ML19329E147

Chapter 14 of AR Nuclear 1 PSAR, Safety Analysis. Includes Revisions 1-18

ML19329E147
Person / Time
Site:	Arkansas Nuclear
Issue date:	11/24/1967
From:	ARKANSAS POWER & LIGHT CO.
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NUDOCS 8005300732
Download: ML19329E147 (150)

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TABLE OF CONTENTS Section Pace lh SAFETY ANALYSIS lk-1 lb.1 CORE AND COOLA:!T EOUNDARY PROTECTION ANALYSIS lb-1 1h.1.1 ABNORMALITIES lk-1 1h.1.2 ANALYSIS OF EFFECTS AND CONSEQUENCES lk-3 14.1.2.1 Unconcensated 0:erating Reactivity Chances lk-3 14.1.2.2 Startuo Accident lb-h 1h.l.2.3 Rod Withdrawal Accident From Rated Power Operation lk-6 1h.1.2.h Moderator Dilution Accident lb-8 14.1.2.5 Cold Water Accident lk-10 14.1.2.6 Loss-of-Coolant Flov lh-10 14.1.2.7 Stuck-Cut, Stuck-In, or Drorted-In Control Rod 1h-12 14.1.2.8 Loss of Electric Power lh-13 14.1.2.9 Steam Line Failure lk-15 14.1.2.10 Steam Generator Tube Failures 1k-18 1 14.2 STANDBY SAFEGUARDS ANALYSIS 1h-20 14.2.1 SITUATIONS ANALYZED A:iD CAUSES lk-20 lk.2.2 ACCIDENT ANALYSES lk-20 14.2.2.1 Fuel Handling Accidents 14-20 l lb.2.2.2 Rod E.fection Accident lk-21 14.2.2.3 Loss-of-Coolant Accident lk-27 lk.2.2.h Maximum Hypothetical Accident 1h-Sh 1h.3 REFERD;CES lb-59 l 3 0297 E lk-i

LIST OF TABLES Table No. , Title Pace 1b-1 Abnormalities Affecting Core and Coolant Boundary lk-1 1h-2 Uncompensated Reactivity Disturbances lh-3 lh-3 Situations Analyzed and Causes lh-20 lk-b Reactor Building Structural Heat Capacitance Segments 1b-3h lk-5 Core Flooding Tank Performance Data 1h-37 14-6 Tabulation of Loss-of-Coolant Accident Characteristics for Spectrum of Hot Leg Rupture Sizes lh-h2 lb-7 Tabulation of Less-of-Coolant Accident Characteristics for Spectrum of Cold Leg Rupture Sizes 1k-h3 lh-8 Reactor Operating Conditiens for Evaluation lh L6 1h-9 Reactor Building Structure Data for Analysis of Time-Dependent Reactor Building Pressure lh-h7 1h-10 Summary of Reactor Building Pressure Analysis for Reactor Building Emergency Cooling (2h0 x 106 Etu/hr) 14-51 14-11 Sensitivity Analysis Showing the Effect of Parameters 3 on the Two-Hour Iodine Dose at the Exclusion Distance Following IEA 14-58 0288 N 1k-li

LIST OF FIGURES (At rear of Section) Figure No. Title 14-1 Startup Accident from 10-9 Rated Power Using a 1.2% Ak/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 10.0% Ak/k; High Flux Reactor Trip Is Actuated 1h-3 Peak Thermal Power versus Rod Withdrawal Rate for a Startup Acci-dent from 10-9 Rated Power lk h Peak Neutron Power versus Rod Withdrawal Rate for a Startup Acci-dent from 10-9 Rated Power 14-5 Peak Thermal Power versus Trip Delay Time for a Startup Accident Using a 1.2% Ak/k Rod Group at 5.8 x 10-5 (Ak/k)/see from 10-9 Rated Power lb-6 Peak Thermal Power versus Doppler Coefficient for a Startup Acci-dent Using a 1.2% Ak/k Rod Group at 5.8 x 10-5 (Ak/k)/see from 10-9 Rated Power 14-7 Peak Thermal Power versus Trip Delay Time for a Startup Accident Using All Rods at 5.8 x 10-h ( Ak/k)/see from 10-9 Rated Power 1h-8 Peak Thermal Power versus Doppler Coefficient for a Startup Acci-dent Using All Rods at 5.8 x 10-h (Ak/k)/see from 10-9 Rated Power 1h-9 Peak Pressure versus AllRodsat5.8x10-{ripDelayTimeforaStartupAccidentUsing (Ak/k)/see from 10-9 Rated Power lb-10 Peak Pressure versus Tripped Rod Worth for a Startup Accident Using All Rods at 5.8 x 10-h (Ak/k)/sec from 10-9 Rated Power lb-ll Peak Pressure versus Doppler Coefficient for a Startup Accident Using All Rods at 5.8 x 10-4 (Ak/k)/see from 10-9 Rated Power 14-12 Peak Pressure versus Moderator Coefficient for a Startuu Accident Using All Rods at 5.8 x 10-h (Ak/k)/sec from 10-9 Rated' Power 1h-13 Rod Withdrawal Accident from Rated Power Using a 1.2% Ak/k Rod Group at 5.8 x 10-5 (Ak/k)/sec; High Pressure Reactor Trip Is Actuated 14-1h Peak Pressure versus Rod Withdrawal Rate for a Rod Withdrawal Accident from Rated Power lh-15 Peak Pressure versus Trip Delay Time for a Rod Withdrawal Acci-dent from Rated Power Using a 1.2% Ak/k Rod Group; High Pres-sure Reactor Trip is Actuated 0299 1h-111

FIGURES (Cont'd) Figure No. Title lk-16 Peak Pressure versus Doppler Coefficient for a Rod Withdrawal Accident from Rated Power Using a 1.2% ok/k Rod Group 1h-17 Maximum Neutron and Thermal Power for an All-Rod Withdrawal Accident from Various Initial Power Levels ' lh-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 Functica of Time after Loss of Pump Power lh-20 Mini =um DNER Which Occurs during the Coastdown for Various Initial Power Levels 1k-21 Reactor System Cooling Rate for a Steam Line Break of h in.2 lb-22 Per Cent Core Experiencing DNB as a Function of Ejected Control Rod Worth at Ultimate Power 1h-23 Zr-H2 O Reaction as a Function of Ejected Control Rod Worth at Ultimate Power lk-2h Reactor Neutron Power Variation with Ejected Control Rod Worth lk-25 Reactor Thermal Power as a Function of Ejected Control Rod Worth 1k-26 Enthalpy Increase to Hottest Fuel Rod versus Ejected Control Rod Worth 1h-27 Effect on Reactor Neutron Power of Varying the Doppler Coefficient - Rod Ejection at 10-9 Ultimate Power 14-28 Effect on Reactor Neutron Pcver of Varying the Moderator Coefficient - Rod Ejection at 10-9 Ultimate Power 1h-29 Effect on Reactor Ther=al Power of Varying the Doppler Coefficient - Rod Ejection at 10-9 Ultimate Power 1h-30 Effect on Reactor Thermal Power of Varying the Moderator Coefficient - Rod Ejection at 10-9 Ultimate Power 14-31 Reactor Thermal Power versus Trip Delay - Rod Ejection at Ultimate Power 14-32 Enthalpy Increase to the Hottest Fuel itad versus Trip Delay Time - Rod Ejection 14-iv 0300 W 1

FIGURES (Cont'd) Figure No. Title 14-33 LOFT Semiscale Blowdown Test No. 5h6 - Vessel Pressure versus Time lh-3h Predicted Per Cent Mass Remaining versus Time - LOFT Test No. Sh6 1h-35 Neutron Power versus Time for a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture at Ultimate Power Without Trip 14-36 Reactivity versus Time for a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture at Ultimate Power Without Trip 1h-37 Integrated Power versus Break Size for a Spectrum of Rupture Sizes 1h-38 Core Flow versus Time for a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture 1h-39 Hot Channel Clad Surface Heat Transfer Coefficient after DN3 versus Time for a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture lk h0 Reactor Vessel Water Volume versus Time for a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture for 600 psig Core Flooding Tank Opera-ting Pressure 14-kl Reactor Vessel Water Volume versus Time for a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture for h00 psig and 1,000 psig Core Flood-ing Tank Operating Pressures 1k h2 Core Flooding Tank Analysis; Maximum Clad Temperature versus Time to Quench for a 36-in. ID, Double-Ended Hot Leg Pipe Rupture lk h3 Maximum Hot Spot Clad Temperature versus Maximum Heat Transfer Coef-ficient after DNB for a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture 14-kh Maximum Hot Spot Clad Temperature as a Function of Time to Reach DNB for a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture 14-h5 Hot Spot Clad Temperature versus Time for a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture and Variable Quench Coefficient lh-h6 Hot Spot Clad Temperature as a Function of Full-Power Seconds Re-sulting from Void Shutdown for a 36-in. ID, Double-Ended Hot Leg Pipe Rupture lh-h7 Hot Spot Clad Temperature versus Time for a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture and Variable Sink Temperature 1k-h8 Mass Released to Reactor Building for the Spectrum of Hot Leg Rup-tures 1k h9 Reactor Coolant Average Pressure for the Spectrum of Hot Leg Rup-tures 1k-v 0301 W '

FIGURES (Cont'd) Figure No. Title 14 Hot Leg Ruptures - Reactor Vessel Water Volume versus Time including Effects of Boiloff and Injection 14-51 Hot Spot Cladding Temperature versus Time for Spectrum of Hot Leg Ruptures 14-52 Reactor Coolant Average Pressure - Spectrum of Cold Leg Rupture Sizes 1k-53 Cold Leg Ruptures - Reactor Vessel Water Volume versus Time including Effects of Boiloff and Injection 1h-Sh -Hot Spot Cladding Temperature versus Time for Spectrum of Cold Leg Ruptures lh-55 Emergency Core Cooling Systems Capability 1h-56 Reactor Building Pressure versus Time in. ID, Double-Ended, Hot Leg Pipe Rupture lh-57 Reactor Building Pressure versus Time for a 36-in. ID, Double-Ended', Hot Leg Pipe Rupture with and without Cooling of the Recir-culated Spray Water 1h-58 Reactor Building Vapor and Sump Coolant Temperatures $ollowing a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture 1b-59 Reactor Building Pressure versus Time after Rupture (8.5 ft2) 1h-60 Reactor Building Pressure versus Time after Rupture (3.0 'ft2) 14-61 Reactor Building Pressure versus Time after Rupture (2.0 ft2) 14-62 Reactor Building Pressure versus Time after Rupture (1.0 ft2) 1h-63 Reactor Building Pressure versus Time after Rupture (0.h ft2) 1h~- 6h Reactor Building Energy Inventory for a 36-in. ID, Double-Ended Hot Leg Pipe Rupture 1h-65 Reactor Building Energy Inventory for 3.0 ft2 Rupture 14-66 Reactor Building Vapor and Su=p Te=peratures versus Time after a 36-in. ID, Double-Ended,' Hot Leg Pipe Rupture 14-67 Reactor Building Vapor and Su=p Temperatures versus Time after a 3.0 ft2 Rupture lh-68 Reactor. Building Pressure versus Time after a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture with Core Flooding Tank (s) j 14-vi ' _ - ~

FIGURES (Cont'd)  ! Figure No. Title 14-69 Reactor Building Pressure versus Time after a 3.0 ft2 Rupture with Two Core Flooding Tanks 14-70 Criteria 17 for a 36-in. ID, Double-Ended, Hot Leg Pipe Rupture lb-Tl Reactor Building Zirconium Reaction Capability for 59 psis Design Pressure lb-72 Thyroid Dese from Loss-of-Coolant Accident Hour, 2h-Hour, and 30-Day Doses lk-73 Maximum Hypothetical Accident Thyroid Dose Assuming Fission Product Release per TID-lh8hh lk-Th Integrated Direct Dose Following E3.A vith 3-3/h-Foot Reactor Build-ing Wall Thickness 0303 g lk-vii

i lh SAFETY ANALYSIS 14.1 CORE AND COOLANT BOUNDARY PROTECTION ANALYSIS 1h.1.1 ABNORMALITIES In previous sections of this report both normal and abnormal operations of the various systems and components have been discussed. This section summarizes and further explores abnormalities that are either inherently terminated or require the normal protection systems to operate to maintain integrity of the fuel and/or the reactor coolant system. These abnormalities have been evalu-ated for rated power of 2,452 MWt. Whenever a fission product release to the environment occurs, the release is based upon the fission pro. duct inventory associated with the ultimate reactor core power level of 2,568 MWt. Fission product dispersion in the atmosphere is assumed to occur as predicted by the dispersion models developed in 2.3. Table 1h-1 su==arizes the potential ab-normalities studied. Table 14-1 Abnormalities Affecting Core and Coolant Boundary Event Cause Effect Uncompensated Operat- Fuel depletion Reduction in reactor system ing 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 terminated by nega-rod (*) with- tive Doppler effect, reactor drawal. trip from short period, high re-actor coolant system pressure, or overpower. No equipment dam-age or radiological hazard. Rod Withdrawal Acci- Uncontrolled Power rise terminated by over-dent at Rated Power rod withdrawal. power trip or high pressure trip. No equipment da= age or radiological hazard. (*) Control rod, rod, and control rod assembly (CRA) are used interchange-ably in this section and elsewhere in the report. A control rod group consists of a sy= metrical arrangement of four or more control rod assemblies. See 7.2.2.1.2. 0304 N lb-1

Event Cause Effect Moderator Dilution Equipment malfunction Slow change of power terminated Accident or operator error. by reactor trip on high tempera-

             ,                                 ture or pressure. During shut-down a decrease in shutdown mar-gin occurs, but criticality does not occur. No radiological haz-zard.

Lors of Coolant Flow Mechanical or None. Core protected by reactor electrical low-flow trip. No radiological failure of re- hazard. actor coolant pump (s). Stuck-Out, Stuck-In, Mechanical or None. Suberiticality can be or Dropped-In electrical achieved if one rod is struck-out. Control Rod failure. If stuck-in or dropped-in, con-tinued operation is permitted if effect on power peaking not se-vere. No radiological hazard. Loss of Electric Miscellaneous Possible power reduction or re-Power faults. actor trip depending en con-i dition. Redundancy provided for safe shutdown. w Steam Line Failure Pipe Failure.

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Reactor automatically trips if rupture is large. No damage to ! reactor system. Integrated l doses at exclusion distance are 4.3 x 10-3 rem whole body and 0.60 rem thyroid. Stram Generator Tube Tube failurc Reactor automatically trips if Feilures leakage exceeds normal makeup capacity to reactor coolant sys-tem. Integrated doses at exclu-sion distance are 0.83 rem whole body and 1.1 x 10-2 rem thyroid.

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6-5-68 [ 14-2 Supplement No. 4

14.1.2 ANALYSIS OF EFFECTS AND CONSEQUENCES 14.1.2.1 Uncerrensated Oreratine Reactivity Chanzes 14.1.2.1.1 Identification of Cause During normal operetion of the reactor, the overall reactivity of the core changes because of fuel depletion and changes in fission product poison concen-tration. These reactivity changes, if left either uncompensated or overcom-pensated, can cause operating limits to be exceeded. In all cases, however, the reactor protection system prevents safety limits from being exceeded. No damage occurs from these conditions. 1k.1.2.1.2 Analysis and Results During normal operaticn, the automatic reactor control system senses any reac-tivity change in the reactor. Depending on the direction of the reactivity change, the reactor power increases or decreases. Correspondingly, the reac-tor coolant system average temperature increases or decreases, and the auto-matic reactor control system acts to restore reactor power to the power demand level and to reestablish this temperature at its set point. If manual correc-tive action is not taken or if the automatic control system malfunctions, the reactor coolant system average temperature changes to compensate for the reac-tivity disturbance. Table lk-2 summarizes these disturbances. Table 1k-2 Uncomrensated Reactivity Disturbances Maximum Rate of Average Reactivity Rate, Temperature Change Cause (Ak/k)/sec (Uncorrected), F/see Fuel Depletion -6 x 10-9 -0.0006 Xenon -3 x 10-8 -0.003 These results are based on +6 x 10-5 (Ak/k)/F moderator coefficient and -1.lh x 10-5 (Ak/k)/F Doppler coefficient. The nominal value of +6 x 10-5 (Ak/k)/F is representative of the moderator coefficient at the beginning of core life for an equilibrium cycle. This value is also valid at BOL for the first cycle after 15 days. A higher value (+10 x 10-5 (Ak/k)/F) exists at the start of the first core cycle. However, the effect of this slightly higher valae has been shown to be of minor importance 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. 0306 14-3

1k.1.2.2 Startup Accident 1h.1.2.2.1 Identification of Cause The objective of a normal startup is to bring a suberitical reactor to the crit-ical or slightly supercritical condition, and then to increase power in a con-trolled manner until the desired power level and system operating temperatures are obtained. During a startup, an uncontrolled reactivity addition could cause a nuclear excursion. This excursion is terminated by the strong negative Doppler effect if no other protective action operates. The following design provisions minimize the possibility of inadvertent contin-uous rod withdrawal and limit 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 25 per cent over-lap in travel between two successive rod groups. This overlap oc-curs at the minimum worth for each group since one group is at the end of travel and the other is at the beginning of travel. The max-imum worth of any single control rod group is 1.2% Ak/k when the reactor is critical as specified in 7.2.2.1.3

b. Control rod withdrawal rate is limited to 25 in./ min.

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

            . range.

d. .1 short-period withdrawal stop, alarm, and trip are provided in the

intermediate range.

l e. A high flux level and a high pressure trip are provided in the power range. The reactor protection system is designed to limit (a) the reactor thermal power.to 114 per cent of rated pow'er to prevent fuel damage, and (b) the reac-tor coolant system pressure to 2,515 psia. 1h.1.2.2.2 Methods of Analysis An analog model of the reactor core and coolant system was used to determine the characteristics of this accident. This analog model used' full reactor coolant flow, but no heat transfer out of the system and no sprays in the pres-surizer. The rated-power Doppler coefficient (-1.lk 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 moving along the steepest part of the rod-worth vs rod-travel curve. A reactor trip cn short period was not incor-porated-in the analysis. The nominal values of the principal parameters used were: 0.3 see trip-delay, +6 x 10-5 (Ak/k)/F moderator coefficient, and -1.14 x 10-5 (Ak/k)/F Doppler coefficient. The total worth of all the control rods inserted into the reactor core following any trip is 8.h5 Ak/k without a stuck control rod, or 5.4% Ak/k (the nominal case in this study) with a stuck rod. a 14-h g

14.1.2.2.3 Results of Analysia Figure 1h-1 shows the results of withdrawing the maximum worth control rod group at a rod speed of 25 in./ min from 1 per cent suberitical. This group is worth a maximum value of 1.2% ak/k. This rod velocity and worth result in a maximum reactivity addition rate of 5.8 x 10-5 (ak/k)/sec. The Doppler effect begins to slow the neutron power (*) rise, but the heat to the coolant increases the pressure past the trip point, and the transient is terminated by the high pressure trip. Figure lk-2 shows the results of withdrawing all 69 control rod assemblies (with a total worth of 10.0% ak/k) at the maximum spead from 1 per cent sub-critical. This results in a maximum reactivity addition rate of 5.8 x 10-h (ak/k)/sec. About 15.3 see after passing through criticality, the neutron power peaks at 147 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 ter=inates 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 de-termine the effect of varying several key parameters. Figures lb-3 through 1h-6 show typical results for the single group, 1.2% Ak/k startup accident. Figures 1h-3 and lb-h show the effect of varying the reactivity addition rate on the peak thermal power and peak neutron power. This reactivity rate was varied from one order of magnitude below the nominal single rod group case (1.2% Ak/k) to more than an order of magnitude above the rate that represents all rods (10.0% Ak/k) being " vn at once. The slower rates - up to about 0 5 x 10-3 (ak/k)/sec - vi' in the pressure trip being actuated, whereas only the very fast a actuate the high neutron flux level trip. Figures 1k-5 and 14-6 show the peak thermal power variation as a function of a wide range of trip delay times and Doppler coefficients for the 1.2% Ak/k rod group. Only a small change in power is noted. Figures 14-7 and lh-8 are the corresponding results from the withdrawal of all rods (10.0% ak/k). Since this transient incerts reactivity an order of magnitude faster than does the single control rod group case, there is censiderably mere variation in the peak thermal power over these vide ranges. At high values of the Doppler coeffi-cient, the neutron power rise is virtually stopped before reaching the high flux trip level. Reactor power generation continues until sufficient energy is transferred to the reactor coolant to initiate a high pressure trip. This re-sults in a higher peak thermal power. Figures lb-9 through lb-12 show the peak pressure response to variations in several key parameters for the case where all rods are wi..arawn. It is seen that the safety valve is opened when these parameters are changed considerably from the nominal values, except in the case of the moderator coefficient which has little effect because of the short duration of the transient. Again for a high Doppler coefficient, the high pressure trip is relied upon. (*) Neutron power is defined as the total sensible energy release from fission. g 1h_5 0308

None of these postulated startup accidents, except for reactivity addition rates greater than 2 x 10-3 (ak/k)/see, which is three times greater than for withdrawal of all rods at once, causes a thermal power peak in excess of h0 per cent rated power or a nominal fuel rod average temperature greater than 1,715 F. The nominal 1.27, ak/k rod group withdrawal 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 frem this startup accident. The 10.0f ak/k withdrawal - using all 69 rods - causes a peak pressure of only 2,h65 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 thermal power approach 114 per cent, and the peak pressure never ex-ceeds 2,515 psia. 1h.1.2.3 Rod Withdrawal Accident From Rated Power Oreration 1h.1.2.3.1 Ider 'fication 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 reac-tor is at rated power. As a result of this accu =ed accident, the power level increases; the reacter coolant and fuel rod temperatures increase; and if the withdrawal is not terminated by the operator or protection system, core dam-age would eventually occur. i The following provisions are made in the design to indicate and terminate this accident:

a. High reactor outlet coolant temperature alarms.

b. High reactor coolant system pressure alar =s.

c. High pressuri::er level alarms.

d. High reactor outlet coolant temperature trip.

e. High reactor coolant system pressure trip.

f. High power level trip.

