REACTOR THERMAL POWER VERSUS TRIP DELAY TIME - ROD EJECTION AT ULTIMATE POWER

,,.f.. 02 CRYSTAL RIVER UNITS 3 & 4 l - FIGURE 14-31 1

BOL Parameters

- D

. -1.14 x (ak/k)/F __ /

a g - 6.0 x 10 (ak/k)/F 42 -9 ak/k . 0.5f. (10 Ultimate Power) ,

Sk/k . 0.37. (Ultimate Power) t ' . 5.47 x 10 sse

/

38 _

~ E0L Parameters

(

D . -1.36 x 10 -5(ak/k)/F e a g U1timate j -o Assume Zero Power, -

g ,

BOL and EOL 4 ak/k . 0.3f. (Ultimate Power) 34

",_ - a' . 2.75 x 10' see E

w 2 /

/

e A

$ 26 '

2

= / e j

f 22

/

• Nominal Case p 18

1. #I 10

-9 Ultimate

1. Power. BOL I4 0 0.2 0.4 0.6 0.8 1.0 TripDelay(r),sec ENTHALPY INCREASE TO THE HOTTEST FUEL ROD ,

VER$US TRIP DELAY TIME ROD EJECTION CRYSTAL RIVER UNITS 3 & 4 01 M llllla FIGURE 14 32

O

+ i e i e se i e i 6 . is . i i aiin 3 , , ,,,,, , , , , , . , ,

2400 2000  %- -

~

\ Expe riment al -

1600 \ ~ ~ ~ Predicted

i. N

- \ -

! \\

! 1200 h

_ \

800 W C hh.

~

\ _

400 \

\

N _

0 t i t i i*ii ,

i .,4 ,

0.001 0.01 0,1 1 10 100 Time, se-

]

l 1

LOFT SEMISCALE BLOWDOWN TEST NO. 546 VESSEL PRESSURE VERSUS TIME 0104 cavst^t aivea units 3 a 4 i -

0l:3.,,o., FIGURE 14 33

100 S \

n J

2

?

$ 6:

a. \

n 1 2 (

x a: c h

22 % Me u u r- d

! )

w C .: 2: 3: 40 c; 60 Tite. se:

' ~

0105 PREDICTED PERCENT MASS REMAINING VERSUS TIME LOFT TEST NO. 546 CRYST AL RIVER UNITS 3 & 4 3 FIGUR E 14- 34

O

/

/

/ o r -- o c =

N O

o l

o 8 8 8

• R R 2 3 5 2. , ..

l l

l i

NEUTRON POWER YERSUS TIME FOR A 36 IN. ID, I DOUBLE ENDED, HOT LEG PIPE RUPTURE AT ULTIMATE POWER WITHOUT TRIP v CRYSTAL RIVER UNITS 3 & 4 l

' f, , ,,; .

0106 @E_ eicuRE i4-34-.

AM END.1 (1 15-68)

I

6 5 ^

Density AK 3 ^K e

x Total w

D 2 g

.t l ,

> /

i g 0 / -

,,, N / Doppler AK _.

g ms

-2

-3 0 .4 .8 1. 2 1.6 2.0 Time, see REACTIVITY VERSUS TIME FOR A 36 IN.

ID, DOUBLE ENDED, HOT LEG PIPE RUPTURE AT ULTIMATE POWER WITHOUT TRIP CRYSTAL RIVER UNITS 3 & 4

. 01y G E FIGURE 14-34 b AMEND.1 11568)

'~~'

70 60 50 7 A

• (

x S 40 f f

, 1 <

c o 0 30 N O 20 3

T 10 A

C U 1 2 3 4 5 6 7 8 9 10 Time, sec

! 'N CORE FLOW VERSUS TIME FOR A 36 IN.

