ML19322A763
| ML19322A763 | |
| Person / Time | |
|---|---|
| Site: | Oconee  |
| Issue date: | 12/01/1966 |
| From: | DUKE POWER CO. |
| To: | |
| References | |
| NUDOCS 7911210786 | |
| Download: ML19322A763 (119) | |
Text
TABLE OF COIiTEt!TS Section Page 14 SAFETY AIIALYSIS 14-1 14.1 CORE AND C00LAliT BOUIIDARY PROTECTICII ANALYSIS 14-1 14.1.1 ABNORMALITIES 14-1 14.1.2 ANALYSIS OF EFFECTS AND CONSEQ,UENCES 14-2 14.1.2.1 Unco =pensated Operating Reactivity Changes 14-2 14.1.2.2 {tartup Accident 14-3 14.1.2 3 Rod Withdrawal Accident Frc= Full Power Operation 14-5 14.1.2.4 Moderator Dilution Accident 14-7 14.1.2 5 Cold Water Accident 14-9 O
14.1.2.6 Loss-of-Coolant Flov 14-10 14.1.2 7 Stuck-Out, Stuck-In, or Dropped-In Control Rod 14-12 14.1.2.8 Loss of Electric Power 14-13 14.1.2 9 Steam Line Failure 14-16 14.1.2.10 Steam Generator Tube Failures 14-18 14.2 STAIGBY SAFEGUARDS AITALYSIS 14-19 14.2.1 SITUATIONS ANALYZED AfD CAUSES 14-19 14.2.2 ACCIDENT ANALYSES 14-21 14.2.2.1 Fuel Handling Accidents 14-21 14.2.2.2 Rod Ejection Accident 14-22 14.2.2 3 Loss-of-Coolant Accident v+-C 14.2.2.4 Maximum Hypothetical Accident 14-4(y -
14.3 REFERENCES
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O LIST OF TABLES Table No. Title g 14-1 Abnormalities Affecting Core and Coolant Boundary 14-1 14-2 Uncompensated Reactivity Disturbances 14-3 14-3 Situations Analyzed and Causes 14-20 14-4 Reactor Building Structural Heat Capacitance Segments 14-33 14-5 Core Flooding Tank Performance Data 14-37 14-6 Reactor Operating Conditions for Evaluation 14-39 14-7 Reactor Building Structure Data for Analysis of Time-Dependent Reactor Building Pressure 14-40 14-8 Sununary of Reactor Building Pressure Analysis for Three Reactor tailding Emergency Cooling Units 14 .43 -f /
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14-11 (Revised 4-18-67) 321
p LIST OF FIGURES (At Rear of Section)
Figure No. Title 14-1 Startup Accident from 10-9 FullPowerUsinga1.2%ok/kRod Group; High Pressure Reactor Trip is Actuated.
14-2 Startup Accident From 10-9 Full Power Using All Rods With a Worth of 10% Ak/k; High Flux Reactor Trip Is Actuated.
14-3 Peak Thermal Pove AccidentFrom10gVersusRodWithdrawalRateForAStartup Full Power.
14-4 Peak Neutron Pove Accident From 10-b"Full Versus Rod Withdrawal Rate For A Startup Power. .
14-5 Peak Ther=al Power Versus Trip Delay Time For A Startup Ace'.ien UsingA1.2%ok/kRodGroupAt5.8x10-5(ak/k)/Sec.From10-g Full Power.
14-6 Peak Thermal Power Versus Doppler Coefficient _For A Startup Atci-deng Using A 1.2% ak/k Rod Group At 5.8 x 10-2 ( ak/k)/Sec. From lo- Full Power.
14 'l Peak Thermal Power Versus T Time For A Startup Accident UsingAllRodsAt58x10gipDelay/Sec.From10-9 (ak/h) Full Power.
14-8 Peak Thermal Power Versus Dopple dentUsingAllRodsAt5.8x10gCoefficientForAStpupAcci-(ak/k)/Sec.From10 Full Power.
14-9 Peak Pressure Versus .ip Delay Time For A Startup Accident Using All Rods At 5 8 x 10- ( ak/h)/Sec. From 10-9 Full Power.
14-10 Peak Pressure Versus Trippe UsingAllRodsAt5.8x10gRod.WorthForAStartupAccident (ak/h)/Sec.From10-9 Full Power.
14-11 Peak Pressure Versus Dopple UsingAllRodsAt58x10gCoefficientForAStartupAccident
( ok/k)/Sec. From 20-9 Full Power.
14-12 Peak Pressure Versus Modera artup Accident UsingAllRodsAt58x10"gorCoefficientForA( ok/k)/Sec. From 10- Full Power 14-13 RodWithdrawalAccidentFromFullPowerUsingA1.2)ak/kRod Group a:; 5 8 x 10-> ( ok/k)/Sec; High Pressure Reactor Trip Is Actuated.
14-14 Peak Pressure Versus Rod Withdrawal Rate For A Rod Withdrawal Accident From Full Power.
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14-111}
FIGURES (Cont'd) Figure No. Title 14-15 Peak Pressure Versus Trip Delay Time For A Rod Withdrawal Acci-dent From Full Power Using A 1.2% ak/k Rod Group; High Pressure Reactor Trip Is Actuated. 14-16 Peak Pressure Versus Doppler Coefficient For A Rod Withdrawal Accident From Full Power Using A 1.2% ak/k Rod Group. 14-17 Per Cent Reactor Coolant Flow As A Function of Time After Loss Of Pu=p Power. 14-18 Mini =um DNBR Which Occurs During The Coastdown For Various Initial Power Levels. 14-19 Reactor System Cooling Rate For 4 in2 Steam Line Break. 14-20 Per Cent Core Experiencing 'DNB' As A Function Of Ejected Control Rod Worth At Full Power. 14-21 Zr-H2 O Reaction As A Function Of Ejected Control Rod Worth At Full Power. 14-22 Reactor Neutron Power Variation With Ejected Control Rod Worth. 14-23 Reactor Ther=al Power As A Function Of Ejected Control Rod Worth. 1h-24 Enthalpy Increase To The Hottest Fuel Rod Versus Ejected Control Rod Worth. 14-25 The Effect On Reactor Neutron Power Of Varying The Doppler Coef-ficient - Rod Ejection At 10-9 Full Pover. 14-26 The Effect On Reactor Neutron Power Of Varying The Moderator Co-efficient - Rod Ejection At 10-9 Full Power. 14-27 The Effect On Reactor Thermal Power Of Varying The Doppler Coef-ficient - Rod Ejection at 10-9 Full Power. 14-28 The Effect On Reactor Thermal Pgver Of Varying The Moderator Co-efficient - Rod Ejection At 10 > Full Power. 14-29 Reactor Ther=al Power Versus Trip Delay Ti=e - Rod Ejection At Full Power. 14-30 Enthalpy Increase To The Hottest Fuel Rod Versus Trip Delay Time - Red Ejection. 14 Reactor vessel Water Level For Rupture of 10 In. ID Pipe. Q 14 Reactor Vessel Coolant Inventory. After 36 In. ID Pipe Double Erded Rupture, 14-iv .. - 323
O V FIGURES (Cont'd) Figure No. Title 14-31 LOFT Semiscale Blowdown Test No. 546 - Vessel Pressure Versus Time. 14-32 Predicted Per Cent Mass Remaining Versus Time - LOFT Test No. 546. 14-33 Core Flow Versus Time For A 36 In. E Double-Ended Pipe Rupture. 14-34 Hot Channel Clad Surface Heat Transfer Coefficient After DNB Versus Time For A 36 In. E Double-Ended Pipe Rupture. 14-35 Reactor vessel Water Volume versus Time For 36 In., Double-Ended Pipe Rupture - For 600 Psig Core Flooding Tank. 36 Reactor vessel Water Volume Versus Time For 36 In., Double-Ended Pipe Rupture For 400 Psig And 1,000 Psig Core Flooding Tank. 14-37 Maximum Clad Temperature Versus Time To Quench For A 36 In. E Double-Ended Pipe Rupture. 14-38 Maximum Hot Spot Clad Temperature Versus Maximum Heat Transfer Coefficient Af ter DNB For A 36 In. E Double-Eitded Pipe Rupture. 14-39 Hot Spot Clad Tenperature Versus Time For 36 In. E Double-Ended Pipe Rupture And Variable Quench Coefficient. 14-40 Hot Spot Clad Temperature versus Time For 36 In. E Double-Ended Pipe Rupture And Variable Sink Temperature. 14-41 Mass Release To Reactor Building For The Spectrum Of Hot Leg Ruptures. 14-42 Reactor Coolant Average Pressure For The Spectrum Of Hot Leg Ruptures. 14-43 Reactor Building Pressure Versus Time - 36 In. E Double-Ended Rupture. 14-44 Reactor Building Pressure Versus Time Af ter Rupture - 8.5 Ft Rupture. 14-45 Reactor Building Pressure Versus Time After Rupture - 3 Ft Rupture. 14-46 Reactor Building Pressure Versus Time Af ter Rupture - 2 Ft Rupture. 14-47 Reactor Building Pressure Versus Time After Rupture - 1 Ft Rupture. ("] U 14-v (Revised 4-18-67) {f
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O FIGURES (Cont'd) Figure No. Title 14-48 Reactor Building Pressure Versus Time Af ter Rupture - 0.4 Ft Rupture. 14-49 Reactor Building Energy Inventcry For 36 In. ID Double-Ended Rupture. 14-50 Reactor Building Energy Inventory For 3 Ft Rupture. 14-51 Reactor Building Vapor And Sump Temperatures As A Function Of Time Af ter Rupture - 36 In. ID Double-Ended Rupture. 14-52 Reactor Building Vapor And Sump Temperatures As A Function Of Time Af ter Rupture - 3 Ft2 Rupture. 14-53 Criterion 17 Case For 36 In. ID Double-3ded Rupture.
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14-54 Reactor Building Zr Reaction Capability For-M Psig Design Pressure. 14-55 2 Hour Dispersion Model. 14-56 24 Hour And 30 Day Dispersion Models. 14-57 Thyroid Dose From Loss Of Coolant Accident - 2 Hour Dose. 14-58 Thyroid Dose From Loss Of Coolant Accident - 24 Hour And 30 Day Doses. 14-59 Maximum Hypothetical Accident Thyroid Dose Assuming 1007. Core Meltdown - 2 Hour Dose. 14-60 Maximum Hypothetical Accident Thyroid Dose Assuming 1007. Core Meltdown - 24 Hour And 30 Day Doses. 14-61 Integrated Direct Dose Following MHA With 3% Foot Reactor Building Wall Thickness.
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325-14-vi (4-18-67)
14 SAFETY ANALYSIS 14.1 CORE AND C00IANT BOUNDARY PROTECTION ANALYSIS 14.1.1 ABNORMALITIES In previous sections of this report both normal and abnormal operations of the various systems and components have beea discussed. This section summm izes and further explores abnormalities that are either inherently terminated or re-quire the normal protective systems to operate to maintain integrity of the fueland/orthereactorcoolantsystem. These abnormalities have been evaluated for a rated power of 2,452 mwt. Whenever a fission product release to the en-vironment occurs, the release is based upon the fission product inventory as-sociated' with the ultimate reactor core power level of 2,568 mwt. Table 14-1 su=marizes the potential abnor=alities studied. Table 14-1 Abnormalities Affecting Core and Coolant Boundary Event Cause Effect s Uncompensated Operating Fuel depletion Reduction in reactor system aver- , Reactivity Changes or xenon build- age temperature. Automatic reac-up. tor trip if uncompensated. No equipment damage or radiological hazard. Startup Accident Unc9ntro11ed Power rise terminated by negative rod ( )with- Doppler effect, reactor trip drawal, from short period, high reactor coolant system pressure, or over-power. No equipment damage or ra-diological hazard. Rod Withdrawal Accident Uncontrolled Power rise terminated by over-at Full Power rod withdrawal. power trip or high pressure trip. No equipment damage or radiologi- l cal hazard. ) Moderator Dilution Equipment mal- Slow change of power terminated Accident function or by reactor trip on high tempera- l operator error, ture or pressure. During shut- l down a decrease in shutdown mar- ' gin occurs, but criticality does not occur. No radiological hazard. l (*) Control rod, rod, and control cluster assembly (CCA) are used inter-changeably in this section and elsewhere in the report.
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326 i 14-1 I
Table 14-1 (Cont 'd ) Event Cause Effect Los s -of- Coolant Flow Mechanical or None. Core protected by reaccor electrical low-flow trip. No radiological failure of re- hazard. actor coolant Pump (s ) . Stuck-Out or Stuck-In Mechanical or None. Subcriticality can be or Dropped-In Control electrical achieved if one rod is stuck-o t. Rod failure. If stuck-in or dropped-in, con-tinued operation is permitted if effect on power peaking not severe. No radiological hazard. Loss of Electric Power Miscellaneous Possible power reduction or reac-faults. tor trip depending on condition. Redundancy provided for safe shut-down. Radiological hazard within limits of 10 CFR 20. Steam Line Failure Pipe failure. Reactor automatically trips if rupture is large. No damage to reactor system. Integrated doses at exclusion distance are 0.001 rem whole body and 0.210 rem thy-roid. Radiological hazard is within limits of 10 CFR 20. Steam Generator Tube Tube failure. Reactor automatically trips if Failures leakage exceeds normal makeup capacity to reactor coolant sys-tem. Integrated doses at exclu-sion distance are g.064 rem whole body and 2.3 x 10" rem thyroid. Radiological hazard is within limits of 10 CFR 20. 14.1.2 ANALYSIS OF EFFECIS AND CONSEQUENCES 14.1.2.1 Uncompensated Operating Reactivity Changes 14.1.2.1.1 Identification of Cause During nornal operation of the reactor, the overall reactivity of the core changes because of fuel depletion and changes in fission product poison con-centration. These reactivity changes, if lef t either uncompensated or over-compensated, can cause operating limits to be exceeded. In all cases, however, the reactor protective system prevents safety libits from being exceeded. No damage occurs from these conditions. i O
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14-2 (Revised 4-18-67 ) l
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O 14.1.2.1.2 Analysis and Results During normal operation, 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 de-mand level and to reestablish this te=perature at its aet point. If manual corrective action is not taken or if the automatic control system malfunctions, the reactor coolant system average temperature changes to compensate for the reactivity disturbance. Table 14-2 summarizes these disturbances. Table 14-2 Uncompensated Reactivity Disturbances Maximum Rate of Average Rate, Te=perature Change Cause Reactivity (Ak/k) /sec (Uncorrected),F/see Fuel Depletion -6 x 10~9 -0.0006 Xenon Buildup -3 x 10 -0.003 These results are based on +6 x 10~0 bk/k)/Fmoderatorcoefficientand-1.14 x 10-5 hk/k)/F Doppler coefficient. These reactivity changes are extremely slow and allow the operator to detect and compensate for the change. 14.1.2.2 Startup Accident 14.1.2.2.1 Identification of Cause The objective of a normal startup is to bring a suberitical reactor to the critical or slightly supercritical condition, and then to increase power in a controlled manner until the desired power level and system operating temperatures are obtained. During a startup, an uncontrolled reactivity addi-tion 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 overlap in-travel between two successive rod groups. This overlap occurs at the min 4=m worth for each group since one group is at .the end of travel-and the other is at the beginning of travel. The taximum vorth of any single control rod group is never greater than '1.2% Ak/k when the y reactor is critical as specified in 7 2.2.13 C (~ l- 14-3 328 W
- b. Control rod withdrawal rate is limited to 25 in./ min.
- c. A short-period withdrawal stop and alarm are provided in the source ran6e-
- d. A short-period withdrawal stop, alarm, and trip are provided in the inter =ediate range.
- e. A hi 6h flux level and a high pressure trip are provided in the power ranSe.
The reactor protective system is designed to limit (1) the reactor thercai power to 114 per cent of full power to prevent fuel i'amage, and (2) the reac-tor coolant system pressure to 2,515 psia. 14.1.2.2.2 Methods of Analysis An analog model of the reactor core and coolant system was used to detemine the characteristics of this accident. This analog codel used full reactor coolant flow, but no heat transfer out of the system and no sprays in the pressurizer. The full-power Doppler coefficient (-1.1h x 10-2 (ak/k)/F) was used although the Doppler is =uch 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 on short period was not in-corporated in the analysis. The nominal values of the principal parameters used were: 0 3 see trip delay, +6 x 10-5 (ak/k)/F moderatui coefficient, and
-1.lk x 10-3 (ak/k)/F Doppler coefficient. The total worth of all the control rods which are inserted into the reactor core following any trip is 8.6% ok/k without a stuck centrol rod or 5.6% ok/k (the nominal case in this study) with a stuck rod.
14.1.2.2 3 Results of Analysis Figure 14-1 shows the results of withdrawin6 the maximum worth control rod group at a rod speed of 25 in./ min from 1 per cent suberitical. This group is worth 1.2% ok/k when fully removed. This rod velocity and worth result in a maxi =um reactivity addition rate of 5.8 x 10 5 ( ok/k)/sec. The Doppler effect beS ins 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 14-2 shows the results of withdrawing all 69 control rod assemblies (with a total worth of 10% ok/k) at the maximum speed from 1 per cent sub-critical. This results in a maxite reactivity addition rate 'of 5.8 x 10 4 (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 ne6ative Doppler effect. The high neutron flux trip takes effect 0.25 sec. after the peak power is reached and terminates the transient. The peak therLal heat flux is only 16 per cent of the full power heat flux. A sensitivity analysis was performed on both of these startup accidents to de-temine the effect of, varying several key parameters. Figures 14-3 through 14-6 show typical res' u lts for the single group,1.2% ok/k startup accident. l L. '; 14-4 G 1 ..
Figures 14-3 and 14 4 show the effect of varying the reactivity addition rate on the peak ther=al power and peak neutron power. This reactivity rate was varied from one order of magnitude below the nominal sind le rod group case (1.2% ok/k) to more than an order of magnitude above the rate which represents all rods 10% ak/k) being withdrawn at once. The slower rates - up to about 0 5 x 10- ( ak/k)/sec - will result in the pressure trip bein6 actuated, where-as only the very fast rates actuate the hi 6h neutron flux level trip. Figures 14-5 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 few per cent difference is noted. F16ures 14-7 and 14-8 are thecorrespondingresultsfromthewithdrawalofallrods(10%ak/k). Since this transier+ inserts reactivity an order of magnitude faster than the single control rod group case, there is considerably more variation in the peak ther-mal power over these wide ranges. At hi Sh values of the Doppler coefficient, the neutron power rise is virtually stopped before reaching the hi $h flux trip level. Reactor power generation continues until sufficient energy is trans-ferred to the reactor coolant to initiate a high pressure trip. This results in a higher peak thermal power. Figures lk-9 through 14-12 show the peak pressure response to variations in several key parameters for the case where all rods are withdrawn. It is seen that the safety valve is opened when these parameters are changed considerably from the nominal values, except in the case of the moderator coefficient which p has little effect because of the short duration of the transient. Again for a ( high Doppler coefficient, the high pressure trip is relied upon. None of these postulated startup accidents, except for reactivity addition rates of 2 x 10-3 ( ak/k)/sec, which is three times greater than for withdrawal of all rods at once, causes a ther=al power peak in excess of 40 per cent full power or a nominal fuel rod avera6e te.~perature greater than 1,715 F. The nominal 1.2% ok/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 =aximum rate of coolant expansion resulting fro = this startup acci-dent. The10%ak/kwithdrawal-usingall69 rods-causesapeakpressureof only 2,465 psia because the flux trip is actuated prior to the pressure trip. It is concluded that the reactor is completely protected against any startup accident involving the withdrawal of any or all control rods, since in no case does the thermal power approach 114 per cent, and the peak pressure never ex-ceeds 2,515 psia. 14.1.2 3 Rod Withdrawal Accident From Full Power Operation 14.1.231 Identification of cause A rod withdrawal accident presupposes an operator error or equipment failure which results in accidental withdrawal of a control rod group while the re-actor is at full power. As a result of this assumed accident, the power level increases; the reactor coolant and fuel rod te.peratures increase; and if the withdrawal is not terminated by the operator or protective system, -, core damage would eventually occur. %.d x 1u-5 330
The following provisions are =ede in the design to indicate and terminate this accident:
- a. High reactor outlet coolant te=perature alar s.