14.1.2.3.2 Methods of Analysis An analog computer model was used to determine the characteristics of this ac-cident. A complete kinetics model, pressure model, average fuel rod model, steam demand coastdown model to 15 per cent of rated thermal load, coolant transport model, and a simulation of the instrumentation for pressure and flux trip were included. The initial conditions were normal rated power operation without autcmatic control. Only the moderator and Doppler coefficient of reactivity were used as feedback. The nominal values used for the main para-metersvege0. see trip delay time, -1.14 x 10-5 (dk/k)/F Doppler coefficient,

 +6 x 10 - (ik k)/F modertter coefficient, 25 in.'/ min control rod speed, and 1.2%o k/k control rod group worth. The total worth in all the control rods 1k-6                           02 L

inserted into the reactor core following any trip is 8.h% ak/k without a stuck control rod, or 5.h% ak/k (the nominal value used) with a stuck rod. The foregoing rod speed and group rod worth give a maximum reactivity addition rate of 5.8 x 10-5 (ak/k)/sec, which is the nominal case. The reactor protec-tion system is designed to limit (a) the reactor power to 114 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. 1L.1.2.3.3 Results of Analysis Figure 1h-13 shows the results of the nominal red withdrawal frca rated power using the 1.2". k/k rod group at 5.8 x 10-5 (ak/k)/sec. The transient is ter-minated by a high pressure trip, and reactor power is limited to 108 per cent, much less than the design overpower of llh per cent of rated pcVer. The changes in the parameters are all quite small, e.g., 5 F average reacter cool-ant temperature rise and 200 psi system pressure change. A sensitivity analysis of important parameters was performed around this nomi-nal case, and the resultant reactor coolant system pressure responses are shown in Figures 14-lh through lh-16. Figure 1h-1h shows the pressure variation'for a very vide range of rod with-drawal rates - more than an order of magnitude smaller and greater than the nominal case. For the very rapid 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 slover transients. In no case does the thermal power exceed 108 per cent rated power. An analysis has been perfor=ed extending the evaluation of the red withdrawal accident for various fractional initial power levels up to rated power. This evaluation has been performed assuming simulated withdrawal of all 69 control rods giving a maximum reactivity addition of rate of 5.8 x 10-h (ak/k)/sec. This rate is a factor of ten higher than used in the cases evaluated at rated power. The results of this analysis are shown in Figures 1h-17 and Figures 14-18. As seen in Figure IL-lT the peak ther=al power occurs for the rated power case and is well below the maximum design power of llh per cent. The peak neutron power for all cases is approximately 117 per cent of rated power and repre-sents a slight overshoot above the trip level of 114 per cent. Figure 1h-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 temperature only increases by 26 F. It is therefore readily con-cluded that no fuel damage would result from simultaneous all-rod withdrawal from any initial power level. Figures 14-15 and 1h-16 show the pressure response to variations in the trip delay time and Doppler coefficient. For the higher values of the Doppler coef-ficient, the pressure trip is always actuated, and, therefore, the pressure levels off. 0310 M 1h-7

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. 1h.1.2.4 Moderator Dilution Accident 14.1.2.4.1 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 several interlocks and alams to prevent improper operation. These are as fol-lows:

a. Flow 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 timing device to limit the integrated flow has been set. Dilution water is added at flow rates up to 70 gym.

b. Flow 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).

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

The makeup and purfication system nomally has one pump in operation which supplies up to 70 gpm to the reactor coolant system and the required flow to the ) reactor coolant pump and control and drive seals. Thus, the total makeup ficw 'l available is limited to 70 gpm unless the operator takes action to increase the l amount of makeup flow to the reactor coolant system. When the makeup rate is greater than the maximum letdown rate of 70 gpm, the net water makeup will cause the pressurizer level control to close the makeup valves. The nominal moderator dilution event considered is the pumping 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-sients when the system pressure is lower than the nomininal value and the pres-surizer level is below nomal. This flow might be as high as 100 gpm. In addition, with a combination of multiple valve failures or maloperations, plus more than one makeup pump operating and reduced reactor coolant system pressure, the resulting inflow rate can be as high as 500 gpm. This consti-tutes the maximum dilution accident. A reactor trip would teminate unborated water addition to the makeup tank, and total flow into the coolant system would be terminated by a high pressurizer level. The criteria of reactor protection for this accident are:

a. The reactor power will be limited to less than the design overpower of _114 per cent rated power to prevent fuel damage.

cau. M 14-8 D 2-8-68 Amendment No. 1

b. The reactor protection system will limit the reactor coolant system pressure to less than the system design pressure of h 500 psig.

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

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

14.1.2.4.2 Analysis and Results The reactor is assumed to be operating at rated power with an initial boron concentration (1,800 ppm) in the reactor coolant system. The dilution water is uniformly distributed throughout the reactor coolant volume. Uniform distribution results frcm a discharge rate of 70 - 500 gpm into a reactor cool-ant flow of 88,000 gpm. A change in concentration of 100 ppm produces a 1% ak/k reactivity change. The effects of these three dilution rates on the reac-ter are as follows: Average Reactor Dilution Water Reactivity Rate, Coolant System Flow, gpm (ak/k)/sec Temp. Change, F/sec 70 +2.5 x 10-6 0.3 100 +3.6 x 10-6 03 500 +1.8 x 10-5 0.4 The fastest rate of dilution can be handled by the automatic control system, which would insert rods to maintain the power level and reactor coolant system temperature. If an interlock failure occurred while the reactor was under manual control, these reactivity additions would cause a high reactor coolant temperature trip or a high pressure trip. In any event the thermal power will not exceed 114 per cent rated power, and the system pressure will not exceed the design pressure of 2,500 psig. Therefore moderator dilution accidents will not cause any damage to the reactor system. During refueling or maintenance operations when the reactor closure head has been removed, the sources of dilution water makeup to the makeup tank--and therefore to the reactor coolant system--are locked closed and the makeup pumps are not operating. At the beginning of core life when the boron concen-tration is highest, the reactor is about 9.5% ak/k suberitical with the maxi-mum worth rod stuck out. To demonstrate the ability of the reactor to accept moderator dilution during shutdown, the consequences of accidentally filling the makeup tank with dilution water and starting the makeup pumps have been evaluated. The entire water volume from the makeup tank could be pumped into the reactor coolant system (assuming only the coolant in the reactor vessel is diluted), and the reactor would still be 6.5% ak/k suberitical. 0312 14-9 2-8-68 Amendment No 1

1k.1.2.5 Cold Water Accident The absence of individual Joop isolation valves eliminates the potential source of cold water in the reactor coolant system. Therefore, this accident is not credible in this reactor. 14.1.2.6 Loss-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 reac-tor coolant pu=ps should fail. A pumping failure can occur from mechanical failures or from a loss of electrical pcuer. With four independent pumps avail-able, a mechanical failure in cne pump vill not affect operation of the others. Ench reactor coolant pump receives electrical power from one of the two elec-trically separate busses of the 6,900 volt system discussed in 8.2.2.2. Loss of the unit auxiliary transformer to which the 6,900 volt busses are normally connected vill initiate a rapid transfer to the startup transformer source with-out loss of coolant flow. Faults in an individual pump motor or its power sup-ply could cause a reduction in flow, but a complete loss of flow is extremely unlikely. In spite of the low probability of a ec=plete loss of power to all reactor ecol-ant pumps, the nuclear unit has been designed so that such a failure would not lead to core damage. The reactor protection criterion for loss-of-coolant-flow conditiens starting ,) . ct rated power is that the reactor core vill not reach a Departure from Nucleate Boiling Ratio (DNBR) smaller than the DNBR in the hot channel at the steady state d; sign overpower. This corresponds to a DNER of 1.38 at llh per cent rated power (Table 3-1). 14.1.2.6.2 Methods of Analysis The loss-of-coolant-flow accident is analyzed by a combination of analcg and digital computer programs. Analog simulation is used to determine the reactor flow rate following loss of pumping power. Reactor power, coolant ficv, and inlet temperature are input data to the digital program which determines the core thermal characteristics during the flow coastdown. The analog model used to determine the neutron power following reactor trip in-cludes six delayed neutron groups, control rod worth and rod insertion char-ceteristics, and trip delay time. The analog model used to determine flev coastdown characteristics includes description of flow-pressure drop relaticns in the reactor coolant loop. Pump ficv characteristics are determined frcm 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 used to compute channel DNBR continually during the coastdown transient. System flow, neutron power, fission product decay heat, and core entering enthalpy are varied as a function of time. The program maintains a transient inventory of stored heat which is determined from fuel and clad temperatures beginning with the initial s,_j

                                             ~

R 0313

steady state conditions. The transient core pressure drop is determined for average channel conditions. The representative hot channel flows and corres-ponding DNBR are obtained by using the average core pressure drop. The hot channel DNBR as a function of time is compared with the design DNBR at maximum overpower to determine the degree of heat transfer margin. 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 lov flow condi-tion until initial downward movement of control rod, is 300 milli-seconds.

d. The per cent of beginning-of-life neutron 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-ft2, 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 evaluation are pre-sented in Figures 1k-19 and 14-20. Figure 14-19 shows the per cent reactor flow as a function of time after loss of all pu=p power. Figure lh-20 shows the mini-mum DNBR's which occur during the coastdown for various initial power levels.

The degree of core protection during coastdown is indicated by comparing the DNBR for the coastdown with the design value of 1,38 at 11h per cent rated power. This DNBR (1.38) in the representative hot channel corresponds to a 99 per cent con-fidence that 99 5 per cent of the core vill not experience a departure from nu-cleate boiling under steady state conditions at the design overpower (3.2.3.1). 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 reactor instrumentation). This power level represents an allowance of plus 2 per cent rated power for trsnsient overshoot. This power level also rep-resents the maximum power demand that v311 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) 4 per cent nuclear instrument ation. The true power level could be as high as

 '108 per cent at 102 per cent indicated power. As shown in Figure 1k-20, how-ever, the DNBR at 108 per cens is 1.44, whien is signifiesntly larger than the design DNER.

1 The reacter coolant system is capable of providing natural circulation flow after l the pumps have stopped. The natural circulation characteristics of the reactpr coolant systen have been calculated using conservative values for all resistance and form loss factors. No voids are assumed to exist in the core or reactor 0314 1h-ll ,

outlet piping. The folleving tabulation and Figure 9-8 show the natural cir-culation flow capability as a function of the decay heat generation. Time After Decay Heat Natural Circulation Flow Required for Loss of Core Power, Core Flow Available, Heat Removal, Power, see  % Full Flow

                           %                                        % Full Flev 0.36 x 102             5                   4.1                      2.3 2.2 x 102              3                   3.3                      1.2 1.2 x 10k              1                   1.8                      0.36 1.3 x 105             1/2                  1.2                      0.20 The flows above provide adequate heat transfer for core cooling and decay heat removal by the reactor coolant system.

The reactor is protected against reactor coolant pump failure (s) by the protec-tien system and the integrated control system. The integrated control system initiates a power reduction on pump failure to prevent reactor power from ex-ceeding that permissible for the available flow. The reactor is tripped if in-sufficient reactor coolant flow exists for the power level. The operating lim-its for less than four pumps in operation have been presented in h.3.7. 14.1.2.7 Stuck-Out, Stuck-In, or Drceted-In Control Rod 14.1.2.7.1 Identification of Cause The control rod drives have been described in 3.2.h.3. The results of continu-ous control rod withdrawal have been analy::ed in 14.1.2.2 and Ih.l.2.3. In the event that a control rod cannot be moved because of electrical faults or mechan-ical seizure, localized power peaking and suberitical margin must be censidered. 14.1.2.7.2 Analysis and Results Adequate hot suberitical =argin is provided by requiring a suberiticality of 1% Ak/k suberitical with the control rod of greatest worth fully withdrawn frca the core. The nuclear analysis reported in 3.2.2 de=enstrates that this criterien can be satisfied. In the event that an unmovable control rod is partially cr fully inserted in the core or a single rod is dropped during operation, its location and effect on local power distribution determine whether continued power operation is permis-sible. The location of a stuck rod in the core vill be studied further to de-fine permissible conditions of operation. The criteria for these studies are (a) operation with a stuck rod vill not increase the DNB probability above the probability specified for design conditions, and (b) a hot suberitical =argin of 1% Ak/k vill be maintained with the stuck rod in its inoperative position and the operating rod of greatest reactivity worth in the fully withdrawn posi-tion. If a control rod is dropped into the core during power operation, the same con- 3 sideration of localized power peaking as for a stuck rod vill apply. Q 0315 1k-12

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14.1.2.8 , ' Loss of Electric Power

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14.1.2.8.1 , Identification of Cause O The Russellville Nuclear Unit is designed to withstand the effects of less of clectric load or electric power. Two types of power losses are considered:

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

b.

         '          A hypothetical condition resulting in a complete loss of all plant power.                       '

The reactor protection criteria for these conditions are that fuel damage vill not occur from an excessive power-to-flow ratio and that the reactor coolant Gystem pressure vill not exceed design pressure. lk.1.2.8i2 Results of "Blackcut" Conditions Analysis The net effect of a " blackout" condition on the nuclear unit vould he opening of all 161 and 500 kv breakers, thus disconnecting the plant from the entire transmission system. When this occurs en the nuclear unit, a runback signal on the integrated master controller causes an automatic power reduction to 15 per c;nt reactor power. Other actions that occur are as follows:

a. All vital electrical loads, including reactor coolant pumps, condenser circulating water pumps , condensate and condensate booster pumps , and other auxiliary equipment, vill continue to obtain power from the unit generator.

' Feedvater is supplied to the steam generators by steam-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 frequency vill peak at less than the overspeed trip

                . point and decay back to set frequency in h0-50 sec.

c. Folleving closure of the turbine Covernor valves and the reheat stop and interceptor valves, steam pressure increases to ;he turbine bypass valve se,t point and may increase to the steam syster. safety valve set point. Steam is relieved to the condenser and to J.te atmosphere.

JSteam venting to the atmcsphere occurs for about 9 min following black-out from 100 per cent rated power until the turbine bypass can handle all excess steam generated. Approximately 166,000 pounds of steam 3 are vented to the atmosphere during this time. The capacity of the modulating turbine bypass valve is 15 per cent of the valves vide open (VWO) steam flow, and that of the safety valves is 100 per cent of VWO steam flow. The first safety valve banks are set at 1,050 psig with additional banks set at pressures up to 1,104 psig (5 per cent above design pressure as allowed by code). Steam venting permits 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 auxiliary load. This allows sufficient steam flow for regu-lating turbine speed control. Excess power above the unit auxiliary load is rejected by the turbine bypass valve to the condenser. ~~~ J 03.1.6 5 lh-13 5-3-68

d. During the short interval while the turbine speed is high, the vital electrical loads connected to the unit generator will under-

              'go speed increase in proportion to the generator frequency increase.

All motora and electrical gear so connected are designed for the in-creased frequency.

e. 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 damage or excessive pressures on the reactor coolant system. There is no resultant radiokgical 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 1 per cent failed fuel and 1 gpm steam generator tube leak-age is shown to be safe by the analfsis presented in 11.13.6 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 primarily due to the release of Xe and Kr. Release of these gases is not increased by the steam relief because even without relief, all of these gases are released to the atmos-phere through the condenser vacuum pump exhaust. The rate of release of iodine during the approximately 2 minuter of relief would be increased by almos; a factor of 10 , because the iodine 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 0.004 MPC at the 0.65-mile exclusion distance. 14.1.2.8.3 Analysis Results of Complete Loss of All Plant Power The second power loss considered is the hypothetical case where all plant power except the plant batteries is lost. The sequence of events and the evaluation of consequences relative to this accident are given below:

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

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 feedwater pump is provided to supply feedwater any time the main feed pumps cannot operate. The emer-gency feed pump takes suction from the condensate storage tank. 17 The emergency pump supplies feedwater to the steam generators.

The emergency feed pump is driven by steam from either or both m j steam generators. ' 14-14 NI.7 hup ent No. 17

d. (Cont'd)

A full capacity electric-driven emergency feedwater pump serves as a backup to the turbine-driven emergency feedwater pump. Its power supply is from the 4.16 KV bus (A2), which can be energized 17 by closing the tie breaker from the 4.16 KV engineered safegus.rd bus (A4) and then manually starting the pump from the Control ,1 Room. ' 14-14-a 5-4-70 Supplement No. 17 L, W 0318

The controls and auxiliary systems for theemergency feed pump operate on d-c pcwor fro.c the plant batteries. A recirculation line from the emergency pump discharge back to the 17' condensate storage tank is provided to per=it periodic testing.

e. The condensate storage tank provides cooling water in the unlikely event that all power is lost. The minimum condensate inventory is 250,000 gal. This inventory provides sufficient water for decay heat cooling (assuming infinite irradiation at 2,568 MWt) for a period of approximately one day.

The features described above per=it decay heat cooling of the nuclear unit for an extended period of time following a complete loss of electric power. The foregoing evaluation demonstu tes the design features incorporated in the dtsign to sustain loss of power ceniiticns with just the plant batteries to oper-ate system controls. Immediate operation of the emergency feedwater pump is not of critical nature. The reactor can sustain a complete electric power loss with-out emergency cooling for about 25 min. before the steam volume in the pressurizer is filled with reactor coolant. These 25 min. are derived as fo11cvs:

a. Steam generators evaporate to dryness 10 min.

b. Pressurizer safety valves open 5

c. Pressurizer fills with water (due to reactor coolant system expansion) 10 25 min.

B: yond this time reactor coolant will boil off, and an additional 90 min. will have clapsed before the boiloff will start to uncover the core. The emergency feedwater pump can be actuated within this period of time. Accordingly, core protection is insured for the unlikely condition of total loss of plant electric power. 14.1.2 9 Steam Line Failure 14.1.2 9 1 Identification of cause Analyses have been performed to determine the effects and consequences of loss of secondary coolant due to failure in the steam line between the steam generators and the turbine. Tb; criteria for plant protection and the reldase of fission products to the cnvironment are as follows:

a. The reactor shall trip and remain suberitical without boron addition until a controlled rate of system cooldown can be effected.

b. There will be no fuel amge as a result of the transient.

c. No steam generator tube da= age will occur due to the loss of secondary side pressure and resultant temperature gradients.

d. Doses will be within acceptable limits.