')~' j ID, DOUBLE ENDED PIPE RUPTURE 0108 CRYSTAL RIVER UNITS 3 & 4 ROG44 PJ" FIGURE 14 35

1500

u. -

I d A c 1400 1

" I DNB at 3"' 1300 0.25 see J J E 1200 \

Calculated by Quinn's

}

[ Modified Sieder-Tate Equation

[ 1100 0  % ,

j 1000

• y I Ns \

900

% \ '

I 800 U '

• 7 Slump Model a 700 - S imul ation y 7 3

o 600

\

2 \ G

500

\

o o b

= 400

\ s 300

\

200

\\1 3

L \

100  %

0 0 2

\

1 3 4 5 6 7 8 9 10 Time,cee HOT CHANNEL CLAD SURFACE HEAT TRANSFER COEFFICIENT AFTER DNB VERSUS TIME FOR A 36 IN. ID, DOUBLE ENDED PIPE RUPTURE CRYST AL RIVER UNITS 3 & 4 0109 b, FIGURE 14-36

A

() .

24 Core Top 20 -- - - -- -- -

l

} 14" Pipe 50", N -

' ' ' 2] L' v4" Pipe 335 N2 -g , , 12", . Pipe-33% N2

- r i i i

/ 12" Pipe-505 N2

~. l f i i2 / /

i

////

--~- -

Core Bottom

-U 8

{

O i

/

e l

=4

/

J O I O 5 10 15 20 25 30 35 40 45 50 55 60 Time, sec REACTOR VESSEL WATER VOLUME VERSUS TIME FOR l 36 IN. lD, DOUBLE-ENDED PIPE RUPTURE FOR 600 PSIG CORE FLOODING TANK OPERATING PRESSURE

~

CRYSTAL RIVER UNITS 3 & 4

.h '

)

1 ~' -

- """'d'7 0

AMEND. 2 (2 7-68)

O 24 Core Top 20 12" Pipe-505 N.2

?= -

@ 1000 psig yF -

14" Pipe-33",N2

@ 400 psig 16 / / r 12" Pipe-33r E,

_ N2 @l000 psig -

f 14" Pipe-50% N2

[

3 12 / @ 400 psig E

O E f Core Bottom

/

E o

7 [

/

Z cc -

4 1 g 0 '~

0 5 10 15 20 25 30 35 40 45 50 55 60 Time, see i

i REACTOR VESSEL WATER VOLUME VERSUS TIME FOR 36 IN. ID, DOUBLE-ENDED PIPE RUPTURE FOR 400 PSIG AND 1,000 PSIG CORE FLOODING TANK OPERATING PRESSURES CRYSTAL RIVER UNITS 3 & 4 j\ gl Em FIGURE 14-38 AMEND. 2 (2-7-68) l l

l

2500 2400

/

Hot Spot 2300

/ II ~

t 2200 600 lb-12"-50% N / 7 i

2100

\ / / $_

600 lb-14-33% N

[ /

j / #

n

$ 1800 g 04 '

1700 01- - 3 N 400 lb-14"-50% N

/ [ [ 1000 lb-12"-33% N

< - 1000 lb-12"-50% N 10 12 14 16 18 20 22 24 26 28 30 32 34 36 Ouench Time, sec CORE FLOODING TANK ANALYSIS; MAXIMUM CLAD

' *. TEMPERATURE VERSUS TIME TO QUENCH FOR A

) , 36 IN. ID, DOUBLE-ENDED PIPE RUPTURE CRYSTAL RIVER UNITS 3 & 4

@s_ riouRe 24 39

. ~ . . . . - . - . . - - . . - - . -

0 2800 2600 m 1

. 2400 0

B 2 2200 k e

1 2000

• I \

\

U '

u 1800 Nominal '

E.

Design N N Point o 1600 '

=

0 j 1400 1200 0 200 400 600 800 1000 1200 1400 1600 1800 2000 2200 Maximum Heat, Btu /hr-ft 2_y MAXIMUM HOT SPOT CLAD TEMPERATURE YERSUS MAXIMUM HEAT TRANSFER COEFFICIENT AFTEP DNB ',

i FOR A 36 IN. ID, DOUBLE ENDED PIPE RUPT,RE g) CRYSTAL RIVER UNITS 3 & 4

                      =~       FIGURE 14-40

O i 2000 - Nominal Design Point

                              '      1900    -

i b 1800 -

                              %u 1    1700     -

5 s

                              ,     1600     -
                             $      1500     -

a

                             ,      1400    -

E e, 1300 -

j 1200 - iroo , , , i 0 1 2 3 4 5 Nucleate Boiling Period, sec MAXIMUM HOT SPOT CLAD TEMPERATURE AS A FUNCTION OF TIME TO REACH DNB FOR A 36 IN. ID, DOUBLE ."NDED, HOT LEG PIPE RUPTURE