- b. High reactor coolant system pressure alar s.
- c. High pressurizer level alarm.
- d. High reactor outlet coolant temperature trip.
- e. High reactor coolant system pressure trip.
- f. High power level trip.
14.1.2 3 2 Pethods of Analysis An analog computer model was used to determine the characteristico of this ac-cident. A complete kinetics =cdel, pressure model, average fuel rod model, steam demand model with turbine coastdown to 15 per cent of full load, cool-ant transport model, and a simulation of the instrumentation for pressure and flux trip were included. The initial conditions were normal full power opera-tion without autoca*,1c control. Only the moderator and Doppler coefficient of reactivity were used as feedback. The nominal values used for the =nin pa-remeters were: 0 3 see trip delay time, -1.14 x 10-5 (/_,k/h)/F Doppler coef-ficient, +6 x 10-5 (ak/k)/F moderator coefficient, 25 in./ min control rod speed, and 1.2% ak/k control rod group worth. The total worth in all the con- ' trol rods which are inserted into the reactor core following any trip is 8.6% Ak/k without a stuck control rod or 5.6% ak/k (the nominal value used) with a stuck rod. The above rod speed and group rod worth give a maxi =um reactivity addition rate of 5.8 x 10-5 (ak/k)/see, which is the nominal case. The reactor protec-tive system is designed to limit (1) the reactor power to 114 per cent of full power to prevent fuel damage, and (2) the coolant system pressure to 2,515 psia to prevent reactor coolant system damage. 14.1.2 3 3 Results of Analysis Figure 14-13 shows the results of the nominal rod withdrawal from full power using the 1.2% ak/k rod group at 5.8 x 10-5 (ak/k)/sec. The transient is terminated by a high pressure trip, and reactor power is limited to less than i the design overpower of 114 per cent of full power. The changes in the pare.m-I eters are all quite small, e.g. 5 F average reactor coolant temperature rise and 200 psi system pressure change. A sensitivity analysis of important paraceters was performed around this nom-inal case and the resultant reactor coolant system pressure responses are shown in Figures 14-14 through 14-16. Figure 14-1h shows the pressure variation for a very wide range of rod with-drawal rates - = ore than an order of magnitude s= aller and greater than the nominal case. For the very rapid rates, the neutron flux level trip is bb, , 14-6
,y
) actuated. This is the pri=ary protective device for the reactor core; it also protects the system against high pressure during fast rod withdrawal accidents.
The high pressure trip is relied upon for the slower transients. In no case does the thermal power exceed 109 per cent full power. Figures 14-15 and 14-16 show the pressure response to variations in the trip delay time and Doppler coefficient. For the higher values of the Doppler co-efficient, the pressure trip is always actuated, and, therefore, the pressure levels off. This analysis shows that the high pressure trip and the high flux level trip adequately protect the reactor against any rod withdrawal accident from full power. 14.1.2.4 Moderator Dilution Accident 14.1.2.4.1 Identification of Cause The reactor utilizes boric acid in the reactor coolant to control excess re-activity. The boron content of the reactor coolant is periodically reduced to compensate for fuel burnup. The dilution water is supplied to the reactor coolant system by the high pressure injection and purification system. This system is designed with several interlocks and alarms to prevent improper operation. These include: (- a. Flow of dilution water to the letdown storage 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 posi-tion (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 gpm.
- b. Flow of dilution water is automatically stopped when either the flow has integrated to a preset value or when the rods have been insert-ed to a preset position (at about 75 per cent full stroke).
- c. A varning light is on whenever dilution is in progress.
The high pressure injection and purification system nor= ally has one pump in operation which supplies up to 70 gpm to the reactor coolant system and re-quired flow to the reactor coolant pump seals. There are two additional pumps 72er anh~whn.:. - M r t',ra ;- M ' rXwM --# = ?--> Thus, the total makeup flow available is limited to 70 gpm unless the operator takes action to increase the 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 cakeup 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 letdown storage tank to the reactor coolant system by the high pressure injection pump. O It is also possible, however, to have a slightly higher flow rate during tran-sients when the system pressure is lower than the nominal value, and the 2: ----- 332
G pressuriser level is below non::al. This flov might be as high as 100 gp . In addition, with a co=bination of =ultiple valve failures or caloperations, plus = ore than one high pressure injection pu=p operating and reduced reac-tor coolant syste= pressure, the resulting inflow rate ca.n be as high as 500 sp. This constitutes the =cxi=u= dilution accident. A reactor trip would terminate unborated water addition to the letdevn storace tank and total flow into the coolant syste= vould be ter=inated by a high pressuriser level. The criteria of reactor protection for this accident are:
- a. The reactor power vill be limited to less than the design overpower of 114 per cent full power to prevent fuel da= age.
- b. The reactor protective syste= vill limit the reactor coolant system pressure to less than the syste= design pressure of
?,500 psig.
- c. The reactor =inimum suberiticality =argin of 1% Ak/k will be maintained.
14.1.2.4.2 Analysis and Results The reactor is assu=ed to be operating at full power with the maxi =u= ini-tial boron concentration (2,200 ppn) in the reactor coolant system. The dilute water is unifor=ly distributed throughout the reactor coolant vol-u=e. Achargeinconcentrationof100pp=producesa1%Ak/kreactivity change. The followin6 tabulation shows the effects of the above three dilution rates on the reactor. Average Reactor Coolant Dilute Water Reactivity Rate, Syste: Te=perature Flev, t;p= (Ak/k)/sec Chance, F/see
~70 + 3 0 x 10-6 03 100 + 4.4 x 10-6 03 500 + 2.2 x 10-3 05 e
$i 14-8
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(q The fastest rate of dilution can be handled by the auto =atic control system, which would insert rods to maint?.in the power leve:. and reactor coolant sys-tem temperature. In the event of an interlock failure and the reactor is under manual control, these reactivity additions would cause a high reactor coolant temperature trip or a high pressure trip. In any event the ther=al power vill not exceed 114 per cent full power, and the system pressure vill not exceed the design pressure of 2,500 psis. Therefore moderator dilution accidento vill not cause any damage to the reactor system. During refueling or maintenance operations when the reactor closure head has been ren.oved, the sources of dilute water =akeup to the letdown storage tank--and therefore to the reactor coolant system--are locked closed, and the high pressure injection pu=ps are not operating. At the beginning of core life when the boron concentration is highest, the reactor is about 10% ak/ksuberiticalwiththe=aximumworthrodstuckout. In order to demon-strate the ability of the reactor to accept =oderator dilution during shut-down, the consequences of accidentally filling the letdown storage tank 4 with dilute water and starting the high pressure injection pu=ps have been evaluated. The entire water volu=e from the letdown storage tank could be pumped into the reactor coolant system (assuming only the coolant in the J reactor vessel is diluted), andthereactorvouldstillbe77 oak /ksub-critical. 14.1.2 5 Cold Water Accident f'
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The absence of individual loop isolation valves eliminates the potential source of cold water in the reactor coolant system. Therefore, this accident is not credible in this reactor. t e v th.' 14-9 __ 334
14.1.2.6 Loss-of-coolant Flow 14.1.2.6.1 Identification of Cause A reduction in the reactor coolant flow rate occurs if one or more of the re-actor coolant pumps fail. A pu= ping failure can occur from cechanical fail-ures, or from a loss of electrical power. With four independent pumps a ce-chanical failure in one pump will not affect operation of the others. Each reactor coolant pump receives electrical power from one of the electri-cally separate busses of the h,160 volt system discussed in 8.2.2 3 A sire - lar loss of neither outside power nor the unit generator will cause a loss of electrical power to the pumps. Faults in an individual pump motor or its power supply could cause a reduction in flow; however, a complete loss of flow is extremely unlikely. In spite of the low probability of a complete loss of power to all reactor coolant pumps, the unit has been designed so that such a failure would not lead to core damage. The reactor protection criterion for loss-of-coolant flow conditions starting at full 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 design overpower. This corresponds to a DNBR of 138 at 114 per cent full power (Table 3-1). 14.1.2.6.2 Methods of Analysis The loss-of-coolant-flow accident is analyzed by a combination of analog and digital computer programs. Analog simulation is used to determine the reac-tor flow rate following loss of pumping power. Reactor power, coolant flow, 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 includes six delayed neutron groups, control rod worth and rod insertion char-acteristics, and trip delay time. The analog model used to determine flow coastdown characteristics includes description of flow-pressure drop relations in the reactor coolant loop. Pump flow characteristics are determined from the manufacturer's zone map. Flow-speed, flow-torque, and flow-head relation-ahips are solved by affinity laws. A transient, thermal-hydraulic, B&W digital computer program ia used to compute
~
channel DNBR continually during the coastdown transient. System flow, neutron power, fission product decay heat, anc core encering enthalpy are varied as a function of time. The program =aintains a transient inventory of? stored heat which is determined from fuel and clad temperatures beginning with the initial steady state conditions. The transient core pressure drop is determined for . i.,
}h 14-10
C'\ h average channel conditions. The representative hot channel flows and corre-sponding DIGR are obtained by using the avera3e core pressure drop. The hot channel DIGR as a function of ti=e is compared with the design DIGR at caxi-mum overpower to determine the degree of heat transfer =argin. The loss-of-coolant-flow analysis has been carried out in the power range be-tween 303 and n4 per cent fun power. conditions utili::ed in the analysis are:
- a. Initial core inlet te=perature for given power level is assu=ed 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 con-dition until initial downward =ovement of control rod, is 300 -
milliseconds.
- d. The per cent of initial reactor power as a function of ti=e after loss of pu=p is as shown in Figure 3-7
- e. The pu=p inertia is 70,000 lb-ft ,2 14.1.2.6 3 Results of Analysis The results of this analysis show that the reactor can sustain a loss-of-coolant-flow accident without damage to the fuel. The results of the evalua-tion are presented in Figures 14-17 and 14-18. Figure 14-17 shows the per cent reactor flow as a function of time after loss of all pump power. Figure 14-18 shows the mini =um DIGR's which occur during the coastdown for various initial power levels. The degree of core protection during coastdown is indi-cated by comparing the DIGR for the coastdown with the design value of 138 at 114 per cent power. This DIGR (138) in the representative hot channel corre-sponds to a 99 per cent confidence that 99 5 per cent of the core vin not experience a departure from nucleate boiling under steady state conditions at the design overpover (3 2 31).
Under nor=al conditions, the maximum indicated power level from which a loss-of-coolant accident could occur is 103 per cent full power (as indicated by reactor instru=entation). This power level represents an allowance of plus 3 per cent full power for transient overshoot. This power level also represents the maximum power demand which vill be permitted to the reactor control system. The 103 per cent run power is an instrument-indicated value and is subject to the following =aximu:s errors: (1) i 2 per cent heat balance and (2) t 4 per cent nuclear instrumentation. The true power level could be as high as 109 per cent at 103 per cent indicated power. As shown in Figure.14-18, however, the DIGR at 109 per cent power is 1.42 which is significantly larger than the design DIGR. \ 14-11 _, 336
The reactor coolant system is capable of providing natural circulation flow after the pu=ps have stopped. The natural circulation characteristics of the reactor coolant system have been calculated using conservative values for all resistance and form loss factors. I;o voids are assumed to exist in the core or reactor outlet piping, The following tabulation and Figure 9-7 show the natural circulation flow capability as a function of the decay heat genera-tion. Decay Heat ITatural Circulation Core power, n Core Flov, % Full Flov 1/2 1.2 1 1.8 3 33 5 4.1 The above flows provide adequate heat transfer for core cooling and decay heat re= oval by the reactor coolant system. The reactor is protected against reactor coolant pump failure (s) by the pro-tective system and the integrated control system. The integrated control sys-tem initiates a power reduction on pump failure to prevent reactor power from exceeding that permissible for the available flov. The reactor is tripped if insufficient reactor coolant flov exists for the power level. The operating limits for less than four pumps in operation have been presented in 4 3 7 1k.1.2 7 Stuck-Out, Stuck-In, or Dropped-In Control Rod 14.1.2 7 1 Identification of Cause The control rod drive assemblies have been previously described in 3 2.4 3 The results of continuous control rod withdrawal have been analyzed in 14.1.2.2 and 14.1.2 3 In the event a control rod cannot be =oved due to either elec-trical faults or mechanical seizure, consideration must be given to localized power peaking and suberitical cargin. 14.1.2 7 2 Analysis and Results Adequate hot suberitical margin is provided by requiring a suberiticality of 1%Ak/ksuberiticalwiththecontrolrodofgreatestworthfullywithdrawn from the core. The nuclear analysis reported in 3 2.2 demonstrates that this criterion can be satisfied. In the event an un=ovable control rod is partially or fully inserted in the core or a single rod is dropped luring operation, its location and effect on local power distribution detemine whether continued power operation is per-missible. Further studies of the locatio:2 of a stuck rod in the core vill be 14-12
O made to define permissible conditions of operation. The criteria for these studies are: (1) operation with a stuck rod vill not increase the IEiB prob-ability above the probability specified for design conditions, and (2) a hot subcritical margin of 1% M/k vill be =aintained with the stuck rod in its inoperative position and the operating rod of greatest reactivity worth in the fully withdrawn position. If a control red is dropped into the core during power operation, the same consideration of localized power peaking as for a stuck rod vill apply. 14.1.2.8 Ioss of n ectric Power 14.1.2.8.1 Identification of Cause The Oconee Nuclear Station is designed to withstand the effects of loss of electric 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 which results in a complete loss of all station power.
O The reactor protection criteria for these conditions are that fuel damage vill not occur from an excessive power-to-flow ratio nor vill the reactor coolant system pressure exceed design pressure. 14.1.2.8.2 Results of " Blackout" Conditions Analysis The net effect of a " blackout" condition on the station vould be openin6 Of the main generator breakers, thus disconnecting the station from the entire transmission system. When this occurs, a runback signal on the integrated
= aster controller causes an auto =atic power reduction to 15 per cent power.
Other actions which occur include:
- a. All vital electrical loads, including reactor coolant pumps, con-denser circulating water pu=ps, condensate and condensate booster pu=ps, and other auxiliary equipeert, vill continue to obtain power from the unit generator. Feedvater is supplied to the steam gen-erators by steam-driven feed pu=ps.
- b. As the electrical load is dropped, the turbine generator accelerates and closes the governor valves and reheater intercept valves. The unit frequency vill peak at less than the overspeed trip point and decay back to <tet frequency in 40-50 sec.
- c. Following closure of turbine governor valves and reheat intercept valves, steam pressure increases to the turbine bypass valve set
() point, and may increase to steam system safety valve set point. A
~
Stea= is relieved to the condenser and to the atmosphere. Stea= venting to the atmosphere occurs for about two minutes following i 14-13 33g
blackout from 100 per cent initial power unt,il the turbine bypass can handle all excess steam generated. The capacity of the modulating tur-bine bypass valve is 25 per cent of the :pives vide open (WO) steam flov, and that of the safety valves is W per cent of WO steam flow. The first safety valve banks are set at 1,050 psig with additioral banks set at pressures up to 1,104 psig (5 per cent above design pres-sure as' allowed by code ). Steam venting pemits energy removal from the reactor coolant system to prevent a high pressure reactor trip. The initial power runback is to 15 per cent power which is greater than the unit auxiliary load. This allows sufficient steam flow for regulating turbine speed control. Excess power above unit auxiliar/ load is rejected by the turbine bypass valve to the condenser.
- d. During the short interval while the turbine speed is high, the vital electrical loads connected to the unit generator vill undergo speed increase in proportion to the generator frequency increase. All motors and electrical gear so connected are designed for the increased fre-quency.
- e. After the turbine generator has been stabilized at auxiliary lead and set frequency, the station operator =ay reduce reactor power to the auxiliary load as desired.
The blackout accident does not produce any fuel da= age or excessive pressures on the reactor coolant system. There is no resultant radiological hazard to station operating personnel or to the public from this accident as only secondarf system steam is discharged to the at=osphere. J Unit operation with failed fuel and steam generator tube leap.a.3e was shown to be safe by the analysis presented in 11.1.2 5 2. For the sane conditions, the steam relief accompanying a blackout accident would not change the whole body dose. The whole body dose is pri=arily due to the release of Xe and Kr. Re-lease of these gases is not increased by the steam relief because even with-out relief, all of these gases are released to the at=osphere through the con-denser air ejector. The rate of release of iodine during the approximately two minutes of relief would be increased by almost a factor of 104 , because the iodine is released directly to the at=osphere rather than through the con-denser and station vent. However, the quantity released during this short time is so s=all that it would only increase the annual release by less than 4.0 per cent. Since the iodine concentration from operation with failed fuel and tube leakage was only 0.001 MPC, the added contribution from steam relief is insignificant. 14.1.2.8 3 Results of co=plete Loss of All Station Power Analysis The second power loss considered is the hypothetical case where all station power except the station battery is lost. The sequence of events and the evaluation of consequences relative to this accident are:
- a. Ioss of electric power deenergizes the holding ma$ net circuits on the control rod drive assemblies. A loss of power results in gravity insertion of the control rods.
14-14 339 9 ;.
x
- b. The steam generator safety valves actuate after the turbine trips and prevent excessive temperatures and pressures in the reactor coolant system. These safety valves will close after 20 sec when the flow through the turbine bypass valve is sufficient to relieve excess steam and provide for decay heat re= oval. The environmental consequencesofthisaccidentareonly1/6ofthosediscussedunder the blackout conditions in 14.1.2.8.2 as the safety valves relieve for only 20 see compared to two minutes for a blackout condition.
- c. The reactor coolant system flow decays without fuel damage occur-ring. Decay heat removal after coastdown of the reactor coolant pumps is provided by the natural circulation characteristics of the system. This capability was discussed in the loss-of-coolant-flowevaluation(14.1.2.6).
- d. A turbine-driven emergency feedwater pu=p is previded to supply feedwater any time the =ain feed pumps cannot ope ate. The emer-gency feed pump takes suction from the condenser hotwell and the uppersurgetank(condensatestorage). The emergency pump supplies feedvater to the steam Generators. The emergency feed pump is driven by steam from either or both steam 62nerators.
The controls and auxiliary systems for the emergency feed pump op-erate on d-c power from the station battery. O A recirculation line from emergency pump discharge back to the con-denser is provided to permit periodic testing.
- e. The unit condenser cooling water system is arranged to preride cooling water even in the unlikely event that all power is lost.