W 14-15 {}3,j,9 5_h.7o Supplement No. 17

12.1.2 9 2 Analysis and Results The rate of reactor system cooling following a steam line break accident is a function of the area of the failure and the steam generator water inventory tvailable for cooling. The steam generator inventory increases with power level. 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 release effects. The i==ediate 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 gen-cration. 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 feedwater flow. In addition, the turbine control valves will open to maintain power gen-Gration. Increased feedvater flow causes the average reactor coolant tempera-ture to decrease, and the resulting temperature error calls for control rod withdrawal. The limiting action in this condition is the 102 per cent limit on power demand to the rod drive control system. If the moderator temperature coefficient of reactivity is small or slightly positive, the reactor power will decrease when the control system reaches the power demand limit because of con-tinuing temperature decrease. The reactor will then trip on low reactor cool-ant system pressure. A reactor trip will initiate a reduction in the feedwater flow to the steam generators. When the moderator temperature coefficient is negative, the reactor power will t:nd to increase with decreasing average coolant temperature. This will cause control rod insertion to limit reactor power to 102 per cent. With power limi-t d 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 blow down to the atmosphere. The stimum cooldown of the reactor coolant system would be that resulting from the blowdown frcm one steam generator. A typical cooling rate following re-actor trip for a steam line rupture of 4 in.2 is shown in Figure 14-21. The tabulation below lists the approximate time required to blow down the con-t:nts of the steam generator with a ruptured steam main. Ieak Area, in.2 Blowdown Time, sec 4 860 32 110 128 27 A steam line failure of large area results in high steam flow with resulting rapid pressure decrease in the reactor coolant system and steam system. The 14-16 / ~ 2-8-68 Amend =ent No. 1 MEO

reactor trips on low reactor coolant system pressure or high flux. Reactor i trip causes turbine trip and reduction in feedvater flow to decay heat level. The turbine trip closes the turbine stop valves which isolate the steam lines and prevent blovdown of the stesm generator whose secondary side does not have a pipe rupture. The steam generators are designed to maintain reactor system integrity upon loss-of-secondary-side pressure. Therefore, this accident vill not lead to a reactor coolant system failure. Assuming the blowdown from one steam generstor results from a secondary steam system rupture, the maximum cooling rate during this accident occurs during the first 10 see after the break. The maximum cooling rate is approxfr.ately 3 F/sec, and a low pressure or high flux trip occurs. The net cooldown of the reactor coolant system, assuming total blowdown of one steam generator and accounting for transfer of core stored heat and decay heat, is less than 50 F. This results in an average coolant temperature of 530 F which is about 10 F lower than the nor=al zero power average coolant temperature. The minimum shutdown cargin at 540 F vith the most reactive rod stuck out is 2 9% ak/k, The reduction in reactivity shutdown margin associated with cool-ing the moderator temperature 10 F below its nor=al shutdown temperature of 540 F vould be 0 30% ak/k. Using tJe maximum negative value for the moderator temperature coefficient (-3 0 x 10 (ak/k)/F), the shutdown margin at 530 F vould be 2.6% ak/k, which is adequate to prevent return to criticality. In addition, high pressure injection vill be actuated during the cooldown period following a large area stean line failure if the pressure drops belov 3 1800 psig. This system supplies borated water to the reactor coolant system to increase the shutdown margin further. Boron addition to the reactor cool-ant during the controlled cooling of the system to atmospheric pressure vill prevent criticality at lover temperatures. The effect of a steam line rupture inside the reactor building has been eval-uated by conservatively assuming an instantaneous release to the reactor building of the enert associated with this accident. The mass and energy releases per steam gerurator 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 56 Reactor Coolant System Energy Transferred 17.6 Total 58,800 51.2 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 below the reactor building design pressure of 59 psig. 14-17 OW. 5-3-68 Supplement No. 3

The environmental consequences from this accident are calculated by assuming that the nuclear unit has been operating with 1 gpm steam generator tube 3 leakage. The reactor coolant activity assu=es prior operation with 1 per cent filed fuel rods. With these assumptions, the steam generators contain a total of 0.09 equivalent curies of iodine-131. It is further assumed that steam generator leakage continues for three hours before the nuclear unit can be cooled down and the leakage terminated. This additional leakage cor-responds to 3.h equivalent curies of iodine-131. The iodine is assumed to be released-directly to the atmosphere where it mixes in the wake of the re-actor building. With these assumptions an integrated dose to the thyroid at the exclusion distance of 0.60 rem is obtained. The corresponding dose to h the whole body, dose due to krypton and xenon, is h.3 x 10-3 rem. 14.1.2.10 Steam Generator Tube Failures 1h.1.2.10.1 Identification of Accident In the event of a reactor coolant leak to the secondary system, such as a complete severance of a steam generator tube, the activity contained in the coolant would be released to the secondary system. Radioactive gases and some of the radioactive iodine would be released to the atmosphere through the condenser air removal system. 1h.1.2.10.2 Analysis and Results In analyzing the consequences of this failure, the following sequence of events is assumed to occur: s

a. A double-ended rupture of one steam generator tube occurs with ,

unrestricted discharge from each end,

b. The initial leak rate, approximately h30 gpm, exceeds the normal makeup of 70 gpm to the reactor coolant system, and the system pressure decreases. No operator action is assumed, and a low reactor coolant system pressure trip will occur in about 8 min.

c. Following reactor trip, the reactor coolant system pressure cen-tinues to decrease until high pressure injection is actuated at a pressure of 1,800 psig. The capacity of the high pressure in-jection is sufficient to compensate for the leakage and maintsins both pressure and volume control of the reactor coolant system.

Thereafter, the reactor is conservatively assumed to be cooled down and depressurized at the normal rate of 100 F per hour,

d. Following reactor trip, the turbine stop valves will close. Since a reactor coolant to secondary system leak has occurred, steam line pressure vill increase, opening the steam bypass valves to the condenser. Each bypass' valve actuates at a lover pressure than do the safety valves. The reactor coolant that leaks as a a result of the tube failure is condensed in the condenser. Only the fission products that escape from the condensate are released to the atmosphere.

0322 -) 4 S 6-5-68 lh-18 Supplement No. 4

e. The affected steam generator can be isolated by the turbine stop valves 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 gen-erator until the te=perature is reduced to 250 F. Thereafter, cooldown to ambient conditions is continued using the decay heat re= oval system. f. At the design cooling rate for the pressurizer of 100 F/hr, de-pressurization to 1,050 psig r quires approximately 1 7 hr. Dur-in6 this time period 1.68 x 10 cc (5,650 ft3) of reactor cool-ant leaks to the secondary system. This leakage corresponds to approxicately 32,100 curies of xenon-133 if the reactor has 3 been operating with 1 per cent failed fuel. The radioactivity released during this accident is discharged through the turbine bypass to the condenser and then out the plant vent. A partition factor of 10 4 is assu=ed for iodine in the condenser. (1,2) Noble 6ases are assu=ed to be released directly to the plant vent. The total dose to the whole body from all the xenon and krypton released is only 0.83 rem at '3 the 0.65-mile exclusion distance. The corresponding dose to the thyroid at the same distance is only 1.1 x 10-2 rem. This calculation conserva-tively assu=es that the plant vent discharge mixes in the wake of the build-ing structures rather than remaining at its elevated release height. 14-19 1 0323 k 5-3-68 Supplement No. 3

lb.2 STANDBY SAFEGUARDS AHALYSIS 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 en the ulti= ate pcVer level of 2,568 MWt rather than the rated power level of 2,h52 MWt. Table lk-3 su=marizes the potential accidents studied. Table lk-3 Situaticns Analvred and Causes Event Cause Effect Fuel Handling Mechanical damage Integrated dese at exclusion dis- 3 Accidents during transfer. tance is 0.5h ren thyroid and 0 5h 5 rem whole body. Rod Ejection Failure of centrol Scme clad failure. Thirty-day dose Accident rod drive pressure at exclusion distance is 1.3 rem housing thyroid. Los s-o f-Coolant Rupture of reactor No clad melting. Thirty-day dose a: Accident coolant system exclusion distance is 12.h rem thy-roid.

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Maximum Release of 100% rare Two-hour dose at exclusicn distance Hypothetical gases, 50% iodine, is 63 rem thyroid. Thirty-day dose Accident and 1% solid fissicn at h-mile low population ene is products. 22.h rem thyroid. 1h.2.2 ACCIDENT ANALYSES 1h.2.2.1 Fuel Handline 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 centrol rods removed, the k of a core is no greater than 0.98. In the spent fuel storage pool, the f6$[ assemblies are stored under water in storage racks having an eversafe geometric array. Under these conditions, a criticality accident during re-fueling 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 refuel-ing operations. kl,3d. ' g 7-3-68 Supplement No. 5 14-20

ik.2.2.1.2 Analysis and Results The fuel asse=bly is censervatively assu=ed to have crerated at 29 MWt, twice the power level of an average fuel asse=bly. The reactor is assumed to have been shut down for 2h hr, which is the minimum time for reactor cooldown, re-actor elesure head re= oval, and re=cval of the first fuel asse=bly. It is further assumed that the entire outer row of fuel rods, 56 of 208, suffers l3 damage to the cladding. Since the fuel pellets are cold, only the gap activity is released. Ihe fuel rod gap activity is calculated using the escape rate coafficients and calculational methods discussed in 11.1.2.1. The gases released from the fuel assembly pass through the spent fuel storage pool water prior to reaching'the auxiliary building atmosphere. As a minimum, the gases pass through 10 ft of water. Although there is experimental evidence that a portien of the noble gases vill remain in the water, no retention of noble gases is assumed. Based on the data in References 3 and h, 99 per cent of the iodine released frcm the fuel assembly is assu=ed to remain in the water. The total activity released to the building atmosphere is therefore Iodine 27.5 dose equivalent curies I-131 3 Noble gases 2.69 x 10h curies The auxiliary building is ventilated and discharges through filters to the plant vent. The discharge frc= the plant vent is assumed to mix in the wake of the building structures rather than remain at its elevated release point. This assumption produces less favorable dilution and, therefore, higher ground con-c;ntrations at the exclusion distance. The activity is assumed to be released as a puff from the plant vent. Atmo- 3 spherge dilution is sec/m developed in calculated using the 2-hour dispersion factor of 3.8 x 10-4 2.3. The total integrated dose to the whole body at the 0.65-mile exclusion distance is 0.54 rem, and the thyroid dose at the same dis-tance is 5.h rem. 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 in-creased by a factor of 10. The dose from this increased iodine release is still a factor of more than 5 belev 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 red withdrawal (lk.1) were shown to be safely terminated without damage to the reactor core or reactor coolant system integrity. In order for reactivity to be added to the core at a more rapid rate, physical failure of the control rod drive housing or con-trol rod drive no::le must occur. Failure in the drive upper pressure housing can cause a pressure differential to act on a control rod asse=bly and rapidly oj2ct the asse=bly from the core region. The power excursion due to the rapid incrcase in reactivity is limited by the Doppler effect and terminated by re-Ector protection system trips.

U 0325 lk-21 g 5-3-68 Supplement No. 3

The criterion for reactor protection, should this condition occur, is that the reactor will be operated in such a manner that a control rod ejection accident vill not further damage the reactor coolant system.

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

Worth of ejected rod 0 3% Ak/k Rod ejection time 0.150 see Ultimate power level 2,568 Et Reactor trip delay 0 3 see The severity of the rod ejection accident is dependent upon the worth of the ejected rod and the reactor power level. The control rod group of greatest worth is the first of the entire rod pattern to be withdrawn from the core. The vorth of the ejected rod can be as high as 30 per cent of the total pattern worth of 10.0% Ak/k, i.e. , 3% A k/k. However, the 3% 4k/k value exists only when the reactor is suberitical. The details of control rod vorth calcu-lations 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 main-tained at a level whereby the reactor is at least 1 per cent sub-critical with the control rod of greatest worth fully withdrawn - from the core. Therefore, rod ejection, when the reactor.is sub-critical and all other rods are in the core, does not cause a nuclear

                                                                                  )

excursion. As criticality is approached, the worth of the re=aining control rc.is decreases. At criticality, rod ejection would result in a maximum reactivity addition of 0 56% A k/h. At rated power, but before equilibrium xenon is established, the total rod pattern worth remaining in the core is 2.8% Ak/k. At equilibrium xenon the pattern vorth is 1.8% Ak/k. Before estab-lishing equilibrium xenon, the greatest single control rod vorth is 0.46% A k/k. A single rod worth of up to 0 7% has been used in the analysis of this accident. In order for any one rod to have this much worth, it would neces-sarily be fully inserted in the core. Assuming that a pressure housing failure occurs in such a manner that it no longer offers any restriction for rod ejection, the time and therefore the rate of reactivity addition can be calculated. Further assuming that there is no viscous drag force limiting the rate of ejection, con-trol 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 time of 0.150 sec (75 per cent of active core height) is used in the analysis.

b. Fuel Rod Damage Criteria Power excursions caused by reactivity disturbances of the order of ma6nitude occurring in rod ejection accidents could lead to three s

                                                                                 ')

14-22 REVISED,.'2-8-68

potential modes of fuel rod failure. First, for very rapid and large transients in which there is insufficient time for hest transfer from fuel to cladding, fuel melting followed by vaporization can generate destructive internal pressures without increasing cladding tempera-tures significantly. The second mode occurs when the internal vapor pressure is not sufficient to cause cladding rupture, but subsequent heat transfer raises the temperature of the cladding and weakens it Lntil failure occurs. The third mode occurs when the nuclear excur-sion has insufficient energy to cause significant melting of the fuel, but subsequent heat transfer to clad from fuel may cause excessive cladding temperatures. In all three cases there is a possible occur-rence of metal-water reactions. 'However, only very rapid and large transients will generate a rapid pressure buildup in the reactor cool-and system. The energy required to initiate UO2 fuel meltin gm, based on an initial temperature of 68 F.(5)gThe is heat 220 to of 225 cal / fusion requires an additional 60 cal /gm. Any further energy addition va-porizes the fuel and produces a buildup of vapor pressure within the fuel rod. The effect of the vapor pressure is dependent upon the temperature and ultimate strength of the cladding. Energy additions of up to h20 cal /gm have been calculated to be necessary before the bursting pressure of cladding is exceeded. The lower limit for pro-ducing significant fuel vapor pressure (1h.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. At a lover limit, the potential for bursting of cladding and release of molten fuel to the reactor coolant is conservatively set at a fuel enthalpy of 280 cal /gm in this evaluation. For power excursions with energy bursts below 280 cal /g=, zirconium-water reactions are possible. A correlation of the TREAT experiments presents a method of correlating the potent $al zirconium-water reac-tion as a function of fission energy input.lT) These data are based on initially cold (room temperature) 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 temperature or by adding the initial steady state fuel enthalpy to the nuclear energy burst and obtaining an equivalent final fuel enthalpy. Accordingly, a zirconium-water reaction requires a minimum fuel enthalpy of 125 cal /gs. Increasing fuel enthalpies cause a linear increase in the percentage of the reaction, which may be approximated by

          %Zr-H 2 O Reaction = 0.125 (Final Fuel Enthalpy - 125).

It is assumed that DNB will take place when the clad reaches a heat flux of 6.36 x 105 Btu /hr-ft 2 . At this heat flux the hot fuel rod enthalpy would be approximately 140 cal /gm at EOL and 130 cal /gm 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 assumed to fail from overheating and releases the gap activity of that fuel rod. M 14-23

14.2.2.2.2 Method of Analysis The hypothetical control rod ejection accident was investigated using the exact 1-dimensional WIGL2 digital computer program. (8) It was found that.the point kinetics analog model results a5 reed with the WIGL2 results to within 10 per cent for rod worths up to 0 75% A k/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 the the average power has reached a level just a few per cent above the initial power level, the flux shape has almost no perturbation in the region previously occupied by the ejected rod. The entire reactor flux then rises uniformly until the Doppler effect teminates the excursion. Thus by applying the peak-to-average flux factors of 2 92 for EOL and 3 24 for BOL to the point kinetics results, the peak and integrated flux at any point in the reactor can be accurately assessed. 14.2.2.2 3 Analysis and Results

a. Source Power A sensitivity study at source level has been done around a single rod worth of 0 5% Ak/k. This analysis was perfomed with the core 0 5% k/k suberitical so that a total rod worth of 1% Ak/k was,g withdrawn in 0.150 sec. The reactor power was initially at 10 of the ultimate power level. The low pressure trip occurs at 1 7 see after the ejection starts, and the reactor power is terminated at a peak value of 39 per cent ulti= ate power. This peak neutron .

Power value is not reached until about 15 see after the rod is j ejected because Doppler feedback controls the rate of rise and m35nitude of the neutron power. Therefore, a low pressure trip will teminate the accident before significant power is generated owing to the loss of coolant through the rupture. An analysis was perfomed for the accident above without a low pressure trip to de=onstrate the capability of the reactor to accept the accident. In this case the neutron power reaches 1,000 Wt (39 per cent ulti-

        = ate pover), and the peak fuel temperature is 990 F. This is far below the melting temperature of UO , and the resultant thermal powerisonly16percentofultimakepower. Hence, no fuel anmge would result from the rod ejection accident at source power level.

b. Ultimate Power A sensitivity study at ult hate power level hac been done around an assumed single rod worth of 0 3% Ak/k. The analysis includes rod worths from 0.1 - 0 7% Ak/k. For the ult hate power case at beginning-of-life (BOL), the ejection of a single control rod worth 03%Ak/kwouldresultinvirtuallynoZr-HOreactionandapprox-2 imately 1% of the core experiencing DUB (see Figures 14-22 and 1h-23).

Thehotfuelrodwouldreachapeakenthalpyofabout166 cal /gs. For the end-of-life case (EOL), the reactor neutron power peaks at 6,190 st, 200 milliseconds after the start of ejection of a o.3% h

                     $                14-24 REVISED, 2-8-68 03'88

Ak/k control rod. The prompt negative Doppler effect terminates the power rise, and control rod insertion from high flux signal ter-minates the excursien. The total neutron energy burst during the transient is approximately 3,200 IN-sec. The final fuel enthalpy of the nominal rod is 113 cal /gs, i.e., the enthalpy of the hot rod is 163 cal /gm. This enthalpy is considerably below the minimum range (220:to 225 cal /gs) 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 distri-bution of the energy of the excursion. With this distribution of fuel enthalpies, and using the TREAT correlation, 0.53 per cent of the zirconium cladding may react (see Figure 1h-23) to contribute an addi-tional 677 IG-see of energy. The resultant temperature increase is 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 system. As a result of the postulated pressure housing failure, which produces a rupture size of 0.0h ft2, reactor coolant is lost fron the system. The rate of mass and energy input to the reactor building is consider-ably lower than that for the 3.0-ft2 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 3.0-ft2 rupture. The environmental consequences from this accident are calculated by conservatively assuming that all fuel rods that undergo a DNB result in clad failure ar.d subsequent release of the gap activity. Actually, most of the fuel rods vill recover from the DNB, and no fission prod-uct release vill cecur. 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 11.1.2.1. Using the environmental models and dose rate methods discussed under the 3 loss-of-coolant accident, the total integrated dose to the thyroid at the exclusion distance from this accident is only 1.3 res in 30 days , which is more than a factor of 200 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 14-2h through 14-32. Figure l 1h-24 shows the variation in the peak neutron power as a function I of the worth of the ejected control rod. For the nominal 0.3% Ak/k l case from ultimate power, the peak neutron power is less than 300 per I cent, again assuming that a low pressure trip does not occur. The rod ejection from source level results in a Doppler turn-around be- l (' fore the flux trip is reached. Figure 14-25 shows the variation in l the corresponding thermal power with control rod vorth. ' 5-3-68 Supplement No. 3

Figura 1k-26 shows th7 corrzsponding cnthalpy incr;asa of tha hot fu;l rod versus control rod worth. Note the very s=all spread in values for the SQL and EOL ultimate power conditions. As expected, the en-thalpy increases with rod vorth. Figures 14-27 through 14-30 show the peak reactor neutron and ther=al powers as a function of changes in the positive moderator te=perature coefficient and negative Doppler coefficient for the nominal 0 5% Ak/k control rod ejection from source level. There was insignificant varia-tion of the peak neutron and cher=al power with changes in the two re-activity feedback coefficients. Figure 14-31 shows the change in nominal ther=al power with variations in the trip delay tine for the nominal 0.3% Ak/k rod ejection fro = 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 1k-32 9hovs the corresponding change in the total enthalpy increase of the hot fuel rod versus the trip delay. The ther=al power never exceeds llL per cent ulti= ate power for any of the variations studied using the no=inal rods (0.3% ak/k for ulti-

 = ate power and 0.5% ak/k for source level). The hot fuel rod average te=perature never increases by more than 310 F above the ultimate power peak value (h,090 F). It is therefore concluded that each of these parameter variations has relatively little effect on the nominal results.
                                                                                  's 14-26                           0330 4

t _ _ _ _ .