                          ..                                 CRYSTAL RIVER UNITS 3 & 4 3    FIGURE 14-40 e AMEND. I t1 15-68)

2800 h=0 2600 # 2400

                                               /
                                                    /

2200

      ,,,,                          [               7   h . is stoih,_,,2_,

w 1800 g [ h - 20 8 E 1600 / \ 3 / \ l 1400

                    ,                 x
  !"::        H                           x                      g h = 100 800                                                 '

t 600 N 400 200

                '           '          '             '         '            I 0

O 10 20 30 40 30 60 Time, see HOT SPOT CLAD TEMPERATURE VERSUS TIME FOR 36 IN. ID, DOUBLE ENDED PlPE RUPTURE AND VARIABLE QUENCH COEFFICIENT CRYSTAL RIVER UNITS 3 & 4 O\ h E FIGURE 14-41

                                                                                    ~
                                                                                   ,s 4

O 2500 2400 2300 w e 5 k 2200 5 r [ o 3 2100 2000 Nominal Value l

                           !900 l

1 0 1 2 3 4 Full-Power Seconds HOT SPOT CLAD TEMPERATURE AS A FUNCTION OF FULL POWER SECONDS RESULTING FROM VOID SHUTDOWN FOR A 36 IN. ID, DOUBLE ENDED, HOT LEG PIPc RUPTURE 0116 CRYSTAL RIVER UNITS 3 & 4 b FIGURE 14-41-a AMEND.1 (1-15 68)

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HOT SPOT CLAD TEMPERATURE VERSUS TIME FOR 36 lH. ID, DOUBLE ENDED PIPE RUPTURE AND VARIABLE SINK TEMPERATURE CRYSTAL RIVER UNITS 3 & 4 g s'hs_PioVRei42 od

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(+* t i - , - - - - n i e s a r a u t I p u R O i e e n a i t L f i e 4 i g r 0 u S 02

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m - I m

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z O HPl - HIGH PRESSURE INJECTION Q LPI - LOW PRESSURE INJECTION

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              -4 l HPl PUMP + 2 CORE FLOODING TANKS . I LPI PUMP 2 CORE FLOODING TANKS + 1 LPI PUMP P res su ri zer Surge a                                                                                      3 in.               6 in. Line                                           36 in
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p REACTOR BUILDING PRESSURE VERSUS TIME Q "'-

                                   '-          0iN                                         " ' ' ' ~ " " ' """'"""

CRYSTAL RIVER UNITS 3 & 4 no

                                                                                        ~.".,         FIGURE 14-45

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       < O Cm 14 ft Rupture 3,000 gpm R.B. Sorays

• E .m 4vu 51.9 6,500 gpm injection n " O cCmMC 49.6 O C r 50 @ without Spray Cooler m

z- ym C m ,g< @ With Spray Cooler m . f m -- M - I. w zd u P m>ym M i a e u az a 40

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          "                                                                                                                                                     f N                                                              Time after Rupture, sec.                                                               !

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          .-   '                                                       COOLANT TEMPERATURES FOLLOWING A 36 IN.
           ,        ,                                                                l.D. DOU BL E-END ED_ RUPTU R E CRYSTAL RIVER UNITS 3 & 4
                            -..            0128                                              ne
                                                                                             =         FIGURE 14-47

T eoj 4il 45 64I II3II I iiI64 I i i I l ) I i34 ii{ 614Ii4 Ii i 3 4I I&4i i

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AFTER RUPTURE 8.5 FT2 RUPTURE , I CRYSTAL RIVER UNITS 3 & 4 . . ' 1 Fo. ". FIGURE 14 40

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                                             ,                                       REACTOR BUILDING PRESSURE VERSUS TIME