This is accomplished by the arrangement shown in F16 ure 9-8. Condenser cooling water intake is obtained from a point below minimum level in the Little River branch of Lake Keovee. Circu-lating water pumps are provided to overcome line friction and dis-charge condenser flow to the Lake Keovee. An emergency discharge line, nor:: ally closed by a power-to-close valve is provided to ob-tain circulation even if power is lost. This line discharges to the tailrace of the Keowee Dam. These provisions insure a supply of cooling water at all times and provide extended use of condenser hotwell and upper surge tank inventory for cooling purposes. The features described above permit decay heat cooling of the unit for an ex-tended period of time following a complete loss of electric power. The above evaluation demonstrates the design features incorporated in the de-sign to sustain lor.,s of power conditions with just the station battery to operate system controls. Immediate operation of the emergency feedwater pump and the emergency condenser cooling water system is not of critical nature. Each reactor can sustain a complete electric power loss without emergency cooling for about 25 min before the steam volume in the pressurizer is filled with reactor coolant. These 25 min are derived as follows: O 340
~*^'
[ l3 14-15 J
a. O Stea= generators evaporate to dryness 10 min
- b. Pressurizer safety valves open 5
- c. Pressuriser fills with water 10 25 min Beyond this ti=e reactor coolant vill boil off, and an additional 90 min vill have elapsed before the boiloff vill start to uncover the core.
The e=ergency feedvater pump and the e=ergency condenser cooling vater system can be actuated within this period of time. Accordin6l y, core protection is insured for the unlikely condition of total loss of sta-tion 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 secondar/ coolant due to failures in the steam lines between the steam 3enerators and the turbine. The criteria for station protection and the release of fission products to the environment are:
- a. The reactor shall trip and re=ain suberitical without boron )
addition until a controlled rate of system cooldown can be effected.
- b. The potential environmental consequences from radioactivity in the secondar/ coolant system shall not exceed those specified by 10 CFR 20.
14.1.2 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 inventor / available for cooling. The steam generator inventor / increases with power level. The inventory at full power is 46,000 lbs and decreases linearly to I 20,000 lbs at 15 per cent power. The stea= line break accident analysis is performed at full power in order to deten::ine maxi =um cooling and inventor / release effects. l 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 ini-tially cause the reactor control system to act to restore steam pressure and j load generation. A steam line rupture of a small area causes a relatively slow decrease in l steam pressure. This places a demand on the control system for increased l feedvater flov. In addition, the turbine control valves will open to e 14-16
l l l 1 l 4 s O maintain power generation. Increased feedvater flow causes the average reactor coolant temperature to decrease, and the resulting temperature error calls for control rod withdrawal. The limiting action in this condition is the 103 per cent limit on power demand to the rod control system. If the moderator tem-perature coefficient of reactivity is small or slightly positive, the reactor power will decrease when the control system reaches the power demand limit be-cause of continuing temperature decrease. The reactor will then trip on low reactor cociant system pressure. A reactor trip will initiate a reduction in the feedwater flow to the steam Benerators. When the moderator temperature coefficient is negative, the reactor power will tend to increase with decreasing average coolant temperature. This will cause control rod insertion to limit reactor power to 103 per cent. With power 11=- ited at 103 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 steam lines are isolated by the turbine stop valves as shown in Figure 10-1. Following isolation cf the steam generators, the unit with the ruptured steam line continues to blev down to the atmosphere. The max 1::um cooldown of the reactor coolant system would be that resulting from the blowdown from one steam generator. A typical cooling rate following reactor trip for a 4 in.2 steam line rupture is shown in Figure 14-19 The following tabulation lists the approximate time required to blow down the Os contents of the steam generator with a ruptured steam main. Leak Area, in.2 Blowdown Time, sec h 860 32 110 128 27 A steam line failure of large area results in high steam flow with resultin6 rapid pressure decrease in reactor coolant system and steam system. The re-actor trips on low reactor coolant system pressure. Reactor trip causes turbine trip and reduction in feedwater flow to decay heat level. The tur-bine trip isolates the steam lines and prevents blowdown of the other steam generator whose secondary side does not have a rupture. The steam generators are designed to maintain reactor system integrity upon loss of-secondary-side pressure. Therefnre, this accident will not lead to a reactor coolant system i failure. Assuming blowdown from both steam generators resulting from secondary steam system ruptures, the maximum cooling rate durin6 this accident occurs durin6 - the first 10 see after the break. Thecoolingrateisapproximately3F/see at the peak. The net cooldown of the reactor coolant system, assuming total i blowdown of the steam generators and accounting for transfer of core stored heat and decay heat, is approximately 60 F. This results in an average cool-ant temperature of 520 F which is about 20 F lower than the normal zero power y average coolant temperature. Since control rod worth is provided for an
\ -
'h' 1u-17 342
average reactor coolant temperature change of 40 F (no load to 15 per cent load), the reduction in reactivity shutdown margin would be equal to the re-activity associated with a 20 F moderator temperature decrease. Using the maximum value for the moderator temperature coefficient (-1.7 x 10-4 ok/k/F), the reductior would be 0.34 per cent ak/k. The minimum shutdowa margin at 540 F with the most reactive rod stuck out is 1.1 per cent ak/k. The shut-down margin at 520 F would be 0.76 per cent ak/k which is adequate to pre-vent return to criticality. In addition, the high pressure injection system will be automatically actuated on low reactor coolant pressure following a large area steam line failure. This system supplies borated water to the reactor coolant system to further increase the shutdown margin. Boron addition to the reactor coolant during the controlled cooling of the system to atmospheric pressure will prevent criticality at lower temperatures. The environmental consequences from this accident are calculated by assuming that the unit has been operating with steam generator tube leakage. The re-actor coolant activity assumes prior operation with 1 per cent failed fuel rods. With these assumptions, the steam generasors 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 unit can be cooled down and the leakage terminated. This additional leakage corresponds to 3.4 equivalent curies of iodine-131. The iodine is asa2med to be released directly to the atmosphere where it mixes in the wake of the building structures. The total integrated dose to the thyroid and the whole body dose at the one mile ex-clusion distance as a result of this accident are given in Table 14-1. The , total release of all activity when averaged over a year is less than the ' allowable limits of 10 CFR 20. 14.1.2.10 Steam Generator Tube Failures 14.1.2.10.1 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 air ejector. 14.1.2.10.2 Analysis and Results In analyzing the consequences of this failure, the following sequence of events is assumed to occur:
- a. A double-ended rupture of one s team generator tube occurs with unrestricted discharge from ea'ch end.
- b. The initial leak rate, approximately 435 gpm, exceeds the normal makeup of 70 gpm to the reactor coolant system, and system pressure decreases. -
O ( , 14-18 (Revised 4-18-67 ) -
(D 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 con-tinues to decrease until the high pressure injection system is actuated at a pressure of 1,800 psig. The capacity of the high pressure injection system is sufficient to compensate for the leak-age and maintains 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 valve will close. If a reactor coolant to secondary system leak hac occurred, steam line pressure will increase, opening the steam bypass valves to the con-denser. Each bypass valve actuates at a lower pressure than do the safety valves. The reactor coolant that leaks as a result of the tube failure is condensed in the condenser. Only the fission prod-ucts that escape from this condensate are released to the atmosphere.
- e. The af fected steam generator will be completely isolated and the leakage terminated when the reactor coolant system pressure falls below the setpoint of the steam bypass valves, i.e. , 975 psig.
Cooldown continues with the unaf fected steam generator until the temperature is reduced to 250 F, af ter which cooldown to ambient k)% conditions is provided by the decay heat removal system.
- f. At the design pressurizer cooling rate of 100 F per hour, depres-surizationto975psigreguiresapproximately1.7 this time period 1.6 x 10 cc (5,650 f t3 ) of reactor coolant leaks hours. During i
to the secondary system. This leakage corresponds to 45,800 curies of xenon-133 if the reactor has been operating with 1 per cent fuel defects. The radioactivity released during this accident is discharged through the tur-bine bypass to the condenser and then out the station vent. It is conserva-tively assumed that the station vent discharge mixes in the wake of the build-ing structures tition factor of rather 104 is than remaining assumed at itsinelevated for iodine release (goint. the condenser. A par-
, 2) Noble gases are assumed to be released directly to the station vent. The total integrated thyroio dose and the whole body dose as a result of this accident are lis ted in Table 14-1.
I 14.2 STANDBY SAFEGUARDS ANALYSIS I 14.2.1 SITUATIONS ANALYZED AND CAUSES In this section accidents are analyzed in which one or more of the pro-tective barriers are not effective and standby safeguards are required. All
- accidents evaluated in this section are based upon the ultimate power level of 2,568 mwt rather than the design power level of 2,452 mwt as the high ,
power level is used as the basis for design of the reactor building and l ( f)S engineered safeguards system. Table 14-3 summarizes the potential accidents studied. 14-19 (Revised 4-18-67) _, 1
O Table 14-3 Situations Analyzed and Causes Event Cause Effect Fuel Handling Accidents Mechanical damage Integrated dose at ex-during transfer. clusion distance is 0.98 rem thyroid and 0.099 rem whole body. - Rod Ejection Accident Failure of control Some clad failure. rod drive assembly Thirty-day dose at ex-pressure housing. clusion distance is 1.1 rem thyroid. Loss-of-Coolant Accident Rupture of reactor No clad melting. Thirty-coolant systen. day dose at exclusion distance is 0.015 rem thyroid. Loss-of-Coolant Accident Rupture of reactor No clad melting. Thirty-Effect of Purging coolant system. day dose at exclusion distance is 0.075 rem thyroid. Maximum Hypothetical 100 per cent melting Two hour dose at exclu-Accident of reactor core. sion distance is 227 rem thyroid, 1.7 rem whole body. Thirty-day dose at low population dis-tance is 75 rem thyroid, 3.0 rem whole body. i Maximum Hypothetical 100 per cent melting Two hour dose at exclu-Accident, Engineered of reactor core. sion distance is 1.55 rem Safeguards Leakage thyroid. Thirty-day dose at low popule; ion dis-tance is 12 rem thyroid. l l TABLE 14-3
,^ C . 14-20 (Revised 4-18-67) -
345
14.2.2 ACCIDENT ANAI,YSES f 14.2.2.1 Fuel Handling Accidents 14.2.2.1.1 Identification of Accident ! Handling of spent fuel assemblies is performed entirely underwater. Prior to refueling, the reactor coolant and the fuel transfer canal water above the re-actor are increased in boron concentration so that, with all control rods re-moved, the kefg of a core is no greater than 0.98. In the spent fuel storage pool, the fuel assemblies are stored underwater in storage racks having an eversafe geometric array. Under these conditions, a criticality accident dur-ing refueling is not considered credible. Mechanical damage to the fuel assem-blies during transfer operations is possible but improbable. This type of acci-dent is considered the maximum potential source of activity release during re-fueling operations. 14.2.2.1.2 Analysis and Results The fuel assembly is conservatively assumed to have operated at 29 mwt, twice the power level of an average fuel assembly. The reactor is assumed to have been shut down for 24 hr, which is the minimm time for reactor cooldown, re-actor closure head removal, and removal of the first fuel assembly. It is further assumed that the entire outer row of fuel rods, 58 of 208, suffers damage to the cladding. Since the fuel pellets are cold, only the gap activ-ity is released. The fuel rod gap activity is calculated using the escane rate coefficients and calculational methods discussed in 11.1.1.3. j l The gases released from the fuel assembly pass through the spent fuel storage pool water prior to reaching the Auxiliary Building atmosphere. As a minimm, 1 the gases pass through 10 ft of water. Although there is experimental, evidence j that a portion of the noble gases will remain in the water, no retention of l
- noble gases is assumed. For the iodine, 99 per n of that released from the j fuel assembly is aasumed to remain in the water. ' The total activity re- i leased to the building atmosphere is therefore Iodine 28.4 curies l Noble gases 2.79 x 104 curies The Auxiliary Building is ventilated and discharges to the station vent. The !
discharge from the station vent is assumed to mix in the wake of the building l structures rather than remain at its elevated release point. This assumption produces less favorable dilution and, therefore, higher ground concentrations at the exclusion distance. Also, with this assumption, the doses at the exclu-sion distance are essentially the same whethe'r or not the ventilation system is operating. The activity is assumed to be released as a puff from the station vent. Atmo-spheric dilution is calculated using the 2-hour meteorological model, and a breathing rate of 3.47 x 10-4 m3/sec is assumed. The total integrated dose to the thyroid and the whole body dose at the one rdile exclusion distance is given in Table-14-3. In evaluating the sensitivity of this analysis, the thyroid dose at the one mile exclusion distance is directly [iroportional to the quantity of g iodine released.
,[/.
'~
14-21 (Revised 4-18-67) 346 1
14.2.2.2 Rod Ejection Accident 14.2.2.2.1 Identification of Accident Reactivity excursions initiated by uncontrolled rod withdrawal (14.1) were shown to be safely 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 assembly housing or control rod drive nozzle must occur. Failure in the drive assembly upper pressure housing can cause a pressure differential to act on a control rod cluster assembly and rapidly eject the assembly from the core region. The power excursion due to the rapid increase in reactivity is limited by the Doppler effect and terminated by reactor protecti"e system trips. 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 acci-dent will not further damage the reactor coolant system.
- a. Accident Bases The rod ejection accident is based on the following:
Worth of ejected rod 0.27. Ak/k Rod ejection time 0.150 see Reactor full power level 2,568 mwt Reactor trip delay 0.3 sec . 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 ot greatest worth is the first of the entire rod pattern to be withdrawn from the core. The worth of this rod can be as high as 30 per cent (37. Ak/k) of the total pattern worth (107. Ak/k) . The details of control rod worth calculations, 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 subcritical, 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 i from the core. Therefore, rod ejection when the reactor is sub-critical, and when all other rods are in the core, does not cause l a nuclear excursion. As criticality is approached, the worth of the remaining control rods decreases. At criticality, rod ejection would result in c maximum reactivity addition of 0.57.Ak/k. At full power, but before equilibrium xenon is established, the total rod pattern worth remaining in the core is 2.47.A k/k. At equi-librium xenon the pattern worth is 1.47. Ak/k. Prior to establishing 14-22 (Revised 4-18-67) 347 w ~
4
,r )
{ / V , equilibrium xenon, the greatest single control rod worth is 0.15 per cent ak/k. A single rod worth of 0.2 par cent Ak/k is assumed for analysis of this accident at this time. f 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, control rod travel time to the top of the active region of the core is calculated to be 0.176 seconds. To account for the S-shaped reactivity worth versus, position of the rod, an ejection time of 0.150 seconds (75 per cent of active core height) is used in the analysis,
- b. Fuel Rod Damage Criteria j
Power excursions caused by reactivity disturbances of the order of magnitude occurring in rod ejection accidents could lead to three potential modes of fuel rod failure. First, for very rapid and large transients in which there is insufficient time for heat transfer from fuel to cladding, fuel melting followed by vaporiza-tion can generate destructive internal pressures without increasing p cladding temperatures significantly. The second mode occurs when i d the internal vapor pressure is not sufficient to cause cladding rupture, but subsequent heat transfer raises the temperature of the cladding and weakens it until failure occurs. The third mode occurs when the nuclear excursion has insufficient energy to cause significant melting of the fuel, but subsequent heat trans-fer to clad from fuel may cause excessive cladding temperatures. In all three cases there is a possible occurrence of metal-water reactions. However, only very rapid and large transients will generate a rapid pressure buildup in the reactor coolant system. The energy required to initiate UO fuel meltin gm, based on an initial temperature 2 of 68 F.(5)gThe is heat 220 to of 225 cal / fusion requires an additional 60 cal /gm. Any further energy addition vaporizes 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 420 cal /gm have been calculated to be necessary before the bursting pressure of cladding is exceeded. The lower limit for producing significant fuel vapor pressure (14.7 psi) is a25 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 lower limit, the potential for bursting of clad-ding and release of molten fuel to the reactor coolant is conserv-atively set at a fuel enthalpy of 285 cal /gm in this evaluation. fS For power excursions with energy bursts below 285 cal /gm, zirconium-(d 6 water reactions are possible. A correlation of the TREAT experiments presents a method of correlating the potential zirconium-water reac-tion as a function of fission energy input.(7) These data are b sed 34$ L ,i 14-23 (Revised 4-18-67) 2
on initially cold (room temperature) fuel rods, but are alco corre-lated as a function of peak adiabatic core temperature. This cor-relation can be used either by co=puting the core temperature, or by adding the initial steady state fuel enthalpy to the nuclear energy burst and obtaining an equivalent final fuel enthalpy. Ac-cordingly, a zirconium-water reaction requires a mini =um fuel en-thalpyof125 cal /gm. Increasing fuel enthalpies cause a linear in-crease in the percentage of the reaction, which may be approximated by the following for=ula:
%Zr-E20 Reaction = 0.125 (Final Fuel Enthalpy - 125)
It is assu=ed that DNB vill take place when the clad reaches a heat 2 flux of 6 36 x 105 Btu /hr-ft. At this heat flux the hot fuel rod enthalpy would be approx 1=ately 140 cal /gm at EOL and 130 cal /gm at BOL. App 41ng 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 above values can be calculated. Any fuel rod exceeding the above enthalpy values is assu=ed to fail from overheating and re-leases the gap activity of that fuel rod. 14.2.2.2.2 Method of Analysis The hypothetical control rod ejection accidenth vas investigated using the exact 1-dimensional WIGL2 digital ec=puter program. w): OIt was found that the point kinetics analo5 model results agreed with the WIGL2 results to within 10 per cent for rod worths up to 0 75% Ak/k. The point kinetics model assumes an initial flux distribution which is undisturbed by local control rod cluster asse=blies. The space-dependent model, however, has significant flux depressions in the vicinity of control rods. Although the flux throughout the core begins to increase shortly after the start of the rod ejection, the flux increase in this depressed region rises more quickly such that by the time 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 eject-ed red. The entire reactor flux then rises unifomly until the Doppler effect teminates the excursion. Thus by applying the peak-to-average flux factors (2 92 maximum) as discussed in 3 2 3 to the point kinetics results, an accurate assessment can be cade of the peak and integrated flux at any point in the re-actor. 14.2.2.2 3 Analysis and Results
- a. Source Power The ejection of the control rod at source power vill result in a power excursion which is teminated by a low reactor coolant system pressure trip. This analysis was perfomed with the core 0 5% Ak/k suberitical so that a total rod worth of 1% Ak/k was withdrawn in 0.150 sec. The reactor power was initially at 10-9 of full power.
The ice pressure trip occurs at 1 7 see after the ejection starts, and the reactor power is teminated at a peak value of 39 per cent full power. This peak neutron power value is not reached until a-bout 15 see after the rod is ejected because Doppler feedback con-trols the rate of rise and magnitude of the neutron power. There-pp fore, a low pressure trip will teminate the accident before 14-24 hk9
significant power is generated, due to the loss of coolant D through the rupture. An analysis was perfor=ed for the above accident without a low pres-sure trip to demonstrate the capability of the reactor to accept the accident. In this case the neutron power reaches 1,000 mwt (39 per cent fun pover), and the peak fuel temperature is 990 F. This is far below the melting temperature of UO2 , and the resultant ther=al power is only 16 per cent of full power. Hence, no fuel damage would result from the rod ejection accident at source power level.
- b. Full Power For the full power case at beginning-of-life (BOL), the ejection of a single control rod vorth 0.2% ok/k would result in virtual 4 no Zr-H2O reaction and no DIG (see F1 6 ures 14-20 and 14-21). The hot fuel rod would reach a peak enthalpy of about 130 cal /gm.
