ML14342A557

Davis-Besse Unit 1 Updated Final Safety Analysis Report, Revision 30, Section 15, Accident Analysis

ML14342A557
Person / Time
Site:	Davis Besse
Issue date:	11/21/2014
From:	FirstEnergy Nuclear Operating Co
To:
Shared Package
ML14339A821	List: L-14-305, Drawing E-52B, Sht. 45A, Rev. 9 ML14339A813 ML14342A512 ML14342A513 ML14342A516 ML14342A517 ML14342A553 ML14342A555 ML14342A556 ML14342A557 ML14342A558 ML14342A560 ML14342A683 ML14342A685 ML14342A693 ML14342A845 ML14342A846 ML14342A847 ML14342A848 ML14342A849 ML14342A850 ML14342A852 ML14342A853 ML14342A854 ML14342A856 ML14342A876 ML14342A877 ML14342A878 ML14342A879 ML14342A880 ML14342A881 ML14342A882 ML14342A883 ML14342A884 ML14342A885 ML14342A910 ML14342A911 ML14342A912 ML14342A913 ML14342A914 ML14342A916 ML14342A917 ML14342A918 ML14342A920 ML14342A924
References
L-14-305
Download: ML14342A557 (264)

Text

Time. s (After CROM Release) DAVI S- BESSiE NUCt EAR POWER STAT1 ON NORM A1 I ZED ROO MIRTH VERSUS TINE FIGURE 15.1-1 REVISION 0 JULY 1982

neutron Power. % Tne rma I Power. O, Fuel Temoerature Change. F Average Core Mooerator lemuerature Cnange. F Seac tor Sys tern Pressure Psla MVIS-BESSE NJCLEAR WR STATIC)*([

STARTUP ACCIENT FROA 10-9 RATED PCXR FCR A REACTIVITY ADOITION RATE OF 1.65 X 10" [A K/K 11s: HIW PRESSURE REACTOR TRIP IS ACTUATED FIGURE 15.2. l -I 2E'JISICN I;! JULY 1590 f hr rm8 l 80 4 0 Power, % 1# I 308 Fue I 231 tamprrtturr 200 Change.f 150 DAVIS-BESSE WCLEAR PER STATIW STARTUP ACCIDENT FRM 10-9 RATED PMR FOR A REACTIVITY ADOITIQJ RATE OF 7.19 X 10-4 tAK/Kl/s [SIMJLATANEWS WITHDRAW OF ALL CRA'S I; HIW FLUX REACTOR TRIP IS ACTUATED FIGURE 15.t.l-2 REVISION 9 JULY 1989 b I, - d - I I I 1 ..I 1 1 1 m-8 m-' REACTIVITY ADOITION RATE (A K/K Vm DAVIS-BESSE NUCLEAR POUIER STATIW PEAK THERW PWER VERSUS REACTIVITY MOITION RATE FCR A STARTUP ACCIDENT FROM I0 -9 QATED POmR FlGURE 15.2.1-3 REVISION 9 JULY 1989 REACTIVITY ADOITIOJ RATE [A K/K l/s WIS-BESSE NUCLEAR PCmR STAT I ON EM wUTRC1J POWER VERSUS REAClIVITy AM)ITIm RATE FPR A STARTUP ACCI#NT FRaM 10-9 RATED PMR FIGURE 15.2.1 -4 REVISION 9 JULY 1 989 DAVIS-BESSE NUCLEAR POWER STATION PEAK THERMAL POWER VERSUS DOPPLER COEFFICIENT FOR A STARTUP ACCIDENT WITH A CONSTANT REACTIVITY ADDITION RATE OF I .65 x la-4 [AK/K FROM 10-9 RATED PMR FIGURE 15.2.1-5 REVISION 9 JULY 1989 MVIS-BESSE NUCLEAR POWER STATION PEM THERMAL VERSUS MCMRATCR COEFFICIENT A STARTUP ACCIDENT WITH A CONSTANT - - SEACTIVITY ADOITICN RATE CF I .65 X 10-4 t n K/K 11s FRO4 10-9 RATED PmR FIGURE 15.2.1-6 REVISION 9 JULY 19/39 DAVIS-BESSE NUCLEAR POWER STATION PEAK THERMAL POWER VERSUS DOPPLER COEFF1:CIENT FOR A STARTUP ACCIDENT WITH A REACTIVITY ADOITION RATE OF 7.19 X 10-4 [ h K/K [ SIWLATMWS WITHDRAW OF ALL CRA 'S I: FROM 10'9 RATED POWER FIGURE 15.2.1-7 REVISION 9 JULY 1989 MVIS-B*S*

NICLEAR PUER STATION PEAK TERWl POYR VERSUS KERATm CEFFICIENI Fm A STMW KTIEM WITH A REACTIVITY PCOITIPl RATE Q 7.19 X I 0-4 [A K/K Ih L SIMllTCN015 WITHaZAWa W ALL CRA'SI FRH 10-9 RATED KMR FIGURE 15.2.1-8 REVISION 9 JULY 1989

..., , . j I... , . . . . L, ... . ..... . , .. .. . . ca m .. 30 Ti (s) -

., r .............

.....-. - -

ku tron Power X Therms I Power X Fue I uo Te-ers turm 30 kdmrs tor I. s Tenpure ture IaO Change F 0.5 0.0 Resc tor DAVIS-ESSE UEPR PWER STATION CRA WITHCRAW&

ACCIENT FRaM RATED PCER FCR A REACTIVITY ADDITI~J RATE F 2.3 x I 0-4 r a K/K I/.: HIW FLUX REACTOR TRIP IS NTTUATED FIGURE 15.2.2-1 ME-BESSE UEAR PWER STATION PEAK PREsSE ERWS REACTIVITY POOITIPl RATE FOR A CRA WITHDRAWAL AOCIIENT FROH RATED PWER FIGURE 15 .t .2-2 REVISION 9 JULY 1989 Trip 001 Tim., r DAVIS-BESSE WCLEM PWER STA,TION PEAK PRESSURE VERSUS TRIP ELAY TIIE FOR A CRA WITHDRAWAL ACCIDENT FRCN RATED POYYER WITH A COSTANT REACTIVITY AOOITION RATE OF 2.3 X 10.4 (A K/K)/r FIGURE 15.2.2-3 REVISION 9 JULY I 989 4.8 -1.0 -1.2 - 1.4 bpptrr ~..ffici rt. (&IL)/I s tos W1S-BESSE WCLEAR PMR STATIOJ W.iPAESSURE VERSUS OQPPLER CUEFF I CI EM FOCI' A CRA WITHORAWAL ACCIDEM' FROM RATED mR WITH A CQJSTANT REACTIVITY MOITIDN RATE IF 2.3 X 10-4 (A K/K)/s FIWRE 15.2.2-4 REVISION !3 JULY 1 989 DAVIS-BESSE NUCLEAR PmR f TATI,m PEAK PRESSURE VERSUS MOOERATOR CEFI' I CI ENT FDFi A CRA WITHDRAWAL ACClCENT FRM RATED POWER WITH A CWTANT REKTIVITY mI:TIO.( RATE [r 2.3 x 10-4 tAKIKl/m FIGURE 15.2.2-5 RE'I'ISICN 12 JULY 1590 Initial Power, % of Rated Pow~er DAVl S-BESSE NUCLEAR POWER ST AT l ON MAXIMUM NEUTRON AND THERMAL POWER FOR AN ALL-CRA Wl THDRAWAL ACCI DENT FROM VARIOUS IN1 TI AL POWER LEVELS FIGURE 15.2.2-6 REVISION Cl JULY 1982 In Hottest Fuel Rod At The Hot Spot 0 20 4 0 6 0 80 100 Initial Power, !I of Rated Power DAVI S-BESSE NUCLEAR POWER STAT l ON PEAK FUEL TENPERATURE IN AVERAGE ROD AND HOT SPOT FOR AN ALL-CRA WITHDRAWAL ACCIDENT FROB VARIOUS INITIAL POWER LEVELS FI SURE 15.2.2-7 REVISION Cl JULY 1982 Initial Power, % of Rated Power DAVIS-BESSE NUCLEAR POWER STATION PEAK REACTOR CWUWT PRESSURE VERSlJS POWER FOR All. CRA GROUP WIrnWAL FIGURE 15.2.2-8 REVISIDN 0 JULY 1982 Weu t ron Power, % T h erm al Power, % Average Core Wodt!ratar Temperature Chagt, F Reactor Sy st an Pressure, psi a DAVI S-BESSE NUCLEAR POWER ST AT l ON 0.65% Ak/k CRA DROP FROM RATED POWER AT EOL CON0 l TI ON FIGURE 15.2.3-1 REVISION 0 JULY 1982 0 4 8 12 16 Time, s DAVI S-BESSE NUCLEAR POWER STAT l ON PERCENT REACTOR COOLANT FLOW AS A FUNCTION OF TIME AFTER LOSS OF PUMP POWER Fl GURE 15.2.5-1 REVISION 0 JULY 1982:

0 1 2 3 4 Time, s DAV I S-BESSE NUCLEAR POWER STAT I ON - PERCENTNEUTRONPOWERVERSUSTIME FOLLOW IN6 REACTOR TRl P FIGURE 15.2.5-2 REVISION 0 JULY 1982 I0 2 I04 06 108 1 10 112 Overpomr at whi ch Coastdon 8trgi ns, % DAVl S-BESSE NUCLEAR POWER STAT l ON MINIMUM DNBR WHICH OCCURS DURING A FOUR PUMP COASTDOWN FROM VAR l OUS IhI I T I AL POWER LEVELS F I GURE 15.2.5-3 REVISION 0 JULY 1982 r Neutron Porer 3 Reactor Systan - - - - - - 0 I 2 3 5 Time, s DAVI S-BESSE NUCLEAR POWER STAT l ON NEUTRON POWER, FLOW, AND REACTOR SYSTEM PRESSURE FOR A LOCKED ROTOR ACCIDENT, BOL PARAMETERS FIGURE 15.2.5-4 REVISION 0 JULY 1982 0.0 . 0. V 0.8 1. 2 1.6 2.0 2. 4 Time, s DAVIS-BESSE NUCLEAR POWER STATION DNB RATIO VERSUS TIME FOR LOCKED ROTOR ACCIDENT FROM 102% OF RATED POWER FIGURE 15.2.5-5 REVISION 0 JULY 1982 Neutron Power, X Thermal Power, X Average Core

~oeer ator Temper atu re Change. F Reactor Pressure, psi a DAVI S-BESSE NUCLEAR POIER STAT l ON TWO PUMP STARTUP FROM 60% POWER AN0 49% FLOW FIGURE 15.2.6-1 2YOO & REVISION 0 JULY 1982 2300 2m 2 100 , - I 0 2 4 I ~jyF 6 8 Time. 110 s I2 IY 16 18 20 t00ltu?ril l0 10ft1, l0 r13 n 0 Tmrul ltril, I f00 t0 t0 f0 il 0 mrgTot 'l0 ioorili i r rfttr$ c t[f?.ciil&. ' t 0p -10.IOI COm ot3l ilTtTT?ffssum 0 gmmt, .f 0e?t ll .i0!Ittrr strurrTn dlnfl PffSTUrT cilnl,

?3lr srtrr SIrIlrr0l otrril llp.ctrfit, 0p IO Note: Figure 15.2.7-1 represents the original design of the plant. This is the ICS runback to the steam generator low level limit. Analysis for the turbine trip when the runback is not successful is given in Reference 67.r0 rn t.0 tllu I s -8[sst ifuu. Em Pilffi sTrT t$fL OSS OF TITIRIIIL ELI gTR I CIt LOAO tT nfit0 Pof[R ; rTlf luT0mT r cPlIrtR Rulf 8rH(Fleun[ 15.2.1-lRevision 30October 2014 n0 n0 til r0.l I r0 il 0 tt al+

a IM IIl3f.f Iil ilf tfl tll$cmt$0=.---L-..E\L.-r L-_.t L Tr--..-L..-\!\r: f^/F 4 -Tit-./\\-..4 E--\-1 n/\J Jl-\--;Fr.l-\*a-4 7\hlr/I /#rr lrrr-^^ --:.L/\\TWU\U rf J\\\\C\\r L Meu tron Powr, % Thermal Power* % Average Core +6 - Moderator

+Y . Temperatu re Change, F +2 2 - -4 I Reactor System Pressure, psi a DAVI S-BESSE NUCLEAR POIIER STAT l ON LOSS OF ALL FEEDWATER FROM RATED POWER FIGURE 15.2.8-1 REVISION 0 JULY 1982 DAVIS-BESSE NUCLEAR POWER STATION STEAM GENERATOR COLLAPSED LEVEL (LOOP 1) FIGURE 15.2.8-lA REVISION JULY I98 DAVIS IBESSE NUCLEAR POWER STAT I ON HOT LEG TEMPERATURE I LOOP 1 I FIGURE 15.2.8-IB REVISION 9 JULY I909 DAVIS-BESSE NUCLEAR POWER STATION HOT LEG PRESSURE [LOOP I I FIGURE 15.2.8-lC REVISION 9 . JULY 1989 DAVIS-BESSE NUCLEAR POWER STATION PRESSURIZER COLLAPSED LIQUID LEVEL FIGURE 15.2.0-10 REVISION 9 JULY 1989 Time After Rupture!, 9 DAVIS-BESSE NUCLEAR POWER STAT1 FEEDWATER LINE BREAK WITH OFFS ITE POWER AVAI LPIBLE FIGURE 15.2.8-2 REVISION C JULY 1982 Tim After Rapture, s OAVI S- BESSE NU& EAR POWER STAT I ON FEEDWATER L INE BREAK W l TH OFFSl TE POWER AVAILABLE FIGURE 15.2.8-3 REVISION 0 JULY 1982 Time After Rupture OAVI S-BESSE NUCLEAR POWER STATION FEEDWATER LINE BREAK Wl TH OFFSl TE POIER AVAILABLE FIGURE 15.2.8-4 REVISION 0 JULY 1982 Tine After Rupture, s OAVI S-BESSE NUCLEAR POWER STAT1 ON FEEDWATER L lNE RUPTURE Wl TH OFFSI TE POWER AVAIL brBLE-CASE I FIGURE 15.2.$8-5 REVISION 0 JULY 1982 Time After Rupture, r DAV I S-BESSE NUCLEAR POWER STAT l ON f EEDWATER 1, l NE BREAK W I TH OFFSl T E POWER AVAl LABLE FIGURE 15.2.8-6 REVISION 0 JULY 1982 Tim After Rupture DAVI S-BESSE NUCLEAR POWER STAT1 ON FEEOIATER LINE BREA I( ll TH LOSS OF OFFSITE POWER AT RUPTURE FIGURE 15.2.8-7 REVISION 0 JULY 1982 Time After Rupture, s DAVI S-BESSE NUCLEAR POWER STATION FEEDRATER LINE RUPTURE WITH LOSS OF OFFSITE POWER AT RUPTURE CASE I I FIGURE 15.2.8-8 REVISION 0 JULY 1982 DAVI S- BESSE NUCLEtLR POWER Sl AT l ON FEEDIATER LINE BREAK 'I I TH LOSS OF OFFSITE POIER AT RUPTURE 1200 L J n I rl] 1000 - - - 8n0 L