.q'o AFTER RUPTURE - 3 FT RUPTURE 2 Qj}} CRYSTAL RIVER UNITS 3 & 4

                                                                                                         -                 FIGURE 14-49

1 1 1 I I lasisiisi islas1861 1314: 56si isatissia 136133166 ililit ill ', , l 3 .,, C. e. 5 C

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                 't                     $                 7            M             r3 C'

Sisd 'aanssaJd PUTT *Tng 2 s: sad REACTOR BUILDING PRESSURE VERSUS TIME AFTER RUPTURE - 2 FT2 RUPTURE CRYST AL RIVER UNITS 3 & 4 0131 O a Th FIGURE 14- 50

I O jairisais sas646.is 44asias4 4 asi464 44 66i6466 i # Q) 4 s i t i 61 i i t o _ ~ p _ d m _ o i _ N 2 .v _ m E 3 - 3 _ g a 2

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Sisd 'aanssa2d Sulpting 2ossesig 1 i REACTOR BUILDING PRESSURE VERSUS TIME l k[L ] AFTER RUPTURE - 1 FT2 RUPTURE l CRYSTAL RIVER UNITS 3 & 4 013? . no i c"o'". ". o FIGURE 14- 51

J6isissai isisasias asssisisi insaisisi sisisis sinisiiic 4 o m 3 M n O I

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                 -            =                                                                         -

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g .. viserinii siirrivin viiiiiiii visitt ii, initivist .iittisil 'o S S S  % 2  % Si sd 'aznsss2d Suipting Joasesg 003 REACTOR BUILDING PRESSURE VERSUS TIME - l AFTER RUPTURE - 0.4 FT2 RUPTURE ' CRYSTAL RIVER UNITS 3 & 4

              =            FIGUR E 14-52

O 3 10

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                  ~

Steam Air W ure] 9 j,# Liquid ,l i ^

                                 /                            .

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Structures Removal Cooler - 3 10' . g 3 Emergency [ _~ Cooling Units - 0 , , ,,,,,, 10 , , , , , , , , , , , , , , ,,, 1 , , , , , , , , , O 10 10' 10 10 10 Time af ter Ruoture, sec S 1 l REACTOR BUILDING ENERGY INVENTORY

  .s'
                        -                                                                   FOR 36 IN. ID, DOUBLE ENDED RUPTURE
                     'u                     - -       0134                                       CRYSTAL RIVER UNITS 3 & 4
                                                                                                           -                FIGURE 14-53
~

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                     -    3 Emergency Cooling Units                                                                              '
                     ~                                                       Total
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                                                          /                       Liquid g                                                                                                                                 -
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3 j -

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Heat - l [j Structures Removal 10 000I0' _ w 3 Emergency [ Cooling Units . o , , ,,,,, , , , , , , , , , ,, , , , , in ~ , ,, 10 10 10 to 10 10 Time af ter Rwature, see REACTOR BUILDING ENERGY INVENTORY FOR 3 FT2 RUPTURE CRYSTAL RIVER UNITS 3 & 4 g \ be lllll%~ FIGURE 14 54

p 300 - . . . . . . , i i , , . . . . i ,iii., , , . , , , , , , , , , , , , b l 5

               ~

250 Reactor Building Vapor Temperature

I m Reactor Building Sump Temperature N i i  ;

                             /                                                               ,                                                  :
   }      150          g                                                                     i f          :

g 3 Emergency Cooling Units -

                                                                                                                                                ~

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   . ..._ 60 l                                                                                                                                 ~

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               -1                       0                            1                         2                           3                        4 10                       10                         10                         to                           10                     10 Time after Rupture, see P

REACTOR BUILDING VAPOR AND SUMP TEMPERATURES FOR 36 IN. ID, DOUBLE ENDED RUPTURE f AS A FUNCTION OF TIME AFTER THE RUPTURE CRYSTAL RIVER UNITS 3 & 4

                                            ..         0136                                      .-

c' - .m FIGURE 14 55

300 . . ..i.. . . ... i , , . . . . . . i i ..is ii ,ii>1 3 Emergency  : Cooling Units b 250 . ,