For the end-of-life case (EL), the reactor neutron power peaks at 4,150 cvt 200 milliseconds after the start of ejection of a 0.2% Ak/k control rod. The minimum reactor period during the transient is 160 milliseconds. The prompt negative Doppler effect teminates the power rise, and control rod insertion from high flux signal ter-minates the excursion. The total neutron energy burst during the transient is approximately 2 The final fuel enthalpy of
% ./ the nominal rod is 112 cal /gm,500 mw-sec.
i.e., the enthalpy of the hot rod is 157 cal /gm. This enthalpy is considerably below the minimum range (220to225 cal /gm)forcentralfuelmelting. As a result of the excursion, approximately 10 per cent of the core would have' DIG (see Figure 14-20). The power distribution at the beginning of core life, with the higher power peaking factors shown in 3 2 3, was used to de;emine the dis-tribution of the energy of the excursion. With this distribution of fuel enthalpies, and using the TREAT correlation, 0.4 per cent of the circonium cladding may react (see F1 6 ure 14-21) to contribute an ad-ditional 510 mv-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 da=a61n8 pressure pulses are produced in the system. As a result of the postulated pressure housing failure, . reactor cool-ant is lost'from the system. The rate of mass and energy input to the reactor building is considerably lover than that for the 10 in. ID rupture which is discussed in 14.2.2 3 This lover rate of energy input results in a lover reactor building pressure than that obtain-ed for the 10 in rupture. The environmental consequences from this accident are calculated by conservatively assuming that all fuel rods which undergo a DIG result n in clad failure and subsequent release of the gap activity. Actually, t
} :ost of the fuel rods vill recover from the DIG, and no fission prod-
\_/ uct release vill occur. For the case of a 0.2% Ak/k rod ejection C' 14-25 350
from full power at the end-of-life, 10 per cent of the fuel rods are O assumed to fail, releasing 133,000 equivalent curies of I-131 to the Reactor Building. Fission product activities for this accident are calculated using the methods discussed in 11.1.1.3. The total inte-grated dose to the thyroid from this accident is calculated using the environmental models and dose rate calculations discussed under the loss-of-coolant accident. The resultant dose given in Table 14-3 is more than two orders of magnitude below the guideline valses va.lves of 10 CFR 100.
- c. Sensitivity Analysis A sensitivity analysis was performed on this control rod ejection accident, and the results are shown in Figures 14-22 through 14-30.
Figure 14-22 shows the variation in the peak neutron power as a function of the worth of the ejected control rod. For the nominal 0.27. ok/k case from full power the peak power is only in the order of 150 per cent, again assuming a low pressure trip does not occur. The rod ejection from source level results in a Doppler turn-around before the flux trip is reached. Figure 14-23 shows the variation in the corresponding thermal power with control rod worth. Figure 14-24 shows the corresponding enthalpy increase of the hot fuel rod versus control rod worth. Note the very small spread in values for the BOL and EOL full power conditions. As expected, the enthalpy increases with rod worth. Figures 14-25 through 14-28 show the peak reactor neutron and ther-mal powers as a function of changes in the positive moderator , temperature coefficient and negative Doppler coefficient for the l nominal 0.57. ok/k control rod ejection from source level. There was only insignificant variation (< 27.) of these parameters with changes in the two reactivity feedback coefficients in the nominal l 0.27.ak/k rod ejection from full power. Figure 14-29 shows the change in nominal thermal power with varia-tions in the trip delay time for the nominal 0.27.ak/k rod ejection from full power (the variation from zero power is negligible). The trip delay time does not affect the peak neutron power. Figure 14-30 shows the corresponding change in the total enthalpy increase of the hot fuel rod versus the trip delay. The thermal power never exceeds 112 per cent full power for any of I the variations studied using the nominal rods (0.27. ak/k for full power and 0.57.ak/k for source level). The hot fuel rod average l temperature never increases by more than 220 F above the full power I peak value (4,090 F) . It is therefore concluded that each of these parameter variations has relatively little effect on the nominal results. O
;5 351 o - 14-26 (Revised 4-18-67) l l
l l l
"N ) .14.2.2.3 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 amount would be boiled of f owing to residual heat, fission product decay heat and possible heat from chemical reactions unless an alternate means of cooling were available. In order to prevent 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 Bases All components of the reactor coolant system have been designed and fabri-cated 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 j described in Section 4. In addition to the high-integrity system to minimize the possibility of a
- loss of coolant , emergency core cooling is provided to insure that the core does not melt even if the reactor coolant system should fail and release the
'g coolant. This emergency core cooling is provided by the core flooding system,
, ,,) the high pressure injection system and the low pressure injection system. ! These systems are described in detail in Section 6, and their characteristics are summarized in the following paragraphs. The core flooding system has two independent core flooding tanks, each of which is connected to a different reactor vessel injection nozzle by a line containing two check valves and a normally open, remotely operated stop valve. Since these tanks and associated piping are missilc-protected and are connected directly to the reactor vessel, a rupture of reactor coolant system piping will not affect their performance. 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 thermal shield. The core is flooded from the bottom upward. The combined contents of the two tanks (1,880 ft 3 ) 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, supplies coolant at pressures up to the design pressure of the reactor coolant system and at a rate up to 1,500 gpm. Low pressure injection actuated by low reactor coolant system pressure supplies coolant at pressures below 100 psig and at a rate up to 9,000 gpm. Both of these systems can operate at full ; capacity from the on-site emergency electrical power supply and will be in i l operation within 25 sec after the accident. In the reactor vessel, decay heat is transferred to the injection water.
)
i i [,4 ti 14-27 (Revised 4-18-67) __ .- j I
Injection water is supplied from the borated water storage tank. When this tank empties, water is recirculated 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 pres-sure in the building. Reactor Building sprays, actuated on a high building pressure signal of 10 psig, deliver 3,000 gpm to the Reactor Building atmo-sphere. This spray water reaches thermal equilibrium within the building atmosphere during its passage ftam the nozzles to the sump. Spray water is supplied from the borated water storage tank until it is emptied. Thereafter, water collected in the sump is recirculated to the sprays. Cooling is also provided by the Reactor Building emergency cooling system in which recircu-lating 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 cooling system. The heat removal capacity of either of these two Reactor Building cooling systems is adequate to prevent overpres-surization of the building during a loss-of-coolant accident. This analysis demonstrates that in the unlikely event of a failure of the reactor coolant system, both of the other two boundaries that prevent fission product release to the atmosphere, ie, the reactor core and the Reactor Building, are protected from failure. Accordingly, the public would be pro-tected against potential radiation hazards. In order to evaluate this accident, a range of rupture sizes from small leaks up to the complete severance of a 36 in. ID reactor coolant system line has been evaluated. A core cooling analysis is presented for the complete severance of the 36 in. ID reactor coolant piping. Since the large rupture removes the least amount of stored energy from the core, this represents the minimum temperature margin to core damage and, therefore, places the most stringent requirements on the core flooding system. The Reactor Building pressures have been evaluated for a complete spectrum of rupture sizes without the action of core flooding tanks and, therefore, are conservative. The peak pressure occurs for a 3 ft2 rupture rather than for a 36 in. ID (14.1 f t2) rupture. Rupture sizes smaller than the 36 in. ID leak result in longer blowdown times, and the amount of energy transferred to the Reactor Building atmosphere is increased. 14.2.2.3.3 Accident Simulation
- a. Hydraulic Model Blowdown of the reactor coolant system following an assumed )
rupture has been simulated by using a modified version of the l FLASH (9) code. This code calculates transient flows, coolant mass and energy inventories, 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. 14-28 (Revised 4-18-67) 353 i 1
I-4
)
,/
j. Modifications were made to FLASH to make the model more applicable i 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.
I (2) Core flooding tanks have been added. Water flow from the j core flooding tanks is calculated on the basis of the pres-sure difference between the core flooding tanks and' the I point of discharge into the reactor coolant system. The line resistance and the inertial effects of the water in the pipe
;l are included. The pressures in the tanks are calculated by assuming an adiabatic expansion of the gas above the water I level in the tank. Pressure, flow rate and mass inventories ! are calculated and printed out in the computer output.
(3) Additions to the water physical property tables (mainly in the subcooled region) have also been made to improve the accuracy of the calculations. (4) A change in the steam bubble rise velocity has been made from i the constant value in FLASH to a variable velocity as a func-
! tion of pressure. . The bubble velocity term determines the 4 amount of water remaining in the system af ter depressurization 4
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 FLASE code. For smaller ruptures, an appreciable difference exists. The variable bubble velocity i is based on data in Reference 10 and adjusted to correspond j to data from the LOFT semiscale blowdown tests.
Test No. 546 from the LOFT semiscale blowdown tests is a j typical case for the blowdown through a small rupture area. A comparison of the predicted and experimentally observed pressures is shown in Figure 14-31. Figure 14-32 shows the per cent mass remaining in the tank versus time as predicted i by the code. At the end of blowdown, the predicted mass re-maining is 13 per cent. The measured mass remainingLis approx-
- imately 22 per cent.
l The FLASH code describes the reactor coolant system F. the use of i two volumes plus the pressurizer. The system was grouped into two t volumes on the basis of the temperature distribution in the system i as follows: < Volume 1 includes half of the core water' volume, the reactor
! outlet plenum, the reactor outlet piping and approximately 55 per cent of the steam generators.
i Volume 2 inclodes half of the core water volume, the reactor inlet plenum and downcomer section, the reactor inlet piping, i pumps and 45 per cent of the steam generators.
') - - -
~
.}
14-29 (Revised 4-18-67)
,.i..%.-. ,. . - - -. . , p- -
%,, g.i%-. - q w e
Volume 3 represents the pressurizer. O The resistances to flow were calculated by breaking the reactor ecolant system into 24 regions and calculating the volume-weighted resistance to flow for a given rupture location based on normal flow resistances. 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 deter-mined by the CHIC-KIN code in conjunction with the FLASH output. Steam generator heat removal was assumed to cease when the rupture occurred. While the modified FLASH code now has the capability of simulating injection flow from the core flooding tanks, the calculations shown in this report were made prior to the time that the core flooding simulation was added to FLASH. Core flooding was calcu-lated by taking the reactor vessel pressure' as predicted by FLASH without core flooding and using this pressure as input to a sepa-rate program to get the flow from the core flooding tanks. Reac-tor 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 steam by the cold injection water is not taken into account. A more recent analysis using tne FLASH code confirms that conservatism was used in this analysis. 3 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 (CONIJMPT).
- b. Core Thermal Model The core heat generation and heat transfer to the fluid are de-pendent upon the blowdown process. The FLASH program includes a core thermal model and the feedbacks of heat transfer and flow on each other. While the FLASH thermal model is acceptable for determining the effect of core heat transfer on the blowdown pro-cess, a more extensive simulation is necessary for evaluation of the core temperature transient.
Additional analytical models and a digital computer program were i developed to simulate the core thermal transient for the period beginning with the initiation of the leak and ending after the core temperature excursion had terminated. The model includes the effects of heat generation from neutrons : l before reactor trip, neutron decay heat, and fissi.n and O
- s. 9 355
~'
14-30 (Revised 4-18-67)
l
, I i
l' activation product decay heat; the exothermic zirconium-water i 3 reaction based on the parabolic rate law; heat transfer within the fuel rods, limited heat convection from the fuel clad sur- l face to any fluid within the core region, heat transfer from i 3 reactor vessel walls and internals to the coolant, and heat , l transfer from fuel rods to the steam necessary to sustain a l metal-water reaction; and emergency injection flow and boiloff.
- The basic model structure provides 50 equal-volume core regions l with input provisions to allow any choice of power distribution.
; The model may be used to simulate the entire core or any sub-division of the core. ~herefor e, the core geometry may be de-tailed to the degree consistent with the results desired.
The following parabolic law for the zirconium-water reaction l equation (ll) with the following constants is simulated for each of the regions: 1 f lexp O_3\ A= E dt (ro - r) RT where: r = radius of unreacted metal in fuel rod s r, = original radius of fuel rod t = time i K = rate law constant (0.3937 cm2 /sec) i LE = activation energy (45,500 cal / mole) 4 R = gas constant (1.987 cal / mole K) T = temperature, K i The zirconium-water reaction heat is assumed to be generated com-pletely within the clad node. The heat necessary to increase the steam temperature from the bulk temperabtre to the reaction tempera-ture is transferred from the clad at the point of reaction. The above equation implies no steam limiting. However, the program does have provision for steam rate limiting to any degree desired, but no steam limiting of the reactions has been assumed. All heat i from neutron, beta, and gamma sources is assumed to be generated . I within the fuel according to the preaccident power distribution and infinite irradiation. Within each of the regions there is a single fuel node and a single clad node with simulation of' thermal resistance according to the normal fuel rod geometry. Provision is made to simulate four differ-I\ ent modes of heat transfer from the clad node to the fluid sink node
\s_s/ by specifying the time-dependent surface coefficient.
14-31 _(Revised 4-18-67) --- -
\Ji
+bL
- .,c - -
O The surface heat transfer coefficient input data are determined from calculations which are based on flow and water inventory as furnished from the blowdown and the core flooding tank 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 unless otherwise specified): Average fuel temperature of each region. Average clad temperature of each region. Per cent metal-water reaction in each region. Time for the clad of each region to reach the metal-water threshold, the beginning and end of melting, and the slump temperature. ' Heat transferred to the Reactor Building from the core. Heat generation by hydrogen and oxygen recombination. , Total zirconium-water reaction. Total heat stored in metal sinks.
- c. Reactor Building Pressure Model The Reactor Building pressure-temperature analysis is performed us-ing the digital computer code " CONTEMPT" developed by Phillips Petroleum Company in conjunction with the LOFT project. This pro-gram and its capabilities are described in Reference 12. With minor modifications this program was adapted for use on The Babcock
& Wilcox Company's computer.
In this model, the Reactor Building is divided into two regions: the atmosphere (water vapor and air mixture) and the sump region (liquid water). Each region is considered to be well mixed and in thermal equilibrium, but the temperature of each region may be different. The Reactor Building and its internal structures are subdivided into five segments, as shown in Table 14-4, and treated as slabs with 1-dimensional heat transfer. Each segment, divided into several heat conducting subregions, may act as a 14-32 (Revised 4-18-67) 357
heat source or sink. The program includes the capability of j cooling the Reactor Building atmosphere by air coolers (Reactor Building emergency cooling units) and spray cooling (Reactor Building spray system), and cooling the liquid region by sump l coolers (low pressure injection coolers). l
)
During blowdown, mass and energy are added directly co the atmo-sphere where the liquid water present is assumed to fall to the 1 liquid region. After blowdown is over and emergency injection has been initiated, mass and energy are also added directly to I 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 cf sump water is simulated. The Reactor Building calculations are begun by computing steady-state results using initial atmcspheric conditions as the input. Following the rupture, the mass and energy addition is determined from the energy input rates for each time step. Heat losses or gains due to the heat-conducting slabs are calculated. Then the pressure and temperature of the liquid and vapor regions are cal-culated from the mass, volume, and energy balance equations. x- Table 14-4 Reactor Building Structural Heat Capacitance Segments Segment Description 1 Reactor Building Walls and Done 2 Refueling Cavity (Type 304 SS Liner - One Side) 3 Reactor Building Floor 4 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 safe'- guards. Therefore, the effectiveness of any given arrangement can be analyzed. L i fe'N , k-- i h[ 14-33 (Revised 4-18-67) 358
O 14.2.2.3.4 Core Flooding Tank Design Base Accident Analysis The 36 in. ID, double-ended pipe rupture produces the fastest blowdown and lowest heat removal from the fuel. This case therefore represents the most e'ri.ngent emergency core cooling requirements. Results from the modified ve2sion 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 cooling and condensing effect which reduces overall leakage. For the present analysis, no credit has been taken for the extra accumulation of water due to the condensing effect. The blowdown was analyzed using the version of FLASH without core flooding tank simulation. This resulted in higher transient reactor vessel pressure than would have occurred if core flooding tank flow feedback effects were included. The core flooding tank transient analysis was then performed using reactor vessel back pressure which was provided by the FLASH analysis. To determine the fluid remaining in the reactor vessel at any point in time during the blowdown transient, the integrated flow from the core flooding tank is added to the fluid remaining which is predicted by FLASH. The inventory obtained by this method is conservative because it neglects the condensing effect which leads to an additional accumulation of water. The SLUMP digital computer program, as described in 14.2.2.3.3.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 c >re was simu-lated in segments of L/4 of one per cent each. The hottest . egment was ' assigned a peaking factor of 3.1 times the average of the n tal core power density. The reactor is assumed to be initially at the ultimate power level (2,568 MWt). The core analysis simulates reactivity effects which correlate with the results which were obtained from a detailed analysis of void shutdown without control rod insertion. The detailed analysis was made with the digital computer program CHIC-KIN which included positive moderator and initially positive void coefficients. The results from the void shutdown calculation yield a total integrated neutron energy generation of approxi-mately 2.1 full-power seconds. 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 .nitial blowdown period pro-vide 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 approach was adopted by O V~' 359 14-34 (Revised 4-18-67) -
O assuming DNB at 0.25 sec. Nucleate boiling surface coefficients at high flow rates may exceed 50 000 Btu /hr-ft2-F. A nucleate boiling surface coefficient of 25,000 Btu /hr-ft -F b was used in the analysis. However, the series heat transfer from the clad node to the fluid sink is limited to 6,500 Btu /hr-ft2 -F by 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) corre-lation: 0.8 I 0.14 I 3 - B "B yN = 0.023 kD (N e)0.8(NPr) 1+ X h _ ( p pj_. ( lagj where bPF = two-phase film heat transfer coefficient, Btu /hr-ft -F k = fluid conductivity, Btu /hr-ft -F D = hydraulic diame*er, ft h N = eyn s n d er Re Np = Prandtl number x = quality P = density
= viscosity 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.
2 Figure 14-33 shows the core flow vs time for the 14.1 ft leak as calculated by FLASH. Figure 14-34 shows the clad surface heat transfer coefficient versus time based on the flow of Figure 14-33 and the modified Sieder-Tate equation. The straight line in Figure 14-34 indicates the surface heat transfer values which were used in SLUMP, and which are conservative as compared to the re-
. suits obtained from the Sieder-Tate equation.
n U t . . v1 14-35 (Revised 4-18-67) .