• C d 600 E - d Y LOO J d CI 200 b cTJ C 3 0 FIGURE 15.2.8-9 I I REVISION 0 JULY 1982 0 5 10 15 2 0 2 5 Time After Rupture, s The After Ruptwce, s DAVI S-BESSE NUIXEAR POWER STAT1 ON FEEDWATER LINE GREAK Wl TH LOSS OF OFFSITE POWER AT RUPTURE FIGURE 15.2.8-10 REVISION 0 JULY 1982 Time After Rupture, s DAVIS-BESSE NUCLEAR POWER STLTION FEEDWATER LIME RUPTURE WITH LOSS OF OFFSITE POWER AT TRIP-CASE I I I FIGURE 15.2.8-11 REVISION 0 JULY 1982 Time After Rupture, s 6 30 J B 620 610 I 600 590 580 570 0 1 C 2 0 30 4 0 50 Time After Rupture DAVIS-BESSE NUCLEAR POWER STATlON FEEDIATER LINE BREAK WITH LOSS OF OFFSITE POWER AT TRIP FIGURE: 15.2.8-12 REVISION 0 JULY 1982 Time After Rupture, s DAVI S-BESSE NUCLEAR POWER STAT1 ON FEEDIATER LINE BREAK UlTH LOSS OF OFFSI TE POWER AT TRIP FIGURE 1'5.2.8-13 REVISION 0 JULY 1982 Time After Rupture, s DAV I S- BESSE NUCLEAR POWER STAT l ON FEEDWATER LINE BREAK WITH LOSS OF OFFSITE POWER AT TRIP FIGURE 15.2.8-14 REVISION 0 JULY 1982 0 It ACTOR . 5 cooLArl -10 rrtnru TED? CWU#. -29 Or -25 DAVI S-BESSE NUCLEAR POWER STAT l ON LOSS OF A.C. POWER WHILE POWER OPERAT I NG AT RATED- POWER FIGURE 15.2.9-1 REVISIOFI 0 JULY 19'62 Neutron Power, .b Average Core Moderator 5 90 Temperature.

580 F Reactor Sys tem Pressure, ps I a Steam Generator Outlet Pressure, psla Minimum Hot Channe l ONBR 30 Time, s DlrV I S-BESSE NUCLEAR POlER STAT l ON RESPONSE OF REACTOR COOL ANT SYSTEM TO FEEDWATER TEMPERATURE DECREASE Fl GURE 15.2.10-1 REVISION 0 JULY 1982 Neu t ran Power, % Aver age Co re Moderator Tmperatu re, F :I Ppp- Reactor Systm Pressure, psi a no0 Steam Generator OU tl et Pressure, psi a 900 DAV I S-BESSE NUCLEAR POWER STAT l ON RESPONSE OF REACTOR COOLANT SYSTEM TO FEEDWATER FLOW INCREASE TO NO LOAD CON01 TI ON FIGURE 15.2.10-2 REVISION 0 JULY 1982 Feedwa te r Flow, % Steam Generator Downcomer Level. ft Steam Generator 801 I lng Length. tt Turb ~ne Bypass Flow. % Steam Satety Va I ve Flow, % Max I mum Hot Channe l ONBR Pressur ~zer Level Above Sensor. f t lime, DAVI S-BESSE NUCLEAR POWER STAT1 ON liECO-14 EXCESS1 VE HEAT RIEMOVAL DUE TO 115% FW FLOW FIGURE 15.2.10-3 REVISION 0 JULY 1982 Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.3-1 UFSAR Rev 30 10/2014 15.3 CLASS 2 - EVENTS LEADING TO SMALL TO MODERATE RADIOACTIVE RELEASES AT EXCLUSION AREA BOUNDARY

Class 2 events are off-design operational transients or accidents which may result in the

following:

a. Fuel failure in excess of those expected during normal operation.

b. A breach of the fuel cladding (which leads to fission product release) or of the Reactor Coolant System boundary.

c. Operation of engineered safety features or the use of the containment to limit the consequences of a transient.

d. Offsite radiation exposures in excess of the limits permitted during normal operation.

The consequences of Class 2 events are not of such severity as the required interruption or restriction of public use of areas beyond the station exclusion area boundary. Furthermore, these events do not in themselves lead to the occurrence of the more serious Class 3 events.

Table 15.3-1 summarizes the accidents categorized as Class 2 events.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.3-2 UFSAR Rev 30 10/2014 TABLE 15.3-1 Class 2 Events Event Analysis assumptions Effect Loss of reactor coolant from small ruptured pipes or from cracks in large pipes which actuates emergency core cooling Reactor coolant leakage through a spectrum of areas

smaller than for the design basis LOCA is considered. Environmental effects are based on the release of all the gap activity from the fuel. The accident results are discussed in Chapter 6.

Environmental effects are discussed in Subsection 15.4.6. Minor secondary pipe break The rupture of a steam line of small area is considered. The reactor is assumed to be operating with 1% defective fuel and 1 gpm steam generator tube leakage.

Reactor coolant leakage into the steam generator continues until the Reactor Coolant System is cooled down and depressurized to ambient conditions. The consequences of this accident and its environmental effects are discussed in Subsection

15.4.4. Inadvertent loading of a fuel assembly into an improper position Fuel assemblies are loaded into improper core positions.

Also fuel assemblies with incorrect fuel enrichments are loaded into their normal core positions. Conditions which could produce power

maldistributions and the core protection against a

maldistribution of power is presented in Subsection 4.3.4.3.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.3-3 UFSAR Rev 30 10/2014 15.3.1 Loss of Reactor Coolant from Small Ruptured Pipes or from Cracks in Large Pipes Which Actuates Emergency Core Cooling

15.3.1.1 Accident Analysis

The small break analysis is reported in Reference 51. This report, in accordance with the RELAP5-based LOCA evaluation model (BAW-10192PA) presents the results of the spectrum of cold leg breaks from 0.01 ft 2 in cross-sectional area up to and including the full double-end break of the cold leg piping. Double-ended breaks of the HPI and CFT lines were analyzed. Hot leg breaks were not included in the analysis. Sensitivity studies provided with the EM concluded that since their location would prevent a direct loss of the emergency injection fluid out the break, all of the ECC fluid injected by the Core Flooding Tanks, the HPI pump, and the LPI pump must enter the core before being lost out the break.

With SG heat transfer available, the consequences of the small break transient decrease with decreasing break size. Depending on the break location and imposed boundary conditions, a break area can be identified for which the HPI or normal makeup system is capable of matching the leak rate ensuring an orderly shutdown. For example, the leak rate resulting from the rupture of a 3/4" schedule 160 instrument line (0.002 ft

2) is matched by the normal makeup system by 1000 seconds without a complete loss of the pressurizer liquid level. The pressure at 1000 seconds is approximately 7 psi above the HPI actuation setpoint. For larger break areas, the HPI system will be actuated during the transient and will supply borated water to the Reactor Coolant System at a sufficient rate to maintain continuous core cooling. Most break sizes result in a calculated core mixture level below the top of the core, resulting in a temperature excursion. The peak clad temperatures, however, were less than 1800°F and all of the acceptance criteria of 10CFR50.46 were met. Other nonradiological aspects of this accident are discussed in

Chapter 6.

The environmental consequences of this accident would be less than the environmental consequences discussed as a part of Subsection 15.4.6, Major Rupture of Pipes Containing Reactor Coolant Up to and Including Double-Ended Rupture of the Largest Pipe in the Reactor Coolant System (Loss-of-Coolant Accident). Discussion of the methods of detecting small leaks and the time required to evaluate the occurrence is found in Subsection 5.2.4.

15.3.1.2 Effects of Plant Changes

To accommodate a twenty-four (24) month cycle the SFAS RCS Low Pressure Trip analytical setpoint was revised to 1515 psia. This setpoint was included in the analyses contained in Reference 51. The analyses demonstrated that the accident acceptance criteria will be met.

15.3.1.3 Impact of Replacement Steam Generators

As part of the Steam Generator Replacement Project, AREVA performed an evaluation (reference 65) to determine the impact of the replacement Steam Generators on the analyses of record. That evaluation concluded that the existing small break analysis remains applicable with the replacement Steam Generators installed.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.3-4 UFSAR Rev 30 10/2014 15.3.2 Minor Secondary System Pipe Break

15.3.2.1 Accident Analysis

This accident is defined as the rupture of any steam line in the secondary system of small area.

This accident is discussed as a part of Subsection 15.4.4, Steam Line Break.

15.3.2.2 Impact of Replacement Steam Generators

As part of the Steam Generator Replacement Project, AREVA performed an evaluation (reference 64) to determine the impact of the replacement Steam Generators on the analyses of record. The net effects of the replacement Steam Generator's secondary mass limit, its slightly greater heat transfer capacity, and smaller steam outlet nozzles result in the minor secondary system pipe break remaining bounded by the double-ended Main Steam Line Break. Therefore, the existing analyses remain applicable with the replacement Steam Generators installed.

15.3.3 Inadvertent Loading of a Fuel Assembly Into an Improper Position

15.3.3.1 Identification of Causes

The arrangement of assemblies with different fuel enrichments in the core will determine the power distribution of the core during normal operation. The loading of fuel assemblies into improper core positions or the incorrect preparation of the fuel assembly enrichment could alter the power distribution of the core.

The following fuel misloadings are considered in this accident analysis:

a. Misloading a fuel pellet with an incorrect enrichment in a fuel rod.

b. Misloading a fuel rod with an incorrect enrichment in a fuel assembly.

c. Misloading a fuel assembly with an incorrect enrichment into the core.

Enrichment errors in fuel pellets or pins beyond normal tolerances will result in local power shapes which vary from those calculated with nominal enrichments. Assembly enrichment errors or loading errors may cause gross power shapes which are peaked in excess of reference design values. The Incore Instrumentation System is designed to monitor assembly power distributions as discussed in Section 7.8, and is capable of detecting assembly misplacement. Fuel pellet and pin enrichment loading errors in excess of manufacturing tolerances are prevented by extensive loading controls and procedures. One such manufacturing process to assure that fuel pellets have been properly loaded, is by in-process gamma-scanning. Also, gadolinia rods within an assembly are loaded with approved templates that identify their location. Gross fuel assembly misplacement in the core is prevented by administrative loading procedures and the prominent display of identification markings on each fuel assembly upper end fitting. During fuel loading, these identification numbers are compared to the loading diagram by at least two persons working independently.

Following each refueling, an incore power distribution is taken during startup testing and compared to calculated power distributions. Gross fuel assembly misplacement would be detected by the incore detectors during this phase by the fact that a radial power tilt is present or developing. Similarly, the out-of-core detectors will indicate quadrant tilt conditions.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.3-5 UFSAR Rev 30 10/2014 15.3.3.2 Analysis Section 4.3 presents the core protection analysis for the accident.

Power distributions resulting from enrichment loading errors in pellets, rods, and assemblies have been analyzed. The thermal-hydraulic conditions resulting from the perturbed power shapes have been determined and compared to design values. The enrichments analyzed are conservative and are the greatest possible enrichments.

The following cases have been analyzed:

Case 1 A 3.40 wt% Uranium-235 fuel pellet was loaded in the center of a 2.70 wt% U-235 fuel rod. The nuclear analysis was performed using a one-dimensional axial representation of the fuel rod.

Case 2 A 3.40 wt %

235U fuel rod was loaded into the high flux region of a 2.70 wt % assembly. The nuclear analysis was performed in two dimensions.

Case 3A The center assembly of an equilibrium fuel cycle core was replaced by a 3.40 wt % assembly. This was an enrichment increase of 0.55 wt %

235 U.

Case 3B An equilibrium cycle symmetrical assembly (near the outer edge of the core) was replaced by a 3.4 wt % 235U assembly. This was an enrichment increase of 0.55 wt %

235 U. The power distributions from cases 3A and 3B were obtained from a two-dimensional, x-y plane, PDQO7 analysis.

Results The power distributions for case 3A are presented in Figure 15.3.3-1. Only power peaks in the central region of the core are appreciably altered by a misloaded center assembly.

Similarly, misplacement of the center assembly by a higher enriched assembly does not cause a radial power tilt. The maximum radial-local power peak occurs in the center assembly, which is a detector assembly. The incore instrumentation would detect an assembly power increase of this magnitude.

The power distribution for case 3B is presented in Figure 15.3.3-2. A significant power tilt results, which would be detected by the incore instrumentation. Misplaced assemblies in other core orientations would introduce radial power tilts which would be more easily detected than

this case.