1h.2.2.3 Loss-of-Coolant Accident 14.2.2.3.1 Identification of Accident Failure of the reactor coolant system would allow partial or complete 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 a=ount would be boiled off owing to residual heat, fission product decay heat, and possible heat frca chemical re-actions unless an alternate means of cooling vere available. In order to pre-vent significant chemical reactions and destructive core heatup, emergency core cooling equipment rapidly recovers the core and provides makeup for decay heat removal. 14.2.2.3.2 Accident Eases 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 h. 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 melt even if the reactor coolant system should fail and release the coolant. This et. rgency core cooling is provided by the core flooding system, the makeup and purification system (high pressure injection), and the decay heat re= oval system (low pressure injection). These systems are described in detail in Sec-tion c, and their characteristics are su==arized in the following paragraphs. The performance criterion for the emergency cooling equipment is to limit the temperature transient below the clad melting point so that fuel geometry is maintained to provide core cooling capability. This equipment has been conser-vatively sized ta limit the te=perature transient to 2,300 F or less as temper-atures in excess of this value promote a faster zirconiu=-vater reaction rate, and the termination of the transient near the melting point would be difficult to demonstrate. The fuel rods may experience cladding failure during the heatup in the less-of-coolant accident. This could be due to fission gas internal pressure and weak-ening of the clad due to the increase in clad temperature. The mechanical strength of the Zircaloy cladding is reduced as the temperature exceeds 1,000 F auch that the highly irradiated fuel rods , with high fission gas internal pres-sure, may fail locally and relieve the gas pressure when the temperature 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 submerged, and cross-channel flow around the ballooned area vill cool the rod. At worst a local hot spot may occur. l It is calculated that a small number of fuel rods operating at peak power will experience a cladding temperature transient to 1,950 F in about 18 sec. The j injection of emergency coolant, at a time when the cladding is at a temperature ' ' of about 1,950 F, =ay also cause distortion or bowing between supports. As a result some of the fuel rods may crack and allow relief of internal pressure. 1h-27 k) $b$I-l

                                                                                   ]

However, the cladding is expected to re=ain sufficiently intact to retain the solid fuel material and to prevent gross fuel shifting. The tr msient would be limited to regions of the core which operate at peak pcwer. The major portion of the core will 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 temperatures will 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, e;en at the maximum expected temperatures. The supporting stainless steel structure will therefore retain sufficient strength to assure spacing between fuel rods to allow emergency coolant to reach them, 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 tanks, each of which is connected to a different reactor vessel injectior. 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, a rupture of reactor coolant system piping will not affect their perfor=ance. These tanks provide for automatic flooding when the reactor coolant system pressure decreases below 600 psi. The flooding water is injected into the reactor vessel and directed to the bottom of the reactor vessel by the ther=al shield. The core is ficoded from the bottom upward. The combined contents of the two tanks (1,880 fta 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 1000 gpm. Low pressure injection actuated by low reac-tor coolant system pressure supplies coolant at pressures below 100 psig and at a rate up to 6,000 gpm. Both of these systems can operate at full capacity from 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 spray water reaches thermal equilibrium within the build'.ng atmosphere during its passage from the nozzles to the sump. Spray a ter ia supplied from the borated water storage tank until it is emptied. Thereafter, water collected in the sump is recirculated 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 coolers, where steam is condensed. Heat absorbed in the emergency coolers is rejected to the service water _ g -28 0332 2-8-68 Amendment No. 1 x _

system. The heat removal capacity of either of these two reactor building cool-ing systems is adequate to prevent cverpressurization of the building during a loss-of-coolant accident. This analysis demenstrates 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 3.0-ft' rupture rather than for .a 36 in. ID (ll.1 ft2) rupture. Rupture sizes smaller than the 36-in. ID leak result in longer blowdown times, and the amount of energy transferred to tr. t reactor building atmosphere is increased. As a result the intermediate leak size results in a reactor building pressure greater than that produced by the 36 in. ID rupture. 1h.2.2 3.3 Accident Si=ulation

a. Hydraulic Model Blowdown of the reactor coolant system following an assumed ru has been si=ulated by using 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 muse the model more applicable

         .        to this system. The changes are as follows:

(1) The calculation of reactor coolant pump cavitation was based on the vapor pressure of the cold leg instead of the hot leg water. (2) Core flooding tanks have been added. Water flow frem 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 l pressures in the tanks are calculated by assuming an adiabatic 1

 .,                     expansion of the gas above the water level in the tank. Pres-sure, flow rate, and mass inventories are calculated and printed dllllP out in the computer output.

0333 l s rs

(3) Additions to the water physical property tables (mainly in'the subcooled regicn) have also been made to improve the accuracy of the calculations. (b) A change in the : team bubble rise 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 semiscale blevdown tests. Test No. 5k6 from the LOFT semiscale blowdown tests is a typical case for the blevdown through a small rupture area. A comparison of the predicted and experimentally observed pressures is shown in Figure 1k-33. Figure lk-3h shows the per cent mass remaining in the tank versus time as predicted by the code. At the end of blevdown, the predicted mass remaining is 13 per cent. The mea-sured mass remaining is approximately 22 per cent. 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: 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. 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-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. The modified FLASH code has the capability of simulating injection flow from the core flooding tanks. The core flooding transient anal-ysis was performed using the reactor vessel pressure as predicted by FLASH to get the flow from the core flooding tanks. Reactor vessel filling was calculated by adding the mass remaining in the vessel as predicted by FLASH to the mass injected from the core flooding tanks. This method of calculation is conservative in that condensation of ._) 1h-30

ateam by the ccid injection water is not taken into account. A more recent analysis using the FLASH code with condensation effects confir=3 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 (CONTD&T).

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 end the feedbacks of heat transfer and flow on each other.

While tha FLASH thermal model is acceptable for determining the ef-feet of core heat transfer on the blevdown process, a more extensive simulation is necessary for evaluation of the core te=perature tran-sient. Additional analytical models and a digitc.1 computer program (SLU:P) were developed to simulate the core thermal transient for the period beginning with the initiation of the leak and ending after the core tenperature excursion had terminated. The model includes the effects of heat generation from neutrons be-fore reactor trip, neutron decay heat, and fission and activation product decay heat; the exothermic zirconium-water reaction based on the parabolic rate law; heat transfer within the fuel rods, limited heat convection frem 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 flow 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-water reaction equation (11) with the following constants is si=ulated for each of the re-gions:

                           " (r - r)    **E  ~

c 0335 M 14-31

where r = radius of unreacted metal in fuel rod r = original radius of fuel rod t = time K = rate law constant (0.3937 cm2 f,,c) AE = activation energy (h5,500 cal / mole) R = gas cor;stant (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 does 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 ga==a sources is assumed to be generated witnin the fuel according to the preaccident porer distribution and infinite irradia-tion. Within each of the regions there is a single fuel node and a single 's 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 clad node to the fluid sink node by specifying the time-dependent surface coefficient. The surface heat transfer coefficient input data are determineu from calculations which are based on flow and water inventory as furnished from the blowdown and the core flooding tank performance 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. Average clad temperature of each region. Per cent metal-water reaction in each region. gg 1h-32

Time for the clad of each region to reach the metal-vater thresh-old, the beginning and end of meltin6, and the slump temperature. Heat transferred to the reactor building from the core. Heat generation by hydrogen and oxygen recombination. Total zirconium-vater 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 capatilities are described in Reference 12. With minor modifications this program was adapted for use on the B&W Philco-2000 co=puter.

In this model, the reactor building is divided into two regfons: the atmosphere (water vapor and air mixture) and the sump regicn (liquid water). Each region is considered to be well mixed and in thermal equilibrium, but the temperature of each region may be diiferent. The reactor building and its internal structures are subdivided into five segments, as shown in Table lb-h, and treated as slabs with 1-dimensional hest transfer. Each segment, divided into several heat conducting subregions, may act as a heat source or sink. The pro- j gram includes the capability of cooling the reactor building atmo-sphere by air coolers (reactor building emergency cooling units) and spray coolers (reactor building spray system), and cooling the liquid region by sump 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- i gion. After blowdown 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. l l The reactor building calculations are begun by computing steady-state l 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 j the heat-conducting slabs are calculated. Then the pressure and I temperature of the liquid and vapor regions are calculated from the mass, volume, and energy balance equations. s

   -                                                           0337                                I L_.                                        D-

Table lk-h Reactor Building Structural Heat Capacitance Segments Segment Description 1 Reactor Building Walls and Dome 2 Refueling Cavity (Type 30h SS Liner - One Side) 3 Reactor Building Floor b 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. 14.2.2.3.h Accident Analysis

a. Core Floodine Tank Design Base Accident The-36 in. ID, double-ended pipe rupture produces the fastest blow-down and lowest heat removal from the fuel. This case therefore i represents the most stringent emergency core cooling requirements. /

Results from the modified version of FLASH indicate that the core 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. The SLUMP digital computer program, as described in lh.2.2.3.3.b above, is used to evaluate core 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 1/4 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 with the digital computer program CHIC-KIN. This program accounts for changes in flow, pressure, enthalpy, and void fraction. It also computes axially weighted Doppler and moderator coefficients of re-activity for the kinetics calculation. The Doppler coefficient is input as a nonlinear function of fuel temperature, and the moderator void coefficient is input as a function of void fraction. The para =- eters describing the coolant were obtained from the digital computer program FLASH, which in turn used the neutron power output from CHIC-KIN. The: core is assumed to be initially at the ultimate power level ) of 2,568'MWt. M 0338 1h-3h

Figures 1h-35 and 14-36 show the results of the hot leg, lb.1-ft rupture simulation without trip action. Figure lh-35 is the neutron power trace, and Figure lk-36 shows the various components of the re-activity feedback. Figure 1k-37 shows the total energy generated for the spectrum of leak sizes in both the hot and cold legs. Above a 3-ft2 hot leg rup-ture the blevdown forces on the control rod are greater than those resulting from the normal core pressure drop so that control rod in-sertion is not as rapid for the larger break sizes. The dashed por-tion of the curve represents an estimate of degraded control rod in-sertion velocity for the intermediate rupture sizes. The blevdown forces on the control rods during cold leg ruptures do not inhibit rod drop velocity for the complete spectrum of leak sizes. Accord-ingly, the data presented for the spectrum of cold leg ruptures are based upon reactor trip characteristics. The results of this study have been used for determination of hot spot clad temperatures for the loss-of-coolant accident spectrum analysis presented in the fol-loving pages under 1k.2.2.3.4.b. The transient core flow from the FLASH analysis of the 36 in. ID, double-ended rupture was used to determine the core cooling mechanism used in SLUMP. The very high flow rates during the initial blowdown period provide nucleate boiling conditions. However, the time for Departure from Nucleate Boiling (DNB), especially for the hot regions, is extremely difficult to determine. Therefore, a conservative ap-proach was adopted by assuming DNB at 0.25 sec. Nucleate boiling surface coefficients at high flow rates may exceed 50,000 Btu /hr-ft 2 -F. A nucleate boiling surface coefficient of 25,000 Btu /hr-ft 2 - F was used in the analysis. However, the series heat transfer fro = the clad node to the fluid sink is limited to 6,500 Btu /hr-ft -F by 2 the relatively low conductance of the clad. After DNB the surface heat transfer was calculated using the flow provided by FLASH results and Quinn's modified version of the Sieder-Tate (13) correlation:

                                                               .,0.8   0.lh h                                             1-TPF  = 0.023        (NRe)    (u Pr)l/3 1 X

h TPF = two-phase film heat transfer coefficient, Btu /hr-ft2-F k = fluid conductivity, Btu /hr-ft -F D = hydraulic diameter, ft II = eyn lds number Re Np ,= Prandtl number X = quality p = density 0339 s p = viscosity 14-35

subscript B = " Bulk" subscript F = " Film" subscript W = " Wall" With this correlation, bulk steam properties are used in the basic form, and the last two bracketed terms are modifiers which correct for quality and different conditions at the wall. Figure 1h-38 shows the core flow vs time for the 1h.1 ft2 leak as calculated by FLASH. Figure lh-39 shows the clad surface heat transfer coefficient versus time based on the flow of Figure lh-38 and the modified Sieder-Tate equation. The straight line in Figure 1h-39 indicates the surface heat transfer values which were used in SLUMP, 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 stea= quality and temperature corresponding to the most conservative as-sumptions were used. A sensitivity analysis was made for maximum coefficients in SLUMP ranging from h00 to 2,000 Btu /hr-ft2 -F initially and decreasing to zero at the end of blevdown. Results are discussed below. After blowdown no core cooling is assumed until after core recover-ing starts. When the water level reaches the core bottom and starts to rise up on the core, the submerged portion vill be cooled by pool boiling, and any steam thus produced will 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. At the point of initial contact of cool water against hot cladding the heat flux and temperature differences will be such that film boiling 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 Etu/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 core flooding tanks and the reactor vessel (12 in, and lh in. ID). Figure 14-h0 shows the reactor vessel water 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 1h-36 0 W

pressure and low pressure injection beginning at 25 see with emer-gency power available. Similar curves for h00 psig and 1,000 psig core flooding tanks are shown in Figure lk kl. Figure lh h2 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 time. The quench time for a given percentile is taken as that time when 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 some lower point in the core provides an inherent conservatism in the data as plotted. The axial peaking factor profile for the beginning of core life was used as it imposes the most stringent requirements on the core flooding tank design. Peak temperatures for the core flooding systems described above are also shown on Figure 14-h2. These curves demonstrate that all of the systems presented are capable of keeping the peak temperature at the hot spot more than 1,000 F below the melting temperature of the clad. The amount of circonium-water reaction which occurs for each of these core flooding systems is shown in Table 14-5 While this preliminary analysis indicates some difference in the performance of the systems, it is not considered to be a significant difference since the analysis was performed without considering the effects of condensation by the core flooding coolant or of pcssible bypass to the leak of part of the coolant. The preliminary core flooding tank design selected is for a 600 psi charge pressure, 9h0 ft3 of water, h70 ft3 of nitrogen, and a lh-in. supply line. The performance of this system is limiting core tem-peratures is approximately in the center of the range for the sys-tems described. The parameters selected for the final system design will be based on the results of core melting analyses to be conducted as part of the final design of the reactor. For this 600 psi -harge pressure, Figure lk-h2 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 would 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 1k-5 Core Flooding Tank Performance Data Line Nitrogen Total Metal Size, Volume, Water 3eaction, Pressure in.  % of Total  % h00 14 33 .022 400 lh 50 .009 600 lh 33 .005 600 lh 50 .002 600 12 33 .022 600 12 50 .010 1,000 12 33 .003 1,000 12 50 20 S lL-37 0341

Additional analysis was performed to evaluate the sensitivity of the maximum clad temperature to three important thermal parameters. All cases discussed below have in common the following parameters: Leak size: 14.1 ft2 Time of DNB: 0.25 see Time at ultimate power: 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 lb k3 shows the clad maximum te=perature sensitivity to the initial surface heat transfer coefficient after the 0.25 see nucleate boiling period. The coefficient is linearly decreased to zero at 9.5 sec. Zero cooling is maintained until quenching is initiated with a clad surface coefficient of 20 Btu /hr-ft 2 -F. Previous discussion in-dicated justification for assuming 1,000 Btu /hr-ft2 -F for the clad sur-face at 0.25 sec. Figure 14-h3 shows that a value of 1,000 is not on s the most sensitive part of the curve and a 20 per cent decrease in h / vould only result in increasing the peak clad temperature 120 F. The assumption that DUB occurs at 0.25 see is quite conservative. 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-kh shows the effect of variation of time to reach a DUB. 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 lh h5 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-ft 2-F at 0.25 sec, decreasing to zero at 9.5 sec. Heat removal is then zero until the effect of injection cooling is simulated. Figure lk-h5 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-onds resulting from a void shutdown using the maximum positive mod-erator temperature coefficient of +1.0 x 10-4 (Ak/k)/F. The effect of variation of the integrated power on hot spot clad temperature is shown in Figure lk-h6. The resultant integrated power before a void shutdown occurs could increase to 3.h full-power seconds before the hot spot clad temperature would reach 2,300 F, the temperature at which 1.0 per cent 2r-water reaction occurs. m oa42 4 - 1h-38 /

An h value of 15 stops the fast temperature excursion and allows only a low rate of increase thereafter. Since the continuously in-creasing depth of coverage provided by the flooding tanks and the pumped flow injection systems provide additional cooling capability with time, an initial cooling value as lov as 15 is probably adequate. An h value of 20 provides immediate quenching action and a slow cool-ing rate thereafter. An h value of 100 provides very fast cooling. Even though the 100 is a realistic value for film boiling in a pool - the probable mode for the submerged portion of the core - a more conservative value of 20 has been used as the reference for evaluating core flooding tank performance. Figure 14-hi shows hot spot clad temperature transients for a range of pool fluid sink temperatures. Parameters for heat transfer prior to 18 see are the same as discussed in the preceding paragraph. At 18 see a surface coefficient of 20 Btu /hr-ft 2-F was applied with sink temperatures as indicated. All results reported herein previ-cusly have had a sink temperature of 280 F during the quenching pe-riod. Prior to quenching, the sink temperature in all cases is based on the transient fluid pressure which results from the FLASH analysis. Figure 1k-h7 shows that any sink temperature below ap-proximately 500 F is adequate for holding or reducing the clad ten-perature which existed at 18 sec. The core flooding tanks will pro-vide a high flow 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 contact with 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 that the preliminary design of the core flooding system vill provide for covering approximately 80 per cent of the 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 terminated 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,h00 F below the clad selt-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 formation between dissimilar core ma-terials exists. Considering the general area of eutectic formation in the entire core and reactor vessel internals, the following dis-similar metals are present, with major elements being in the approxi-mate proportions shown. . \ - 14-39

Tyre 30h Stainless Steel 19 per cent Chromium 10 per cent Nickel Balance Iron Control Rod 80 per cent Silver 15 per cent Indium 5 per cent cadmium Zircaloy-h 98 per cent Zirconium 1-3/h per cent Tin UO2 All these elements have relatively high melting points, i.e., greater than 2,700 F, except those for silver, cadmium, and indiu= which, in the case of indium, is as lov as approximately 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 Zircalcy-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 h per cent of the cladding would ever exceed the zirconium-iron eutectic point. Since the spacers are located at 21-in. intervals along the assembly and each grid has a very small con-tact area, only a fraction of the h per cent would be in contact with stainless steel. The approximate time period that the h 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 contact is that of a zirconium guide tube with the stainless steel cladding of the control rod. Following blevdcun, heat can be generated in the control rods by absorption of gamma rays. Beta decay heat will be deposited in the fuel rods where generated. Since gamma 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 vill be less than one-half the rates in the fuel. As a result, the peak heat genera-tion rate in control rods adjacent to hot spot fuel vould not exceed _, lk-h0 03M...

an estimated one-half times the rate in these fuel rods which have a 31 Power ratio. The contribution from radiant heat transfer from higher powered fuel rods would be relatively small. The analysis of core melting shows that, with core flooding tanks, fuel rods with a 1.5 power ratio will not exceed 1,500 F. This is well below the eu-tetic 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 (2,270 ppm) to hold the reactor suberitical at reduced temper-atures. The most 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 keff of 0 99 (see Table 3-6, soluble Baron Levels and Worth.) The concentration existing in the reactor building sump after a loss-of-coolant accident from operating power at the begin-ning of core life is 2,174 ppm boren. This concentration represents a boron margin of 354 ppm above the suberiticality design value mar-gin of 1 per cent.

b. Core Cooling Analysis for Spectrum of Leak Sizes 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 and is reported according to the following grouping: (1) hot leg ruptures, (2) cold leg ruptures (3) injection line failures, and (4) injection system capability.