O In applying the Siedar-Tate equation constant values of bulk steam quality and temperature corresponding to the most conservative assumptions were used. A sensitivity analysis was made for maximum coefficients in SLUMP ranging from 400 to 2,000 3tu/hr-f 2t -F initially and decreasing to zero at the end of blowdown. Results are discussed below. After blowdown no core cooling is assumed until after core recovering starts. When the water level reaches the core bottom and starts to rise up on the core, the submerged portion will 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 determining 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. How-ever, in evaluating the core flooding tank design a conservative approach was used by assuming a value of 20 Btu /hr-f t 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 (400 to 1,000 psig), (b) ratio of nitrogen gas volume to total volume (1/3 and 1/2), and (c) size of piping between the core flooding tanks and the reactor vessel (12 in. and 14 in. ID). Figure 14-35 shows the reactor vessel water volume versus time for core flooding tanks operat-ing at 600 psig with different combinations of volume ratio and line size. This figure includes an allowance for boiloff and also shows the effect of the flow provided by high pressure and low pressure injection begintiing i at 25 see when emergency power is available. Similar curves for 400 psig and 1,000 psig core flooding tanks are shown in Figure 14-36. Figure 14-37 shows the maximum clad temperature reached by the hot spot and by the 1, 2, 3, 4 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 t he. highest point in the core at which the peaking factor corresponding to that percentile exists. The fact that the same i peaking factor may e:;ist at some lower point in the core provides an l inherent conservatism in the data as plotted. The axial peaking factor l profile for the beginning of core life was used as it imposes the most l stringent requirements on the core flooding tank design. b 14-36 (Revised 4-18-67)
O Peak temperatures for the core flooding systems described above are also shown on Figure 14-37. These curves demonstrate that all of the systems presented are capable of keeping the peak temperature at the hot spot more than 1000 F below the melting temperature of the clad. The amount of zirconium-water reaction which occurs for each of these core flooding systems is shown in Table 14-5. Thecogefloodingcankdesignselectedisfora600psichargepressure, 3 940 f t water, 470 ft of nitrogen, and a 14 in. supply line. The per-formance of this system in limiting core temperatures is approximately in the center of the range for the systems described. For this 600 psi charge pressure, Figure 14-37 indicates that the hot spot clad temperature would reach 1950 F at 17.5 sec and that less than 5 per cent of the core would exceed 1690 F. For this same case calculations indicate less than 0.005 per cent total zirconium-water reaction for the whole core. Table 14-5 Core Flooding Tank Performance Data Line Nitrogen Total Metal Size, Volume, Water Reaction Pressure In. % of Total % 400 14 33 .022 7- g 400 14 50 .009 g N-) 600 14 33 .005 ' 600 14 50 .002 600 12 33 .022 600 12 50 .010 1000 12 33 .003 1000 12 50 =0 Additional analysis was performed to evaluate the sensitivity of the maximum clad tempera ture to three important thermal parameters. All cases discussed below have in common the follow #.ng parameters: Leak size: 14.1 ft2 Time of DNB: 0.25 see Time at full power: 2 see Time that blowdown cooling ends: 9.5 see Core region: Hot spot Time to initiate quenching: 18 see Dependent variable examined: Clad temperature for hottest 5 per cent of core. Figure 14-38 shows the clad maximum temperature sensitivity to the initial surface heat transfer coefficient after the 0.25 see nucleate boiling period. The coefficient is linearly decreased to zero at 9.5 sec. Zero cooling is maintained until quenching is initiated with a clad surface coefficient of f-ss - 20 Btu /hr-ft 2 -F. Figure 14-38 shows that the assumed value of 1000 Btu /hr-f t2 -F is not on the most sensitive part of the curve and a 20 per cent decrease ( ) v 14-37 (Revised 4-18-67 ) )b .
in h would only result in increasing the peak clad temperature 120 F. Figure 14-39 shows hot spot clad temperature transients for a range of injection cooling coefficients. All cases have a clad surface coefficient of 1000 Btu /hr-ft2 -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 14-39 shows that without any cooling the temperature reaches the melting point in approximately 50 sec. An h value of 15 stops the fast temperature excursion and allows only a low rate of increase thereafter. Since the continuously increasing 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 low as 15 is probably adequate. An h value of 20 provides immediate quenching action and a slow cooling rate thereafter. An h value of 100 provides very fast cooling. Even though the 100 is a real-istic 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 re-ference for evaluating core flooding tank performance. Figure 14-40 shows hot spot clad temperature transients for a range of pool fluid sink temperatures. Parameters for heat. transfer prior to 18 sec are the same as discussed in the preceding paragraph. At 18 sec a surface coeffi-cient of 20 Btu /hr-f 2t -F was applied with sink temperatures as indicated. All results reported herein previously have had a sink temperature of 2SO F during the quenching period. Prior to quenching, the sink temperature in all cases is based on the transient fluid pressure thich reruits from the FLASH analysis. Figure 14-40 shows that any sink temperature below approxi-mately 500 F is adequate for holding or reducing the clad temperature which existed at 18 sec. The core flooding tanks will provide a high flow of water at 90 F. Although some heating will 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 cor; acted the core. Using a reference value of 280 F provides an added conservatism to the analysis. In conclusion, the analysis has shown that the design of the core flooding system will provide for covering approximately 80 per cent of the core at 25 see af ter the double-ended rupture of the 36 in. ID pipe. The high pres-sare and low pressure injection will then provide for continued increase in the water level. The clad hot spot temperature excursion is terminated at 1950 F and less than 5 per cent of the total cladding exceeds 1690 F. Only a minute amount (0.005 per cent) of zirconium-water reaction occurs, and the maximum temperature is at least 1400 F below the clad melting point. O c 363 3 14-38 (Revised 4-18-67) -
14.2.2.3.5 Reactor Building Accident Analysis - d
- a. Design Base Accident A range of leak sizes between 0.4 f t2 and 14.1 ft2 has been evalu-ated. The 14.1 ft2 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 14-6.
table 14-6 Reactor Operating Conditions for Evaluation Parameter Value Reactor Coolant System Pressure, psig 2,185 Reactor Coolant Average Temperature, F - 584 Reactor Power Level, mwt 2,568 Reactor Coolant System Mass, Ib 519,173 Initial Reactor Building Temperature, F 110 Initial Reactor Building Relative Humidity, % 0 Initial Reactor Building Pressure, psig 0 O Under normal conditions at rated reactor power, the nominal temper-ature of the steel and concrete in the Reactor Building is 104 F. In calculating the Reactor Building pressure, it was conservatively assumed that the average temperature of the building atmosphere and structural materials was 110 F. Upon release of hot reactor coolant, the steel and concrete act as heat sinks which reduce the Reactor Building pressure. The heat sinks considered in this analysis are specified in Table 14-7. This analysis assumes that the high pressure injection system had one of the three pumps available for operation and that the low pressure injection system had two of the three pumps available for operation. These systems are assumed to operate on emergency power and can be in operatian to deliver a total injection flow of 6,500 gpm lhin 25 sec after the accident occurs. In accordance with curru : practice, the core flooding tanks are not considered in this analysis. Including the core flooding tanks would decrease peak pressures. During blowdown,ttass and energy releases to the Reactor Building are calculated by FLASH. Figure 14-41 is a plot of mass released to the Reactor Building, and Figure 14-42 is a plot of reactor coolant average pressure, each calculated by FLASH for the spectrum of hot leg ruptures. Following blowdown, a 20-
~h (O
,- 14-39 (Revised 4-18-67)
\ _
O Table 14-7 Reactor Building Structure Data for Analysis of Time-Dependent Reactor Building Pressure Parameter Value Reactor Building Free Volume, ft3 1,900,000 Exposed Liner Plate Surface, ft2 85,340 Mass, Ib 864,900 Dome and Wall Liner Thickness, in. O.25 Refueling Cavity Liner Thickness, in. 0.1875 Reactor Building Concrete Enclosure Consisting of a 3-ft, 3-in. Thick Dome and 3-ft, 9-in. Thick Walls and a 2-ft Thick Floor Vall and Dome Surface, ft 2 78,570 Wall and Dome Mass, lb ho,888,000 Exposed Floor Surface, ft2 8,200 Exposed Floor Mass, lb 2,820,000 Structural and Miscellaneous Steel Exposed to Reactor Building Atmosphere Surface, it2 22,200 Mass, lb 8hh,000 Internal Concrete Surface, ft 2 81,230 Mass, lb 16,192,000 l Refueling Cavity Concrete l Surface, ft 2 13,540 Mass, lb 4,008,000 i E O 14-ho (Revised 5-25-67) 2
region SLUMP model was used to simulate the core thermal transient. This simulation includes fuel heat deneration, metal-water reaction, and quenching when the injection water provided cooling by contact with the core. As any given segment reached 4800 F it was assumed to drop int y water below the core and release all heat down to a datum of. m F. Also, it was assemed that 10 per cent additional zirconium-water reaction occurred. When the water covered approximately 25 per cent of the core, the surface heat transfer coefficient from all the core clad to the water was assumed to be 100 btu /hr-ft 2 -F. The determination of water level was based on 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 period, a 2 surface heat transfer coefficient of 1000 beu/hr-ft -F was used. 2 Af ter blowdown this coefficient was changed to 100 btu /hr-f t -F for the metal below the leak and 5 btu /hr-ft -F above the leak. The coolant sink temperature gas, provided by FLASH for the blowdown
'N period and assumed to be M E F thereafter. The internal heat trans-d fer of the metal was based on a multilayer finite difference model.
The whole process of react 'r coolant system metal heat transfer was simulated with a digital computer program. All heat transferred from the core and the reactor coolant system metal was assumed to generate steam without taking credit for the subcooled condition of the injection water (except for that water which was boiled off*y 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 Building atmosphere terminated. No delay was assumed in l transporting steam to the Reactor Building. The heat from hydrogen burning was added directly to the Reactor Building as hydrogen was evolved from the metal-water reaction. Both reactor inlet (cold) and reactor outlet (not) line breaks were analyzed with FLASH. However, a comolete analysis was made only for the hot line breaks since they provided fo: the most rapid heat - transport from the core. This was true because the hot line breaks had longer blowdown and better heat transfer during blowdown than did the cold line breaks. The results of calculations of fluid and heat transport to the . Reactor Building as determined by FLASH, SLUMP and other analytical models were used as input to the Reactor Building; pressure analysis program, CONTEMPT. [ D l \ 0 14-41 (Revised 4-18-67)
} l l
l
Heat transfer from the Reactor Building atmosphere to the steel liner O was calculated using a condensing coefficient of 620 btu /hr-f t27 until a total heat input of 110 btu /ft2gadbeenachieved. Thereafter, a condensing coefficient of 40 btu /hr-f t -F was used. For heat transfer from the Reactor building atmosphere to the concrete, a condensing coefficient of 40 btu /hr-f 2t -F was used. For heat trans- ! l fer from the sump water to the concrete floor a coefficient of 20 btu / hr-ft2 -F was used. No credit was taken for heat transfer to reinforc-ing steel in the internal concrete structures. For structural and miscellaneous steel, one heat transfer section with an equivalent thickness of 0.93 in, was used. Condensing coefficients of 620 and 40 btu /hr-ft2-F were . sed. Following a loss-of-coolant accident, the Reactor Building is cooled by three Reactor Building emergency cooling units and a spray system. The heat removal ennability of the spray system is at least equal to that of the thre. actor Building emergency cooling units. Each system is designed so that it alone can protect the Reactor Building against overpressure. Initially coolant for the building sprays and for injection to the core is pumped from the borated water storage tank. When water from the borated water storage tank is depleted, the water collected in the Reactor Building sump is recirculated through the Reactor Building sprays and through the low pressure injection coolers to supply the low pressure injection water. The result is an increased injection and spray water temperature. No boiling of the injection water results from this decrease in subcooling. The Reactor Building spray effec-tiveness will decrease. The net result is a decrease in the energy removal rate from the Reactor Building atmosphere. Three Reactor Building emergency cooling units provide heat removal capability of 240 x 10 6btu /hr at a vapor temperature of 281'F. Two cooling units plus 1500 gpm sprays, or 3000 gpm sprays, provide cooling that is at least equivalent to the three Reactor Building emergency cooling units. Each system was assumed to operate on emergency power and was delayed until 25 seconds after the rupture occurred. i ! Figure 14-43 shows the Reactor Building pressure for c lete sever-ance of a 36 in. ID reactor coolant system pipe (14.1 f t rupture area) with 6500 gpm of borated water injection into the reactor coolant system beginning 25 seconds after the rupture. Reactor Building cool-ing is provided by three emergency cooling units. The peak pressure resulting f . this accident occurs l'e4 seconds after the rupture at a value of . psig. Figures 14-44 through 14-48 show the Reactor Building pressure for the other rupture sizes analyzed with the same cooling capability as the 14.1 ft2 rupture above. A summary of the input parameters and results for the spectrum analysis are tabulated in Table 14-8. 14-42 (4-18-67) 0l
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A3f 2 rupture area results in the highest postaccident Reactor Building pressure (see Figure 14-45). Figuras lk k9 and 14-50 show '.he Reactor Building energy inventory as a fune-tion of time after rupture far 14.1 and 3 ft2 rupture areas with three e=er-gency coolers operating. These curves show the effectiveness of the Reactor Building structures and e=ergency cooling units. Figures 14-51 and 14-52 show the Reactor Building vapor te=pe.atures and su=p temperatures following 14.1 and 3 ft2 ruptures. The peak Reactor Building pressure shown in this evaluation for the spectrum of leak sices results in 56.8 psig and is the result of a 3 fta rupture in the reactor outlet piping. The Reactor Building design pressure is 59 psig, and a desi,3a margin of 2 psi exists. With core ficoding tark operation this car;in would be increased further. Criterion 17 of the AEC 3eneral Design Criteria states that the containment (Reactor Building) be designed to acco==cdate the largest credible energy re-lease including the effects of credible metal-water reactions uninhibited by active quenching systems. Although the evaluation of the e=ergency injection systems demonstrates that only a.small a=ount of =etal-water reaction can oc-cur, the case of no injection flow has been evaluated in response to the above criterion. This case assumed that, after blowdown, the reactor vessel vould have water up to the bottom of the core. The core w allowed to heat up by decay heat and =etal-water reaction heat. Stea= flow rate-limiting of the reaction was not considered so long as any water was assumed to be in the vessel. If and when the clad reached the =elt-ing te=perature, it was assumed that the whole region slumped into the bottom of the vessel with the attendant reaction of 10 per cent = ore of the re=ain-ing zirconium and with the release to the Reactor Building of all sensible I and latent heat above 286 F. Upon co=pletion of boiloff, heat input to the Reactor Euilding was assumed to cease. Figure lk-53 shows a Reactor Building pressure of 56.7 psig at 220 ; seconds, the ti=e at which the reactor vessel boils dry. This peak pressure ! is below the 59 psig design pressure of the Reactor Building. i l l l l l l l l l l l O 6 14 kk (Revised 5-25-67) i l
9
)
- b. Reactor Bufl. ding Zirconium Reaction Capability In order to determine the theoretical ultimate zirconium reac-tion capability of the Reactor Building a series of hypotheti-cal accidents was investigated.
Blowdown was based on the 14.1 ft2 leak case. Heat transfer from the core and all reactor coolant system metal below the leak height was assumed to transfer to a ikkf F sink based on a surface coefficient of 50,000 Btu /hr-f t2-F. For reactor cool-ant system metal above the leak height 5 Btu /hr-ft2-F was used. Available core heat consisted of the initial stored heat, the equivalent of two rated power seconds, decay heat, and metal-water reaction 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 superheat at the rate of hy-drogen production from the zirconium-water reaction.
A series of calculatibns for each of the various cooling capa-cities was made varying the energy input rate, i.e., Zr-H 2O reaction rate. A 1 per cent per second zirconium-water reac-g tion produces 1.173 x 106 Btu /sec of metal-water energy and O.902 x 106 Btu /sec hydrogen recombination energy. In all csses the energy was input at a linear rate beginning 10 sec after the rupture. The emergency cooling units and spray coolers were started 25 sec after the rupture. The " time to complete reaction" is the time it takes to reach Reactor Building design pressure (46 psig). j l The results of this study are presented in Figure 14-54. This , amount of allowable zirconium reaction at any time af ter blow-down depends upon the amount of Reactor Building cooling in
]
operation. The capability curves show that' at approximately 10 sec, when the blowdown pressure peak occurs, the Reactor Bu(Iding ould accept an instantaneous zirconium-water reaction of*/33 per cent. This capability increases greatly after the blowdown pressure peak with Reactor Building cooling equipment in operation. With three emerg.ency cooling units in operation a 100 per cent reaction in 3' 44
' B0 see will not exceed the des!.gn pressure o4-55 psig. With three emergency cooling. units and two sprays oper-ating, a 100 per cent reaction in Fr430 seconds will not exceed the design pressure.
l l C) ...- - 7 I 14-45 (Revised 4-18-67)
14.2.2.3.6 Environmental Analysis of Loss-of-Coolant Accidents The safety injection systems are designed to prevent significant core melt-ing in the event of a loss-of-coolant accident. The analyses in the preced-ing sections have demonstrated that these systems will prevent core melting for loss-of-coolant accidents resulting from reactor .oolant system ruptures ranging in size from small leaks to the complete severance of a 36 in. ID main coolant pipe. Without core failure, only the radioactive material in the coolant at the time of the accident is released to the Reactor Building. The environmental consequences from a loss-of-reactor-coolant accident are ana-lyzed by assuming that 1 per cent of the fuel rods are defective prior to the release of reactor coolant to the Reactor Building. Table 11-3 lists the total activity in the coolant. Half of the iodine that is 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 Icakage. No i:redit is taken for removal of airborne iodine by the Re-actor Building spray system. The Reactor Building pressure his tory for this accident is shown in Figures 14-43 to 14-48. While the Reactor Building leakage rate will decrease as the pressure decays, the leakage is assumed to remain constant at the rate of 0.5 per cent per day for the first 24 hours. Thereaftar, since the Re-actor Buildir , has returned to nearly atmospheric pressure, the rate is assumed to be reduce 2 to 0.25 per cent per day and remain at this value for the dura-tion of the accident. The proposed station site and its characteristics are described in Section
- 2. The proposed meteorological models have been developed in Appendix 2B of PSAR and in the answer to Question 2.5 included in Supplement No. 2 (4-18-67 ).
Figure 14-55 presents the downwind dilution as a function of distance from the source for the first 2 hours following the accident. The Reactor Site Criteria, 10 CFR 100, does not require calculation of a 24 hour dose following the accident; however, this calculation has been re-quested for other construction permit applications and is included here. The following formula is used to determina the downwind dispersion: Fft i
" C Bj dG t ( g
, )
Where: f = Fraction of time wind blows from one direction F = Fraction of time a given meteorological condition exists 3 = Wind sector, radians d = Downwind dis tance, meters , G = Average wind speed, meters per second f rz = Vertical dispersion coefficient, meters !