The thermal analysis of cases 1 and 2 (misloaded pellet and pin) resulted in localized DNBR reductions which are limited to the misloaded pellet or pin.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.3-6 UFSAR Rev 30 10/2014 Conclusion

Strict administrative controls will prevent enrichment errors during fuel fabrication and during fuel loading. In the unlikely event that gross core loading errors occur, the incore instrumentation is designed to detect it.

15.3.3.3 Impact of Replacement Steam Generators As part of the Steam Generator Replacement Project, AREVA performed an evaluation (reference 64) to determine the impact of the replacement Steam Generators on the analyses of record. That evaluation concluded that steam generator performance does not impact this accident in any way. Therefore, the existing analyses remain applicable with the replacement Steam Generators installed.

MI DDDDDCJ DDDDD DOD r;;;-i NOMINAL POWER OISTRIBL Ml SLOAOEO ASS94BLY PO,. DI STRIBUTIOM DAY I S *BESSE NUCl. EAR PO IER SU Tl ON RADIAL X LOCAL ASSE*LT PHER D ISTRI llTll*

• CASE JA FI GU RE l 5 . 3 .

3-l REVISION 0 JULY 1982 NOMINAL POWER DISTF!IBUT:Oh MISLOADED ASSEMBLY POWER DISTRIBUTION DAVIS-BSSE IILIClEAR )OVER STAT ION RAOl K X LOCAL ASSElLY POWER OlSlrRlbUTlOW - CASE 30 FIGURE 15.3.3-2 REVISION 0 JULY 1 982 Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-1 UFSAR Rev 30 10/2014 15.4 CLASS 3 - DESIGN BASIS ACCIDENTS Class 3 events are accidents of very low probability, but are postulated because the conservatively calculated potential offsite doses resulting from these accidents is significant.

This will have a bearing on the design and performance of the station to ensure that fission product release to the station environment will not result in undue risk to the public health and safety. These postulated accidents may require operation of engineered safety features. Potential offsite doses resulting from design basis accidents must be less than the guideline values given in 10CFR100. Table 15.4-1 summarizes the accidents categorized as Class 3

events.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-2 UFSAR Rev 30 10/2014 TABLE 15.4-1 Class 3 Events

Event Analysis assumptions Effect Waste gas tank rupture A tank is assumed to contain the gaseous activity evolved from degassing all of the reactor coolant following operation with 1% defective

fuel. Environmental results are shown in Table 15.4.1-1. Steam generator tube

rupture The reactor has been operating with 1% defective fuel and 1-gpm steam generator tube leakage.

Following rupture of the

steam generator tube, isolation of the affected generator is not achieved until the Reactor Coolant

System is cooled down and

depressurized below the lowest pressure set point on

the main steam safety valves. Reactor trips on low Reactor Coolant System pressure. Environmental effects are described in Table 15.4.2-3. CRA ejection accident All fuel rods that experience DNB are assumed to release their total gap activity to the reactor coolant (following operation with 1% defective fuel). Reactor trip occurs on high flux or high pressure.

Some fuel clad failure.

Table 15.4.3-6 presents environmental effects.

Steam line break The reactor has been operating with 1% defective fuel and 1 gpm steam Generator tube leakage.

Reactor coolant leakage into the steam generator continues until the Reactor Coolant System is cooled down and depressurized to ambient conditions. Reactor trips following a large rupture on high flux or low coolant pressure.

Environmental effects are indicated in Table 15.4.4-4.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-3 UFSAR Rev 30 10/2014 TABLE 15.4-1 (Continued)

Class 3 Events

Event Analysis assumptions Effect Break in instrument line or lines from primary system that penetrate containment Double-ended rupture of 2-1/2 in. letdown line outside

Reactor Building. The reactor has been operating with 1% defective fuel. System isolates and reactor trips on low pressure. Flashing of coolant and partial release of activity. Table 15.4.5-2 presents environmental

effects. Loss-of-coolant accident Environmental effects are based on the release of all

the gap activity with the reactor operating with 1%

defective fuel. Release of 100% noble gases, 50% iodine, and 1% solid fission products considered as maximum hypothetical accident. See Table 15.4.6-1 for environmental effects.

Table 15.4.6-2 presents environmental effects of maximum hypothetical accident. Fuel handling accident Gap activity is released from 56 fuel rods in one assembly while in spent fuel storage pool. See Tables 15.4.7-2a and 15.4.7-3 for environmental effects.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-4 UFSAR Rev 30 10/2014 15.4.1 Waste Gas Decay Tank Rupture

15.4.1.1 Identification of Causes

The waste gas decay tank is used in the radioactive waste disposal system to store radioactive gaseous waste from the station until such time that the radioactive decay renders the gas safe for release to the site environment. Rupture of a waste gas tank would result in the premature release of its radioactive contents to the station ventilation system and to the atmosphere through the station vent. Although it is not considered credible, this accident was analyzed in order to evaluate the resultant dose at the site boundary.

15.4.1.2 Analysis of Effects and Consequences 15.4.1.2.1 Safety Evaluation Criteria

The safety evaluation criteria for this accident is that resultant Exclusion Area Boundary and Low Population Zone doses shall not exceed a small fraction of the 10CFR100 limits and the Control Room doses shall not exceed the limits of General Design Criteria 19.

Beginning with cycle 5, the fuel cycle length was extended to 18 months. The plant Technical Specifications limit the RCS activity to a value which is significantly less than the iodine activity associated with 1% failed fuel, which was assumed in the original evaluation of this accident given in section 15.4.1.2.2. Therefore, the Waste Gas Decay Tank activities presented in Table 15.4.1-2 bound the activities associated with any fuel cycle. For this reason no additional evaluation was performed for this accident to support extended 24 month fuel cycles.

15.4.1.2.2 Methods of Analysis

A waste gas tank is assumed to contain all noble gases in one reactor coolant volume at the end of the third cycle and the iodine from one reactor coolant volume after a DF of 10

5. This DF is a conservative addition of 100 for the purification demineralizers and 10 4 for the degasifier. Operation with 1 percent defective fuel is assumed. In addition, the following assumptions are made: a. The accident duration is 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />, that is, 99.9 percent of all airborne activity is released over two hours.

b. The Control Room Ventilation System is isolated upon receipt of a high radiation signal in the Auxiliary Building exhaust stack (station vent). Isolation requires a maximum of 11 seconds (including 6 seconds maximum for instrument response). The air delivery rate for the five seconds after the fan is shut down was calculated using the following assumptions:

1. One supply fan with its related return fan is operating.

2. The supply and return fans are de-energized simultaneously.

3. The system dampers are fully open during the 5 second time interval.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-5 UFSAR Rev 30 10/2014 c. The tank is assumed to rupture and release its contents to the Auxiliary Building. As the release is vented to the atmosphere by the Emergency Ventilation System, it passes through charcoal filters that remove iodine with an efficiency of 95 percent. However, no iodine filtering is credited.

The results are as follows:

Air Delivery Rates Following Fan Shutdown Shutdown Time Intervals Flow Rate (cfm) Running cfm 21,920 One Second 20,824 Two Seconds 19,471 Three Seconds 17,839 Four Seconds 16,440 Five Seconds 15,227 (NOTE: A conservative flowrate of 22,000 cfm for 0 seconds to 15 seconds was analyzed.)

4. The dispersion factors for the fuel-handling accident and waste gas tank rupture are equal since the release point is the same for both accidents.

The release point (station vent) is 160 feet horizontal distance from the control room intake and 180 feet vertical distance.

15.4.1.2.3 Results of Analysis

The rupture of a waste gas decay tank would release the entire contents of the tank to the auxiliary building atmosphere. The Auxiliary Building is ventilated and discharged to the station vent. In the analysis, however, the activity is assumed to be released from the waste gas decay tank to the atmosphere over a two-hour time period. Table 15.4.1-2 lists the isotopic release to

the atmosphere.

Atmospheric dilution for the site and low population zone boundary doses is calculated using the 2-hour dispersion factor developed in Section 2.3. The two-hour integrated doses at the exclusion area boundary and the 30-day doses at the outer boundary of the low-population zone, as shown in Table 15.4.1-1, are well bel ow the limits of the 10CFR100 guideline. The Control Room doses provided in Table 15.4.1-1 are well below the limits of General Design Criteria 19.

See Section 15.4.1.2.1 for the evaluation to support extended fuel cycles.

15.4.1.2.4 Impact of Replacement Steam Generators

As part of the Steam Generator Replacement Project, AREVA performed an evaluation (reference 64) to determine the impact of the replacement Steam Generators on the analyses of record. That evaluation concluded that steam generator performance does not impact this accident in any way. Therefore, the existing analyses remain applicable with the replacement Steam Generators installed.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-6 UFSAR Rev 30 10/2014 TABLE 15.4.1-1 (1) Resultant Doses From Waste Gas Tank Rupture Exclusion Area Boundary 0 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> Low Population

Zone 0 to 30 days Thyroid dose, rem 2.20 x 10-3 1.14 x 10-4 Whole-body dose, rem 0.317 1.65 x 10-2 Operator in Control Room

0 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> Thyroid dose, rem 1.72 x 10-2 Beta-skin dose, rem 2.45 Total body gamma, rem 8.09 x 10

-2 (1)See Section 15.4.1.2.1 for the evaluation to support extended fuel cycles.

TABLE 15.4.1-2 (1) Activity Released Due to Waste Gas Tank Rupture (Ci)

I-131 1.70 X 10-2 I-132 1.59 X 10-2 I-133 1.52 X 10

-2 I-134 1.79 X 10

-3 I-135 7.29 X 10

-3 Xe-131m 8.15 X 10 2 Xe-133m 1.04 X 10 3 Xe-133 9.48 X 10 4 Xe-135m 1.17 X 10 2 Xe-135 2.45 X 10 3 Xe-137 0 Xe-138 1.64 X 10 2 Kr-83m 1.33 X 10 2 Kr-85m 6.34 X 10 2 Kr-85 7.05 X 10 3 Kr-87 3.43 X 10 2 Kr-88 1.04 X 10 3 Kr-89 0 (1)See Section 15.4.1.2.1 for the evaluation to support extended fuel cycles.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-7 UFSAR Rev 30 10/2014 15.4.2 Steam Generator Tube Rupture

15.4.2.1 Identification of Causes

The environmental effects associated with the complete severance of a steam generator tube are evaluated. For this occurrence, activity contained in the reactor coolant would be released to the secondary system. Some of the radioactive noble gases and iodine would be released to the atmosphere through the condenser air removal system and the steam line safety valves.

15.4.2.2 Accident Analysis

15.4.2.2.1 Safety Evaluation Criteria The safety evaluation criteria for this accident are:

a. Resultant doses shall not exceed 10CFR100 limits.

b. Additional loss of reactor coolant boundary integrity shall not occur due to a loss of secondary side pressure and resultant temperature gradients.

Beginning with cycle 5, the fuel cycle length was extended to 18 months. The plant Technical Specifications limit the RCS activity to a value which is significantly less than the iodine activity associated with 1% failed fuel, which was assumed in the original evaluation of this accident given in section 15.4.2.2.4. Therefore the resultant doses presented in Table 15.4.2-3 bound the radiation doses for any fuel cycle (References 54 and 55). The plant Technical Specification limits are based on the NRC evaluation as documented in the NRC SER. For this reason no additional evaluation was performed for this accident to support extended 24 month fuel cycles.

15.4.2.2.2 Methods of Analysis

In analyzing the consequences of this failure, the following sequence of events is assumed to occur (input parameters are shown in Table 15.4.2-1 and results are summarized in Table

15.4.2-2).

a. Reactor Coolant System Response

1. A double-ended rupture of one steam generator tube occurs with unrestricted discharge from each end to the secondary side of the steam generator.

2. The initial leak rate exceeds the normal makeup to the Reactor Coolant System, and system pressure decreases. No initial operator action is assumed, and a low Reactor Coolant System pressure trip will occur.

3. After reactor trip, the Reactor Coolant System pressure continues to decrease until high pressure injection is automatically actuated. The capacity of high pressure injection is sufficient to compensate for the leakage and thereafter, it is assumed that the operator has properly diagnosed the problem and takes action by initiating a reactor coolant system cooldown and depressurization.

After the Reactor Coolant System reaches 500°F and 1065 psia the affected steam generator is isolated.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-8 UFSAR Rev 30 10/2014 4. Following isolation of the affected steam generator, the cooldown is continued at 100°F/hr with the unaffected steam generator until the Reactor Coolant System temperature reaches 280°F. Thereafter, cooldown to ambient conditions is continued using the Decay Heat Removal System.

b. Secondary System Response

1. Following reactor trip, the turbine stop valves close, and under normal conditions, the ICS changes the pressure setpoints on the turbine bypass system from 920 psig to 995 psig on the turbine bypass valves to the condenser. However, it has been assumed in this analysis that this function

has failed.

2. Following closure of the turbine stop valves, the secondary system pressure will increase, causing the turbine bypass valves and the main steam line safety valves to open. Steam relief directly to the atmosphere will continue until secondary system pressure drops below the main steam line safety valve setpoint.

3. Thereafter, the turbine bypass valves will continue to relieve steam to the condenser. The operator initiates Reactor Coolant System cooldown and depressurization by further opening the turbine bypass valve on the unaffected

steam generator.

4. When the Reactor Coolant System pressure has fallen below the 1050 psig low steam safety valve setpoint, the operator closes and latches the turbine bypass valve to the condenser to complete final isolation of the affected steam

generator.

The distinguishing characteristic of this event is the buildup of activity levels in the secondary steam system. The condenser offgas monitor will detect an increase in secondary steam system noble gas release to the station vent and alert the operator via an alarm when the activity level exceeds normal operating limits. This alarm, coupled with dropping RC pressure and pressurizer level indications, provides sufficient information for the operator to diagnose the occurrence of this accident in comparison to other possible events. In addition, N-16 radiation monitors located on each steam line will provide the operator with rapid identification of the affected steam generator.