(1) Hot Leg Ruptures In 14.2.2.3.4a 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 accident. This was therefore the basis for design of the core flooding equipment. A decrease in the rupture size assumed results in decreased maximum clad temperature during a loss-of-coolant accident. l Cor2 cooling evaluations have been performed for a spectrum of four additional rupture sizes using the same basic calculational i technique and assumptions as for the large rupture case. These rupture sizes are 8.5, 3.0, 1.0, and 0.4 ft2 . The reactor coolant system pressure-time history for these rupture sizes is shown in Figure 14-49. The reactor vessel water volume as a function of time after the rupture for the various rupture sizes is shown in Figure 14-50. These water volume curves were generated utilizing the flow available from core flooding tanks, one high pressure injection pump, and one low pressure injection pump. The pumping sys-tems were assumed to have a combined capacity of at least 3,500 gpm with the high pressure pu=p running on emergency power <s within 25 see after the rupture, and the low pressure pump 0345 14-41 2-8-68

delivering 3,000 gpm when the pressure ha.s decayed to 100 psi, or at 25 see, whichever occurs later. Figure 14-51 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 temperaturas which are considerably lower than those resulting from the larger sizes. The results of this study are shown in the following Table 14-6. Table 14-6 Tabulation of Loss-of-Coolant Accident Characteristics for Spectrum of Hot Leg Rupture Sizes Rupture Min. Water Level Below Hot Spot Sizg, Full-Power Bottom of Core, Max. Temp., ft Seconds ft F 14.1 2.1 -6.8 1,950 8.5 3.4 -5.2 1,916 30 -2.2 1,235 1.0 1.5 ((*)) +4.7 1,075 0.4 1.5 15 (**) +12.0 1,015 (*) Blowdown forces on control rods are equal to, or less than, - nor=al pressure drop, and control rods will insert with normal velocities. These values are for trip shutdown i rather than for a void shutdown, but include void reactivity effects. (2) Cold Leg Ruotures A simihr analysis of a spectrum of rupture sizes hs.s been made for tha 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 312,3, The reactor coolant system average pressure for this spectrum of rupture sizes as a function of time is shown in Figure 14-52. The water level as a function of time is shown on Figure 14-53. The water level calculation has been based upon 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 14-54. The results of this analysis are shown in the following Table 14-7 Nd6 . ); 14-42 2-8-68 Amendment No. 1

1 Table 14-7 Tabulation of Loss-of-Coolant Accident Characteristics for Spectrum of Cold Leg Rupture Sites Rupture Min. Water Level Below Hot Spot Size Full-Power (*) Bottom of Core, Max. Temp., ft2 Seconds ft F 8.5 -6.7 1,785 3.0 0.4 ((*) -4.8 1,575 1.0 1.0 (**) 1.8 ) +3.6 1,250 0.4 1.3 (*) + 7. 0 1,090 (*) Blowdown forces on control rods are equal to, or less than, nor=al pressure drop, and control rods will insert with nor-mal velocity. These values are for trip shutdown rather than void shutdown, but include reactivity effects. (3) Evaluation of Emergency Coolant Injection TMne Failure 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 vessel was assumed to fail adjacent to reactor vessel and be-fore the first check valve. (See Figure 6-1.) Thic pipe has aninternaldigeterof11.5in.,andtheresultantrupture area is 0.72 ft . Interpolation of avilable 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 will never be uncovered, and peak cladding temperatures will be slight 1,y less than that shown in Figure 14-54 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 3000 gym of water to the reactor vessel and provide coolant to keep the core cooled. The combined capacity of the two high pressure pumps is 1000 gpm which is in excess of the boiloff rate (680 gpm) due to decay heat immediately after blowdown. With a single 500 gpm high pressure injection pump the excess water above the core is adequate to prevent the core from being uncovered below the three quarter evaluation and beyond 300 sec. the water level will begin to increase. F s 0347 h 14-43 2-8 Amendment No. 1

The high prassura injection system has two imiep ndant chains of flow to supply borat d coolant to the system, If a rupture of high prassure 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 xr; against this small leak. In the event of a component failure in the second high pressure injection loop, the ruptured flow path can be monitored by the operator and spillage flow can be stopped by isolation of the affected piping. The entire capacity of one pump can then be utilized : o handle the sman rupture and protect the core. (4) Evaluation of Emergency Core Injection System Performance for Various Rupture Sites The loss-of-coolant analysis is based upon the operation of one high pressure injection pump (500 gpm), one low pressure injection pump (3,000 gpm), and the operation of the core flooding tanks. The capability of other combinations of engineered safeguards to provide core protection has been evaluated in a preliminary analysis. This capability is shown on Figure 14-55. In this evaluation the core is considered protected if the combination of emergency cooling systems considered will prevent core anmge which would interfere with further core cooling. The high pressure injection equipment with one pump operating can acco=modate leaks up to approximately 3 in. in diameter. The preliminary analysis upon which this conclusion is based indicates that one pump will prcbably have the capability to protect the core for leaks somewhat larger. A combination of one high pressure pump and one Igv pressure in- ') jection pump will protect the core up to a 0.4-ft leak. This is equivalent to the rupture of a pressurizer surge line. One high pressure injection pump plus two low pressuge injection pumps can protect the core up to leak sizes of 3.0 ft . This is considerably in excess of any of the piping connecting to the reactor coolant system. High pressure injection, plus the core 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. Thecorefloodingtanksandonelowpressureinjectionpumpcan protect the core from about a 3-in. leak up to the 14.1-ft leak. Figure 14-55 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 slow. For inte mediate leak sizes, either the core flooding tanks or low pressure injection protects the core following the loss-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 immediate protection and can protect the core for several minutes following the rupture. Due to their limited volume, they must be supplemented by the high flow from the low pressure injection pumps within several minutes following the leak in order to prevent the core from again becoming uncovered as a result of boiling off the core flooding tank coolant. 0348 _) 14-44 2-8-68 s Amendment No. 1

This cvaluation of em:rgency cora cooling capability demonstrates that the core is protected for the entire spectru= of leak sizes in both hot and cold leg piping.

c. Reactor Building Design Base Accident 2 2 Arangeofgeaksizesbetween0.4ft and 14.1 ft has been evaluated.

The 14.1-ft leak is equivalent to a double-ended rupture of the 36-in. ID reactor outlet piping. The reactor operating conditions used in this analysis are listed in Table 14-8. During blowdown mass and energy releases to the reactor building are calculated by FIASH. Figure 14-48 is a plot of mass released to reactor building and Figure 14-49 is a plot of reactor coolant average pressure, each calcuhted by FIASH for the spectrum of hot leg ruptures. Following blowdown a 20-region SLUMP model was used to simuhte the core thermal transient. This simulation includes fuel heat generation, metal-water reaction, and quenching when the injection water provided cooling by contact with the core. The basis for this analysis is that only the =akeup and purification system and the decay heat removal system are working to provide core cooling. It was assumed that the makeup and purification system (high pressure injection) had one of the three pu=ps available for operation and that the decay heat removal system (low pressure in-jection) had both of the two pumps available for operation. These systems are assumed to operate on emergency power and can be in oper-ation to deliver a total injection flow of 6,500 gpm within 25 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 will provide a lower energy release rate and peak reactor building pressures than those resulting from the 6,500 gym flow. As any given segnent reached 4,800 F it was assumed to drop into water below the core and release all heat down to a datum of 280 F. Also, it was assumed that 10 per cent additional zirconium-water re-action occurred. When the water covered approximately 25 per cent of the core, the surface heat transfer coefficieng from all the core clad to the water was assumed to be 100 Btu /hr-ft'-F. The determination of water level was based on injection flow and included the effects of boiloff. Assuming a pool boiling coefficient of 100 for the whole core when only 1/4 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 blowdown pericd a sur-face heat transfer coefficient of 1 # blowdown this coefficient was change,000 Btu /hr-ft gto100 Btu -F wgs used. After

                                                               /hr-ft-Fforthe metal below the leak and 5 Btu /hr-ft -F above the leak. The coohnt s

sink temperature was provided by FIASH for the blowdown period and assumed to be 280 F thereafter. The internal heat transfer of the metal was based on a multilayer finite difference model. The whole process of reactor cochnt system metal heat transfer was simulated with a digital computer program. O.3_pg. l 14-45 2-8-68

All heat transferred from the core and the reactor coolant system metal was assumed to generate steam without taking credit for the subcooled condition of the injection water (except for that portion which was boiled off) until the reactor vessel was filled to the leak 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-vater 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 blevdown than did the cold line 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 reactor building pressure analysis program, CONTEMPT. Table 1h-8 Reactor Oterating Conditions for Evaluation

                                                                              )

Parameter Value Reactor Coolant System Pressure, psig 2,185 Reactor Coolant Average Temperature, F 584 Reactor Power. Level (ultimate), MWt 2,568 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 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 lb-9. _) g 0350 1h h6 -

1 Tsble 14-9 Reactor Building Structure Data for Analysis of Time-Dependent Reactor Building Pressure Parameter Value Reactor Building Free Volume, ft 3 1,900,000 Expcsed Liner Plate Surface, ft 2 85,340 ~ Mass, lb 864,900 Dome and Wall Liner Thickness, in. 0.25 Refueling Cavity Liner Thickness, in. 0.1875 Reactor Building Concrete Enclosure Consisting of a 3-ft, 6-in. Thick Dome and 3-ft, 9-in. Thick Walls and a 2-ft Thick floor Wall and Dome Surface, ft 2 78,570 Wall and Dome Mass, lb 40,888,000 Exposed Floor Surface, ft 2 8,200 Exposed Floor Mass, lb 2,820,000 Structural and Miscellaneous Steel Exposed to Reactor Building Atmosphere Surface, ft2 22,200 Mass, lb 844,000 Internal Concrete Surface, ft2 81,230 Mass, lb 16,192,000 Refueling Cavity Concrete Surface, ft2 13,540 Mass, lb 4,008,000 O.351 14-47

Heat transfer from the reactor building atmosphere to the steel liner was calculated using a condensing coefficient of 620 Btu /hr-ft d-F until a total heat input of 110 Btu /ft2 had been achieved. There-after, a condensing coefficient of 40 Btu /hr-ft -F 2 was used. For heat transfer from the reactor building atmosphere to the con-crete, a condensing coefficient of 40 Btu /hr-ft2-F was used. For heattransferfgomthesumpwatertotheconcreteflooracoefficient 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 0.931 in. vas used. Condensing coef-2 vere used. ficients of 620 and 40 Etu/hr-ft -F Folleving a loss-of-coolant accident, the reactor building is cooled by three reactor building emergency cooling units and a spray systgm. Each' cooling arrangement has a heat removal capability of 2h0 x 100 Btu /hr at a vapor temperature of 286 F. Two cooling units plus 1,50) gpm sprays, or 3,000 gpm sprays, provide cooling that is at least equivalent to the three reactor building e=ergency cooling units. Each system is designed so that it alone can protect the reactor building against overpressure. Each system was assumed to operate on emergency power and was delayed until 35 see after the rupture occurred. Figure 14-56 shcws the reactor building pressure for complete sever- ) ance of a 36-in.'ID reactor coolant system pipe (14.1-ft2 rupture area) with 6,600 gpm of borated water injection into the reactor cool-ant system beginning 25 see after the rupture. Reactor building cool-ing is provided by three emergency cooling units. The peak pressure resulting from this accident occurs 161 see after the rupture at a value of 55.9 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 demonstrate the effectiveness of this system. Initially coolant for the building sprays and for injection to the core is pu= ped from the borated water storage tank. When water from the borated water stor-age tank is depleted, the water collected in the reactor building sump is recirculated through the reactor building sprays and th:.ough the decay heat removal coolers to supply the low pressure injection water. ature. The result is an increased injection and spray water tenper-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 re-actor building atmosphere. The requirements for ecoling the water recirculated from the reactor building sump to the reactor building spray syste= are set by the de-sign basis of this system. The design basis is to maintain the post-accident reactor building pressure below the design value. This x ' criterion can be met by spraying the su=p water directly into the re- -- J actor building atmosphere without additional cooling, other than that provided by the decay heat removal system. ()[l[j2l 14-48 REVISED, 2-8-68

9 e The water temperature in the reactor building sump during the recir-culation phase of a loss-of-coolant accident is maintained below saturation te=perature 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 trans-fer surface of these coolers is set by the normal operating con-ditions under the decay heat removal coolers. These coolers re-duce 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 nor=al operating conditions under the decay heat removal operation mode. The cooling capat-ility of this mode of operation will maintain the reactor coolant at 140 F or less at 20 hours after extended rated power operation and is in excess of that required under accident conditions. The per-for=ance of these coolers at various inlet temperatures is shown in Figure 6-4. Figure 14-57 shcws that the reactor building pressure decays to less than 5 psig in 24 hours. For comparison purposes and to show that the effect of spraying cooler water into the reactor building is a ll, a second curve is presented on Figure 14-p? which is based upon a spray recirculation cooling rate of 100 x 100 Btu /hr (approxi-mately equivalent to two decay heat removal coolers) at a sump tem-perature of 195 F. (This is the temperature of the sump when recir-culr tion to the sprays begins.) Figure 14-58 shows the temperature 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 minor ef-fect on the rate of pressure decay during the first 24 hours. Tests 3 conducted in Ger=any and at OPl1 have demonstrated that the effect-iveness of spray water for iodipe pemoval is not affected by the temperature of the spray water.\ l0 1 Accordingly, it is concluded that no cooling of the recirculated spray water is required for this accident. Figures 14-59 through 14-63 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"my of the input parameters and results for the spectrum analysis are tabulated in Table 14-10. A 3.0-ft2 rupture area results in the highest post-accident reactor building pressure (see Figure 14-60). Figures 14-64 and 14-65 show the reactor building enerEy 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 ef-fectiveness of the reactor building structures and emergency cool-ing units. Figures 14-66and14-67showthereactorbuildiggvaportemperatures i and sump temperatures following 14.1 and 3.0-ft ruptures, i 14-49 03M , l 5-3-68 Supplement No. 3

Tha p:ak reactor building pr;ssura shown in this cvaluation for the spectrum of leak sizes results is 56.8 psig and is the re-sult of a 3.0-ft2 rupture in the reactor outlet piping. The reactor building design pressure is 59 psig and a design margin of about 2 psi exists. The above analyses conservatively assume that the hydrogen liber-ated will burn at the rate formed, and that no core flooding tank operation occurs. The zirconium-water reaction begins at 40 sec and stops at 130 see, by which time the 6,600 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 hydrogen generated. The resultant concentration of hydro-gen (at time of r n h m metal-water reaction rate) in the steam leaving the reactor vessel is 7 2 volume per cent. This concen-tration is 'oelow the flam= ability limit. Further dilution will occur as the steam enters the reactor building, and combustion will not occur, even as the reactor building is depressurized. The effect of core flooding tanks on the reactor building pressure is shown in Figures 14-68 and 14-69 In this analysis the minimum injection flow (6,600 gpm) and the three reactor building emergency cooling units start at 25 seconds. Each core flooding tank con-tains 940 ft3 of water. For the 36-in. ID, double-ended pipe rup-ture two core flooding tanks limit the zirconium-water reaction to 0.063 per cent, and the potential hydrogen energy release is ap- 3 proximately 57,000 Btu. For this case a peak building pressure of 53.1 psig is reached. One core flooding tank will limit the zir-conium-waterreactionto0.45gercent,andthepeakbuildingpres- ) sure to 52.8 psig. The 3.o-ft rupture with two core flooding tanks operating results in a reactor building pressure of 53.3 psis (see Figure 14-69). Criterion 17 of the AEC General Design Criteria (*) requires that the containment (reactor building) and engineered safety features be designed to accom=odate the largest credible energy release in-cluding the effects of credible metal-water reactions. Although the evaluation of the emergency injection systems de=enstrates that only a m il a= cunt of metal-water reaction can occur, the case of no injection flow has been evaluated. This case assumed that, after clowdown, 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-water 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 I reached the melting temperature, it was assumed that the whole re-gion slumped into the bottom of the vessel with the attendant reac-tion of 10 per cent more of the re=aining zirconi'um and with the re-  ! lease to the reactor building of all sensible and latent heat above l 280 F. (*)The criteria as proposed by the AEC in its press release H-252 of Nove=ber 22, 1965

                                                                                 \

v 14-50 -Q$d,- 5-3_68 Supplement No. 3

1 Upon completion of boiloff, heat input to the reactor building was assumed to cease. Figure 14-70 shows a reactor building pressure of 56.7 psig at 220 seconds, the time at which the reactor vessel l boils dry. This peak pressure is below the 59 psig design pres-sure of the reactor building.

     /
   /
 /

S 0355 14-50a 5-3-68 Surgyglemcat NA 3

Table lh-10 Su= mary of Reacter Building Pressure Analysis for Reactor Building E=ergency Cooling (240 x 106 Etu/hr) Rupture Size, ft 2 14.1 8.5 3.0 2.0 1.0 0.4

             ~

Reference Figure No. 14-56 14-59 14-60 14-61 1k-62 1h-63 Time Blowdown Ends, see 15 20 48 68 141 351 , Time Low Pressure In-jection Begins, sec 25 25 39 59 121 321 Fraction of Core Zr-reacted 0.08 0.05 <0.01 zero zero zero Time Zr-reaction Begins, sec 40 50 130 -- -- -- Time Zr-reaction Ends, see 130 130 131 -- -- -- Time to Reach Peak Pressure, see 161 20 40 67 181 261 Peak Building Pressure, l psig 55.9 55.6 56.8 55.1 52.5 h6.0 Vapor. Temperature at Peak Pressure, F- 283 282 28h 282 279 270 Sump Temperature at Peak Pressure, F 236 231 226 220 216 203 Conditions for All Cases

a. 600 gpm high pressure injection

b. 6,000 gpm low pressure injection

c. Reactor hot leg rupture

d. No core flooding

e. No reactor building sprays

f. Three emergency cooling units start 35 see after the rupture.

k 0356 I $

                                           -1k-51                    REVISED, 2-8-68

d. Peactor Building Zirconium Reaction Carability In order to determine the thecretical ultimate zirconiu= reaction capability of the reactor building a series of hypothetical accidents was investigated.

Blcwdown was based cn the lh.1-ft2 leak case. Heat transfer from the core and all~ reactor coolant system metal below the leak height was assumed 50,000 Btuto/hr-f transfer 2-F. toFor a 280 F sink based on a surface coefficient of reactor coolant system metal above the leak height 5 Btu /hr-ft -F 2 was used. Available core heat consisted of the initial stored heat, the equiv-alent'of two full power seconds, decay heat, and metal-water reac-tion heat, which was added at arbitrary linear rates. The total heat transferred from the core and reactor coolant system metal was assumed to produce steam from water initially at the saturated condition. Hydrogen recombination energy was added to the reactor building as cuperheat at the rate of hydrogen production from the zirconium-water reaction. A series of calculations for each of the various cooling capacities was made varying the enrgy input rate, i.e., Zr-H2 O reaction rate. For exampla, 1.per cent per second zirconium-water react 1.173 x 10b Btu /see of metal-water energy and 0.902 x 10gon produces Btu /see hydrogen recc=bination energy. In all cases the energy was input at a linear rate beginning 10 see after the rupture. The emergency s cooling units and spray coolers were started 35 see after the rup-ture. -) The " time to ecmplete reaction" is the time it takes to reach reactor building design pressure (59 psig). The results of this study are presented in Figure 14-71. This amount of allowable zirconium reaction at any time after blowdown depends upon the amount of reactor building cooling in operation. The capa-bility curves show that at approximately 10 sec, when the blowdown pressure peak occurs, the reactor building could accept an instantan-ecus zirconium-water reaction of 2 per cent. This capability increases greatly after the blcudown pressure peak with reactor building cooling equipment in operation.

             .With three' emergency cooling units in operation a 100 per cent reac-tion in 3,100 see will not exceed the design pressure of 59 psig.

With three emergency cooling units and two sprays operating, a 100 per sure.cent reaction in 1,200 seconds will not exceed the design pres-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

 'dtmonstrated that safety injection will prevent clad melting for loss-of-cool-ant. accidents resulting from reactor coolant system ruptures ranging in size from'small leaks to the ecmplete severance of a 36-in. ID. main coolant pipe.
                                                                                               }

g 0357 lh-52 REVISED, 2-8-68 e

I 1 l l l 9 Without clad melting, only the radioactive material in the coolant at the time " of tne accident plus some gap activity is released to the reactor building. The environmental censequences 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. While perforation of fuel cladding vill require some time, it is conservatively as-sumed that all of the fuel rods release their gap activity to the reactor building. Half of the iodine released is assumed to plate out on 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. 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 showed 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 showed that the equation below can be used to calculate the mass transfer co- j efficient when the rate of transfer is controlled by the gas film resistance: DOM 7

                                   /        Idvo11/2 g 31/3 '

kg= dP 2 + 0.6 where 2 kG = gas film mass transfer coefficient, gm/cm -sec-atmos D = diffusivity of iodine in air, em 2/see p = density of air, gs/cm3 My = molecular weight of iodine, gm/gm-mole M = mean molecular weight of the air-iodine mixture in the boundary layer P = partial pressure of air in the gas film, atmos d = drop diameter, em v = relative velocity between the drop and the gu phase, or approximately the teminal velocity of the drop, em/see u = viscosity of the air u Since the mass transfer of iodine is gas-film-controlled, k gis approximately cqual to KG (below), and the foregoing equation can be rewritten in terms of 14-52a OM 5-3-68 Sumlement NR._5)

3 the velocity of deposition, V g:

                                                   /                           s RT          RT EMIo               Idvol 1/2' fu i1/3 V={K3 g      = My M d    3 PI 2 + 0.61    y         g where V = overall velocity of deposition, em/see R = universal gas constant = 82.057, atmos-cm fg_g=_     3     cole T = absolute temperatule, K Kg = overall mass transfer coefficient, gm/cm2-see-atmos Since the maximum possible iodine concentration in the large volume in the re-actor building is less that 10-7 g=/ce,             the partial pressure of air in the gas fib:1, P, can be taken as the total pressure, and the mean molecular weight,
     !( , can be taken as the molecular weight of air', M . If the gas equation is ustd, the +quation may be simplified somewhat by sbstituting MA /RT for p/P, as follows:

V =D 2 + 0.61 Pdvo 1/2l ur )1/3' l 6 d Lp gDog ,

 \
   \

Tho surface area of drops available for iodine absorption can be calculated s from the next equation, which is based on the assumption that all the drops > ara spherical and have the same diameter. S= ndD0 , @0 , M 43 d dv 6 wh re 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, c= H = drop fall height, em v = drop fall velocity or terminal velocity, em/see 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 dI fV,S g= y

                                             ,Lc}

I = -A I s 1h-52b d 5-3-68 Supplement No. 3

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Ve = free volume of reactor building, cm3 or ft3 As = iodine removal time constant, hr -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:

                                         = e-As t where h=fractionofinitialinventoryremaining o

t = spray time, hr When the specific parameters for %e Russellville Nuclear Unit are used: F = 3,000 gym v = 397 cm/sec H = 99 ft Vs = h.84 cm/see V = 1.9 x 106 ft3 6VRFH d = 1,000 microns A S

                                                       =

V dy = 27.9 hr-1 e These iodine removal calculations have conservatively corrected the iodine d.eposition velocity (Vg) 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 Ih.2.2.h.3 in terms of the 2-hour iodine dose at the exclusion distance following an IGA. Although the reactor building leakage rate will decrease as the pressure de-cays, the leakage is assumed to remain constant at the rate of 0.20 per cent p r day for the first 24 hours. Thereafter, since the reactor building will have returned to nearly atmosphere pressure, the rate is assumed to be reduced to 0.10 per cent per day and remain at thf s value for the duration of the ac-cident. l The atmospheric dispersion characteristics of the Plant site are described in 2.3 A breathing rate of 3.h7 x 10-h m3/see is assumed for the 2-hour exposure. For the first24-hour 8 hoursexposure, a breathing

                 , and a rate of 1.Th xrate 10-gf 3.47 x 10-4 m3/see is assumed for the m3/see is assumed for hours. Forthe30-dayexposure,abreathingrateof2.32x10gheremaining16      m3/see is assumed.