- = Horizontal dispersion coefficient, meters y
14-46 (Revised 4-18-67 ) h A
O\ A = Cross-sectional area of structures normal to wind, square meters c = Building shape factor Values of 7z and 7 obtained y from Gifford, et. al. (14 ) Appendix 2B of PSAR plus answer to Question 2.5, Supplement No. 2 (4-18-67 ) recommends the following for use in the above equation: 12 hours, Pasquill F, u = 1 mps; 40 per cent frcm a single wind sector (22 degram) with the plume averaged over the entire sector plus the remaining 60 per cent from the single sector and the two coterminous sectors, with the plume averaged over the entire three sectors (67 degrees ): and, 12 hours, Pasquill E, G = 2 mps; 60 per cent from a single sector (22 degrees) with the plume averaged over that sector plus the remaining 40 per cent from the single sector and the two coterminous sectors, with the plume averaged over the three sectors (67 degrees). Figure 14-56 also shows the downwind dilution based upon the 24 hour model. For the 1-30 day model at the proposed site, the following conditions from Appendix 2B of PSAR and the answer to Question 2.5, Supplement No. 2 (4-18-67) are used:
)
\s./ Atmospheric Frequency of Average Wind Condition Occurrence, Per Cent Speed, m/sec Pasquill F 25 2 Pasquill E 10 3 Pasquill D 20 4 Pasquill C 45 5 B = 22 degrees (0.393 radians) f = 0.30 The values are used in the equation shown on the previous page for the 24 hour model and the downwind dilution as a function of distance from the source - is shown in Figure 14-56. The dose ta the thyroid per curie inhaled is obtained from TID-14844. I-131 1.48 x 106 rem per curie I-132 5.35 x 10* rem per curie I-133 4.0 x 103 rem per curie I-134 2.5 x 104 rem per curie I-135 1.24 x 105 rem per curie A breathing rate of 3.47 x 10' m3 /sec is assumed for the 2-hour and 24-hour exposure and 2.32 x 10-4 m 3/sec for the 30-day exposure. It is assumed 'that 50 per cent of the Reactor Building leakage will go into the penetration rooms which are maintained at a negative pressure as [O~') 14-47 (Revised 4-18-67) -
- n. - - - - -,
described in 6.5. The atmosphere in these rooms is discharged through char-coal filters to the station vent. The charcoal filters are assumed to be 90 per cent efficient for iodine removal. The remaining 50 per cent of the Re-actor Building leakage is assumed to escape directly to the atmosphere. By this method only 55 per cent of the iodine released from the Reactor Build-ing is ultimately released to the atmosphere where it mixes in the wake of building structures. No credit is taken for the elevated release from the station vent. Figure 14-57 shows the total integrated dose to the thyroid as a function of distance from the Reactor Building for the 2-hour exposure. Figure 14-58 shows the same information for the 24-hour and 30-day exposure. These doses are several orders of magnitude lower than the guideline values of 10 CFR 100, and there is no hazard to the general public. The direct dose from this accident is insignificant due to the shielding provided by the concrete Re-actor Building. 14.2.2.3.7 Effects of Reactor Building Purging At times during the normal operation of the reactor, it may be desirable to purge the Reactor Building. in the event a loss-of-coolant accident were to occur during purging operations, activity would be released to the environ-ment. The purge valves will be completely closed in 5 seconds. During this time period, assuming a 36 in. dcuble-ended rupture, approximately 75 per cent of the reactor coolant has been blown down. The activity in the Reactor Building is due to the reactor coolant activity af ter operation with 1 por cent failed fuel. For this case, 0.3 per cent of the Reactor Building atmosphere will escape through the purge valves prior to closing, corresponding to a release of 1.7 equivalent curies of iodine-131. This dose, when added to th( thyroid dose for a loss-of-coolant accident wi th-out purging is well below the 10 CFR 100 guidelines. Therefore, purging operations can be performed during reactor operation. The dose at the one-mile exclusion distance as a result of this accident is given in Table 14-3. 14.2.2.4 Maximum Hypothetical Accident 14.2.2.4. 1 Identification of Accident The analyses in the preceding sections have demonstrated that even if a loss-of-coolant accident were to occur, no significant core melting will occur. However, in order to demonst rate that the operation of a nuclear power station at the proposed site does not present any undue hazard to the general public, a hypothetical accident involving a complete core meltdown for one unit is evaluated. No mechanism whe reby such a core meltdown occurs is postulated since it would reonire a multitude of failures in the engineered safeguards systems provided to prevent its occurrence. As a result of the meltIown, fission products are assumed to be relea ed trom the core as stated in TID-14844, namely 100 per ce,t of the noble gases, 50 per cent of the halogens and 1 per cent of the solids. Further, 50 per cent of the iodines released to the Reactor Building are assumed to plate out. Other parameters tuch as meteorological conditions, iodine inventory of the fuel, Reactor Building leak rate, etc, are the same l 14-48 (Revised 4-18-67 )
e as assumed for the loss-of-coolant accidcat in 14.2.2.3.6. The average iodine inventory, in terms of equivalent curias 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 14.2.2.4.2 Analysis and Results of Environmental Analysis Figure 14-59 presents the total integrated dose to the thyroid as a function of distance from the Reactor Building for 2-hour exposures. Figure 14-60 presents the same information for the 24-hour and 30-day exposures. It can be seen that the thyroid dose at the exclusion distance of 1 mile is Icss than the guideline values of 10 CFR 100. The direct dose to the whole body following the accident is shown in Figure 14-61. No significant dose acists from this source at the exclusion distance. The dose to the whole body from the passing cloud has been calculated with the same meteorological conditions as used for determining the thyroid dose. 1 These doses are listed in Table 14-3 and are well below the guideline values of 10 CFR 100. 14.2.2.4.3 cffect of Rainout During the Maximum Hypothetical Accident To provide a further evaluation of the suitability of the site, the effects of rainout on surrounding drinking water reservoirs following the maximum hypothetical accident were analyzed. Calculations were made for the case of continuous rain lasting 24 hours covering the general area of the reservoir and the site. The maximum rainout rate as a function of distance is cal-culated from the following equation:(15) 9 o , Qg e-(y /2 7 y) xe c y M where: u , = Maximum rainout rate, curies per sec per m2 x = Downwind distance, .acters 7 = 1 f=Horizontaldispersion, Crosswind distance frommeters plume axis, meters ! Qo = Release rate, curies per see i The above equation is conservative since the results do not consider the l
, wind speed or vertical distribution in the cloud. The wind direction is assumed to remain towards Lake K2owee for the 24 hour period with the plume center lines uniformly distributad over this section. Rainout is assumed to occur under neutral conditions, Pasquill D, which is typical for a rainy day.(15) s 14-49 (Revised 4-18-67 )
The average release rate from the Reactor Building during the 24 hour period following the accident is 0.74 equivalent curies of iodine-131 per second. Using the above equation, the maximum iodine rainout is calculated by assum-ing that all of the iodine that has rained out remains in the surrounding reservoir and is not affected by runoff. The average number of curies in the reservoir during a one year period is reduced by a factor of 0.0318 due to the natural decay of iodine. Assuming that this activity mixes in the reservoir and that a child with a 2 gram thyroid continually drinks 300 ml 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 1200 ml per day is used. The nearest drinking water intake is approximately six miles from the site. At this distance, the total integrated dose to a child's thyroid is 0.64 rem and to an adult's thyroid is only 0.26 rem. These doses are well below the limits of 10 CFR 100. 14.2.2.4.4 Effects of Engineered Safeguards Systems Leakage During the Maximum Hypothetical Accident An additional source of fission product leakage during the maximum hypotheti-cal accident can occur from leakage of the engineered safeguards systems ex-ternal to the Reactor Building during the recirculation phase for long-term core cooling. A detailed analysis of the potential leakage from these sys-tems is presented in 6.4. The enalysis demonstrated that the maximum leak-age is about 6000 cc/hr and that about 2500 cc/hr of this leakage will flash into steam. It is assumed that the water being recirculated from the Reactor Building sump through the external system piping contains 50 per cent of the core saturation iodine inventory. This is the entire amount of iodine release from the re-actor coolant system. The 50 per cent escaping from the reactor coolant sys-tem is consistent with TID-14844. The assumption that all of the iodine escaping from the reactor coolant system is absorbed by the water in the Reactor Building is conservative since much of the iodine released from the fuel will be plated out on the building walls. The activity in the recir-culation water is equal to 0.048 equivalent curies of I-131 per cc of water. It is assumed that all the iodine contained in the 2500 cc/hr of water which is flashed is released to the Auxiliary Building atmosphere. Leakage from the Auxiliary Building occurs due to exfiltration. The most re-strictive case for a ground release occurs during inversion conditions. It is assumed that the building leaks at the rate of 100 per cent per day with atmospheric dilution occurring in the wake of the building. For this build-ing leak rate and the inversion condition, the iodine in 2500 cc/hr will produce an integrated dose to the thyroid in 2 hours at the one mile exclu-sion distance as lis ted in Table 14-3. The above analysis assumes that all of the i' dine released from the cere is taken up by the recirculating water in the engineered safeguards systems. With this assumption, no iodine is available for Reactor Building leakage. Comparison of the 2 hour thyroid dose at the exclusion distance from Reactor Building leakage to the dose from safeguard leakage shows that iodine absorption by the spray water will reduce the off-site dose by a factor of about 150. O 14-50 (Revised 4-18-67 ) 12
'. s v
14.3 REFERENCES
(1) Watson, L. C., Bancroft, A. R., and Hoelke, C. W., Iodine Containment by Dousing in NPD-11, AECL-ll30. (2) Styrikovich, M. A. , et al. , " Transfer of Iodine from Aqueous Solutions to Saturated Vapor," Soviet Journal of Atomic Energy 17, July 1964. (3) Dispersion of Soluble Radioactive Material in Water, CF-58-3-109. (4) International Symposium on Fission Product Release and Transport Under Accident Conditions, Oak Ridge, 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 Journal of Chemical Physics, December 1956. (7) Reactor Development Program Progress Report, ANL-6912, June 1964. (8) AEC Research and Development Reports, WIGL2 - A Program for the
, Solution of the One-Dimensional Two-Group, Space-Time Diffusion s Equations Accounting for Temperature, Xenon and Control Feedback, (s)
WAPD-TM-532, October 1965. (9) Margolis, S. G. and Redfield, J. A., FLASH: A Program for Digital Simulation of the Loss-of-Coolant Accident, RAPD-TM-534, May 1966. (10) Grenda, R. J. and Patterson, J. F. , "The Velocity of Rising Steam in a Bubbling Two-Phase Mixture," Transactions of The ANS 5, No. 1, p 151, June 1962. (11) Possible Zirconium Water Reactions in Water Reactors, AEC Regulatory S taf f Sympot '.um, April 27, 1965. (12) Wagner, R. J. and Finnegan, L. J. , "An Analytical Model for Predict-ing the Pressure-Temperature History Within a Containment Vessel in Response to a Loss-of-Coolant Accident," Phillips Petroleum Company, Atomic Energy Division, Idaho Falls, Idaho, Presented at J NS Meeting, Washington, D. C., November 1965. (13) Quinn, E. P., Forced-Flow Heat Transfer to High-Pressure Water Beyond the Critical heat Flux, ASME 66WA/HT-36, November 27, 1966. (14) Hilsmeier, W. F. and Gifford, F. A., Jr., Graphs for Estimating Ai_mo-spheric Dispersion, ORO-545. (15) Culkowski, W. M. , Deposition and Washout Calculations Based on the Generalized Gaussian Plume Model, ORO-599.
! i , i ,! . > < . i i i i
j i ,
- j' j
. i ! ! 1 I t',
1 STARTUP ACCIDENT FROM 10~9 FULL POWER USING A 1.2% dk/k ROD GROUP: HIGH PRESSURE REACTOR TRIP IS ACTUATED. l (PARAMETER VALUES ARE: 0. 3 SEC. TRIP DELAY
-1.14 x 10-3 (Ak/k)/F DOPPLER COEFFICIENT, AND
+ 6 x 10-5 (ak/k)/F MODERATOR COEFFICIENT).
s M mirem OCONEE NUCLEAR STATION FIGURE 14-1 14
J l it i 1 4 I i i i iii iii,ii, NEUTRON 150 -l l ! ' ' , l 7 ll POWER , % 10 0 l' ,
, !ll
l 50 lj; lll l , l 0
- i. i i i j i 20 ;, i ,
! i i ,
i i j /i N i i i 10 ; 3 .f _ i ,x, , i THERMAL 1i i i o .. POWE R ,% ll i lll l i
! ll 1
! 8
! i I
! !< i.
i e i ! l ' ! ! 200 l , j , l,;; ; F UE.L 100 l! 'f l' l TE MPE RATURE # O CHANGE,F ll,' ! l l i i i i ; i ' i i i + i ii i ! i (w I . i ! i ii,ii i 10 ; i , ,i i i ; AVERAGE CORE MODERATOR 5
' _ M b, t .
;f, i ;
i i i i j TEMPE RATURE I i i t i f i l CH ANGE , F ! i ; , i i i i i : i 5 SECM 2500 l'
; ,-v , i 2350 ,
fi , i,i REACTOR SYSTEM E00
! 'I i PRE SSURE , PSIA , l', l l' l
~
1 ji ! ! ii i ! , ,i i i ii I a i STARTUP ACCIDENT FROM 10 '9 FULL POWER USING A' L RODS WITH A WORTH OF 10% Ak/k : HIGH FLUX REACTOR TRIP IS ACTUATED. ARAMETER VALUES ARE: 0. 3 SEC. TRIP DELAY,
- 1.14 x 10 * (ak/k)/F DOPPLER COEFFICIENT, AND
+ 6 x 10 - (ak/k)/F MODERATOR COEFFICIENT).
OCONEE NUCLEAR STATION o { j 15
\ 100 I 6 6 i i i I 4 1 4 i i iii 6 4 i 4 6 e i i I I i ii 6
/
$0
- l e 60 i i
/
2 m 40 Minh Pretsure _ ! _ tt: A rw
'avet Trip Level Irip v.
Y A l / 20 dCo p ,,
/ _
s f f [ All R ds
^" r t t ii1 i i ? , , ,,
0 ' ' ' t 1 i '1 It ! i ? e 10-5 10-4 10*) gy2 ROD WITHDR AW AL R ATE, f ak/h)/SEC l PEAK THERMAL POWER VERSUS ROD WITHDRAW AL RATE FOR A STARTUP ACCIDENT FROM 10-9 FULL POWER OCONEE NUCLEAR STATION FIGURE 14-3 16
4 g ; ,6 10 _ y g g ,gigg g g g gg;ll g g g jijeg i gg 6 - 4 - a: W - y 2 O 4 10 3 2 $ 2 6 _ g _ g 4 - All Rods _ a: W - y 2 O 2 y10 _ Single _
- r. - Control Rod -
Group : $ 6 - W - Z 4 -
/ _
y _ W 1 2 - 10 l I I I IIIII I I ' IIIII I ' I ' I I I I I ' I II 2 4 o 2 4 o Z 4 o 2 4 6 go-5 go-4 10-3 go-2 10-6 ROD WITHDR AW AL RATE, (.ik/h)/SEC PEAK NEUTRON POWER VERSUS ROD WIjHDRAW AL RATE FOR A STARTUP ACCIDENT FROM 10' FULL POWER OCONEE NUCLEAR STATION (nSowin' FIGURE 14-4 17
0 21 e 20
/
/
o 19 r C.
.1 4
18 x N 17 16
- Nominal 15
/ , ,
0 0. 4 0. 8 1. 2 1. 6 TRIP DELAY TIME, SEC l l I 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)/SEC FROM 10-9 FULL POWER soeurswee OCONEE NUCLEAR STATION () v' FIGURE 14-5 18 .
l l l O 1 30 l
. 25 c:
3 9 N
~
a N s
< 15 m h
~
W Nominal H 2: W t
\d 0 0 -0.4 -0.8 - 1. 2 - 1. 6 -2.0 DOPPLER COEFFICIENT x 105 (ak/k)/F PEAK THERhiAL POWER VERSUS DOPPLER COEFFICIENT FOR A ST ARTUP ACCIDENT USING A*1. 2% a k/k ROD GROUP AT 5. 8 x 10-4 (ak/k)/SEC FROh110-9 FULL POWER - mucan OCONEE NUCLEAR STATION L C' FIGURE 14-6 19
O - i 40 gt
. r
/
~
O 30 ' C. s 2 /
/[ !
O 5 / p / f k /-- Nominal 2 // 10 0
/ 0. 4 0. 8 1. 2 1. 6 TRIP DELAY TIME, SEC PEAK THERMAL POWER VERSUS TRIP DELAY TIME FOR A ST ARTUP ACCIDENT USING ALL RODS '
AT 5. 8 x 10-4 (.1k/k)/SEC FROM 10-9 FULL POWER l OCONEE NUCLEAR STATION o t 20
C\ U l I 40
\ High Flux __ _ High Pressure I y \ Level Trip Level Trip W T ,
5 k \ l { i 30 - 3 a
< 1 l N \
f U l 20
~
a
\w l y Nominal s'
10 0 -0.4 -0.8 - 1. 2 - 1. 6 -2.0 DOPPLER COEFFICIENT x 105 (Ak/k)/F PEAK THERMAL POWER VERSUS DOPPLER COEFFICIENT FOR A STARTUP ACCIDENT USING ALL RODS AT 5. 8 x 10-4 (Ak/k)/SEC FROM 10-9 FULL POWER l p (ouu,rcau OCONEE NUCl. EAR STATION L' FIGURE 14-8 21
2600 2550
'[
c: 2500 y j Safety Valve g 2450 Set Point 1
+ No minal Z j l
< 2400 W
2350 0 0. 4 0. 8 1. 2 1. 6 2. O TRIP DELAY TIME, SEC PEAK PRESSURE VERSUS TRIP DELAY TIME FOR A STARTUP ACCIDENT USING ALL RODS AT 5. 8 x 10-4 (Ak/k)/SEC FROM 10-9 FULL POWER m l Som OCCNEE NUCLEAR STATION 't, FIGURE 14-9 22
1 l I l \ 2550 Safety Valve 4 2525 Set Point 3
. 2500
\ 3 W
c: o g 2475 y W No minal N
*~
2450 % N 'n
'A
() 2425 l 2400 4 6 8 10 0 2 TRIPPED ROD WORTH, % A k/k PEAK PRESSURE VERSUS TRIPPED ROD WORTH FOR A STARTUP ACCIDENT USING ALL RODS AT 5. 8 x 10~4 (Ak/k)/SEC FROM 10-9 FULL POWER ouu ronn OCONEE NUCLEAR STATION i, {(N) FIGURE 14-10 23
O %s 2550 i i j ; High Flux _ ~_ High Pressure
< 2525 Level Trip ~ Level Trip
- l t i 2500
\ l u Safety Set Point alve d l c: l O I 8
$ 2475 d No minal N i 2450 -
W k C- .~.4 2 5 2400 0 -0.4 - 3. 8 - 1. 2 - 1. 6 -2.0 DOPPLER COEFFICIENT x 105 (ak/k)/F PEAK PRESSURE VERSUS DOPPLER COEFFICIENT FOR A STARTUP ACCIDENT USING ALL RODS AT 5. 8 x 10-4 (ak/k)/SEC FROM 10-9 FULL POWER ( OCONEE NUCLEAR STATION FIGURE 14-11 24
._ .__ -_.O.. . - ,
O 2500 g I
$ 2490 i 2480 g
- Nominal #
g 2470 j -
/
/
, 2460 g
-4 0 +4 +8 +12 +16 ,
1 MODERATOR COEFFICIENT x 105 (ak/ki/F PEAK PRESSURE VERSUS MODERATOR COEFFICIENT FOR A STARTUP ACCIDENT USING ALL RODS AT 5. 8 x 10-4 (ak/k)/SEC FROM 10-9 FULL POWER m m$om OCONEE NUCLEAR STATION \ ' FIGURE 14-12 1 1 25 l
O t i V i
! !!' .'! ! t i
' I ! ! ' !
NEUTRON 12 0 ,
. ! ' ' f-[{.
POV/E R , % i ! ' ' ao 'll 40
. i i #
'_! f-i ,i'
( l1 . I i i i ! ' i i > ! i !I
.4 1 i ii i l i e THERMAL 120 , , , ; ; ;
POW E R , */. SO l, ie
!ll!
iii
!\'
I 40 ; , , i
' ' ' I 0
i j_ ! i i !! ! ! l! ! ' FUEL 200 , , , i lj l TEMP ER ATU RE O i i -1 N - CH ANGE , F -200 fl f
- 400 ,
lll ; l[ i i i
! i i e ! _
AVER AGE 5
'#N i CORE MoOERATOR i i
m'8
! i i ,N TEMP ER ATU RE ! . ' ' !