The method used to calculate all coolant activities is described in detail in Chapter 11.

15.4.2.2.3 Results of Analysis

The results of this accident are summarized in Tables 15.4.2-2 and 15.4.2-3. The analysis shows that the consequences of this accident are within the established criteria stated in Subsection 15.4.2.2.1. A steam generator tube failure concurrent with partial loss of coolant flow produces effects less severe than the rated power condition described above. This is due to the less severe cooldown transient that occurs.

See Section 15.4.2.2.1 for evaluation to support extended fuel cycles.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-9 UFSAR Rev 30 10/2014 15.4.2.2.4 Environmental Consequences During the venting time of the affected steam generator, it is conservatively assumed that all fission products leaking from the Reactor Coolant System go directly to the atmosphere. Prior to the tube rupture, the unit is assumed to have been operated with a 1 gpm tube leak and 1%

defective fuel rods. Volatile activity that reaches the condenser is released to the atmosphere after passing through the condenser air removal system. An iodine partition coefficient of 10 4 is assumed between the liquid and vapor phases in the condenser (references 1 and 2) during this

accident.

Individual isotopic activities which enter the secondary system in the reactor coolant during this accident are listed in Appendix 15A, Table 15A-4. The doses presented in Table 15.4.2-3 are conservatively calculated assuming that all the iodine and noble gas activity contained in the reactor coolant is released to the affected steam generator (435 gpm for 34 minutes) and subsequently released directly to the environment (i.e. credit for iodine partitioning is not considered). The result and doses are within the10CFR100 guidelines.

See Section 15.4.2.2.1 for the evaluation to support extended fuel cycles.

15.4.2.2.5 Consequences of Less Severe Ruptures

A rupture of a steam generator tube which results in a leak rate equal to the primary makeup capability will be detected by N-16 detectors in the main steam line headers in less than 15 seconds, setting off an alarm in the control room. Upon receipt of the alarm signal in the control room, the operator initiates station shutdown.

Site boundary thyroid and whole body doses are necessarily lower than those listed in Table 15.4.2-3, since the activity released from the primary to the secondary is far less.

See Section 15.4.2.2.1 for the evaluation to support extended fuel cycles.

15.4.2.2.6 Effects of Plant Changes

15.4.2.2.6.1 Twenty-Four Month Fuel Cycles To support the change to twenty-four month operating cycles, it was necessary to revise the SFAS RCS Low Pressure trip analytical setpoint to 1515 psia. Table 15.4.2-1 lists the High Pressure Injection setpoint (which corresponds to the SFAS RCS Low Pressure trip) as 1600 psig. This change in the trip setpoint that actuates HPI does not affect the results of the analysis nor the environmental consequences of the accident.

The water assumed to exit the steam generator until isolation occurs is held at a constant radioactive concentration throughout the analysis. No credit is taken for the dilution of the RCS fluid that would occur when HPI begins to enter. Since the time to isolation of the steam generator is not affected, the same amount of radiation is released to the environment. Therefore, no change in the environmental consequences occurs.

15.4.2.2.6.2 Impact of Replacement Steam Generators

As part of the Steam Generator Replacement Project, AREVA performed an evaluation (reference 64) to determine the impact of the replacement Steam Generators on the analyses of record. That evaluation concluded that since the replacement Steam Generator's tube inside Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-10 UFSAR Rev 30 10/2014 diameter and length are unchanged, the leak flow rate is unaffected. Therefore, the existing analyses remain applicable with the replacement Steam Generators installed.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-11 UFSAR Rev 30 10/2014 TABLE 15.4.2-1 (1) Steam Generator Tube Failure Parameters Initial tube leak rate in affected steam generator, gpm 435

Leak rate in unaffected generator, gpm 1 Normal makeup rate, gpm 70 High pressure injection setpoint, psig 1600 (2) Assumed defective fuel, %

1 (1)See Section 15.4.2.2.1 for the evaluation to support extended fuel cycles.

(2)See Section 15.4.2.2.6 for revised setpoint.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-12 UFSAR Rev 30 10/2014 TABLE 15.4.2-2 Summary of Steam Generator Tube Failure Analysis

Low pressure trip occurs at 8 min. High pressure injection automatically starts at 12 min.(1) Operator takes action by initiating Reactor Coolant System cooldown and depressurization at

20 min. Final isolation of affected steam generator is achieved at 34 min. Initiation of Decay Heat Removal System is achieved at 184 min. Volume of injection water required to compensate for reactor coolant leakage prior to affected steam generator isolation

1978 ft 3 Steam venting time to the atmosphere from affected steam Generator

30 sec. Steam vented to the atmosphere from affected steam Generator 18,667 lb. Steam venting time from unaffected steam generator to atmosphere (through condenser)

145 min. Total steam vented to atmosphere (through condenser) for unaffected steam generator prior to actuation of Decay Heat Removal System 286,000 lb.

(1)See Section 15.4.2.2.6 for evaluation of revised SFAS Low Pressure trip setpoint.

TABLE 15.4.2-3 (1) Resultant Doses From Steam Generator Tube Rupture

Exclusion area boundary 0-2 hours LPZ boundary 0-30 days Thyroid dose, Rem 27.1 1.41 Whole body dose, Rem 0.23 0.012 (1)See Section I5.4.2.2.1 for the evaluation to support extended fuel cycles.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-13 UFSAR Rev 30 10/2014 15.4.3 CRA Ejection Accident

15.4.3.1 Identification of Causes

Reactivity excursions initiated by uncontrolled CRA withdrawal (Section 15.2) were shown to be safely terminated without harming the reactor core or the integrity of the Reactor Coolant System. In order for reactivity to be added to the core at a more rapid rate, physical failure of a pressure barrier component in the Control Rod Drive Assembly must occur. Such a failure could cause a pressure differential to act on a Control Rod Assembly and rapidly eject the CRA from the core region. The power excursion due to the rapid increase in reactivity is limited by the Doppler effect and terminated by reactor protection system trips. No operator action is required.

Since Control Rod Assemblies are used to control load variations and boron dilution is used to compensate for fuel depletion, only a few Control Rod Assemblies are inserted (some only partially) at rated power level. Thus, the severity of a CRA ejection accident is inherently limited because the amount of reactivity available in the form of CRA worth is relatively small.

Accident Bases:

Using an analytical method based on diffusion theory, the worth of the most reactive Control Rod Assembly in each CRA group was determined for different Control Rod Assembly configurations. The maximum CRA worths and other important parameters used in the study are shown in Table 15.4.3-1. The tripped CRA worth corresponds to the minimum worth available with the maximum-worth CRA stuck out at BOL and EOL.

The severity of the CRA ejection accident depends on the worth of the ejected CRA and the reactor power level. The Control Rod Assembly group of greatest worth is the first in the entire CRA pattern to be withdrawn. The maximum worth of a CRA in this group can be as high as 1.3 percent k/k, but a CRA would have this worth only when the reactor is subcritical. The details of the Control Rod Assembly worth calculations are presented in Chapter 4, and the methods of selecting the number of CRA's in each group are presented in Chapter 7.

When the reactor is subcritical, the boron concentration is maintained at a level that ensures that the reactor is at least 1% subcritical with the CRA of greatest worth fully withdrawn from the core. Thus, a CRA ejection will not cause a nuclear excursion when the reactor is subcritical and all the other CRA's are in the core.

A detailed analysis has been performed at power (2772 MWt) and zero power for CRA worths from 0.2 to 0.7 percent k/k.

A maximum CRA worth of 0.65 percent k/k at power (2772 MWt) has been considered as a limiting value to demonstrate the inherent ability of the system to safely terminate this postulated transient.

A CRA must be fully inserted in the core to have the greatest reactivity worth value. Assuming that the failure occurs so that the pressure barrier no longer offers any restriction to the ejection and that there is no viscous drag force limiting the rate of ejection, the CRA travel time to the top of the active region of the core is calculated to be 0.176 second. Since most of the reactivity is added during the central 75% of this travel, only this distance is used in the analysis, resulting in an ejection time of 0.15 second for the analysis.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-14 UFSAR Rev 30 10/2014 Fuel Rod Damage:

The consequences of a CRA ejection accident depend largely on the rate at which the thermal energy resulting from the nuclear excursion is released to the coolant. If the fuel rods remain intact while the excursion is being terminated by the negative Doppler coefficient and by reactor trip, then the energy release rate is limited by a relatively low surface-to-volume ratio for heat transfer. The energy stored in the fuel rods will then be gradually released to the coolant (over a period of several seconds) at a rate that poses no threat to the integrity of the Reactor Coolant System. However, if the magnitude of the nuclear excursion is so great that the fuel rod cladding does not remain intact, then both fuel and cladding may be dispersed into the coolant to such an extent that the heat transfer rate increases significantly.

Power excursions caused by reactivity disturbances of the order of magnitude occurring in CRA ejection accidents could lead to three potential modes of fuel rod failure. Failure by the first mode occurs when internal pressures developed in the fuel rod are insufficient to cause cladding rupture, but subsequent heat transfer fuel to cladding raises the temperature of the cladding and weakens it until local failure occurs. Departure from nucleate boiling (DNB) usually accompanies and contributes to this mode of failure, and little or no fuel melting would be expected. In this mode of failure, fuel fragmentation is usually only minor, and any dispersal of fuel to the coolant would occur very gradually; system contamination would be the worst probable consequence.

The second mode of failure occurs when significant fuel melting causes a rapid increase in internal fuel rod pressure which, combined with a loss of cladding strength at higher temperatures, causes the fuel rod cladding to rupture (ref. 6). Some fuel vaporization may occur, contributing to the pressure buildup. Considerable fragmentation and dispersal of the fuel would be expected in this mode.

The third and most serious mode of fuel rod failure is the occurrence of extensive fuel melting and subsequent vaporization due to a very large and rapid reactivity transient in which there is insufficient time for heat to be transferred from the fuel to the cladding. In this mode, destructive internal pressures can be generated without increasing cladding temperatures significantly.

In evaluating the effects of these modes of failure, two failure thresholds are considered. The first, associated with a gradual and usually minor cladding failure, may be defined approximately by the minimum heat flux for DNB at the cladding surface. The second failure threshold, defined as the enthalpy threshold for prompt fuel failure with significant fragmentation and dispersal of fuel and cladding into the coolant, is used to describe the energy required to cause failure by either the second or third mode of failure.

A correlation of the results of different experiments conducted on Zircaloy-2-clad UO 2 fuel rods at TREAT (ref. 7) has been interpreted by the experimenters to show a threshold at 280 cal/g of fission energy input. That is, below this value the fuel rod can be expected to remain intact, and above this value fragmentation can be expected. The enthalpy corresponding to the melting

point of UO 2 is about 260 cal/g, (ref. 8) and the heat of fusion is at least 78 cal/g (ref. 9). Thus, the 280 cal/g represents a condition in which only part of the fuel is molten. Also of interest as a probable indication of the degree and rapidity of fuel and cladding dispersal are the measurement of pressure rise rates in the autoclave in the TREAT experiments (ref. 7).

Preliminary analysis indicates that there is only a modest pressure rise up to an energy input of 400 cal/g. Above 500 cal/g, however, there is a very definite pressure pulse. Thus, between 400 and 500 cal/g there is a transition, which probably corresponds to the change from the second to the third failure mode discussed previously.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-15 UFSAR Rev 30 10/2014 A fuel failure threshold of 280 cal/g, at the pellet radius corresponding to the average temperature of the hottest fuel pellet, has been used in this study to define the extent of fuel

failure.

In computing the average enthalpy of the hottest fuel pellet during the excursion for the rated power (2772 MWt) cases, it is assumed that no heat is transferred from the fuel rod between the time the accident is initiated and the time when the neutron power returns to the rated power level (2772 MWt). For the zero-power cases, the enthalpy increase was based on the peak

value of the average fuel temperature. In all cases the average enthalpy rise from the integrated energy or the fuel temperature traces is multiplied by the maximum peaking factor to obtain the enthalpy increase in the hottest fuel pellet.

The latest correlation of the ANL TREAT (ref. 7) data from the meltdown experiments on

Zircaloy-2-clad UO 2 fuel rods shows the threshold for the zirconium-water reaction to be 210 to 220 cal/g energy input. A conservative threshold value of 210 cal/g is used in this study.

In calculating the volume of the core experiencing burnout in a given CRA ejection accident, it is assumed that any DNB condition results in burnout for each rod where the DNB occurs. DNB in a CRA ejection transient is assumed to occur whenever the peak thermal power of a given fuel rod exceeds the peak at steady-state conditions that could result in a DNB, which in turn is assumed to occur for a DNBR of 1.3 using the W-3 correlation.

In determining the environmental consequences from this accident, an even more conservative approach is taken in computing the extent of DNB experienced in the core. All fuel rods that undergo DNB to any extent are assumed to experience cladding failure with subsequent release of all the gap activity. Actually, most of the fuel rods will recover from DNB, and no fission product release will occur. The fuel rods that experience DNB at BOL are assumed to have EOL gap activities.

15.4.3.2 Accident Analysis

15.4.3.2.1 Safety Evaluation Criteria

The safety evaluation criteria for this assumed accident are:

a. The effects of a Control Rod Assembly ejection accident shall not further violate the Reactor Coolant System integrity.

b. The resultant doses shall not exceed 10CFR100 limits.