Th2 iodine doses to the thyroid per curie inhaled are obtained from the values given in TID-lh8hh: I-131 1.48 x 106 rem per curie I-132 5.35 x lok rem per curie ~ 14-53 - s 5-3-68 s-u

, y I-133 4.0 x 105 rem per curie I-13h k 25xlo rem per curie I-135 1.24 x 105 rem per curie Figure lk-72 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 0.65-mile exclusion distance is 1 5 rem for a 2- 3 hour exposure, 3.h rem for a 2h-hour exposure, and 12.h rem for a 30-day expo - sure. These doses are vell below the guideline values of 10 CFR 100. The di-ract dose from this accident is insignificant since it is several orders of mag-nitude below 10 CFR 100. 14.2.2.3.6 Effects of Reactor Building Purging At times during the normal 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 valves vill 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 blevn 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 build-ing atmosphere vill escape through the purge valves before they close, corre-sponding to a release of 3 equivalent curies of iodine-131. This analysis as-suces unrestricted flow through the purge line for the full, 5-second closing time. No reduction in flow is assumed as the valve closes, and therefore the results are conservative. The release of this iodine results in a total inte- t grated thyroid dose of .58 rem'at the exclusion distance. This dose, when added - l3 to the thyroid dose for a loss-of-coolant accident without purgin6, is well be-lov the 10 CFR 100 guidelines. Therefore, purging operations can be performed during reactor operation. 14.2.2.h Maximum Hypothetical Accident 14.2.2.h.1 Identification of Accident The analyces in the preceding sections have demonstrated that even in the event of a lost-of-coolt.n;. acci.2ent. no significant ccre meltin;; vill occur. However, to demonstrate that the operation of a nuclear pcVer clant at the proposed site does not present any undue hazard tc the general public, a hypothetical accident involvi.cg a gross releace of fission products is evaluated. No mechanism vnere-by such a release occars is postulated since it veuld require a multitude of failures in the engineered safeguards provided to prevent its occurrence. Fis-sion products are assumed to be ra.leesed frcm the core as stated in TID-lh84h, nhmely,100 per cent of the noble gases , 50 per cent of the halogens , and 1 per cent of the solids. Purther, 50 per cent of the iodines released to the reactor building are assumed to plate out. Other parameters, such as meteorological conditions, icdine in-ventory of the-fuel, reactor building leak rate, etc., are the same as those assumed for the loss-of-coolant accident in 14.2.2.3.5 The average iodine in- .

                           +-

d 14-5h 000.2 Supplement No. 3

ventory, in terms of. equivalent curies of iodine-131 available for leakage at different time periods after the accident, is as follows: 0 to 2 hours 29 x 106 curies 0 to 24 hours 23.1 x 106 curies 1 to 30 days 5.2 x 106 curies lk.2.2.4.2 Analysis and Results of Environmental Analysis Figure 1b-73 presents 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. It can be seen that the 2-hour thyroid dose of 63 rem at the exclusien distance 3 of 0.65 miles, and the 30-day thyroid dose of 22.4 rem at the 4-mile low popu-lation zone, are less than the guideline values of 10 CFR 100. The direct dose to the whole body following the accident is shown in Figure lk-

74. 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 4.5 ren, and the 30-day dose at the h-mile low population zone is 0.95 rem. 14.2.2.h.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 chemica'. sprays in the reacter building. The results are shown followinginanTable MHA.1h-11 in terms of the 2-hour iodine dose at the exclusion distance 14.2.2.4.h Effect of Rainout During the Maximum Hypothetical Accident To further evaluate the suitability of the site, the effects of rainout on sur-rounding drinking-water reservoirs (Figure 2-11) following the maximum hypothet-ical accident were analyzed. 24-hour rainfall covering the general Calculations areawere made of each for the resarv ir. case of a continuous The maximum rain-out rate as a function of distance is calculated from (17 2 2 to = 90 *-(y /2oV) max xeoy /2x wh;re ta max' = maximum rainout rate _, curies per sec per m2 g- x = downvind distance, meters 000.3 oy = horizontal dispersion, meters

5-3-68 45 Supplement No. 3

y = crossuind dictance from plume axis, meters Qo =_ release rate, curies per sec e = 2 718 The equation above is conservative, since the results do not consider the vind speed or vertical distribution in the cloud. The vind direction is assumed to remain toward the respective reservoirs listed in Figure 2-11 for the 24-hour l3 period with the plume centerlines uniformly distributed over this section. Rain-out is assumed to occur under Fasquill "D" conditions, which is typical for a rainy day. (14) The average release rate from the reactor building during the 24-hour period following the accident is 0.028 equivalent curies of iodine-131 per see. Using l3 - the foregoing equation, the maximum iodine rainout is calculated by assuming that all of the iodine that has rained out remains in the reservoir and is not affected by runoff. The average number of curies in the reservoir during a one-year period is re-- duced by a factor of 0.0318 due to the natural decay of iodine. Assuming this activity mixes in the reservoir and that a child with a 2-gram thyroid contin-ually drinks '300' mt per day of the contaminated water, the total dose to the thyroid has been calculated using the methods of TID-14844. For an adult with a 20-gram thyroid, a drinking rate of 1,200 mi per day is used. The nearest drinking water intake, on Russellville Reservoir, is approximately 5 miles from the site. At this distance, the total integrated dese to a child's thyroid is ) 0.16 rem and to an sdult thyroid is only 0.064 rem. These doses are well belov l3 the limits of 10 C1R 100. 14.2.2.4.5 Effects of Engineered Safeguards Leakage During the Maximum Hypothetical Accident An additional source of fission product leakage during the maximum hypothetical accident can occur from leakage of the engineered safeguards external to the re-actor building during the recirculation phase for long-term core cooling. A detailed analysis of the potentiel leekage from these systems is presented in 6.3. That analys ts' demon.itrated that the maximum leakage is about 5,000 cc/hr. It is assumei 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 reacter cooling system. Tne 50 per cent escaping from the reactor coolant system is consistent with TID-148hk. Tna nacumption that all the iodine escaping from the reactor coolant system is absorbed by the water in the reactor building is con-servative since much of the iodine released from the fuel vill be plated out on-the building valls. The activity in the recirculation water is equal to 0.037 equivalent' curies of I-131 per cc of water. Since the temperature of water in the_ reactor building sump is less than 200 F vhen recirculation occurs, the ic-dine release is calculated using a gas / liquid partition coefficient of 9 x 10-3,

                       ,                         ,                                           q)t g               14-56         0004 5-3                                                                            Supplement No. 3

LL2akage frem the auxiliary building is caused by exfiltration. The most restrie-tive case for a ground release oc;urs during inversion conditions. It is as-

   .sumed that the building. leaks .at the rate of-100 per cent per day with atmo-spheric dilution occurring in the vake of the building. For this building leak rate and'the inversion c'ondition, ,the iodine vill produce an integrated dose to the' thyroid of 0.02T. rem in 2 hours ' at the 0.65-mile exclusion distance.                 l3 4

i

0005-j'" g y 1 -57 '

68 Supplement No. 3

                                                                                                                                                                                                 .O             u Table 14-11' Sensitivity Analysis Showing the Effect of Parameters on the                                                       3 Two-IIour Iodine Dose at the Exclusion Distance Following an MHA Iodine (l)                    ~ Iodine (2)-

Removal- Removal

                           -Drop             Drop Fall Velocity of'                                            . Time        Iodine II)       Time           Iodine (2)                               ,

Case- . Size, . Velocity, Deposition, Temp, Press., Constant, Doce, . Constant, Dose,

              -No.         microns-            cm/sec                         cm/see        F-          psig    hr-1            rem           hr-1              rem           Remarks
                 -1         1,000                  397:                         4.8h       286           59     14.0             79           27.9               63     Operation,of trie.reac-tor building spray sys-
       ;                                                                                                                                                                tem at maximum building temperature and pressure.

2 1,000 397; 6.44, 212 25 18.6 71- 37.2' 59 Operation of the reac-tor building spray sys-tem after partial cooling, s about 1 hour. F

          $3                1,000,                397                          11 55       100            0     33.3-            60           66.6               54'    Operation of the reac-tor building spray sys-tem after cooling to ambient conditions.

4 200 76 6.87 286 59 568 h7 1035 47 Effect of small drop size. C5 1,000 20,300 31 5 286 59 1.77 300 3.54 17h Drop velocity required C C "... to give 300 rem.

          @ 6' 4,900                            1,025                           3.86       286           59    1 77             300            3.54             174     Drop size required to give 300 rem.

For all cases, reactor building free volume = 1 9 x 106 ft3 , and drop fall height = 99 ft. Notes: (1) Flow rate of sprays = 1,500 gpm.

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

5$ 2 P.

            !.?

14.3 REFERENCES

(1) Watson, L. C., Bancroft, A. R., and Hoelke, C. W., Iodine Containment by Dousing in NPD-11, AECL-1130. (2) Styrikovich, M. A., et al, " Transfer of Iodine from Aqueous Solutions to Saturated Vapor",'SoWet Journal of Atomic Eneray 17, July 196h. (3) Dispersion of Soluble Radioactive Material in Water, CF-58-3-109. (h) International Symposium on Fission Product Release and Transport Under Accident Conditions, Oak Ridge, Tennessee, April 1965 (5) Liimatainen, R. C., et al., Studies of Metal-Water Reactions at High Temperature, ANL-6250. (6) Ackerman , R. , et al. , "High Temperature Vapor Pressure of UO 2

                                                                              "' # "#""1 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-Dimensicnal Two-Group, Space-Time Diffusion Equations Ac-counting for Temperature, Xenon and Control Feedback, WAPD-TM-532, October, 1965. (9) Margolis, S. G. and Redfield, J. A., FLASH: A Program for Digital Simu-lation of the Loss-of-Coolant Accident, WAPD-TM-53h, 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 R?sponse I to a Loss-lf-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 Beyon_d_ the Critical Heat Flux, ASME 66WA/HT-36, November 27, 1966. (14) Griffiths, V., The Removal of Iodine from the Atmosphere by Sprays, 3 AHSB (S) R h5, 1963. (15) Taylor, R. F., " Absorption of Iodine Vapor by Aqueous Solutions", Chem. Eng. Sa., X. No. 1/2, pp 68-80, April 1959. (16) Ranz, W. E. and Marshall, W. R., Chem. Eng. Proeress, M , 141, 173, 1952. k 0007 14-59 4 5-3-68 Supplement No. 3

(17) Culkovski, W. M., Deposition and Washout Calculaticns Based on the General-3 ized Gaussian Plume I4cdel, CEO-599. (18) crystal aiver Unit 3 Nuclear Generating Plant, Preliminary Safety Analysis Report, Supplement No.1, Question 8.8, Docket No 50-302, February 7,1968.

                                                                                           )

0008 J m 4 1k-60 5-3-68 t SuPPl ement No. 3

                                        -   i       < 1 I         -
                                                                         ! i i       ! ! r ! !

i ' ' ' I  ! ' ' 14 0 i i i . i e i ;iiI i l l NEUTRON 30

                                      '                    I^            '       ! I                   '

POWE R ,$ 20 f

                                      ;              j

[I  ; l l' l 10

                                      '           # '                    i    !     !
                                          /            I     -      t -       i     i              !

O M! i I N ' ' ' ' '

                          '                                  i !

20 I i I !

                                            ! !              ! !                             i THERMAL          10 j     ['N                             l POWER,%

0 j i ii!

                          'i                                            t     6 i      ' i            1 200                 l            l            !

FUEL TEMPERATURE 100 , l f Y% ' O i CHANGE,F  ! AVERAGE 10 _- m l! ~ i I I I - CORE MODERATOR  ! ' "

                                              ~#

TEMPERATURE O CHANGE,F i I I I

                                        *- 5 S EC -*

2500 ,'  ! ' ' ' ' l REACTOR 2350 _

                                                         , /

SYSTEM 2200

                                           '~

PRESSURE, PSIA __ i i STARTUP ACCIDENT FROM 10-9 RATED POWER USING A 1.23 AK/K ROD GROUP; \' HIGH PRESSURE REACTOR TRIP IS ACTUATED. 0009 Figure 14-1

I l NEUTRON POWER,% 150 100 50 0 / i 20 I THERMAL 10 / A POWER,% 0  ! N-200 FUEL TEMPERATURE / ' s , CHANGE,F 0 10 AVERAGE 5 CORE MODERATOR , ,,,,,,_ TEMPERATURE O / CHANGE,F

                                    *- 5 S EC-+-

2500 i I ' iI REACTOR 8 '-- 2350 ' SYSTEM 7 l PRESSURE, # ' 2200 ' PSIA STARTUP ACCIDENT FROM 10-9 RATED POWER USING ALL RODS WITH A WORTH OF 10.0% AK/K; 'v HIGH FLUX REACTOR TRIP IS ACTUATED. 0010 Figure 14-2

100 i i iiii g i i iiii i i ig ii i i i  : i i

                                                                                                                                                                                                                        /

80

                                                                                 =e a

y 60 2 3 U 40 High Pressure  :: -- Hi gh Fl ux ---

                                                                                 -*                                                                            Level Trip                  Level Tri
                  >ET                                                                                                                                                        ._A o-m                                                                                                                                                                              /

OH > Mominal s I /

                 - :r x oO                                                                                                          One Control         \                                I       f 2
                 $"c                                                                                                         Rod Group             y       '                       I f          Mominal C                     r--                                                                                                                                                    /~AllRods Q     w y

_. O mr -- f-

                                 -4 T 0    ' ' '                              '     '    ' ' ' ' ' '              '     '  '                    ' ' ' ' ' ' ' '

A o @ 4 6 -5 2 4 6 -4 2 4 6 inm 10 10 10 10 00:o n :o < Rod Withdrawal Rate, (ak/k)/sec >

                 >>m C               -4                                        ";U C             m Co Ch 7             O --i C (D                            > Ch T :D O -1 :o
   &             EC O I          mT C C4             :o

10 l l l llll

                           ;      I l l l llll           l                       l  l l ll lll      l 1 l Illlg 6    -

y - - g 2 E I 10 a I

                *         ~                                                                                         ~

O 8 _ _ e - ~ c 6 All Rods  ! _ 2 , 8 u

                ;                                     SI"9I *

• 2
• 10 Control Rod Z C e _

,C [ 6 EK5; 4

                         ~~

l ~ O o g; _ o m E

• EG 2 -

     =  #=

y=ox O 10 ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' I ' ' ' ' o_ '" E' -6 2 4 6 -5 2 4 6 -4 2 4 6 -3 2 4 6 -2

       &g g             10                    10                      10                      10                   10 g                                           Rod Withdrawal Rate, ( Ak/k)/sec a
'io yo "[ E
            =
     ,      w
~

4: C o a = =o

21 [ M we. [ C y 19 e E I 18 5 j l

                                                   /

[ 16 r

                                                !\         Nominal l                 I 15 0    0.4                        0.8         1. 2         1.6 Trip Delay Time, sec l

I l PEAK THERMAL POWER VERSUS TRIP DELAY TlHE l FOR A STARTUP ACCIDENT USING A l.27. AK/K R0D GROUP AT 5.8 x 10 (aK/K)/SEC FROM 10 RATED POWER Figure 14-5 O oo:i.a , 1 l l _

30 25 a 20 m

!                x- s
;;                            N   %

15 m 3 %_ 2 r 10 Nom.i n al

                              /                           -

4 E 5 0 0 -0.4 -0.8 - 1. 2 -1.6 - 2. 0 Doppler Coefficient, (Ak/k)/F x 10 PEAK THERNAL POWER VERSUS DOPPLER COEFFICIENT FOR A STARTUP ACCIDENT

                                                                            ~

USING A l.27. aK/K ROD GROUP AT 5.8 x 10 (aK/K)/SECFROM10' RATED POWER . Figure 14-6 k 001.4

40 r

                           /

I f

    =                /

20 1 7

               /[/

Mominal 10

             /

0 0.4 0.8 1. 2 1.6 Trip Delay Time, sec I PEAK THERMAL POWER VERSUS TRIP DELAY TIME FOR A STARTUP ACCIDENT USING ALL RODS AT 5.8 x 10" (aX/K)/SECFROM13' RATED POWER F i gu re 14-7 , bw l 001.5 1

I I I I High Flux High Pressure

            \       Level Trip Level Trip w

N J f M

                \                         f\

3 E

                     ,                  :                 ~

E n 5 f 1 \ ' i Nom;nal . 0 -0.4 -0.8 -1.2 -1.6 -2.0 Doppler Coefficient, (ak/k)/F x 10 PEAK THERMAL POWER VERSUS DOPPLER COEFFICIENT FOR A STARTUP ACCIDENT USING ALL RODS AT 5.8 x 10'

                                                                                   ~

(AK/K)/SEC FROM 10 RATED POWER Figure 14-8 9 0016

2600 .2 2550 E f 2500 j [ Safety Valve 2450 %- e- Nomi nal Set Point 2400 2350 0 0.4 0.8 1.2 1.6 2.0 Trip Delay Time, sec PEAK PRESSURE VERSUS TRIP DELAY TlHE FOR A STARTUP ACCIDENT USING ALL RODS AT 5.8 x 10"

                                                                                           ~

(AK/K)/SEC FROH 10 RATED POWER Figure 14-9 5 0017

l 2550 g I 2525 [ Safety Set Valve Point ia 2500 w 3 l

                                                                              = 2475                     g                                                                    l E                             \

g 2450 Nomi n al -+ \- I  % . 2425 l 2400 0 2 4 6 8 10 Tripped Rod Worth, % ak/k l PEAK PRESSURE VERSUS TRIPPED ROD WORTH FOR A STARTUP ACCIDENT USING ALL RODS AT 5.8 x 10" ( AK/K)/SEC FROM 10' RATED POWER Figure 14-10 6 w t8 1 l l

t' 2550 l I l High Flux i I. I 2525 - High Pressure LevelTrip-[ Level Trip 2500

       ;                                                             S fety alve h                                                   'l         Set Point _
                                                                                        ~

2475 - 2 ' t Nominal 1 \

    .x g    2450 N     I a.

2425 l 2400 ' 0 '

                      -0.4      -0.8                                                  2
                                                - 1. 2                                             \
                                                                - 1. 6         -2.0 Doppler Coef ficient. (ak/k)/F x 10 PEAK PRESSURE VERSUS DOPPLER COEF FOR 5.8 x 10' A STARTUP ACCIDENT USING ALL RO (oK/K)/SEC FROM 10' RATED        POWER Figure 14-11)cjig

2500 1 2490 E 2480 e a Nominal #

a. 2470 p -

% / E s # 2460 e 2450 '

       -4    0          +4          +8           + 12        + 16 Moderator Coef ficient, (ak/k)/F x 10 PEAK PRESSURE VERSUS MODERATOR COEFFICIENT FOR A STARTUP ACCir>ENT USING ALL RODS AT 5.8 x 10     (aK/K)/SEC FROM 10' RATED POWER Figure 14-12 6                                                    0020

I ! *

                                      ; ! i                :            6      i !
                       . i 1          , .      !

i !  :

                                 '                                                    i 4,

NEUTRON l20 _; 5

                                                                            ' l 3              !