!' T
-5 ' ' I ! ' '
' ' ! 5 CHANGE g F i i l *
! . i 1
-10 , l , lll,'l ,
, i l7 5 SEC~
2500 7 l l , T SYSTEM 2350
- - - d- -
' 2200 , i ,
PSIA i i 2050 , ,, y 1900 , l i ll ROD WITHDRAWAL ACCIDENT FROM FgLL POWER USING A 1.2% dk/k ROD GROUP AT 5. 8 x 10" (ak/k)/SEC; HIGH USEDPRESSURE ARE: 0. 3REACTOR SEC. TRIP TRIP DELAY, IS -ACTUATED. 1.14 x ({ARAMETERS (ak/k)/F DOPPLER COEFFICIENT, AND + 6 x LO*(0*(ak/k)/F MODERAID R COEFFICIENT). mtym OCONEF. NUCLEAR STATION () ' FIGURE 14-13 26
a 2450 4 i i 6 i i 6 6 6 4 e i 6 4 6 e i 4 6 4 4 *
~
High Pressure High Flux Level Trip I.evel Trip
-s 2400 $
~
No mi.u . s f 2350 5 -
$ 2300 A i k . / s ;
d 2250 2200 4 o 8 2 4 0 6 4 4 o a Is -5 10-4 10-3 ROD WITHDRAW AL R ATE, (ak/k)/SEC PEAK PRESSURE VERSUS ROD WITHDRAWAL RATE FOR A ROD WITHDRAWAL ACCIDENT FROM FULL POWER OCONEE NUCLEAR STATION FIGURE 14-14 27
O V 2500 g 2475 N 2450 > / b / 5 2425 > /
* /p 2400 Nominal 1 2375 I
I 2350 0 0. 4 0. 8 1. 2 1. 6 TRIP DELAY TIAIE, SEC PEAK PRESSURE VERSUS TRIP DELAY TIh1E FOR A ROD WITHDRAWAL ACCIDENT FROh! FULL POWER USING A 1. 2% ok/k ROD GROUP; HIGH PRESSURE REACTOR TRIP IS ACTUATED l ( OCONEE NUCLEAR STATION l 0 _ 1 1, 28
O 2425 i i i i High Flux High Pressure Trip Trtp ; g 2400
- Jc: 2375
! ( Nominal o /
$ 2350
~
- 4
/
< 2325 O :
2300 0 -0.4 -0.8 - 1. 2 - 1. 6 -2.0 DOPPLER COEFFICIENT x 105 (Ak/k)/F PEAK PRESSURE VERSUS DOPPLER COEFFICIENT FOR A ROD WITHDRAWAL ACCIDENT FROM FULL POWER IISING A 1. 2% Ak/k ROD GROUP 1 I s (ops taan OCONEE NUCLEAR STATION , FIGURE 14-16 l l 29 l l 1
Y l O 100 l
\
I I = 70,000 lb-ft s N z H N i z 60 l o o
\ N l
l N I H N 40 M l 20 0 0 4 8 12 16 TIME, SEC PERCENT REACTOR COOLANT FLOW AS A FUNCTION OF TIME AFTER LOSS OF PUMP POWER l {m00*a OCONEE NUCLEAR STATION FIGURE 14-17
.... 30
O J
- 1. 7 I
Trip Set Point l Maximum 7 (107. 5%) Ove rpowe r J J N W< 1' 6 ' Rated Power With Minimum DNR Ratio {Uh Overshoot and Deadband l in Hot Channel at i14% N (1039 l Power Steady State (1. 38) Fk= N #^ 9Z 15 N l g
;c0 '
N I , N I= 30 lb-ft' O - N l 55 Y
= 8 I % ~
r (~
@s'4 c< l N N j
4 U l N 0s ZE 1. 3 l l 2E - l l I 1 ' ~,10 0 102 104 106 108 110 112 114 OVERPOWER AT WHICH COASTDOWN BEGINS, % MINIMUM DNBR WHICH OCCURS DURING THE COASTUOWN FOR VARIOUS INITIAL POWER LEVELS (ouu ro OCONEE NUCLEAR STATION i O) t
%J FIGURE 14-18 31 1
O 0.06 0.05
\
s 0 Y 0.04 \s 1 r x i
\
l J \ l r j 0.03 N j Z O i U 0.02 l ! l i . I N l 0.01 e 0 100 200 300 400 500 600 TIAIE AFTER BREAK, SEC REACTOR SYSTEht COOLING RATE FOR 4 IN2 STEAh! LINE BREAK h OCONEE ilUCLEAR STATION
) FIGURE 14-19 a
32
a 40 g Full Power BOL Parameters c / r.)/ F D= - 1.14 x 10 - 5 ( 35 og = 6.0 x 10-5 (ak/k)/F r= 0.3 s e- Delay f* = 5. 47 x 10' Sec 30 Full Pawer EOL Parameters EOL 3 a D= . 6 x 100 (64/k)/F ag = Assume Zero Z 25 r= 0.3 see Delay " C f* = 2. 7 5 x 10-5 Sec [ c 3 0 Z E'!
= 20 W
X
- .a
\ O 15
/
10
+ Nom"nal Casa BOL 5
7 0' O.1 0. 2 9. 5 .4 v. 3 0. 6 0. 7 CONTROL ROD WORTH, c Ak/k o PERCENT CORE E..PERIENCING DNB AS A FUNCTION OF EJECTED CONTROL ROD WORTH AT FULL POWER t' (mthta OCONEE NUCLEAR STATION FIGURE 14-20 33
O
- 2. 5 Full Power EOL Parameters EOL a
D = -1.14 x 10 5 (ak/k)/F
- 2. 0 -
"h1 = 6.0 x 10-5 (ak/k)/F T= 0.3 Sec Delay f* = 5. 47 x 10- 5 Sec g Full Power EOL Parameters y 1. 5 -
"D
= -1.36 x 10-5 (ak/k)/F 0 a ,gg = Assume Zero
$ T= 0.3 Sec Delay O f* = 2. 7 5 x 10-5 Sec
] 1. 0 ,,
O :
- 0. 3 BOL
+ Nominal Case 0
O.1 0. 2 0. 3 0. 4 0. 5 "'. 6 0 0. 7 CONTROL ROD WORTH , Lak/k ZR-H 2 O REACTION AS A FUNCTION OF EJECTED CONTROL ROD WORTH AT FULL POWER p pt .Mm OCONEE NUCLEAR STATION tQ FIGURE 14-21 34
A U b EOL Parameters 6 i i, 4 a D
- . 64 0' ($ k / *g )/ F l
_,, / _a,y
- Assume Zero [
f f,=
- 0. 3 Sec D
- 2. 7 5 < 10'glay sec
/
f F ull Power BOL Parameters EOL 3 10 - a =
-1.14 x 10-5 (ak/k)/F ,
D
- a,y = e.0 x 10-5 (3k/g)/r / /
6 - 7
- 0. 3 Sec Delav / /
-f* = 5. 47 x 10 5 3,c / /
4 /
/ Y
/
l s C / / z' Po r / a , No minal BOL 3 Case 10-O s
; s Z I !/
O o 3 / N I /
$ 4
[ z z o g Nominal y ; Case
$ 10"9 Full Power I / SOL 10
/
/
/
o / 4 , o 2 i 10 0.1 0. 2 0. 3 0. 4 0. 5 0. 6 0. 7 0, g ! CONTROL ROD WORTH, ",ok/k REACTOR NEUTRON POWER VARIATION WITH EJECTED CONTROL ROD WORTH l OCONEE NUCLEAR STATION FIGJRE 14-22 35
O
/
Full Power EOL 120 #
/
Full Power DOL C 100 Nominal j, Case is 9 j 80 1 H BOL Parameters
- 0 (3 g og = -i.14 x 10-5 (ak/ki/F o a 0 ,10-5 (Ak/k)/F
< t'gg == 6.
- 0. 3 Sec Delay y la = 5. 7 5 x 10-5 3', c 40 EOL Parameters 10-9 Full Power
= - 1, 3 6 x 10' (ak/k)/F 3OL i Assume Zero r = 0. 3 See Delay e f* = 2. 7 5 x 10*' Sec 20 m
s 0 l Nominal Case
- 0. 0 0.1 0. 2 0. 3 0. 4 0. 5 0. 6 0. 7 CONTROL ROD WORTH % Ak/k l
REACTOR THERMAL POWER ' AS A FUNCTION OF EJECTED CONTROL ROD WORTH l' (on00m OCONEE NUCLEAR STATION 0 FIGURE 14-23 36
i 60 , , , BOL Parameters
- a D
= -1,14 x 10 ' (ok/k)/F 7
a,y= 6. 0 x 10- (ak/k)/F 3 Full Power U 50 - t = 0. 3 Sec Delay
, --t- - -
i / EOL
,. f* = 5. 47 x 10~ 5 Sec l
~ l I
I O - EOL Parameters c: a cD=. 6 x 10U (ak/k)/ F )
$ 40 - a,y= Assume Zero !
/ ! f
' I I
v = 0. 3 Sec Delay h
' Full Power f* = 2. 7 5 x 10-5 Sec - ~~
BOL F M . H l l 9 30 l ! /. c: ! O , p W _ . . _ , d I 10~9 Full
+ Nominal Power h 20 Case l
_ j BOL c. I , l
/ j l
.a '
< 10 -+ - - * -
- -+ . - - -j _--- .. + Nominal
? . i z -
6 e 1 i (4 , _ I I i I l 1 0 0.1 0.2 0. 3 0. 4 0. 5 0. 6 0. 7 CONTROL ROD WORTH, o ok/k ENTHALPY INCREASE TO HOTTEST FUEL ROD VERSUS EJECTED CONTROL ROD WORTH i um OCONEE NUCLEAR STATION FIGURE 14-24 37
O
# 80 BOL Parameters d og = 6. 0 x 10-' (ak/k)/F
'$ ok/k = 0. 5%
O
- 3. r = 0. 3 Sec Delay q
h h f* = 5. 47 x 10" Sec 5 \ l a z x N
% 40 ^
0 % Nominal Case - W N 20 l
- 0. 5 -0.7 -0.9 - 1.1 - 1. 3 - 1. 5 - 1. 7 DOPPLER COEFFICIENT x 105 (ak/k)/F THE EFFECT ON REACTOR NEU1'RON POWER OF VARYING THE DOPP LER COh.FFICI. TNT-ROD EJECTION AT 10-9 FULL POWER
]
N Otro OCONEE NUCLEAR STATION FIGURE 14-25 38 [ 1
O BOL Parameters
- a D = -1.14 x 10-5 (ok/k)/F 3 k/k 0. 5%
0 1 40 - T = 0. 3 Sec Delay >
~
h f* = 5. 47 x 10 Sec Nominal Case
]
z " g 36 O H U W 32 t t ! ! ! ! 0 3 6 9 12 15 18 MODER ATOR COEFFICIENT x 105 (.ik/k)/F THE EFFECT ON REACTOR NEUTRON POWER OF VARYING THE MODERATOR COEFFICIENT-ROD EJECTION AT 10-9 FULL POWER fnuit,roatja\ i OCONEE NUCLEAR STATION FIGURE 14-26 39
O - 24 BOL Parameters W % G ,g = 6. 0 x 10-5 (a k/k)/ F 5 a / k = 0. 5% k A t = 0. 3 Sec Delay 20 % f* = 5. 47 x 10- 5 3,c , a 2 A-s 16
]\
Nominal Case N % l j U C 12
-0.5 -0.7 -0.9 - 1,1 - 1. 3 1. 5 -1.7 DOPPLER COEFFICIENT x 105 (ak/k)/F l
i THE EFFECT ON REACTOR THERMAL POWER OF VARYING THE DOPPLER COEFFICIENT - ROD EJECTION AT 10-9 FULL POWER ( o OCONEE NUCLEAR STATION 40 l
l l O
. 6 - i i i i i 6
$ BOL Parameters 19
- aD = - 1.14 x 10-5 (a k/k)/ F
- Ak/k = 0. 5%
.a ~
r = 0. 3 Sec Delay I
$ f* = 5. 47 x 10~" Sec !
i s x ! W l 1
% Nominal Case = l \
O i 1 ! ! ! W 15
= 0 ; 6 9 12 15 18 MODERATOR COEFFICIENT x 105 (ak/k)/ F THE EFFECT ON REACTOR THERhiAL POWER OF VARYING THE hiODER ATOR COEFFICIENT -
ROD EJECTION AT 10-9 FULL POWER 1 l t , I {mt ro CCONEE NUCLEAR STATION - FIGURE 14-28 I 41 l l l l
l l l O 112 ; ; ; g ; BOL Parameters 111 -a = -1.14 x 10-U (ak/k)/F a = 6. 0 X 10- (ok/k)/F 110 - ok/k = 0. 5%
/ * = 5. 47 x 10- Sec at 109 EOL Parameters '
y c D = - 1. 3 6 x 10-U (ak/k)/F g108 n g = Assume Zero [ A k/k = 0. 5% a 107 f * = 2. 7 5 x 10-5 Sec 2 > ' y 106 Full Power, BOL f g 105 ' > O s k 104 C se J
/
103 / .
' /
/ Full Power, EOL _
102 f 101 100 0 0.1 0. 2 0. 3 0. 4 0. 5 0. 6 0. 7 0. 8 0. 9 TRIP DELAY, SEC REACTOR THERMAL POWER VERSUS TRIP DELAY TIME - ROD EJECTION AT FULL POWER ountrante OCONEE NUCLEAR STATION t FIGURE 14-29 42
l l O 42 BOL Par' aeters a '
~ D* ~*
~~ ~
i u = 6. 0 x 10-5 3,4 /k)/F : ok/k = 0. 4% (10-9 Full Power) l 2 38 0 ok/k s ,o (Full Power) i f 3
'J f* = 5. 4 7 x 10 ' 5 Sec I
< . i U i g
EOL Parameters l O 34 a D = - 1. 3 6 x 10' (ak/k)/F a = Assume Zero ok/k = 0. 2% (Full Power) t o
- k. Full Power g f* = 2. 7 3 x 10-5 Sec BOL and EOL m 30 e . ;
W . ; H ! H , t l 9
~
i I !
- 1 0 26 - -
' I W :
Y -
. , _ . b_
d ! ! U i l i 3 22 - a. _ _ _ _ _ . . . . l' _
- + Nominal Case
$ i y 18 4 10-9 Full Power BOL 14 0 0. 2 0. 4 0. 6 0. 8 1. 0 TFIP DELAY ( r), SEC ENTHALPY INCREASE TO THE HOTTEST FUEL ROD VERSUS TRIP DELAY TI.\1E -
ROD EJECTION m irowfa OCONEE NUCLEAR STATION FIGURE 14-30 43
e e 1 3 3 F I rI I i I i ; e g 2600 2000 - - Yl Iseerisents1 - g * * -Prodisted
. s ,
i - \ ~ i, \ E UM \ I 400 - ah
\ -
.M l \;
N
\
i \ - 0 , , , , .i , N% , , , , 0.M1 0,01 0.1 1 is gon Time, see ORIGINAL 135C8: 4-15-67 1 1 I i ! 1 LOFT SEMISCALE BLOWDOWN TEST NO. 546 - t .i VESSEL PRESSURE VERSUS TIME {ont,$a OCONEE NUCLEAR STATION FIGURE 14-31 44
- - ____ _,. -_ _ _ , , , - - - - y ,e, -
l l l i l l 100 l '
! ! l r
I 80 y s ' 2 E
$ 60
- \(
E . I pJ ' i t
, I l i !
22% Measured l 20 g ; 1 1 i ! ! I i 1 0 0 10 20 30 40 50 60 Time, see PREDICTED PER CENT MASS REMAINING VERSUS TIME - LOFT TEST NC, 546 ORIGINAL ISSUE: 4-18-67 nt ro OCONEE NUCLEAR STATION \j FIGURE 14-32 45
O 70 60 50 7 2 ^ u 4
.a 40
/{
I
-)
o
\ jl i <
C I 1 20 3 1e x<T i 0 I % 0 1 2 3 4 5 6 7 8 9 10 Time, see CORE FLOW VERSUS TIME FOR A 36 IN. ID DOUBLE-ENDED PIPE RUPTLSE ORIGINAL ISSUE: 4-18-67 o ( OCONEE NUCLEAR STATION 9d w,
O U 1500 i l I l l 1400 -
~j w ! '
1300 , - - - - - m .
]m 1200 ..
Calculated by Quinn's y Modified Sieder-Tate Equation f a 1100 _ i _7, __ yl
; I O \ l I T 1000 L ,
0 900 !\ \ ' iY 5 U 800
\ s i
f
=
Slump Model 700 _ Simulation \ ' 8 i j i 0 I'
^
j 600 l I (~ . 1 U 500 I i l 5 400 ' b ' 6 i j \l. l S 300 ' ' i
\ '
200 ! i I
\ '
- I .
^
100 4 I 0 O 1 2 3 4 5 6 7 8 9 10 Time, see HOT CHANNEL CLAD SURFACE HEAT TRANSFER COEFFICIENT AFTER DNB VERSUS TIME FOR A 36 IN ID DOUBLE-ENDED PIPE RUPTURE ORIGINAL ISSUE: 4-18-67 { e OCONEE NUCLEAR STATION i [h s 1 FIGURE 14-34 w./ 47
O 24 i 1 { 1 { i i ____ . _ ~
/~'" ,4 -
14" Pipe-50% N y 14" Pipe-33% 2 l 12" Pipe-33% N 2 l 12" Pipe-50% N i " //// i M/ Core Bottem
/
i . f/ r 0 0 5 10 15 20 25 30 35 40 45 50 55 60 Time, sec ; 1 REACTOR VESSEL WATER VOLUME VERSUS TIME FOR 36 IN. DOUBLE-ENDED PIPE RUPTURE - FOR 600 PSIG CORE FLOODING TANK REV: 4-18-67 FIG 14-32 (12-1-66) REDRAWN E Po OCONEE NUCLEAR STATION g TO REFLECT SYSTEM CHANGES FIGURE 14-35 48
O 24
)
20 - --- - - - p l 12" Pipe-50% N 2 7 l l 3 -
@ 1000 psig l -
I w r
/ 14" Pipe-33% N 2
* @ 400 psig i j ,
/
i 12" Pipe-33% 3 N, @ 1000 psig , S 14" Pipe-50% N 2
@ 400 psig
} ,
- l
- '- ~ ~
8 - t U x 1 ( 0 0 5 10 15 20 25 30 35 40 45 50 55 60 Time, sec REACTOR VESSEL WATER VOLUME VERSUS TIME FOR 36 IN DOUBLE-ENDED PIPE RUPTURE - FOR 400 PSIG AND 1,000 PSIG CORE FLOODING TANK REV: 4-18-67 FIG 14-32 (12-1-66) REDRAWN To REFLECT SYSTEM CHANGES g t rea OCONEE NUCLEAR STATION b$ Q FIGURE 14-36 49
O l 2500 2400 i Hot Spot 2300
; 2200
/
600 15-12"- 50% N
~
/ 1% _
5 / 600 lb-14-33% N
'c
'a:
y #c } yyy7 p o i ::: . MN5'
/xwe e ::::::::: : 400 lb-14"- 50% N
^
f 1000 lb-12"- 33% N 1500 1000 lb-12"- 50% N 1400 10 12 14 16 18 20 22 24 26 28 30 32 34 36 Quench Time, see MAXIMUM CLAD TDIPERATURE VERSUS TIME TO QUENCH FOR A 36 IN, ID DOUBLE-ENDED PIPE RUPTLTE ORIGINAL ISSUE: 4-18-67 - {ont ram OCONEE NUCLEAR STATION FIGURE 14-37 50 t .- - _
\v 2800 . 2600 n. a?
$ 2400 3
a 2200
\,k A .