15.4.3.2.2 Methods of Analysis

A B&W digital computer program has been used to analyze the CRA ejection accident. This program agrees to within a few percent in all cases with CHICKIN (reference 10). The core heat transfer model allows for up to 30 radial mesh points in the fuel and clad, and the mesh size can be different in the two regions. The model accounts for the gap conductivity and film coefficient of heat transfer. Reactivity feedback is calculated in each mesh point and in the coolant and is weighted for inclusion in the kinetics simulation. The thermal properties are input separately for each mesh point but remain constant with time. The loop model includes a simulation of the steam generator which can have a nonlinear hea t demand input on the secondary side. Trip action is initiated on high or low Reactor Coolant System pressure or on high neutron flux.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-16 UFSAR Rev 30 10/2014 Decay heat can be taken into account as well. This code was used to calculate the neutron and thermal power, integrated energy, reactivity components, pressure, and fuel rod and loop temperatures. Six delayed neutron groups are considered. The control rod trip is represented by a multi-segment curve of reactivity insertion during trip versus time, obtained by combining the actual CRA worth curve with a CRA velocity curve. Nominal values for the various nuclear and physical parameters used as inputs are listed in Table 15.4.3-2.

As a check on the point kinetics calculation, the CRA ejection accident was also analyzed for a limited number of cases in support of the Technical Specification CRA worth using the two-dimensional, space-and-time dependent TWIGL digital computer program (reference 11). The point kinetics model assumes that the flux shape remains constant during a transient. This flux shape contains peaking factors which reflect unusual CRA patterns such as the flux adjacent to a position where a high worth CRA has been removed. Therefore, these point kinetics peaking factors are much higher than any that would actually occur in the core during normal operation.

The purpose of using an exact space-time calculation is to find the flux shape during a transient. However, a transient wherein a CRA is ejected from the core must necessarily start with a flux shape that is depressed in the region of the ejected CRA. In fact, the higher the worth of the CRA, the more severe becomes the depression. This flux depression also causes a fuel temperature depression. When the CRA is ejected from this position, the flux quickly assumes a shape that shows some local peaking.

However, when this "exact" peaking is applied to a region initially at depressed fuel temperatures, as it is in the case of the regions adjacent to the ejected CRA, the resultant energy deposited in these regions causes a lower peak temperature and peak thermal power than does applying an arbitrary maximum peaking factor to an undepressed peak power region.

The results from TWIGL were used to calculate the maximum total energy deposited in each region of the core following a CRA ejection; the highest energy is reported in Table 15.4.3-3.

The result is that the hottest region simulated in the TWIGL code actually undergoes a less severe transient than the hottest fuel rod assumed in the point kinetics model. As seen in Table 15.4.3-3 this result is uniformly true for all CRA worths.

15.4.3.2.3 Results of Analysis

Zero Power Level:

The nominal BOL and EOL CRA ejection analysis was performed at 10

-3 of power (2772 MWt), and the results can be seen in Table 15.4.3-4. No DNB and no fuel damage would result from these transients.

A sensitivity analysis has been performed around these two cases in which the Doppler and moderator coefficients, trip delay time, and CRA worth were varied. Figure 15.4.3-1 shows the peak neutron power as a function of ejected CRA worth from 0.2 to 0.7 percent k/k. The curve shows two distinct parts corresponding to worths less than and values near to and greater than. Figure 15.4.3-2 shows the corresponding results for the peak thermal power. It is seen that for CRA worth values near prompt critical, the period is small enough to carry the transient through the high neutron flux trip. For lower values the pressure trip is relied on. No DNB occurs for any of these parameter variations.

Figure 15.4.3-3 shows that the peak enthalpy in the fuel for the CRA worths in the range being evaluated never exceeds 80 cal/g. Therefore, no threshold for damage is approached.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-17 UFSAR Rev 30 10/2014 Figures 15.4.3-4 and 15.4.3-5 show the peak neutron and thermal power as a function of Doppler coefficient from -0.8 to -1.9x10

-5 (k/k)/°F. It is seen that the variation is relatively small. Similar results are shown in Figures 15.4.3-6 and 15.4.3-7 for the variation of the moderator coefficient from -4.0x10

-4 to 2.0x10

-4 (k/k)/°F. The slope of the curve for 10

-3 power (2772 MWt) at BOL is the greatest slope of any of the four curves because this case relies on the pressure trip, which makes it a longer transient. Figure 15.4.3-8 shows the effect of the trip delay time on the peak thermal power. It is seen that there is very little effect.

Rated Power:

An analysis was performed for a BOL CRA ejection at power (2772 MWt). The results of this analysis are shown in Table 15.4.3-4. A sensitivity study was made around this case and around the same CRA worth at EOL. Figures 15.4.3-1 through 15.4.3-8 show these results.

As seen in Figure 15.4.3-2, the peak thermal power shows relatively little change with increased CRA worth. The peak neutron power in Figure 15.4.3-1 does show a marked change with increased worths, but the thermal effect is small because the transients are rapidly terminated by the Doppler effect. As further evidence of this small thermal effect, the peak fuel enthalpies are given in Figure 15.4.3-3. The threshold for the zirconium-water reaction is not reached until values of BOL and EOL ejected CRA worths are above any that are considered feasible. The effects on the peak neutron and thermal powers of varying the Doppler and moderator coefficients and trip delay time are shown in Figures 15.4.3-4 and 15.4.3-8.

The results of the DNB calculation for BOL are shown in Figure 15.4.3-9. For the BOL analysis, ejection of the maximum CRA worth of 0.65 percent k/k at rated power (2772 MWt) results in 45% of the pins in DNB.

15.4.3.2.4 Energy Required to Produce Further Reactor Coolant System Damage

The reactor vessel has been analyzed to estimate the margin that exists between the CRA worths assumed for the calculated CRA ejection accident transients and those worths that could initiate reactor coolant system failure. The pressure vessel material is SA-533 Grade-B steel.

Table 15.4.3-5 lists the values used in this analysis. The radial deformation assumed to represent failure of the vessel is 50% of the total elongation, or 0.13 in./in.

To calculate the weight of an explosive charge required to reach 50% elongation, the vessel was simulated by a single cylinder with the same OD as the actual vessel, but with an increased thickness to account for the thermal shield and core barrel.

Using the formula for the equivalent vessel, the required weight of explosive charge was calculated. The results indicate that 1410 pounds of TNT would strain the mid-meridian ring to the 50%, i.e., 0.13 in./in. The 1410 pounds of TNT have an energy equivalent of 6.74x10 8 cal. Ejected CRA worths higher than those reported in the preceding sections were analyzed to estimate the transient required to obtain an energy release equivalent to 1410 pounds of TNT.

These cases were evaluated to find the amounts of fuel melting and zirconium-water reaction.

Using the conservative assumption that all the fuel that exceeds the melting threshold is fragmented, dispersed into the coolant, and quenched to the coolant average temperature, a total thermal energy release can be determined. The conversion of this energy release to an equivalent deformation energy is dependent upon the duration of the release. TNT has an energy release in microseconds and a deformation conversion efficiency of about 50%. The Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-18 UFSAR Rev 30 10/2014 energy generated during a reactor transient from the zirconium-water reaction and a molten fuel dispersal is in the range from milliseconds to seconds. Thus, the conversion efficiency to deformation energy would be considerably less and is assumed to be 1/5 that of TNT (reference 13). Using these figures, the reactor vessel's capability is 3.37x10 8 cal, and under the foregoing assumptions, a reactivity addition of 1.52 percent k/k is required to release energy necessary to cause deformation of the vessel.

15.4.3.2.5 Conclusions

The hypothetical CRA ejection accident has been investigated in detail at two different initial reactor power levels: nominal power (2772 MWt) and zero power; both BOL and EOL conditions were considered. The results of the analysis prove that the reactivity transient resulting from this accident will be limited by the Doppler effect and terminated by the Reactor Protection System with no serious core damage or additional loss of the coolant system integrity. Furthermore, it has been shown that an ejected CRA worth greater than 1.52 percent k/k would be required to cause a pressure pulse, due to prompt dispersal of fragmented fuel and zirconium-water reaction, of sufficient magnitude to cause rupture of the pressure vessel, whereas the maximum CRA worth shown in Table 15.4.3-1 is about a factor of 2 less.

As a result of the postulated pressure housing failure associated with the accident (Subsection 15.4.3.1), the reactor coolant is lost from the system. The rate of mass and energy input to the Containment Vessel is considerably lower than that subsequently reported for the smallest rupture size considered in the loss-of-coolant analysis (Chapter 6). The maximum diameter hole size resulting from a CRA ejection is approximately 2.76 inches. This lower rate of energy input results in a much lower containment vessel pressure than those obtained for any rupture sizes considered in this loss-of-coolant accident.

15.4.3.2.5.1 Partial Coolant Flow Condition

For partial flow operation, two ejected CRA worths were analyzed at nominal (2772 MWt) and

zero power.

The results of the 0.65% ejected CRA worth case show peak thermal power values of 96% and 126% for two and three pumps, respectively. Calculations for percent pins in DNB and peak enthalpy of hottest fuel rod were made and show that the CRA ejection protection criteria was not exceeded. The worst case was for three pumps and its results showed 6% of the fuel pins were in DNB and the peak enthalpy of the hottest fuel rod was 194 cal/gm.

15.4.3.2.6 Environmental Consequences

The environmental consequences of this accident are calculated by conservatively assuming that all fuel rods undergoing DNB release all of their gap activity to the reactor coolant. Just the activity in the gap is released from the fuel assembly since only the DNB limits are exceeded, and the worst possible consequence of exceeding DNB limits is possible cladding defects. The fuel rods in DNB are calculated for the ejection of the maximum CRA worth at BOL from nominal power (2772 MWt). Subsequently this gap activity and the activity in the reactor coolant from operation with 1% defective fuel is released.

After the CRA ejection occurs (causing a 2.76-inch Reactor Coolant System rupture to the Containment Vessel), the reactor coolant will undergo a subcooled expansion for approximately 50 seconds. As the Reactor Coolant System continues to depressurize, the reactor coolant will be at saturation temperature corresponding to the reactor coolant pressure. When the Reactor Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-19 UFSAR Rev 30 10/2014 Coolant System pressure falls below the setpoint of the turbine bypass valves, the secondary system can be isolated. Moody leak flow rate tables were used to determine the reactor coolant to secondary system leakage during the accident. Using these tables and nominal power (2772 MWt) conditions, a leak flow area was calculated for the 1 gpm leak rate. This area and the Moody tables were then used to determine the reactor coolant to secondary leakage during the Reactor Coolant System depressurization. It is conservatively assumed that choked flow in the leak flow area did not occur, and that the steam generator pressure was low enough to allow critical flow during the accident. This yields a maximum leak flow rate. It is also conservatively assumed that the friction factor associated with the leak flow area was equal to 1.0, and that the discharge coefficient was equal to 1.0. Using these assumptions, it is calculated that 5 gallons is released to the atmosphere from the condenser. A gas-to-liquid partition factor of 10

-4 is assumed for the iodine in the condenser, (refs. 1 and 2) but the noble gases are assumed to be released directly to the atmosphere.

All reactor coolant that is not released to the secondary system is released to the Containment Vessel. Fifty percent of the iodine released to the Containment Vessel is assumed to plate out.

Fission product activities for this accident are calculated using the methods discussed in Chapter 11. Doses resulting from this accident were evaluated using the environmental models and dose rate calculational methods discussed in the section on the loss-of-coolant accident.

Table 15.4.3-6 shows the resulting thyroid and whole body doses for a 2-hour exposure at the exclusion distance and for a 30-day exposure at the low population distance, which include the dose contribution due to the activity released to the atmosphere via the secondary system and that released via containment vessel leakage. Activity released due to normal operation within the Technical Specification Limits were not considered in this accident analysis and are considered to be negligible. The doses resulting from the accident are well below the guideline values of 10CFR100.

15.4.3.2.7 Additional Analyses

Moderator Coefficient Evaluation

As part of the above analyses, sensitivity studies were performed with moderator coefficients as negative as -4.0 x 10

-4 k/k/°F at both Hot Full Power (HFP) and Hot Zero Power (HZP) conditions (Reference 31). As expected, these studies demonstrated that the Control Rod Ejection event is less severe as the moderator coefficient becomes more negative (see Figures 15.4.3-6 and 15.4.3-7). Therefore, a Control Rod Ejection event occurring with a moderator coefficient of -4.0 x 10

-4 k/k/°F at either HFP and HZP conditions will continue to meet the Safety Evaluation Criteria of Section 15.4.3.2.1.

Control Rod Ejection Re-analysis A re-analysis of the Control Rod Ejection Accident (CREA) was performed (See reference 40) using USAR methodology inputs and verified with new methodology, by use of the RELAP5 code. The RELAP5 methodology was within 3% of the USAR's old KAPP code methodology in calculating the resulting effective Full-Power-Seconds (FPS) for the CREA. Using the acceptance criterion of a fuel enthalpy of 210 cal/gm (threshold value for Zirconium-water reaction) the maximum total peak would be 3.43, which is higher than the currently allowed design total peak of 2.96. Therefore, the CREA re- analysis bounds the design peak assumption. BWFC also verified that the 45% of fuel rod failures originally calculated as the consequence of the CREA and assumed in the offsite dose calculation is bounding. This is so because the original calculations were done with a very conservative point (0-D) kinetics model Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-20 UFSAR Rev 30 10/2014 in combination with an adiabatic fuel enthalpy calculation. A 3-D re-analysis with a stronger ejected rod assumption (0.8% k/k) and a more realistic, non-adiabatic model yielded results of approximately 20% of the rods in DNB. The use of a higher total peaking factor (3.43) would result in a proportionately higher number of pins in DNB, but would be well below the 45% fuel rod failures assumed for the radiological dose calculations.

Extended Fuel Cycles

Additional evaluations were performed to support the extended fuel cycle. The assumptions used in the new evaluations are more conservative than the assumptions given in USAR section 15.4.3.2.6, which discusses the environmental consequences due to a control rod ejection accident. The original analysis provided in this section takes credit for the availability of the condenser to reduce the iodine releases (using a gas-liquid partition factor of 10,000) from the secondary side to the environment. This assumption is not as conservative as the NRC's SER because off-site power may not be available following this accident. That is, the evaluations performed by the NRC assumed that off-site power is not available for this accident.

The following assumptions, used in the new evaluations, are more conservative than the assumptions given in the USAR section 15.4.3.2.6.