POWER, % , , h- t i i ; i j {l 80  ; , , i j j i,!  ; ; ; , i i i e i i i i i i 4i i i i i i . i *iit I !\ 1 l 0

                       ' '                     ' ' ' I I i                 I   I' i      !   i            -
                                                               !    ,   j      ! .
                                 '                    '                    ' I '

I f i i  ! i  ! I l 1 THERMAL P OW E R , *, 120 l!!,' , l!  ! I I~ _l

                                                                                   \

80 .__ j i i ' ! ' l *

                                 . i i ! i i                 ! I t                 i N 40 : i          ! !t !!!I                     ' ! '

3 i i

                                     !i*1!' .           
                                                                            ' !           I 0

s i

                       ! i iy   '

t 1i!

                                                 !        I
                                                                    > < i iii i   e FUEL             200    !I' -                  ! - !i'!                     ! !i TEMPERATURE O                                              - -:
                                                                                   ^<

CHANGE, F  ! '!  ! l i ; ' I

               -200    l i i i ,

i , ; i i i ;i i; i- i-N 3-

               -400    I'!'.
                       ! i
                                          '! i i i i i

i l I i ii i l i i  ! ! i i i i . i i 5 , , - AVERAGE CORE MODERATOR 0 j TEMPERATURE

                          , ,       ,         i      j 4
                  -5      I I                 I I I                                              \

CHANGE, F ' ' i

                                          !!         !        i
                -10                       ! '        !        !!
                                          !          I        I I I

6+- 5 SECM t i 2500 j !!I !I lll I Lf- I- f 2350 d -- I I MN ' REACTOR SYSTEM 2200 l;;

                                    ; , ; j j M ' Il
                                                                                        \{;\

PRESSURE. PSIA 2050 '!l'!ll' .

                                                                                             \

1900-t-lll,' l R0D WITHDRAWAL ACCIDENT FROM RATED POWER USING A l.2G AK/K R0D GROUP AT 5.8 x 10-5 (AK/K)/SEC; HIGH PRESSURE h REACTOR TRIP IS ACTUATED. F i g u r e 114- 13 OiN1

2450 i e i i 1 1 I I i I i 1 1 I I I I I l l l 1

                               ~

High Pressure u  ; High Flux n Level Trip

                                                                          -s Level Trip Lj                     2400 3

Nominal

                                                                       /                                               -
                     .2 E  2350 ri ri o    *.

o o m =

om> a I x m o '

          > m :D     =   2300
          -1 o m     a m ocn      E o      cn EC o-m                _
                                                                                                                      ~

o -4 m x1 mo< mmm 2250

             > ::o E cn
              >c r cn             -
              > :o oo oo          2200
                                 '   ' ' '          '     '     '   ' I ' I I                 !   I   I   I I I I I og                  4     6     8      ~
                                                                                                                           ~

m- IO 10 10 z -1 ri -i m Rod Withdrawal Rate, (ak/k)/ sec o O :o C

E-(n >

F C-ee m N N

2500 1 2475

  .?
   $. 2450                                         2 [                       ,
    .                                          f                              1 2

2 2425 g 2 / l l

   ?

a. j / \l is 2400 g e I I I Nominal l 2375 2350 1 0 0.4 0.8 1.2 1.6 l Trip Delay Time, sec l I i l PEAK PRESSURE VERSUS TRIP DELAY TlHE FOR A ROD WITHDRAWAL ACCIDENT FROM RATED POWER USING A 1.2f. AK/K ROD GROUP: HIGH PRESSURE I REACTOR TRIP 1S ACTUATED. l l i Figure 14-15 l !- l 0023 1

i 1 2425 , , ,

  ,         High Flux   :     :   High Pressure Trip
 'a  2400   Trip a

[ 5 2375

                          /                      Nominal v>
                                                                                  \

• 2350 t

i

                     /

2325 E 2300 0 -0.4 -0.8 -1.2 -1.6 -2.0 Doppler Coefficient. (ak/k)/F x 10 l l I l I i l PEAK PRESSURE VERSUS DOPPLER COEFFICIENT I FOR A ROD WITHDRAWAL ACCIDENT FROM RATED POWER USING A 1.2f. AK/K ROD GROUP Figure 14-16 l 0024 l l

i20 h Marimus Neutron Power (l177.) 100 7 a J Maximum Thermal I g Power ,

                                        \[/
                              /

80 /

                  /

70

               /

[ 60 0 20 40 60 80 100 initial Power, f, of Rated MAXIMUM NEUTROM AMD THERNAL POWER FOR AN ALL-ROD WITHDRAWAL ACCIDENT FROM VARI'0US IMITIAL POWER LEVELS Figure.14-17 oozs

i i e i .

             -                                                              r       _

M

             - Maximum Temp. in Hottest Fuel                                       -

l

             ~                                                                     -

Rod (HotSpot)

   ~         _                                                                    _

o u 30 A l w _ f _

c' _ _

3 - - g - _

     ,E g       _                                                                    _

g _

     .1  20 1       -                                                                    _

_ Max. Temp. in Average Fuel Rod

             ~
                                                                            /

y # _ 10 ' ' ' ' ' O 20  % 60 80 ICO Ini ti al Power, y, PEAK FUEL TEMPERATURE IN AVERAGE ROD AND HOT SPOT FOR AM ALL-ROD WITHDRAWAL ACCIDENT FROM VARIOUS INITIAL POWER LEVELS Figure 14- 18 9 00?6

100 h 80 I = 70,000 1 b-f t

           \

m N t 60 - a N a N t N 4 3 i l l 20 0 0 4 8 12 16 Time, sec l 1 PER CENT REACTOR COOLANT FLOW AS A FUNCTION OF TlHE AFTER LOSS OF PUMP POWER Figure 14-19 0027

l.7 l l l 'l Trip Set Point Maximum l _ (107.5%) Overpower

                                                                                                        ~
                                     ~

s wI

       '*0 I%                                                   Minimum DNS Ratio in Hot Channel at i147.

Power Steady State ( l.38) " o.

  ~               N                          I e;    i.5               %    '
                                                                                                     \

l e

- o.                               N        -

l I . 70.000 l b- f t 2 } g' Rated Power with a4 Overshoot and Deadband g3 g, g g ' ( 102%) , a l j 5 ai N .e 5 l i

                                                                                                     \

.5 t xo l I

1. 3 I

I I l.2 I 100 10 2 10 4 106 108 110 112 114 l 1 Overpower at which Coastdown Begins. ', 1 l WIMINUM DN8R wHICH OCCURS DURING THE COASTD0wM FOR VARIOUS INITIAL POWER LEVELS Figure 14-20 9 I 0026

                                                                                \

l 0.06 l i 0.05

         \                                                                      l
           \                                                                    l
             \\

0.04

s O N i

g 0.03

                     \

E \ 3 0.02 N (

                                              \   w
                                                       \     A 0.01 0

0 100 200 300 400 500 600 Time after Break. sec RfACTOR SYSTEM COOLING RATE FOR STEAM LINE BREAK OF 4 IN. Figure 14-21 0029

l BOL Parameters a ( Ak/ k)/ F 0 . .l.14 x to 35 - 6.0 x 10 a (ak/k)/ F M r = 0.3 sec Delay

f. ' = 5. 47 x 10" sec I I I 2 30 i i i EOL Parameters E0L

                                    ~

a . - 1. 36 x 10 (Ak/k)/F o . Assume Zero 5 r . 0.3 sec Delay A. ' . 2.75 x 10 E 5 20 0 w f O 15

                                    /

10 Nominal BOL Case 5 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 Control Rod Worth. %Ak/k PER CENT CORE EXPERIENCING OMB AS A FUNCTION OF EJECTED CONTROL ROD WORTH AT ULTlWATE POWER Figure 14-22

                                                                                      ~

0030

l I l 1 i

2. 5 i i 80L Parameters OL a . -1.14 x 10" (ak/k)/F

2. 0

            *M   = 6.0 x 10' (ak/k)/F
              , = 0.3 see Delay E * . 5.47 x 10          sec 5                        EOL Parameters g  1.5
                                 -5

• /

J . D . - 1. 36 x 10 om . Assume Zero

c .M d _

r . 0.3 see Delay 1.0 '

                                -5 L * . 2. 75 x 10         sec 0.5 BOL
                                        - hominal Case 0                                              .

0.1 0. 2 0.3 0.4 0.5 0.6 0.7 Control Rod Worth. 7.ak/k 2R-H 0 REACTION AS A FUNCTION OF EJECTED CONTROL R00 WORTH AT ULTIMATE POWER Figure 14-23 8 0031

6

           ._             EOL Parameters                                                      f 4

_ . , . .i.36 . i0 (am/h)/F /

           -e     . Assume Zero

_ r . 0.3 sec De er l' 2.75 10 see Ultimate SOL Para.eters E0L 10

. , . .i.14 = 105(a6/6)/r '

5 /

          -. . 6.0          10 (an/=)/F        -            '                         /

6 M / f

          - r . 0.3 sec Delar                           /                       /

4 t' = 5.47 a 10 sec

          -                                 /                       /

2 / s t / / Ultimate e / Nominal / '*** E 10 Case 80L i

                                                                                 /

s' r 6 / 5 4 /

                                                                    /
                                                               /

Nominal 2 Case 10 Ultimate Power. 10'

                                /
                                  /                   BOL
                             /
                          /

6

                 /                                                                                      -

4 2 10 0.3 0.2 0.3 0.4 0.5 0.6 0.7 0.3 Control Rod w o rth. 'ak/h REACTOR NEUTROM P0wER VARIATION WITH EJECTED CONTROL RCD WORTH Figure 14 24 0032

140 / Ultimate Power. E0L 120

                                                                            /

If

                                                                   )            Ul timate Power.

BOL 100 Nominal Case e et

  .                BOL Parameters 6

_p

,I 0 * '               #

80

           . . 6.0 m 10         (ak/h)/F

]. r . 0.3 sec Delay I l' . 5.47 10 sec 60 ECL Parame r,s a .,. -1.36 = 10 (ak/h)/F

           . . Assume Zero 40
            , . 0.3 sec Delay
             *               -5                                                  -9 g . 2.75 s 10         sec                                          10     Ultimate Po-er.

80L 20 m j Nominal l Case 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 Control Roo worth. *ak / h REACTOR THERWAL POWER AS A FUNCTION OF EJECTED CONTROL R00 wCRTH Figure 14-25 03:3

60 i i e i i , BOL Parameters l l a,. -i.14 x 10 -5 (ak/k)/F _ug - 6.0 x 10 (ak/k)/F / Ultimate Power. 50 EOL y e = 0.3 see Delay b ~1* . 5. 47 x 10 sec g j EOL Parameters "O ' a,. -i.36 x 10 -5 (ak/k)/F  ! n

        -u         Assume Zero                                          - Ul t imate Power.

[ g BOL O r = 0.3 sec Delay o 30 -

                                                                                              '/

L'n 2.75 x 10' sec d

                              / i N                           f 2                     /               N-     Nominal Case
                                                                                         , -9 5   M                                                                                Ultimate    -
  =                                                                                   Power. BOL i                                                                  /
 '                                                                f
 $

• Nominal Case 80

                                                    /

s / s # p 0 0.1 0.2 0.3 0.4 0. 5 0.6 0.7 l Control Rod Worth, %Ak/k i i l i ENTHALPY INCREASE TO THE HOTTEST FUEL ROD  ! VERSUS EJECTED CONTROL RCD WORTH Figure 14-26 h

80 BOL Parameters

                                               ~

a . 6.0 x 10 (ak/k)/F J ak/k . 0.5% 60 r . 0.3 sec De ay L ' = 5. 47 x 10 sec g 40 N A Nominal Case g  % 20

         -0.5 -0.7        -0.9          - l .1       - 1. 3         -l.5        -1.7 5                              I Doppler Coef ficient. (ak/k)/F x 10                                    1 1

EFFECT ON REACTOR NEUTRON POWER OF VARYING THE DOPPLER COEFFICIENT - ROD EJECTION AT 10' ULTlHATE POWER Figure 14-27

 ~

g coas

44 i i  : i i g BOL Parameters a,. -I. s4 x IO -5(sk/k)/F 5 ~ Ah / k = 0. Sf.  ; j  % r . 0.3 see Delay

                               ~

L' = 5.47 x 10 8 sec

    }x                                    g Nominal Case a 36 e

E 32 0 3 6 9 12 l5 Moderator Coef ficient. (ak/k)/ F x 10 EFFECT ON REACTOR NEUTRON POWER OF l VARYING THE MDDERATOR COEFFICIENT - ROD EJECTION AT 10 ULTIMATE POWER Figure 14-28 g 0036

24 BOL Parameters a a = 6.0 x 10" ( Ak/ k)/ F 3 g' ' ak/k . 0.57. g 20 r = 0. 3 sec Delay a

                      \      N l '       5.47 x 10'  sec 5                                 \

I . l A S 16 Nominal Case -g a  % N m_ 12

          -0.5   -0.7      -0.9      -l .1         - 1. 3         1.5          1.7 Doppler Coef ficieni, (ak/k)/F x 10 EFFECT ON REACTOR THERMAL POWER OF VARYING THE DOPPLER COEFFICIENT R00 EJECTION AT 10' ULTIMATE POWER Figure 14-29 v

4 003?

1 i i i i i i e i BOL Parameters x

     . 19  -n        -l.1    x 10 -5 (ak/k)/F                                                 "

D j ak/k . 0. 57.

           - r . 0.3 see Delay

2.

• 5. 47 x 10' see Nominal Case 2

g -

             /

2 15 0 3 6 9 12 15 18 Hoderator Coefficient, (ak/k)/F x 10 EFFECT ON REACTOR THERHAL POWER OF VARYING THE H0DERATOR COEFFICIENT - ROD EJECTION AT 10' ULTlHATE POWER Figure 14-30 8 0038

             ~'

112 I i l l l BOL Parameters lli -

                 ,,    . -l.14 x       -5(ak/k)/F 6.0 x 10                                         /

110

             - ,, M                     (ak/k)/F ak/k - 0.37 109   -
                                    -5 g' . 5.47 x 10 sec EOL Parameters 108
   *             ,,    . -1.36 x 10 -5 (ak/k)/F               [

y 107

                 <<g = Assume Zero l

Ak/k . 0.37.

             - a' . 2.75 x 10
                                     ~

sec I i 105 ' [ Ultimate Power. g

   &    104 103
                                 /        >                                                                        1 102                                              Ultimate Power,         -

E0L

         ,0,              /                                  ,    ,          i

(( Mominal Case 100 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 Trip Delay, sec l l l l l . REACTOR THERMAL POWER VERSUS TRIP DELAY TIME - ROD EJECTION AT ULTIMATE POWER m Figure 14-31 0 0039

   -                                                                                                                       j l

BOL Parameters

               < ' , . -1.14 x          (ak/k)/F                             /

u g 6.0 x 10 (ak/k)/F 42 -9 ak/k.0.5%(10 Ultimate Power) , ak/k . 0.3% (Ultimate Power)

           ~ * . 5.47 x 10             sec E0L Parameters 38    .-                                                #

s o 0 = -1.36 x 10 -5 (ak/k)/F Ultimate 2 j - , , Assume Zero Power. _ g BOL and EOL j ak/k . 0.3", (Ultimate Power) ",_ 34 - ,. - 2.75 x 10' see o 1 0* 3 30

-                                           f 2                                          /

s 2 3 26 ' 2 E 22

                        /

• Nominal Case p 18

                                                  #I                -9 Ultimate
                            #p#                                  10 Power. BOL f

I4 0 0.2 0.4 0.6 0.8 1.0 Trip Delay (r) . sec ENTHALPY INCREASE TO THE HCTTEST FUEL ROD VERSUS TRIP DELAY TlHE - ROD EJECTION Figure 14-32(30/10

o

 . i    i            I            i      .           i         6          .'

7 _~  ; n L

                                                                    /
                                                           /- /          -e
                                                                         ~
 ^
                                                          /

_ 3 /  :

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 -                   w a j                   _

i I i f 3

I

                                                                        ~
-                                                l y                       -

g

                                                ]                                '.
-                                                                       ~

( h - a f _

                               /
                             /

o i f i f _ l -

    ,     ,1          .            ,      ,           ,         ,          8 8                    @                $             $

a E - -E

                              .....2...y LOFT SEMISCALE BLOWDOWN TEST NO. 546 -

VESSEL PRESSURE VERSUS TIME g 0041 Figure 14-33

idQ 80 e d

 -{ 60 s                                                                              .

I ( ar O 40 h 22% Measured

      "        \                                                 )

xm _ 0 0 10 20 30  % 50 60 Time, sec PRE 0lCTED PER CENT MASS RENAINING VERSUS TIME - LOFT TEST MO. 546 Figure I4-34 e k 004?,

_9

                          /
                             /
        /
               #                           o r                                      -
                                           - o c                                             =

N N

                         <                 o o'

E N 2 E 2 H S 5 'Jamod uoatnaN NEUTRON POWER VERSUS TIME FOR A 36-IN. ID, DOUBLE-ENDED, HOT LEG PIPE RUPTURE AT ULTIMATE POWER WITHOUT TRIP Figure 14-dh043

6. 5 ^ Density AX 2 3 AK - x Total w D 2 x I I , O E  % s _l \ / Doppler AX _

                           %/
     -2
     -3 0   .4     .8                1. 2         1.6        2.0 Time, sec                                           l l

l REACTIVITY VERSUS TIME FOR A 36-IN.

l. D. , DOUBLE-ENDED, HOT LEG PIPE RUPTURE AT ULTINATE POWER WITHOUT TRIP Figure 14-36 l

                  &                                                          l 0044

5 Hot Leg Ruptures a y / A 2 8 Void Shutdown s E

                                                         /

a 3 b(m 5

                 ~

u f E 0 2 l e I ,s / 5 o5 3 ( f

                            /
                                     \

M L Trip Shutdown "m c / s x

         >2  -

l N mO g 's w s Cold Leg Ruptures - Trip Shutdown o" 2

      =
         's        0 3g m

0 2 4 6 8 10 12 14

         $c O    *

• C -

gg Break Size, ft2

 &    L> wC

< U1 N

• O

70 60 50 7 ^ S ( ' U% Y . I f

                  )

i \ I 30

         )

E

   =                  ,

T i 10 N A 0 0 1 2 3 4 5 6 7 8 9 10 Time, see CORE FLOW VERSUS TlHE FOR A 36-IN. ID, DOUBLE-ENDED, HOT LEG PIPE RUPTURE Figure 14-38 002s

I I i l 1500 w 1400 DNS at 3" 1300 0.25 , sec a' y f Calculated by Quinn's

                                   \ [ Nodified    Sieder-Tate Equation
  ; i100                  f 0                  N O 1000 U
    =

A ' \

                                '\

900 y 1 i

    ,     800
                                   '     \

u Slump Model J 700 -Simulation y 3 u 600

                                             \

i

• 500 \

v  % o b z 400 s 300 200 \\

                                                                  ]   s 100                                                       L1   w 0       1     2      3       4      5     6     7      8    9     10 Time,see HOT CHANNEL CLAD SURFACE HEAT TRANSFER COEFFICIENT AFTER ONB VERSUS TIME FOR A 36-IN. 10, DOUBLE-ENDED, HOT LEG PIPE RUPTURE Figure 14-39
                                                   ~

0047

24 i Core Top 20 14" Pipe-50% N2

   ~           i      i     i               -

14" Pipe-33% N2 g ,, p

   ,,                                       <      i      i      ,       ,

12" Pipe-50% N2 i / / 12  ! / 5 [ s

                          /

a Core Bottom l

    $8 u                       (
    $                    /

4 0' 5 10 15 20 25 30 35 40 45 50 55 60 Time, sec. REACTOR VESSEL WATER VOLUME VERSUS TIME FOR A 36 IN. 10. DOUBLE-ENDED. HOI LEG PIPE RUPTURE FOR 600 PSIG CORE FLOODING (ANK OPERATING PRESSURE . 00@$GURE 14-40 SUPPLEMENT 3

24 f- Core Top 20 7 I2" "I? e-50% n, 2 1 1000 p3gg ]

                                     \n          -r-          -

2 2 4 00 psig 7 12" Aipe 3 3g

               ,2O1000psih d

n - ' S 14" Pipe.50% y7

                             /    f
                                                @ 400 psig U
   ;                           /

n ~

   ~

Core Bottom I

• a~

0

• 4

                 '      i
            ~

I O O 5 10 15 20 25 30 35 40 45 50 55 60 Time, sec REACTOR VESSEL WATER VOLUME VERSUS TIME FOR A 36 IN. 10 DO UB LE- ENDED , HOT LEG PIPE RUPTURE FOR 400 PSIG AND 1.000 PSIG CORE FLO90 LNG TANK OPERATING PRESSURES