4 2000 A
\
- 1800 * " "
Design N j o Point N s 1600
- a 2
$ 1400 1200 0 200 400 600 800 1000 1200 1400 1600 1800 2000 2200 Maximum Heat Transfer Coefficient, Btu /hr-ft 2_p MAXIMLM HOT SPOT CLAD TEMPERATURE VERSLG MAXLMUM HEAT TRANSFER COEFFICIENT AFTER DNB FOR A 36 IN. ID DOUBLE-ENDED PIPE RUPTURE ORIGINAL ISSUE: 4-18-67 t'down OCONEE NUCLEAR STATION U FIGURE 14-38 51
m U. i i . . i
/
/
2800 h,C 2600 2400
/
2200 2,7 2000 [ h = 15 Stu/hr-ft 2 1800
/l i I 7 - 20 8
1 6 1600
; 1400
[V~] U U 1200
\ l 1000 I
f . (h = 100 600 x 400 200 0 i i ! ! 0 10 20 30 40 50 60 Time, sec HOT SPOT CIAD TEMPERATURE VERSUS TIME FOR 36 IN, ID DOUBLE-ENDED PIPE RUPTURE AND VARIABLE QUENCH COEFFICIENT ORIGINAL ISSUE: 4-18-67 'h FIGURE 14-39 52
O 2000 I I [l I l l l 500 F I 1300 w [ 280 F e 1600
! i
%F 3 ! 100 F g, 1400 ' i i
- I l 1 1200 I l G l j 1000 s
= 800 /
I l l 600 i l l 0 10 20 30 40 50 60 Time, see HOT SPOT CLAD TEMPERATURE VERSUS TD1E FOR 36 IN ID DOUBLE-ENDED PIPE RUPTURE AND VARIABLE SINK TE'iPERATURE ORIGINAL ISSUE: 4-18-67 ( OCONEE NUCLEAR STATION O .
- 1'-'
1 1
. .. 5.3 l
1
t i .i I O e : C i 3
- M lach 1.3. Double. ' 2
- Ended Pipe Rupture 8.S$ ft -
* 'E ft 3 gg 2 1ft i i, !
//
t [.6ft !
}-
Surge Line Rupture - l' ,
* - 4 E i l l 5
- / :
1 -
- 3/ j q -
r I M2 g 3 10 13* 10 time, sec ORIC1 MAL ISSUE: 4-1847 MASS RELEASE TO REACTOR BUILDING FOR THE SPECTRUM OF HOT LEG RUPTURES OCONEE NUCLEAR STATION FIGURE 14-41 l 54 I _m. _ _ . , . __. ._, ._ _ - _ _ . _ , , . - _
e O 2'
, , , i i iiii a a ii:- i . ii.ri 2
,4 ft 1 2 f g2 g gg 2 7 t ft2 ~
~
14.1 f t g ,
- \ N g 12 m surge Line Rupture
) =
.4 ft I -
h l , 9.35 ft 2 \ _ I'* I I"
; gg 2
\ .
,% 3 f t' 10*2 to-1 10' I 2 13 10 3 10 Time, see j CRIGINAL !$5UE: 4 1%.67 REACTOR COOLANT AVERAGE PRESSURE FOR THE SPECTRUM OF HOT LEG RUPTURES
& E ren OCONEE NUCLEAR STATION FIGURE 14-42 55
-- _a -_... . __ - _ . _ _ _ _ _ - _ . - - . . _ . . . , - _. . _ _ . _ . - _ _ _ . - -- _ - ,_ . - _ . -
t i t i 5 f 6 I I i 60 , , , , ,,,, , , ,,i,,, , , , , , , , , , , , ,,,,, , , , , , , 15.9
- 56.1 : '
, 30 7
h '
; 3 Energency Coolins ;
l ! 2 - Units - i 1 .o - - l 2 E : : I jo : : E 5 5
.r' ! 1 ** _
J I
- / :
i 10 / t : : I ; ~
- O - , , , , iii, , , , , , , , , , , , , , , , , , , , , , , , , , , ,,,,-
0 1 2 3
'0"I 10 10 10 10 10' time af ter Repture, sec D
ORIGINAL ISSCE 12 1-66. (WAS FIG.14 37) REV. (6 18-67) MADE TO REFLEC* MOCIL CHANCES. REV . (S.23-47)DEC7 GAS REACUR SUILDING VOLLNE TO 1.90 a 10
! /
I l i REACTOR BUILDING PRESSURE VERSUS TIME - 36 IN. ID DOUBLE-ENDED RUPTURE , 4 l l
\
mim, e OCONEE NUCLEAR STATION FIGURE 14-43 56 i
m . . __. , _ _ . _ . m._ . _ _ . - I i 1 1 i 43 ' I I ii;si i a i 4 isi 4 i iie isi i iii sai . i i i i , ; ,_
$$.6 39,3 }
~
50 : , 1
- 3 Erersency Coolina : I e : Cette - 1
~
1 -
= 40 I :
a E 5 I g 30 A : 3 .
~
3
~
3 - 2 20 - 10 l 2 0 i , e iiti i r i eii , , , , , , , , , , , , , , , , , ,,- 3 10*I go 2 A 3 10 10 10 10* Time af ter Rupture, see CRIGINAL !$50Er 4 18-47 REY. ($.2 5-6 7) DECREASED REACTLR BUILDING VOLLHE .3 1.90 m 10* FT! i REACTOR BUILDING PRESSURE VERSUS TIME AFTER RUPTLTE - 8.5 FT2 RUPTURE
! n u rowse OCONEE NUCLEAR STATION FIGURE 14-44
. . . . 57 i
'I
- . - . - - - - . ~ , ,.. - ,, , - - - , , --
l I i 1 i O
, i iiii i i i e i ira i i iiii i i e iai a i iiiiL 54.s :
l 32.7 :
$0 \ _
f 3 Emergency Coeltas e , thit s e' d
}* : : :
l l
- : l
~
3 30
/ :
" ~
g
- 20
~
b b 10 O t t t i i r i i i tii t i e i : i e iiii
- i i ? .-
0 t 2 10 10 lg lO 10 ' 10 Time af ter Ruptune, see ORIGINAL !$$LTI 6-18-47 REY. (5-21-4 7) DECREAS BEACTOR BOI! JING VOLT!ME TO 1.90 a 10# FT , t 1 REACTOR BUILDING PRESSURE VERSUS TD1E AFIER RUPTURE - ! y out row e OCONEE NUCLEAR STATION
' ~
FIGURE M-46
~
/%YF 58,
i i 60 , , , i ia iii i i i i iei i . i a iiii i iiiia a i i ia _
- 53.1 2
?3.4 50
- : i
~
1
- 3 Emergency Cooling Units !
j " a w - 3 2 -
'. 30 2 2
.1 -
3 1 1 as 20 2 13 I 2 O - t i ! tti i t iii ! !t ? t * !(; 1 2 13" 10 10 13 13 10* Tire af ter Rupture, sec. CRIGINAL 155122 4 16-67 j REY. (5-25-67) DECREASED RZACNR SUILDING I voltME TO 1.90 106 FT . 3 REACTOR BUILDING PRESS VERSUS TIME AFTER RUPTURE - RUPTURE
.;2 ff^
sur ma OCONEE NUCLEAR STATION
) FIGURE 14-45
./ if-n 59
i I
~
i I l 60 i i e a iiii 4 i i i i iis i i i iisis- i i : i i i i , i ; 1_ 52.5 12.5 5:
. 3 Erergency Cecitna :
s tuits :
. 41' 1 :
2 : E A 2 : J - I R
- :e 10
- 2 3
, .! .,2 i 13' 1[ 10 1
M2 3 Tite af ter Rupture, see I CRIGINAL 155CI: 4-16-67 31V. ( $- 25-6 7) DECRIA5fD REACTCR BUILDINC VOLIME TO 1.90 a 10' FT3 . REACTOR BUILDING PRESSURE VERSUS H}iE AFTER RUPTURE - 1 FI2 RUPTURE S su rc OCONEE NUCLEAR STATION FIGURE 14-47 60
-o -,- - -.
-,,,v . m---. -.
r- - - - - - - - - -
d I
$0
- ' i ii< is i i ;its i a i e iisi i e iii i isi,
- 44 .3
- 6a I
F [ 3 Erergency Coolint : thits s - E : -
; 10 J
3 : - 0 - 2 2 - 23 l 2 C
~
10 l
- / -
- 2 0 , ,,,,,1 , , , , , ,,, , , , , ,,, , , , , , , , , , , , , , , ,
1 2 I 10~ 10 10 10 10 10' l i Time af ter Rupture, sec. OKICIMAL !$$UE: 12-1-66. (1dAS FIG.14-34) J RIV. (6-16-67) MADE TO REF1.ECT MODEL CHANGES. ] RIV. (5-25-6 7) DECREASEg REACTOR SUICING i WLLHE TO 1.90 x 108 FT . I i REACTOR BUILDING PRESSURE VERSUS TIME AFTER RUPTURE - 0.4 FI2 RUPTURE Opt,.mu OCONEE NUCLEAR STATION FIGURE 14-48
- . 61
l I i i i iisIi e i a a i . i i a 6 a a .i b 6 e 6 i i3 3 i i s
/ p_ -
3 Emergency Cooling t' nits
==='"'
Total , _ - - _
/ \_ Steam-Air-Mixture 10 Liquid - '2 : / _
j E
~
_ / 5 ""* *""' % ,,
~
Pressure -
= -
Injection - h Cooler
$ 10 '
[
- 3
- 3 Emergency '
- Coolers ,
a O 1 ? I f '*f ? I I I ff ff f f f 1 f[t t t i e fgl t y 9 9 9 9 e 9g 0 1 # 13' 10 10 10 10 10* Tism after Rupture, sec. CRIGINAL ISSUE: 4-18-67, REV. (5-25-67) DECREASED REACTOR BUILDINC VOLLHE TO 1.90 m 106 FT3 . REACTOR BUILDING ENERGY INVENTORY FOR 36 IN. ID DOUBLE-ENDED RUPTURE 1 (m l (o OCONEE NUCLEAR STATl0!1 FIGURE 14-49 ) 62
i 10 _ i 6 . .. i i i iiaaI i 6 iia i ' ' '
. pq 6 8
~
3 Emergergency Cooling Units # p Total / f Steam-Air g 2 10 / / M1"E" \
. / -
Liquid
~
l [ Structures -
's s -
Lcw Pressure
- Injection Cooler O !"' b 3 Emergency 1
Coolers _ O I 10 , , , , , ,, , r r i in ei i e , ivii i i i i e i 0 D2 I 10 ' 1 10 10 ;g 3 10' Tin after Rupture, see. CRIGINAL ISSL'E: 4-16-67 RIV. (5-25-61) DECPEASED REACTOR SUILDING VOLLHE *O 1.90 x 108 FT3 . ) I i l REACTOR BUILDING ENERGY INVENTORY FOR 3 FT2 RUPTURE j i 1 ouu ronts OCONEE NUCLEAR STATION D W - FIGURT 14-50 i 4 63
- E e ei * .4 . . . .ss.4 4 6 4i;
- /,
,N .
Reactor Building Vapor Terperature h - 250 ,
- /
f
\~\ _
Reactor Building
./ 5"*P T**P''
200
/ '
s t i
/ / \ :
5 g 150 E
/
/ \
- -\ :_
2 3 Emersency Cooling b
- Units 100
! ie ' ri t , , ,, , , , ?
0 1 2 3 10* 10 10 10 10 10 ' Time af ter Rupture, sec. CRICINAL ISStT 4-16-67, i REY. (5-25-67) DECREASE PEACTOR BUILDINC 1 VOLLHE TO 1.90 x 108 FT . 1 1 l l REACTOR BUILDING VAPOR AND SUMP TDGERATURES AS A MINCTION OF TIME AFTER RUPTURE - 36 IN. ID DOUBLE-ENDED RUPTURE out ronu OCONEE NUCLEAR STATION FIGURE 14-51 64
l 300
. 1 I i i i l ii i 6 1 i l i 66 i e 6 6 6 i 64 6 i i 1 4 i s '_
h 3 Emergency Cooling
- Units
/'%- -- [
~
2 250 -
/ Reactor Building Vapor g
- / Temperature _
- r .
w -
/ ~
i E / 7 k N
\
200 Reactor Building - h ,
" - Sump Temperature -
150 \ - 0 ( 2 2
~
,,,,,,,2
\E N
100 i , , , , , , , , , , , ,. ,, , , , ,,,,, , 0 1 2 3 10 10 10 0 10' Time after Rupture, sec. CRIGINAL ISSUE: 4-18-67 REV. (5-25-67)DECFgASEDREACTORBUILDING 3 VOLUME TO 1.90 x 10 FT . REACTOR BUILDING VAPOR AND SUMP TEMPERATURES AS A FUNCTION OF TIME AFTER RUPTURE - 3 FT 2 RUPTURE endom OCONEE NUCLEAR STATION
~
FICURE 14-52 65
t a 3 J I 4 1 40
. i iia aii a i a a a iii i i i iii i e i iiii i i i i i i i '_
. 5 _6.7 -
- 54.1 :
b 50
- 3 Emergency Cooling
, ; L%1t e :
A ;,0 a, . [ 30 2 : : a : : 1
. 20 E
s : : , 10
/
/ :
~ -
i . 3 - i inn i i i inn i iir i i i sin i iiiiic 1 2 3 10'I 10 10 10 10 10 ' Tise af ter Pupture, sec. ORICINAL ISSUII 12-1-M. (IdAS FIC.14-39) AEY. (4-18-47) MADE TO REFLECT MODEL OIANCES. FEY. (5-25-67) DECREASED RZACTOR BUlt&IWC VOLLHE TO 1.90 a 106 y7), [ CRITERION 17 CASE FOR 36 IN. ID DOUBLE-ENDED RUPTURE l (ent rown OCONEE NUCLEAR STATION t FIGURE 14-53 66
- - , . , , , , ,__.n.--. _ . _ , ,. _. _ , _ _ , . . __ , n ,--. ,
d 100 3 Emergency Cooling Units + 2 Sprays 80 s l l 4
" 60
[
} 3 Emergency Cooling Units 4
a 40
. f 2
3
/ :
20 / Structures 0 ! _ 0 400 800 . 1200 1600 2000 2400 2800 3200 Time to Complete Reaction, sec.
- ORIGINAL ISSUE: 12-1-66. (WAS FIG.14-40)
REV. (4-18-67) MADE TO REFLECT MOCEL CHANGES. REV. (5-25 67) DECREASED REACTOR BUILDING VOLUME TO 1.90 x 106 FT3 . l l REACTOR BUILDING 2R REACTION CAPABILITY FOR 59 PSIC DESIGN PRESSURE y \ g s y OCONEE NUCLEAR STATION O _ nctraz 14-54 67
l O ;
~
10 j j i j j ; g j 6 _ Boundary _ Of Low 4 - Population - Zone Exclusion 2 - Distance __ t l i I
"c: 10-4 f 1m 6 N _
k* 4 - i i I i
; 2 -
u O i E 10-5 _ i i 6 - I _ i 4 - 2 - _ 1 10-6 l l l l [ ; ; ; j 2 4 6 2 4 6 103 104 105 Downwind Dis tance, Feet l 2 HOUR DISPERSION MODEL l founjowti) OCONEE NUCLEAR STATION REV: 4-18-67 Y/ FIGURE 14-55 68 O
- y. , _ . - _ _ , ,
O
~
10 i i l l i i i i 6 - _ 4 - _ 24 Hours Boundary Of Low 2 - Population _ i Zone I I m 10-5 j 6
\ .
6 4 - l l c' I o 2 - A O $ 30 Day ) i l d $10-6 I i o _ 6 - 4 - Exclusion _ Distance . 2 - 10-7 I I I I I I i 2 4 6 2 4 6 10 3 10 4 105 Downwind Distance, Feet 24 HOUR AND 30 DAY DISPERSION MODELS m Shate OCONEE NUCLEAR STATION REV: 4-18-67 FIGURE 14-56 69 a,
O V 10-1 _ i l i i i i i t i l i t i i i iIL 8 - _ 6 - - 4 - _ g _ _
. Exclusion C 2 -
Dis tance - 2 5 10-2 P - _
-3 8 - 1 Boundary -
3 - I Of LW - { o 6 - Population Zone O c
- 4 - -
g _ _ e 1 I 2 - _ 10-3 l ; ! i ; ; ; i ; ; ; ; ; ; ; ; ;; 2 4 6 8 4 10 4 2 6 8 103 105 Dow:. wind Distance, Feet THYRCID DOSE FROM LOSS OF COOLANT ACCIDENT - 2 HOUR DOSE p mgron g OCONEE NUCLEAR STATION REV: 4-18-67 D' FIGURE 14-57 70
__ _ _ l l 1 O ! l 10-1 ' _ 1 i i i i i i i i i i i i i iL I 8 - _ i 6 - Boundary - i _ Of Low _ Population ; Zone - Exclusion
@ _ Distance _ l oc E s g 2 _ i c I l
3 - 8 p 10-2 g I 8 - l 30 Day Dose [ I "3
! I -
;j 6 -
u _ 24 Hour Dose I j H l 1 %Y b
~
i i i I 9 l l 10-3 1 I i i ! ! ! 1I ! ! ! . I t ii 2 4 6 8 ' 4 6 8 103 10+ 105 Downwind Dis tance, Feet THOROID DOSE FROM LOSS OF COOLANT ACCIDENT - 24 HOUR AND 30 DAY DOSES l l l l l 1 l REV: 4-18-67 0 trow OCONEE NUCLEAR STATION FIGURE 14-58 71
, _ _ _ _ , _ s_.-,.-----
1000 __ i i i i i i i ii i i i i i i i iL 800 __ _ 600 _ - 400 _ _ ____________ --_____+-_ ! 10 CFR 100 Limit - Boundary o Of Low j 200 - Population I Zone
; - 1 -
0
^ l
- f. 100 \ ;
y _
= 80 - -
g _ m 60 - Exclusion - y - Distance ! - O e 40 - l o e _ _ l 20 - - 10 l I I I i 1 I II I I I I I ! ! I 103 2 4 6 8 104 2 4 6 8 105 Downwind Distance Feet RXIMIM HYPCTDIETICAL ACCIDENT THYROID DOSE ASSUMING 1007. CORE MELTDOWN - 2 HOUR DOSE OCONEE NUCLEAR STATION REV: 4-18-67 \mehm
@ FIGURE 14-59
O 1000 t i i i i i i i i i i i i i I l_ 800 - 600 - - i - 400 _ - 8 l
----10 CFR 100 Limit -
I
$ l Boundary S 200 - 1 Of Low -
o Population l o - Zone -
.=
H ' 100 i
~
80 - 2 g _ 3 60 - Exclusion ' - M - Distance - Q 40 _ O 30 Day Dose 20 - 24 Hour Dose - 10 I I I I I I II I I i ! t I i i ii 2 4 6 8 2 4 6 103 10 4 8 105 Downwind Distance, Feet MAXIMUM HYPOTHETICAL ACCIDENT THYROID DOSE ASSUMING 1007. CORE MELTDOWN - 24 HOUR AND 30 DAY DOSES out rbm OCONEE NUCLEAR STATION f)i REV: 4 .18-67 Y ' FIGURE 14-60 73
O 10 ;;;,
,;, ,,;; ;,;; ,;,, ,,,, ;,; ;,,; ,,,, ,,;, ,,,3 i ; ; ;_
I 10 0 ( 10 _ _ 2 : W - _ x - _ g I - c 10" _ , _ p - 30 Day Dose - O - - i [ S Hour Dose 2 Hour Dase g U 10' v e
=
3 1_ a -
-1 3 l
\ \ 3 Exclusion -
, _ Distance _
[ l 10~4 _--
\ _
- i :
_ N _ 10"
\
\ '!
0 1000 2000 3.30 4000 5000 SG00 DISTANCE FROM EDGE OF REACTOR BUILDING, FEET INTEGRATED DIRECT DOSE FOLLO%1NG MHA WITH 3-1/ 2 FOOT REACTOR BUILDING WALL THICKNESS
** ~ ~
$b OCONEE NUCLEAR STATION Figure number changed
,] from 14-46 to 14-61 -
FIGURE 14-61 7A
--- , , , - - - n,. , - , , , - - - , . - - - - - -}}