1. The fuel rod gap activity is assumed to be 10% of the iodine and noble gas activity in the fuel.

2. No credit for iodine partitioning at the condenser is assumed. All the iodine and noble gas activity released to the secondary side is released to the environment.

These two assumptions are consistent with Regulatory Guide 1.77 and the NRC Safety Evaluation Report for Davis-Besse.

Furthermore, it is conservatively assumed that 100% of noble gas gap activity and 50% of the iodine gap activity from the fuel rods reaching DNB are available for release from the containment simultaneous with the rod ejection accident. This assumption is conservative because the size of the opening in the RCS due to the ejected rod is very small (2.76 inch diameter) and considerable time would be required to release all the activity to the containment.

Using the above assumptions, the total dose via both release (i.e., steam generators and containment) pathways are as follows:

Thyroid Whole Body Exclusion Area Boundary dose (0-2 hr.) 47 Rem 0.3 Rem

Low Population Zone dose (30 days) 5 Rem 0.03 Rem The USAR acceptance criterion for this accident's calculated doses requires that they be less

than 10CFR100 guidelines. The acceptance criterion for both the NRC SER and the Standard Review Plan (SRP) requires that the calculated doses (at the exclusion area boundary and the Low Population Zone) be well within the guidelines of 10CFR100 (i.e., 25% of 10CFR100 guidelines, or 75 Rem Thyroid, 6 Rem Whole Body Dose). Since the calculated doses are less than 25% of 10CFR100 guidelines, the USAR, th e NRC SER, as well as the SRP acceptance criteria are satisfied.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-21 UFSAR Rev 30 10/2014 Reanalysis of Control Rod Assembly Ejection Accident for Mark-B-HTP Fuel Assemblies

Implementation of the Mark-B-HTP fuel assembly design commenced with fuel cycle 15. The BHTP critical heat flux correlation is utilized to predict DNBR for the Mark-B-HTP fuel assemblies. This correlation is applicable to fuel assemblies containing the M5Ž HTP spacer grids. Because of the fuel assembly design change, a re-analysis of the Control Rod Assembly (CRA) Ejection Accident was performed to incorporate the BHTP critical heat flux correlation. The reanalysis was performed in accordance with the safety criteria and analysis methodology defined and/or referenced by Reference 23.

The acceptance criteria for the CRA ejection accident are, (1) a fuel enthalpy of less than 210 cal/gm (threshold value for zirconium-water reaction) and, (2) a total core fuel rod failure of less than 45 percent. A fuel rod failure rate of 45 percent of the total number of fuel rods in the core is assumed by the offsite dose calculation.

A NEMO-K (3-D kinetics model) reanalysis was performed using the design total peak of 2.97, an ejected rod worth of 0.65 percent k/k, and a non-adiabatic enthalpy model. Results of the reanalysis indicate that approximately 41 percent of the rods achieve DNB and the fuel enthalpy is less than 210 cal/gm during the accident. Therefore, the acceptance criteria of this accident are satisfied.

15.4.3.2.8 Impact of Replacement Steam Generators As part of the Steam Generator Replacement Project, AREVA performed an evaluation (reference 64) to determine the impact of the replacement Steam Generators on the analyses of record. The steam generators play only a minor role in this event, supporting the cooldown and depressurization of the Reactor Coolant System. The evaluation concluded that the replacement Steam Generator design is essentially identical to the original Steam Generator design for these purposes. Therefore, the existing analyses remain applicable with the replacement Steam Generators installed.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-22 UFSAR Rev 30 10/2014 TABLE 15.4.3-1 Control Rod Assembly Ejection Accident Parameters

Maximum worth of ejected CRA, % k/k 0.65 CRA ejection time, sec. 0.15

Rated power level, MWt 2772 (See Note 1)

Reactor trip delay time

High flux trip, sec. 0.4

High-pressure trip, sec. 0.6 CRA drive trip time to 2/3 insertion, sec. 1.4

(1) The reanalysis described in USAR Section 15.4.3.2.7 and in Reference 40 was based on 102% of 2772 MWt.

TABLE 15.4.3-2 Nominal Values of Input Parameters for CRA Ejection Accident Analysis (See Note 2, 4)

BOL EOL Delayed neutron fraction, B eff 0.00689 0.00516 Neutron lifetime, msec.

34.6 33.0 Moderator coefficient, (k/k)/°F 0.13x10-4 -3.0x10-4 (see Note 1)

Doppler coefficient, (k/k)/°F -1.28x10-5 -1.45x10-5 Reactor coolant inlet temperature, °F 555.2 555.2 Initial system pressure, psia 2200 2200 Total nuclear peaking factor, Fq 2.89 2.89 (see Note 3)

Average fuel temperature of average pellet, °F 1200 1304 Average fuel temperature of hottest pellet, °F 2400 2490 (1) Sensitivity studies have shown that a moderator coefficient of -4.0x10

-4 k/k/°F at End of Life (EOL) yields acceptable results (see Section 15.4.3.2.7).

(2) Reference 40 lists the nominal values of Input Parameters for the re-analysis of the CRA Ejection Accident.

(3) A total nuclear peaking factor of 3.43 has been shown in a re-analysis of the Control Rod Ejection Accident (reference 40) to yield acceptable results (See Section 15.4.3.2.7)

(4) Input parameters associated with an EOC T AVE reduction maneuver are analyzed and verified acceptable and results are included in the associated cycle reload report. (Reference USAR Appendix 4B)

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-23 UFSAR Rev 30 10/2014 TABLE 15.4.3-3 Comparison of Space-Dependent and Point Kinetics Results Of Fuel Enthalpy (based on 2772 MWt)

Peak-to-average values Fuel enthalpy cal / g Ejected CRA worth

% k/k TWIGL Point kinetics TWIGL Point kinetics BOL rated power 0.38 3.04 3.24 125 150 0.83 2.67 3.24 174 225 BOL zero power 0.56 4.1 3.24 38 60 0.83 4.4 3.24 48 71 TABLE 15.4.3-4 Summary of Control Rod Assembly Ejection Accident Analysis (based on 2772 MWt)

Initial power level, % rated power Ejected CRA worth, % Dk/k Peak power, Neutron% rated power Thermal 0.1 (BOL) 0.65 76 63 0.1 (EOL) 0.65 982 41 0.1 (BOL) 1.0 6,128 156 0.1 (EOL) 1.0 13,612 149 100.0 (BOL) 0.65 702 165 100.0 (EOL) 0.65 1,545 148 Percent of fuel rods in DNB due to ejection of a 0.65 % k/k CRA worth at 100% power BOL, % 45 Reactor coolant to secondary leakage during reactor coolant system depressurization,

gallons 5 NOTE: The reanalysis described in USAR Section 15.4.3.2.7 and in Reference 40 was based on 102% of 2772 MWt.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-24 UFSAR Rev 30 10/2014 TABLE 15.4.3-5 Reactor Vessel Parameters Vessel temperature, °F 600 Yield strength (0.2% offset), psi 55,000 Ultimate strength, psi 80,000 Ultimate strain (u), % 26 Strain energy (E s) per unit volume up to Strain equal to 1/2 ultimate strain, In.-lb/in.

3 8,000 Strain energy (E s) per unit volume up to Ultimate strain, in.-lb/in.

3 17,000 Equivalent pressure vessel dimensions OD, in. 188.25 ID, in. 166.69 Thickness, in.

10.78 The expression (12) used for the weight of explosive required to strain the vessel a given amount is 0.811 15.0Rt/R0373.047.1w1085.1RRt/R117.041.3E407.1 W i15.0 i85.05 2 i 2ei e where W = charge weight (TNT or Pentolite), lb w = weight density of vessel material, lb/ft 3 , R i = initial internal radius of vessel, ft, R e = initial external radius of vessel, ft, t = initial wall thickness of vessel wall, ft, E = wall strain energy, in.-lb/in.

3.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-25 UFSAR Rev 30 10/2014 TABLE 15.4.3-6 Resultant Doses From a CRA Ejection Accident (1) Exclusion area boundary 0-2 hours LPZ boundary 0-30 days Thyroid dose, Rem 1.36 0.254 Whole body dose, Rem 1.14 x 10-2 4.75 x 10-3 (1)See Section 15.4.3.2.7 to support extended fuel cycles.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-26 UFSAR Rev 30 10/2014 15.4.4 Steam Line Break

15.4.4.1 Identification of Causes

The loss of secondary coolant due to a failure of a steam line between the steam generator and the turbine causes a decrease in steam pre ssure and thus places a demand on the control system for increased feedwater flow. Increased feedwater flow, accompanied by steam flow through the turbine stop valves and the break, lowers the average reactor coolant temperature and pressure. The reactor trips on low pressure or high flux, depending on the break size (see subsection 15.4.4.2.3.1). The operation of the Emergency Core Cooling System provides effective core cooling and the ultimate shutdown of the core through its boron addition.

Analyses have been performed to determine the effects and consequences of a loss of secondary coolant due to a double-ended steam line rupture between the containment vessel and the Main Steam Isolation Valves, since this location will maximize the radiation release to offsite. However, to evaluate the effect on the containment vessel, it was then assumed the energy associated with this accident was released to the containment vessel.

15.4.4.2 Accident Analysis

15.4.4.2.1 Safety Evaluation Criteria

The safety evaluation criteria for this accident are:

a. The core shall remain intact for effective core cooling.

b. Loss of reactor coolant boundary pressure integrity resulting from steam generator tube failure due to the loss of secondary side pressure and resultant temperature gradients shall not occur.

c. Resultant doses shall not exceed 10CFR100 guideline values.

15.4.4.2.2 Methods of Analysis

The rate of reactor system cooling following a steam line break accident is a function of the steam generator water inventory available for cooling. The unfouled inventory as a function of power is shown in Table 15.4.4-1. The largest inventory, at rated power, results in the greatest mass available for cooling. Thus, the fouled steam generator inventory as shown in Table 15.4.4-2 was used in this analysis.

Other conservative assumptions used in the analysis of the steam line failure accident are as follows:

a. A minimum tripped CRA worth with the maximum worth CRA stuck out of the core of 3.5% k/k was used. This worth accounts for the moderator deficit, Doppler deficit and reduction in CRA worth to produce a 1% k/k subcritical margin at hot shutdown end-of-life core conditions. CRA worths greater than this minimum

tripped CRA worth are shown in Chapter 4 to always be available even with the highest worth CRA stuck out.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-27 UFSAR Rev 30 10/2014 b. Conservative end-of-life Doppler and moderator coefficients were used. The large negative values of these coefficients produce the greatest reactivity insertion due to the Reactor Coolant System cooldown resulting from the accident.

c. The reactor is assumed to be operating at 102% (of 2772 MWt) power before the accident. Other parameters used in the analysis are summarized in Tables

15.4.4-1 and 15.4.4-2.

d. Loss of off-site power was not assumed.

e. No operator action is required to mitigate the accident to meet the acceptance criteria.

f. The systems required to function during the transient are as follows:

Function System Reactor trip RPS SG isolation/turbine trip (main steam line and main feedwater lines) SFRCS Auxiliary feedwater initiation SFRCS High Pressure injection initiation SFAS g. The steam line rupture accident was analyzed assuming a complete double-ended rupture of the largest steam line. As indicated in Table 15.4.4-1 the analysis was based on the rupture of a 33.9" ID pipe (36" OD). The results shown in Figure 15.4.4-1 are consistent with this pipe diameter.

h. Credit is taken for the turbine stop valve closure. This provides a more reliable redundant means for main steam isolation of the unaffected steam generator. Stop valves of the stem-sealed type have been used on 50 mW and larger General Electric steam turbines since 1948. Over 300 valve-years of service on nuclear turbine stop valves have been accumulated without a known failure to close.

Based on experience through December 1972, G.E. has predicted a valve sticking rate of 0.26 failures per million hours at a 50 percent confidence.

Incipient sticking conditions have been found only on high temperature fossil-fuel units and have been due to the accumulation of an oxide layer in the stem and bushing. Oxidation is not experienced at the relatively low temperature of water cooled nuclear reactor applications. In addition, G.E. turbine stop valves are not subject to failure due to silica build-up between the stems and bushings, since the stem-sealed design precludes the transport of steam-carried impurities back along the stem. Further, there is a reduced tendency for carryover from a once through steam generator. Periodic full-closure testing of the stop valves will disclose any sticking conditions, so that a shutdown could be made to make the necessary

correction.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-28 UFSAR Rev 30 10/2014 i. The low steam generator pressure SFRCS trip results in MSIV closure within 6 seconds.

Each main steam isolation valve is designed with the capability of closing within 5 seconds with steam flow in the normal direction and a differential pressure across one valve of 910 psi. This psi differential to the atmosphere is the maximum that would occur under rupture conditions.

The expected pressure drop across a main steam isolation valve is tabulated as

follows:

% power P, psi 100 6.9 75 3.5 50 1.6 The maximum allowable pressure drop is 8 psig.

j. The main steam line check or non-return valve is located in the turbine building and is not essential for safe shutdown of the plant.

The integrity of the non-return check valves due to a main steam line break upstream of the valve is addressed in subsection 10.3.3.

The non-return check valves are not required to function during accident conditions as supported by the steam line break evaluation of this section. Furthermore, the closure of the turbine stop valves will provide the same degree of mitigation of the blowdown of the unaffected SG.

k. In the event of a main steam line rupture, it is required that the main feedwater stop valves, control valve and the startup valve which is on the bypass line of the control valve be closed. The closure of the control valve and startup valve is to backup the stop valve to insure that main feedwater is isolated. The need for the closure of the main feedwater line together with the closure of main steam isolation valve is to effectively isolate the affected steam generator.

The feedwater stop valve is Q-listed, while the feedwater control valve and startup control valve are not Q-listed.

The stop valve is designed to close in 17 seconds to achieve containment vessel isolation. [Note: 17 seconds was used in this analysis and is not the actual valve stroke time.]