            }                                                             Reactor Building Vapor Temperature                                        j 200                                           ,
                                                                                                                            \

6  : g } Reactor Building Sump Temperature g . - }

  %    150                                                                                                                                         '

t

                                     /                                                                                                             ~

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            ;                                                                                                                                      ~

50 _ l - 5 0" ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' 10 10 10 10 to 10 Time Af ter Rupture, see REACTOR BUILDING VAPOR AND SUMP TEMPERATURES AS A FUNCTION OF TIME AFTER !!UI'TURE 3 FT2 RUPTURE CRYSTAL RIVER UNITS 3 & 4 O137

               %-               FIGURE 14 56

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                                                                                                                  =            FIGURE 14 57

  ~                                                  .

100 90 3 Emergenrj Cooling Units + / 2 Sprays 80 0 /

                                            /                                                 ./

3 Emergency Coo.ir.g Units == 60 / 7 i 3 50 /

                                                                         /

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$ 30 / 2

                                            /

f b l / 4 20 a f / r 10 / St ructures 0 A , 0 200 LOO 600 800 1000 1200 1k00 1600 A800 2000 2200 2L00 2600 2600 3000 Time to Complete Feaction, see REACTOR BUILDING ZR RE ACTION CAPABILITY FOR 55 PSIG DESIGN PRESSURE CRYSTAL RIVER UNITS 3 & 4

                =,.        FIGURE 14-5"                                                -

0139

10  % 8

                                     .\ \

24 Hour Dose 4 % I 30-Day Dose o

                 ? Hour Dose
  =                                                              Limit of 5-Hile Zone i

e i0 i NN . i 8 Exclusion \\ T Distance A \ O

  =                                                                          N 2

3 I N 10 \ 3 2 4 6 8 10 10 10

                                      . Downwind Distance, feet e
                                                     '"'"'"     5' '" " ' 55 'c       '^"'

f-

                                 -     Ol40        COOL ANT ACCIDENT 2 HOURS,24 HOURS,

( AND 30 DAYS DOSES CRYSTAL RIVER UN;TS 3 & 4

                                                               $_        FIGURE 14 59

0 3 10 8 \ _ l

                                                                                                              ~

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                       \\   \                                                                           l-1 30-Day rose                                         l-4                           <                                                           i N
                                           \ K          .                     10 a r* 100 '111:

24-Hou Dese 2-Hour Ecse 2 10 I: 8 1 Nx \\ -

       $      6
                                           \                         g Limit           of c                                                                        5-Mile Zone m      4                                    \                                 l Exclusion Distance 10                      l           l      l       l                        l       l    h     1
                                         '       ^       E         4          2              4   f,   9 10                                             10                                            10 Dowtv.::n d L i s t an ce , feet l

MAXIMUM HYPOTHETICAL ACCIDENT THYROID DOSE ASSUMING FIS$10N h\ PRODUCT RELEASE PER TfD 14844 CRYSTAL RIVER UNITS 3 & 4 I.5 FIGUR E 14-60

()7 10 g,,,

                            ,,,,    ,,,,   ,,,,  ,,ii    sii,   ,,,,   ,,,,     ,,,,       ....     , , , , . . , _

I 1 10

                                                                                                                   =_

_ i _ 0 k

                          '                                                 l 10 E      \                                                                                          E E

g _- _~ E -l \- 8 10 _ , g

                                      \               30-Day Dose E

g _ 8-Hour Dose i 2-Hour Dose

            -2
         'O t

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\ .

I Exclusion  ! E E  : l Distance  : io

            -3                                         k                                         ,
                ;                                          x       x                             i g

i _ 10 E \ \ g _I I  : s i0 0 1000 2000 3000 14000 5000 6000 Distance from Edge of Reactor Building, ft

                                                                                               ..      0142 g) s u

INTEGR ATED DIRECT DOSE FOLLOWING MH A WITH 3% FOOT REACTOR BUILDING WALL THICKNESS CRYSTAL RIVER UNITS 3 & 4 3, FIGURE 14-61}}