                                      $                             0049 F1GURE I4-4I suvetEMENT  3

2500 240C /

                                                                                     /
                                                                              /Hot Spot 2300 2M 600 l b-12"-50% N2                             /                  ""     I% ~

2l00 ' - 2000 14 - ' U 1900 \ s / / 1800 1700 e0 -3 _ 400 1 o-l 4"-50 % M2 e f 1000 1b-l2*-33% N; 1500

          /e
                                \    1000 l b-12"-50% N 2 14M                                  l       !       i 10   12     14    16  18 20    22      24      26      28      30   32     34       36 Quench Time, see CORE FLOODING TANK AN ALYSIS: MAXIMUM CLAD TEMPERATURE VERSUS TIME TO QUENCH r0R A 36 -I N. 10. DOUBLE-ENDED. HOT LEG PIPE RUPTURE Figure 14-42 g                                                            0050 J

2800

                 \

2600 4 5 2400 i 2200

                    \   '

2000 N ' = 1800 Design N N e Point N  % e 1600 E E I400 1200 0 20 0 400 600 800 1000 1200 1400 1600 1800 2000 2200 Maximum Heat Transfer Coefficient, Btu /hr-f t -F MAXIMUM HOT SPOT CLAD TEMPERATURE VERSUS MAXIMUM HEAT TRANSFER COEFFICIENT AFTER DNB FOR A 36-lN. 10, DOUBLE-ENDED, HOT LEG PIPE RUPTURE Figure 14-43

                 $                                                      0051

2000 - Nominal Design Point 1900 - i 1800 -

      %u 1 1700  -

5 1600 - u 2

      ,  1500   -

1

     ",  1400  -

E g 1300 -

j 1200 - l100 ' ' ' ' ' 0 1 2 3 4 5 Nucleate Boiling Period, sec MAXIMUM H0T SPOT CLA0 TEMPERATURE AS A l FUNCTION OF TIME TO REACH DNB FOR A 36-IN. ID, DOUBLE-ENDED, HOT LEG PIPE RUP1URE l

 ,                                                          Figure 14-44

,v l g oosa

I e i i i i 2800

                                                           \     h=0 2600 2400
                                         /

2200 2 0 [ h . 15 Btu /hr-ft -F w g

    =   1800
                         ,  A                                f - 2o e  1600 3
                    /

J

                              \                                                            ,

i 1%0

   %                                                                                       i
   # 1200         I                                                                        l 1000 h   Y h . 100               l 800                                           \

600 x

         %0 200 0   t          t      I            i            i             I 0       10       10       'O            %           50           60 Time. see NOT SPOT CLAD TEMPERATURE VERSUS TIME FOR 36-IN. 10. DOUBLE-ENDED. HOT LEG PIPE RUPTURE AND VARIABLE QUENCH COEFFICIENT Figure 14-45 1 x 0053

l 2500 240 2300 w i t 2 1 2200 5 s o a

  ~

g 2100 2000 Nominal Value 1900 0 1 2 3 4 Full-Power Seconds HOT SPOT CLAD TEMPERATURE AS A FUNCTION OF FULL-P0hER SECONDS RESULTING FROM VOID SHUTDOWN FOR A 36-IN. 10. DOUBLE-ENDED. HOT LEG PIPE RUPTURE Figure 14-46 t oosa

EM . 1 g # 1800 /

u. -

as \ 200 F g 1600 j

                          ;                f                                                 100 F u

o

                     "     I 1%0 E

M, 1 gS . m

• 1200 j

U E* - N S "a 1000 I

                  %o o "8C    %

58" 800 m m I m 9% EEE

• 600

                    . ";           O       10     20 30        %           50                                                60 2  E
                  $om                                Time, see C              o" er           =,a Ut 2         "%m UI 3           mE
               -    Em f

4 5 m

.; i i n o i i n iiii n i n o iiii i o... . io.... . io n . .. orr.n "e n,

                                                ~
                                                  .             ?

o a i ~ O _ e n_- O

a. _

l i  : \ s \ -

N : N  : -*i

~ 3:

                             ^1                                                                 ~

EE" - ee - . . .' E -

                             .s :
                             =3
                                                                                      \            O,

N - -

p ii.ni,,, ,,,,,,,,, ,,,,,,,,, ,,,,,,,,, ,,,,,,,,, ,,,,,,,,, ,ii,....

                         .o               a             e              n            -

ce

                                           ,01 x 91 *ssen MASS RELEASED TO REACTOR BUILDING g                           FOR THE SPECTRUM 0F HOT LEG RUPTURES 0056            Figure 14-48

24 i i i i sisi i i i iinsi i i i i i s si i e i i s iii i i i s i si ft 2.0 1.0

8. 5 3. 2 7 14.1 ft 2

(2 j ft

                                                                                                            -s
                                                                                                               ^

L2

                                                                                 \                         \

3 Jurge Line Rupture ~~

                  =
                        .                                                                                                                           2             -

0.4 ft

                 &=

O

• 6 C"" m E UpW
• a 5 ft'.

    %Sz o 2

m " 14.1 ft c.s o 4 l-u o mo o r- ft 2 y w 2.0 m x _

3. 0 -

c: --4 2 2 ft p ft o < 0

                               ,     , , iieii              e    ' ' ' ' ' ' '               ' , ,,,,,                  f  n i i              ,    , , iiisi 71 m  >

Q 10 10 _a go o 30 i 60 10 o O

      -1   m n   r-   '8                                                                             Time. see 6-     8A   M c,     MM g      C C OM

~ p

.     =

4 A (.D M

1 i n... ....' ii ni.. 4 2,.00 . _ - _ _ _ _ _ t _ N-. i-_ g P et2 1.m . - -l .- - _ - .. . d - -- - L _ s_ - _ .-W --- --

                       !                    \               -

i.600

                        \         [

f j' i.0 vir

• 1,200 /

i i E .00 - -. . .h - - - - -_-. . - E s.ti., cor.-.-- -- --.-- _-- --- 2.0 vi2~ 8.5 ft2

14. I f t2 0 '

0 20 to 50 80 100 120 140 160 1, . ..e HOT LEG RUPTURES - REACTOR VESSEL WATER VOLUME ,;ERSUS i TIME INCLUDING EFFECTS OF BOILOFF AND INJECTION h FIGURE 14-50 l 0058 REVISED 2-8 68 L

i.soa /\ i i4.i itz t.7ao s.s itz Q f

    ,. i.soo E i.3ao        ^    i OG 3             1 C 1.100                                 g      ,
                                                                   - 3.0 ft2

\\

g i.o tt 2 \ J_ / 1 a4 ft2 x 500 0 10 2a 3a 40 50 60 70 Time, sec. HOT SPOT CLADDING TEMPERATURE VERSUS TIME FOR SPECTRUM OF HOT LEG RUPTURE S. Figure 14-51 s A l 0059

66.ii i ii ijiiisisisi sissisisi sinisi46i sii.sidia tisisisia

       ~
                                                          /
                                                              /                    :*/           o

/  :-

                                                              /             [~:
                                                /                                         :
                                                                                  "g      _

A /

                                                /                                              -
                                                                            '                    ~

3

                                                              /                   ~.
                                                                                   =

i m,

                                                                                   =    _

g o_ I

                                    \                                                            '

o

                                    /                                                   :

ii., ,,,, ,,,,,,,,, ....,, ., ,,,,.,,,, .

                                                               ,,,,,,i,,    ,,,,,,,,,            o E                R           *             "             *

• g Jt x 8!sd *eansseJJ o8eJoAy x.. REACTOR COOLANT AVERAGE PRESSURE -

SPECTRUM 0F COLD LEG RUPTURE SIZES F0FegLih@2

k  ! II l o I. i l b g

                          \

l i , i I l I e  ; i o 9 [

                                                                     ~

l EI i 3I I i %1 I I 8' 2 5

• yI [ SI

      = 1    -_ s                                 i!                 g
                                %,.               5l e                                                                  o 9

3I  : _- i

                   ;  k                                                 --

k, x - s-l l l

                                           )

8 { l l l

                                                    !     )          ,

l l ld N  ; I l i l l N< 1 l i i I 3 s i a  ! $ a a .: .: E)I 'J81TM 18888A Jogosey COLD LEG RUPTURES - REACTOR VESSEL WATER VOLUME VERSUS TIME INCLUDING EFFECTS OF BOILOFF AND INJECTION. i, FIGURE 14-53 0061 REVISED 2-8-68

i 2 1,700 ^ { 8 5 f1 I m 1.500 1.300

x

                                                    \      x 2
        "   i'100                                                                                          N                                    -
                            %/

1 '0 ft 2 M - s00

                                                        ^

b 04 ft2 i 700 , l 500 0 10 20 30 40 50 60 70 Time, sec. HOT SPOT CLADDING TEMPERATURE VERSUS TIME FOR SPECTRUM OF COLD LEG RUPTURES Figure 14-54

   .,..                   O 0062

L EGEND: HPI - HIGH PRESSURE INJECTION LPI - LOW PRESSURE INJECTION I HPI PUMP l HPl PUMP + 1 LPI PUNP 2 HPI PUMP + 2 LPI PUNPS 1 HPl PUMP + 2 CORE FLOODING TAKKS l MPI PUMP + 2 CORE FLOODING TANKS + 1 LPI PUMP 2 CORE FLOODING TAMKS , I LPI PUMP Pressurizer Surge 3 in. . 6 in. Line 36 in i I I I i 1 1 1 0.01 .02 .05 0.1 0. 2 0. 5 1.0 2. 0 5.0 10.0 J Break Size, ft l

                                                                                                                                                                             \

l EMERGENCY CORE COOLING SYSTEMS CAPABILITY

                                      $                                                                                      FIGURE 14-55 REVISED 2 8-68 0063 i                                                             _

I a JilieiiIl Ail 463iil i I4*i4l44 i I44ii33I Ii3iei43 4 &IiIi . 9

                                                      /                                                                   m
      -                                                                                                            _        9 m
           "?

m

     -                                                                                                            _         9   %

c - 2 - E 3 - '

    -                   o
                          =
    -                                                                                                                          ac
              -          E                                                                                       -

m J &

• 30 .

m 5 f ( _ N - o

  -                                                                                                            _           E f

i1 eit I it t .I 1 1 a I 1 I t t ! t 1 1 1 ! t I t i ii1 t ! n i t I f I r 1 I t t if I ff9l t 1 5 S S S R R 9 - Sisd 'aanssaJd ou!pting Jopeay REACTOR BUILDING PRESSURE VERSUS TIME - g 36-IN. ID, DOUBLE-ENDED, HOT LEG PlPE RUPTURE Figure 111-56 0064

i1' ll 5

                    ;:.- .-                ::_~            ::-   .

0 I 1 I w - i i n s y s r m-i o a i r l e s 1 t p o r cS o e _i e C l j g o n n y o aC

                                                                                                                           %i i

i id r p y el r i S a o u r l C B mm p p t o u S h h p N - i 4 0 1 g g t t a a i Wi W Na e s 0 0 s 0 0 6 0 ) ) ) 1 2 s I ( ( ) 2 6 3 g ( . 4 ( ' 8 _ i _ 3 0 c _ t ' 1 e i ' s 8 ' i ' e, m r i ' u t i ' p u R i ' g r

                           .                                                                                                                   e 5                                                                                                                    t f

i 5 ' a e mi T 2 i ' 0 1 l I j ' i 4 ' i ' I ' 1 i . ' 4 5 1 e 01 s ' i ' l ' I ' I ' I ' I ' 1 0

                                                                                                - . _ _ ~              :_-               01 0                       0               60             0                0                 0                      0 6                       5               8              3                2                 1
                                              "_ E 46
                                                .           i      5" ;=,oo2'      "

- 2 s 4 mm>o1o" mcc rc 2o Emcu*c:om n <mWu v Zm moa > "m x* o= 8ccrmemxomo 2o4 m - mo r o

                                                                        - -om =:%1com      :    s -t 42 Ic s 42oc1noof

- 2o oM l

1zm

                                                                        -    o

m9 o

ocr>4mo -  % s><- o 2>4m:o e%v Tgc,m 48mN

l 300 _ i i i ei,ie i i e i i i is i i a i i i si i '3 8

                                                                                                                                                                                      'l        8   8  3 8 8
                                                                                                                                                                                                                 b _

5

                                                                                         /               Reactor Building N                                6,600 gpa Core injection         -
                                                                        ;             /                  Vapor Temperature N                            3.000 gpm Building Sprays        -

250 , / , 5 (1) Without Spray Coolers  ! [

                                                                        ~
                                                                        ~

( 2) With Spray Coolers 2

                                                                        }/

k Reactor Building Sump Temperature

                                                                                                                                                                \                                                    {
                                                                        /-                                                                                                                                           -
                                                            ".* 200 i

I o om a N E

                                              @8Q                                                                                                                               N                                   -

150 ca t- o r- > -4

                                                                                                                                                                                  \                             N     -

mzo - 2 M -4 2

                                                   -i Mcm
                                                                                                                                                                                              \                     :

omc _, oh5

                                             = = 2 l00    -
                                                                                                                                                                                                          ~N o>z                        .-                                                                                                                                          :

H y* - rm <:: -

                                                  $g                        #    t    t 1 1 et i            I   I   I t I 8 If         8    '    I I I 8 8'        I        ' ' ' ' ' ' '         '   f   t f f 'ff e-
                                            , r-o' 1.-"          10                           to                          10 2

10 3 10 4 10 5

                                            -t r" z C                                       co o 2 :E                                                                                    Time after Rupture, sec
   @                                        m-m Q                                              5 E, 3m.

W

 ,                                                m O

e m - CD o

         .i4444   asa   eaiaisi.6         ....; ,                                                           =r
                                                        .. 6..  .   .i.. . .     .   . i

_g. I j "o

       . 3                                                                                         -

m

2

                           .                                                                      -               ,0
                         .5                                                                        -

o -

• 3 -

3 y - 2 sj a

5. -

I m Es i N i o

   ~
                                             . i ,    ,           ,     .,      si, ,1 , i T

o 8 s F R R 2 - 5 ! sd 'a;nssaJd Suspling Jo13 eau t REACTOR BUILDING PRESSURE VERSUS A- 2 TIME _ AFTER RUPTURE (8'.5 FT ) 0 0067 Figure 14-59

f l 60 _ i i i e i eis a i i iiei i i i iiisi i r i aiis i i iit 56.8 i _62.7 .-. 50  :

3 Emergency Cooling -

Units a _-

a 40 _ r .<^4 5 -

t. i = b, v/ _

w -

n. -

m 30 _

                                              .s
                                              =
                                                                                                             /                                                                        -
                                              'a                                                                                                                                     _

o 20 t.. _ H m  : _

                         -m Z >

mo - 4 -

         .m              >o                                                                                                                                                        ..

w m := 10 - - pw H - m a

        @)              *C                            2 GD             ::o r-                         :                                                                                                                            2 co o-eo                           0~              ' ' ' ' ' ' '         '   ' ' ' '            '    ' ' ' '       '      ' ' ' '       '
                      =                               -l                           0 mo                           10                           to                        10 1                        2                       3                           4

= 10 10 10

                   ^ m

_' ?g Time after Rupture, sec e c oc o 9 "T1 (T) O H N< v m m 8 U2 m C O Ch

        .i i i i .insi    . . . ..iii: - i . e.ii.ii s
                                                                   . i i - 4 4 . . ..... .....             .i..t     . i        .         g m
      -                                                                                                                         - e s                 -                                                                                              _
                 .               C m                o
      -        e c

o o N v

2 i O  : g I g _ .

a - t d e.

     - d i                                                                                                                    _

m

                                                                                                                                              =

m e w

    -                                                                                                                       _           o

_ ) _ l i l l o 1

  -                                                                                                                                   o
                                                                                                                                      -                   i 1

l l

  ~

1 1 4 1

                                                                                                                         ~

l l 1l t I t I t t t 1 1 ! I t I 1 ! t I f f I f i 1 1 1 1 I a I t 1 i I I f 1 f i 1 1 ? I t t t l t f I f j l o S S 8 R R S * ~ 6! sd *aanss*Jd Su!pt!ng Jopeau 1 1 REACTOR BUILDING PRESSURE VERSUS s 2 TIE AFTER RUPTURE (2.0 FT ) h 0069 i Figure Iti-61 !

             /

i

            \

60 , , , , , , , , , , , , , , , , , 52.5 52.5 -

                                    ~

50 /

3 Emergency Cooling Units ' ~

                     .?                                                                                                                                        '

E 460 3  : - m

a. -

EM ' ' J  : - m - C 3  : C lt 20  : sJ E _ C - _a =  : Z mO

                 -4                                                                /
          >o             10 r m H                   -

m tn -

o  :

c_. - m r-  : . .- C O u - _ H

• 0 ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' 'I tl

0 5O mu 10

                              -l 10                          10 1

10 2 10 3 10 4 Time after Rupture, sec

-.ir     -

e co gn to O c c ll0 n .c v m M m i (A m C N (A

e

      ,163 4 4 4 i i 5   414iiii36       iiie i i ei6    e e *

• i 6 e i e ii4i i e I, b m

o o _

     -g
         =
                                                                                                "o -    u g

_ q,

                                                                                                       =

6 g

                                                                         \                     ~       3
                                                                                         .        -o c                                                            -

_- g _ o h g _ S. 3 .2 m 5 _ _o iiiiiiiii. > > ,in ii , i.,iii,ri ni iiiiin ,,,,i,,,, - S 9 R

• e o 6!sd *aJnstaJd Su!pting ;olmeay REACTOR BUILDING PRESSURE VERSUS 2

TIME AFTER RUPTURE (0.4 FT ) On m ;' s Figure 14-63

10 - i i i iiila I e a is'oli i i i i iiis i i i i iiis i i 8 i i p_ 1

                                                                                                   ~                                                                                                                             y           _.

3 Emergency Cooling Units

                                                                                                   -                                                                                                      s
                                                                                                  -                         Total                        -   -                     --

Steam-Air-Mixture . 2 / y 2 [ Liquid - { 8m / c S rO m o $ - Structures - r'n z cn 4 S - /

                                                                                                          /                                                                                                       Pressure g                                                        oc           g                                                                                                                           i nj.ection e'-
                                                                        @ r-         w Cooler

[ N 5E

                                                                        -4
                                                                                        'l S

r ^ m - 3 Emergency O g .- Coolers _ o -< ^ mz c" z _ o O cm 10 ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' Il' ' ' ' 'i'It ' ' ' i i isi m" n" 10

                                                                                                 -1 10 0

10 1 2 3 4 o 10 to 10 m

   .7_
                                                                                >                                                                       Time after Rupture, sec ua                                                                            ca C                                                                             CD Q
                                                                                =

f -

  ?                                                                             ?

u- _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ - _ _

3 10 ie i i ii. i e i iiiei i i . . i.. . . . . . . . . . is f 3 Emergency Cooling Units

                                                                                                                                                                                                              /j -

p

                                                                                                                                                                                           /

_,/ - 7, - Total / __ 7 Steam-Air 2 10 _ f j/ Mixture

                                                                                                                                                                                              \

y- - To  ! Liquid 3

                                                                                                                    /
                                                                                                                        /                                 T Structures
                               ~
                                                                                                                  /
                                                      ~

m b ' Low Pressure C 5 ia g lajection Cooler si* 10 dJ  : I h 3 Emergency _ m 0 Coolers 2 0 -

                    -     2 P

o -E o .

                                  ,o                                                               . . .....        .    . . .....           .    .     . . . .

m u -1 . . . ..... . . . . . . .

                          ;                                                                                    0                       1
                       ,                10                                                                   10                                                      2                        3

• 10 10 4 m 10 10

                    % 0 A

W Time after Rupture, sec 0

           ;              5--

G f -* a

           =             2

AM _ i i iitill 6 i i i i n ii i i i s insi a a i i: sil 1 i# I 1415 y __ - .- N I f  ;

                                    ~

Reactor Building \ - B' 250

                                                                       /

vapor Temperature ._ K --

/ \  :

g [- Reactor Building s  : / Sump Temperature

                         % 200 m

m o i e

/ '

                                                                                                                                                                                                              \

a e n o o .. C 0 *  : C7 c m < m [

                                                                                                                                                                                                                   \,
                                               /                                                                                                                                                                           ~

rm c 150 y .

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Los Population Distance _ 2 - _ 100 , , , , , ,,, , , , , , , , , , , , , , , , , 103 2 4 6 8 104 2 4 C 8 105 2 4 6 C 106 Downsind Distance, ft. 1 MAXIMUM HYPOTHETICAL ACCIDENT THYROID DOSE ASSUMING FISSION PRODUCT REl. EASE PER T10-14844 ( Figure 14-73 0081- - Supplement 3

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