The feedwater control valves and startup control valves have been designed to provide a rapid reduction in feedwater flow following a reactor trip. It is required that the flow reduction to the steam generators be approximately 3 to 4 percent of full flow within 6 to 10 seconds. To accomplish this criteria, the feedwater control valve is designed to close in a maximum of 7 seconds, and the startup control valve is designed to close in a maximum of 12 seconds.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-29 UFSAR Rev 30 10/2014 The FLASH 2 digital computer program wa s used to determine the characteristics of this accident. The multinode model included a detailed description of both the Reactor Coolant System and the steam generator.

The model provides simulation of most of the secondary system valves including the main and startup feedwater valve, auxiliary feedwater valves, turbine bypass valves, code safety valves, main steam isolation valves, and turbine stop valves.

The model also includes energy balances for the principal steam generator components, the entire Reactor Coolant System (core, loops, and steam generator), and the pressurizer (with both mass and energy transfer). The reactor kinetics, trip logic and action, and a fuel pin simulation with Doppler and moderator temperature feedback are also features of the model.

15.4.4.2.3 Results of Analysis

15.4.4.2.3.1 Minor secondary pipe break:

Minor steam line breaks include all leak areas up to and including the largest steam line other than the main steam line.

A steam line rupture of small area causes a slow decrease in steam pressure. The reactor power will increase with decreasing average reactor coolant temperature as a result of the negative moderator coefficient. The ICS will then cause Control Rod Assembly insertion in an attempt to limit reactor power to 102% of 2772 MWt. A reactor trip occurs due to low reactor coolant pressure or high neutron flux. The high flux and low RC pressure trip functions ensure core protection over the entire steam line break spectrum. The time from rupture to reactor trip as a function of steam line break size for DB-1 is given in Figure 15.4.4-5 for break areas from 1.0 to 5.4 ft

2. The reactor will trip on the shorter of the two trip times shown. Therefore, for break areas larger than about 1.75 ft 2, the reactor trips on low reactor coolant pressure; and for break areas smaller than about 1.75 ft 2, the reactor trips on high flux. Following reactor trip and turbine trip, the turbine stop valves close. The steam generator in the steam loop associated with the rupture blows dry after steam and feedwater isolation on low steam pressure. Decay heat is removed by the unaffected steam generator by steam flow through the turbine bypass system to the condenser. If condenser vacuum is lost, decay heat will be removed by steam relief through atmospheric vent valves and safety valves.

The results of the analysis for the maximum break size at nominal power (2772 MWt) are similar to those discussed above, however, the maximum break size represents the worst condition for a steam line rupture accident.

15.4.4.2.3.2 Double-ended main steam line break:

Following a double-ended main steam line rupture between the steam generator and main steam isolation valve, both steam generators will blow down causing a Reactor Coolant System cooldown and depressurization. The reactor trips on low reactor coolant pressure. Low steam pressure (600 psia) initiates closure of the main steam isolation valves and closure of the main feedwater isolation (block) valves, startup, and main control valves on both steam generators.

The steam generator with the assumed break will continue to blow down after the main steam isolation valves and feedwater valves are closed. The steam pressure in the unaffected steam generator will increase allowing auxiliary feedwater flow to be initiated.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-30 UFSAR Rev 30 10/2014 After feedwater isolation is achieved, the affected steam generator blows dry and auxiliary feedwater flows to the unaffected steam generator. With closure of the main steam isolation

valves, the unaffected steam generator side will repressurize and open the code safety valves, allowing decay heat removal from the reactor system to continue with auxiliary feedwater flow and steam relief through the code safety valves (nine valves per steam line).

The High Pressure Injection System will be actuated during the cooldown period associated with a steam line failure. This system supplies borated water to the Reactor Coolant System to increase the shutdown margin when cooling down below 550°F. During the controlled cooling to atmospheric pressure the addition of boron to the reactor coolant will prevent criticality at lower temperatures. At temperatures above 550°F, no credit is taken for the negative reactivity inserted due to the HPI injection; credit is taken, however, for both the pressure effect and the cooling effect due to the HPI injection.

Low RC system pressure 1515 psia is the principal analytical setpoint for SFAS initiation following a steam line break both inside and outside containment. In the steam line break analysis described in this subsection (15.4.4.2.3.2) and the analysis described in subsection 15.4.4.2.6.6, the HPI system actuation is needed for RC system makeup and is initiated on low RC system pressure in both cases. The low RC system pressure actuation setpoint is reached about 10 seconds after rupture. The conservatism in the delay for the HPI system is discussed in subsection 15.4.4.2.6.6, item 6.

Figures 15.4.4-1 through 15.4.4-3 show the response of the Reactor Coolant System for a double-ended main steam line rupture. Initial ly, both steam generators blow down until a low reactor coolant pressure trip occurs. The reactor coolant temperature leaving the unaffected steam generator increases after the main steam isolation valves close as a result of the pressure recovery and a reduction of the feedwater flow. The temperature of the coolant leaving the affected steam generator decreases until the unit has blown dry, at which time the temperature approaches the inlet temperature. Since the unaffected steam generator main steam isolation valve is closed and the steam generator with the rupture is dry, the Reactor Coolant System temperature can only be lowered as a result of the steam flow from the isolated steam generator through the code safety valves. Eventually, thermal equilibrium is re-established; i.e., the heat removal rate (steam flow through the code safety valves) is equal to the heat input (core decay heat).

The maximum cooling rate occurs during the first 10 seconds of blowdown, with no resultant return to criticality and a minimum DNBR of 1.42 (W-3) (Table 15.4.4-3); therefore, no fuel damage will occur. There is no danger of the hot channel DNBR exceeding the minimum value of 1.3 for this transient since the reactor trips almost instantaneously while the RC flow remains at rated flow. Table 15.4.4-4 shows the resulting thyroid and whole body doses for a 2-hour exposure at the exclusion distance and for a 30-day exposure at the low population distance.

During the first minute following the break, the average tube temperature in the affected steam generator remains above the shell temperature. Since thermal equilibrium is established, the average reactor coolant temperature will remain near the saturated temperature corresponding to the pressure at which the main steam safety valves or atmospheric vent valve (after blocking SFRCS) is set. Therefore, the tube-to-shell temperature difference will approach zero expect

the tubes wetted by Auxiliary Feedwater Spray.

For tubes wetted by Auxiliary Feedwater Spray the tube tensile load is 3379 pounds. The resultant tube stresses will remain less than the stresses corresponding to the design pressure and temperature conditions of the tubes as discussed in Chapter 5.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-31 UFSAR Rev 30 10/2014 15.4.4.2.3.3 Containment Vessel pressure:

The resultant mass and energy releases for the fouled steam generator are shown in Table 15.4.4-2. The resultant increase in the containment pressure due to a 5.4 ft 2 steam line break is 21.4 psi. The temperature response of the containment vapor region to a main steam line break is shown in Figure 15.4.4-4 along with the surface temperature response of the hottest structural heat sink in the containment. Note that, while the containment vapor temperature exceeds the 264°F equipment qualification temperature, the heat sink (which is a thin steel slab) reaches a maximum surface temperature of only 220.4°F. This behavior is characteristic of equipment exposed to short-term temperature transients in a superheated vapor atmosphere. The major process tending to heat the equipment is the condensing heat transfer mechanism. The total heat transfer rate to equipment and structures can be described by the following relationship:

q = h cond (Tsat - T w) + hconv (T v - T w) (1) where q = surface heat transfer rate

h cond = condensing heat transfer coefficient

h conv = convective heat transfer coefficient

T sat = steam saturation temperature at containment atmosphere steam partial pressure

T w = equipment surface temperature

T v = containment vapor temperature

The first term of Equation (1) becomes identically zero forT w > T sat since condensation heat transfer can occur only if the condensable vapor in the region of the condensing surface can be cooled below its saturation temperature.

The maximum value of T sat that occurs in the transient described in Figure 15.4.4-4 is 224 degree°F. For surface temperatures above this maximum saturation temperature, only the second term of Equation (1) can act to Transfer heat to equipment. Since h conv is generally only 1 to 2 percent of the value of h cond, only long exposures to elevated superheated vapor temperatures can bring the equipment temperature above the maximum value of T sat. Thus, for short-term superheated vapor transients such as those encountered in main steam line breaks, a practical equipment and structure maximum temperature is the saturation temperature at

containment atmosphere steam pressure.

(A detailed discussion of this phenomenon can be found in Bechtel Topical Report BN-TOP-3, Rev. 3, submitted to the NRC in August 1975.)

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-32 UFSAR Rev 30 10/2014 15.4.4.2.4 Environmental Consequences

The environmental consequences from this accident are calculated by assuming that:

a. The unit has been operating with a 1-gpm steam generator tube leak in the affected steam generator.

b. The unit has been operating with 1% defective fuel rods.

c. The steam line break occurs between the Containment Vessel and the main steam isolation valve. All other rupture locations would result in lower doses.

d. Reactor coolant leakage into the steam generator continues for 9.0 hours0 days <br />0 hours <br />0 weeks <br />0 months <br /> until the Reactor Coolant System cools down and the pressure differential disappears. A total of 540 gallons of reactor coolant is assumed to be released to the

atmosphere.

The steam line failure is assumed to result in the release of the noble gas and iodine activity contained in the steam generator inventory, the feedwater and the reactor coolant leakage. The iodine, primarily resulting from reactor coolant leakage in the cooldown period following the steam line break, is assumed to be released directly to the atmosphere. Based on these assumptions, the resultant doses from this accident are given in Table 15.4.4-4.

Beginning with cycle 5, the fuel cycle length was extended to 18 months. The plant Technical Specifications limit the RCS activity to a value which is significantly less than the iodine activity associated with 1% failed fuel, which was assumed in the original evaluation of this accident.

Therefore the resultant doses presented in Table 15.4.4-4 and 15.4.4-4a bound the radiation doses for any fuel cycle. The plant Technical Specification limits for the RCS activity are based on the NRC evaluation as documented in the NRC SER. For this reason no additional evaluation was performed for this accident to support extended 24 month fuel cycles.

15.4.4.2.5 Conclusions

This analysis has shown that the reactor trips and remains subcritical, the integrity of the steam generator is maintained, and the environmental doses are within acceptable limits.

15.4.4.2.6 Additional Analyses

15.4.4.2.6.1 Control Room Habitability

There are main steam lines that pass in close proximity to the control room and mechanical equipment room housing the control room ventilation equipment. All the walls of the mechanical equipment room have been designed to withstand the effects of a steam-line rupture. Access into the room has been provided with pressure-tight doors to keep out the steam atmosphere. Access into the control room is from the Turbine Building. The first access door, which is in close proximity to the main steam lines, opens into the elevator lobby. Two additional doors in series must be passed through before entry into the control room. The Turbine Building bas a very large volume; therefore, no significant pressure buildup will result due to a steam-line

rupture.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-33 UFSAR Rev 30 10/2014 In addition, in order to prevent damage to the control room wall due to the whipping of either main steam line or jet impingement after a postulated rupture, the wall thickness has been increased by 12 inches. Direct jet impingement on the control room door from a break in the main steam line is impossible. Therefore, it is not considered credible that an appreciable amount of steam atmosphere can enter the control room to endanger habitability or safe shutdown of the station.

Area radiation monitors are provided for the control room which continuously give the background radiation level. In case of any abnormal increase in the background level, the operator can manually isolate the normal ventilation system and start the Control Room Emergency Ventilation System if needed. The contr ol room doses following a main steam line break accident are less than those for the loss-of-coolant accident presented in Table 15.4.6-2.

15.4.4.2.6.2 Partial Coolant Flow

The most severe steam line break occurs at rated power. For partial flow the cooldown will be slower, and the effects of steam line break less severe.

15.4.4.2.6.3 Steam Line Break (inside containment) Dual S/G Blowdown

Any postulated single failure which results in the opening of a nonactuated atmospheric vent valve in the steam generator not supplying the broken steam line represents a passive failure in addition to the steam line rupture and is considered to be in excess of design requirements. However, an analysis of a double-ended rupture upstream of the main steam isolation valve has been performed assuming the failure of an atmospheric vent valve to close after actuation on the steam line connected to the unaffected steam generator. The extended opening of any one steam relief valve would not be expected to significantly alter the core thermal conditions presented above, since the relief capacities of individual valves have been sized to prevent such possibility.

With the unaffected steam generator isolated, ten valves are capable of relieving steam. One valve is an atmospheric vent valve having a c apacity of 5% rated steam flow. The other nine are safety valves; seven of these safety valves are relieving approximately 7% of full power each; the remaining two are of a lesser capacity. The most severe postulated active failure would be the failure of the atmospheric vent valve to reseat after actuation. However, to demonstrate the capability of the system to accept an even larger steam relief, failure of a code safety valve to reseat will be assumed for this analysis.

The steam line break with dual SG blowdown analysis originally contained in the USAR contained an error with respect to calculated reactor subcriticality. A reanalysis (ref. 21) was

performed with the TRAP2 computer code utilizing updated analytical techniques and assumptions, and is summarized as follows:

The major distinguishing assumption of this steam line break analysis is the failure of a main steam safety valve (MSSV) to reseat, remain ing fully open. Other assumed parameters are listed in Table 15.4.4-6.

Following the double-ended rupture of the main steam line, low steam pressure (600 psia) on the affected steam line actuates SFRCS. This initiates closure of the main steam isolation valves (MSIVs), closure of the main feedwater stop valves, closure of the turbine stop valves, and initiates auxiliary feedwater to the unaffected steam generator (refer to Subsection 7.4.1). In the analyses, the SFRCS trip of the turbine causes a reactor trip vai ARTS. At the time of the Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-34 UFSAR Rev 30 10/2014 break, main feedwater is assumed to instantly go to run-out flow, until reactor trip. After main feedwater isolation the affected steam generator blows dry.

After MSIV closure on the unaffected steam generator, pressure will rise causing opening of the MSSVs. Although this is when the assumed MSSV failure would occur, for simplicity and conservatism the valve is failed open at break initiation.

Figures 15.4.4-6 through 15.4.4-8 show the system response to this accident. The SFRCS-initiated ARTS trip (including delay time) occurs at 0.77 seconds after rupture. After closure of the MSIVs, unaffected cold leg temperatures increase as associated SG secondary pressure increases. The temperature of the affected cold legs decreases until the SG has blown dry at which time it increases to approximately the hot leg temperature. After affected SG blow down, and with auxiliary feedwater established to the unaffected SG, a steady RCS cooldown rate

occurs as steam flow out the stuck open MSSV continues. Although steam flow rates (dependent on SG pressure) will decrease as core decay heat decreases, for conservatism, a constant bounding RCS cooldown rate was assumed.

With respect to minimum DNBR the reanalysis performed did not specifically recalculate DNB ratios. The original analysis performed concluded minimum DNBR values occurred within five seconds of the steam line break, and were well above the limit of 1.3. Initial conditions and assumptions for the reanalysis are essentially identical during initial portions of the transient, and DNBR will be greater during the subsequent cooldown portion of the transient. As such, the previous conclusion regarding DNBR remains valid. Since this is the case, a plot of the heat flux as a function of the time during the transient is not included. The maximum fuel temperature occurs at the time of the rupture, and fuel temperature continually decreases throughout the transient.

During the initial portion of the transient shown, the minimum subcritical margin is 0.569% k/k calculated at 34 seconds after the break. Addi tionally, conservative extrapolation of TRAP2 results was performed to determine if continued MS SV cooling will cause an eventual return to criticality. The conservative extrapolation assumed continued AFW flow to the unaffected SG and operator action to throttle HPI flow to prevent complete filling of the primary system. The extrapolation estimated a return to criticality at approximately 28 minutes after the break.

Review of reanalysis results and comparison with other USAR analyzed steam line breaks indicate that all acceptance criteria for this accident are met. Namely, the reactor trips and remains subcritical, the integrity of the steam generator is maintained, and the environmental doses are within acceptable limits. Note that while conservative extrapolation indicates a return to criticality at 28 minutes after the break, the key assumption of continued AFW flow for 28 minutes is well beyond design-basis required limitations on operator actions, and, as such, these results indicate sufficient time exists to mitigate a main steam line break with concurrent failure of a MSSV on the unaffected steam generator.

15.4.4.2.6.4 Minimum Reactivity Margin Evaluation

The minimum reactivity margins are shown in Table 15.4.4-5 for the following five main steam line break situations:

Case I - 102% of 2772 MWt. Break is inside containment (36" line). No offsite power.

Case ll - 102% of 2772 MWt. Break is inside containment (36 line). Offsite power is available.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-35 UFSAR Rev 30 10/2014 Case III - 102% of 2772 MWt. Break is outside containment and upstream of isolation valves. No offsite power.

Case IV - 102% of 2772 MWt. Break is outside containment and downstream of isolation valves. Offsite power is available.

Case V - Hot standby or low power operation.

Case II represents the steam line break situation presented in subsection 15.4.4.2.3. To readily assess the resultant effect of Case I on minimum reactivity margin, Case II was reanalyzed with loss of offsite power assumed to occur at reactor trip. As shown in Table 15.4.4-5, loss of offsite power would not significantly effect the result. Case III is identical to Case I since the only effect of the steam lines would be to slow down the Case I transient. The minimum subcritical margin for Case IV will be the same as that for Case II if a single failure of the main steam line isolation valve on the affected steam generator is assumed. Otherwise, the subcritical margin for Case IV will be greater than that for Case II since the length of the blowdown period is reduced for the affected steam generator. The subcritical margin for Case V will be larger than that for Case II since in both of these situations, the steam generator inventories are considerably lower than Case II (about 20,000 lbm, compared to 62,500 lbm). Thus, the overcooling of the primary system is reduced and the minimum subcritical margin will be increased.

In all cases analyzed, single failures have been used that lead to increased overcooling of the primary system. Single failures in the feedwater system and in the main steam system have been considered. The failure of the main feedwater isolation valve and of the turbine stop valve have been postulated in the analysis of the steam line break accident, Subsection 15.4.4.2.3.

Both failures have been conservatively assumed to happen simultaneously. Failure of a steam relief valve has been studied in Subsection 15.4.4.2.6.3.

15.4.4.2.6.5 Comparison of Controlling Parameters With and Without Offsite Power

A steam line rupture accompanied by a loss of offsite power is shown to have a subcritical margin slightly less than the case where offsite power is available (Table 15.4.4-5). Immediately upon loss of offsite power the reactor and the reactor coolant pumps trip. The decrease in the reactor coolant flow will inhibit the normal initial rapid cooling of the primary system (with offsite power) thus, decreasing the potential of a return to power. An immediate reactor trip from the loss of offsite power as opposed to the finite time required to trip the reactor on high flux or low pressure with offsite power, increases the potential for a return to power because the core has not increased its stored energy prior to reactor trip. An immediate reactor trip with loss of offsite power also results in less mass release since the feedwater control valves close with reactor trip. Therefore, the variation in the controlling parameters resulting from a steam line rupture with loss of offsite power will be similar to those parameters shown in Figure 15.4.4-1.

15.4.4.2.6.6 Steam Line Break Concurrent with Operator Error Allowing Continued Feedwater

The occurrence of a steam line break accident concurrent with an operator error allowing feedwater to be admitted to the affected steam generator has also been investigated using the following assumptions:

1. The plant was initially operating at nominal power (2772 MWt).

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-36 UFSAR Rev 30 10/2014 2. Operator error allowing a conservative main feedwater pump runout to 135 percent of rated feedwater flow to the affected generator was assumed. This conservative approach bounds an operator error which would allow auxiliary feedwater to be admitted to the affected steam generator.

3. The maximum negative moderator coefficient corresponding to end-of-life conditions of the equilibrium cycle was used, although this is an instantaneous value occurring only at the last moment of operation before refueling.

4. The minimum tripped CRA worth corresponding to the Technical Specification limit for the minimum shutdown margin was used; this was a very conservative

approach.

5. The maximum-worth CRA was assumed to stick out, although the combining of this assumption with 2 through 4 above is clearly an unrealistic approach.

6. The increased capacity of the High Pressure Injection System with decreased Reactor Coolant System pressure has been neglected; design flow rates have been used throughout the transient. Also, the High Pressure Injection System was

assumed to have 35-second delay after the steam line break. The conservatism of this assumption is demonstrated by the sequence of events listed below:

Case I - Loss of offsite power at the instant of the break Sequence of Events Elapsed Time Steam line break/loss of offsite power 0 sec Diesel starts 0.5 sec SFAS setpoint reached 10 sec Diesel up to speed 10.5 sec SFAS time delay 15 sec Diesel sequence step 2/HPI pumps starts 20 sec HPI pump accelerates to speed/HPI pump discharge valve opens 30 sec Total Elapsed Time 30 sec (Time after SFAS setpoint reached) (20 sec)

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-37 UFSAR Rev 30 10/2014 Case II - No loss of offsite power Sequence of Events Elapsed Time Steam line break 0 sec SFAS setpoint reached 10 sec SFAS time delay/HPI pump starts 15 sec HPI pump accelerates to speed/HPI pump discharge valve opens 25 sec Total Elapsed Time 25 sec (Time after SFAS setpoint reached) (15 sec) 7. The boron injection is assumed to be perfectly mixed with all the reactor coolant before entering the core, although the injection occurs at the reactor vessel inlet and so would have the highest concentration in the core region.

8. Perfect heat transfer is assumed in the affected steam generator after the initial part of the transient; that is, the time constant for heat transfer is zero with no stored energy accounted for.

The steam line rupture causes an increase in the heat transfer from the reactor coolant to the feedwater. As Figures 15.4.4-1 through 15.4.4-3 show, this initiates a cooldown of the Reactor Coolant System, such that the reactor trips on low pressure at about 1.13 sec after the rupture (includes a total trip delay of 0.6 second). A main steam pressure reduction to 600 psig trip point initiates an isolation signal that actuates valves isolating both the steam side and the feedwater side of both steam generators. For the cooldown part of the calculations, it is assumed that the main feedwater flow (at 135 percent of rated flow) continues to the affected steam generator. With the above assumptions, the resulting coolant system temperature decrease causes high pressure injection actuation at 35 seconds after the steam line break.

This injection of boron will keep the core subcritical during cooldown below 550°F.

15.4.4.2.6.7 Moderator Coefficient Evaluation Although the Steam Line Break Event is initiated from Hot Full Power (HFP) conditions, it immediately produces a reactor trip which results in the reactor being at least one percent shutdown when Hot Zero Power (HZP) (532°F) conditions are reached. Therefore, since it is the continuing cooldown below HZP conditions that is of concern, the value of moderator coefficient at HZP and colder conditions will determine the reactor response to the Steam Line Break.

Although the original (double-ended rupture of 36 inch pipe diameter) Steam Line Break analyses for Davis-Besse Unit 1 used a constant moderator coefficient of -3.0x10

-4 k/k/°F over all temperatures, later analyses (Reference 31) performed using the TRAP2 computer code instead employed a more realistic reactivity-versus-moderator temperature model, which accounted for both moderator and Doppler effects. Evaluating this reactivity-versus-moderator temperature function produced an average temperature coefficient (combination of moderator and Doppler coefficients) of -3.1x10

-4 k/k/°F over the range of temperatures from HZP down to the lowest RCS temperature that occurred during the Steam Line Break transient. In other words, if a constant temperature coefficient of -3.1x10

-4 "k/k/°F were used for all temperatures Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-38 UFSAR Rev 30 10/2014 below HZP, the resulting reactivity insertion would be identical to that actually calculated in the TRAP2 analyses. Further, since the Doppler coefficient used in the Steam Line Break event was -0.177x10

-4 k/k/°F, the resulting moderator coefficient of -2.923x10

-4 k/k/°F is bounded by the constant value of -3.0x10

-4 k/k/°F assumed in the original analysis. Therefore, the later analyses using TRAP2 are bounded, in terms of reactor response, by the original Steam Line Break analysis, and the temperature coefficient of -3.1x10

-4 k/k/°F for HZP conditions and colder can be considered to be bounding for the Steam Line Break event.

It should be noted that this value of temperature coefficient assumes temperatures of HZP and colder and also assumes that all control rods are fully inserted in the core with the exception of the maximum worth stuck rod, which is fully withdrawn. Limiting moderator coefficients for HFP all rods out conditions may be determined from this temperature coefficient of -3.1x10

-4 k/k/°F, but must account for the differences in temperatures and rod configuration.

15.4.4.2.6.8 Reanalysis of Steam Line Break in Containment

The Containment Vessel's response during a Main Steam Line Break was reanalyzed with mass and energy release data generated by RELAP5/MOD2-B&W computer code. The reanalysis maximized the mass and energy release from the break for the purpose of predicting a conservative peak temperature and peak pressure of the Containment Vessel. Additional details are provided in Section 6.2.1.3.2.

15.4.4.3 Plant Changes and Effects

15.4.4.3.1 Post June 9, 1985 Loss of Feedwater Event

Following the June 9, 1985 loss of feedwater event additional analyses were performed to

demonstrate that the Auxiliary Feedwater (AFW) will not be isolated to both steam generators under assumed limiting single failure conditions following a main steam line break between the steam generator and the associated MSIV. These analyses assumed that the SFRCS low pressure trip on the affected steam generator (steam generator with the break) will trip the turbine and align the auxiliary feed water to the unaffected steam generator. Two cases were

analyzed:

Case A - Turbine Stop Valves (TSVs) close within 1 second of initial SFRCS trip.

Case B - TSVs fail to close (single failure) and MSIVs close within 6 seconds of initial trip.

These analyses show that (Figures 15.4.4-9, 15.4.4-10) the unaffected steam generator pressure would not drop below the SFRCS low pressure trip setpoint provided that the TSVs close in 1 second. Consequently, the SFRCS will not isolate the unaffected steam generator.

The Case B results show that the unaffected steam generator pressure could fall below 600 psid due to failure of the TSVs to close. Since failure of TSV to close represents a single failure, no additional failure (e.g., failure of the AFW isolation valve) needs to be postulated.

It is noted that in order to improve overall reliability, the SFRCS logic was modified such that low pressure in one steam generator would continue to isolate that steam generator, and initiate auxiliary feedwater to other steam generator. A subsequent low pressure trip on the other steam generator is blocked and will not have any effect on the AFW system. If the pressure on the first steam generator recovers above 600 psig, the SFRCS will respond according to Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-39 UFSAR Rev 30 10/2014 conditions in the second steam generator. If both steam generators recover, then SFRCS will respond based upon other plant conditions. Thus the potential for the total loss of AFW to the steam generators under multiple failure conditions is significantly reduced.

15.4.4.3.2 Lowering the RPS Low Pressure Trip Setpoint (Reference 37)

The limiting accident in Section 15.4.4 is the double-ended steam line break (MSLB). For this accident, the reactor trips on low RCS pressure. Consequently, the reduction in the RPS low pressure trip setpoint to 1900 psig would delay the time of reactor trip following the double-ended MSLB. Based upon the rate of depressurization of the RCS prior to the rector trip, lowering the low pressure trip setpoint to 1900 psig would delay the reactor trip by approximately 1.0 second. This delay would have an insignificant effect upon the accident analyses related to RCS overcooling and containment pressurization as discussed below.

For a double-ended steam line break, the delay in reactor trip due to a reduced RCS low pressure trip setpoint does not have a significant impact upon the minimum sub-critical margin or the minimum moderator temperature. In actuality, a delayed reactor trip causes more energy to be added to the RCS, thereby minimizing the overcooling associated with the steam line break. Consequently, using the original RCS low pressure trip setpoint is conservative for analyzing overcooling effects.

A delayed reactor trip does have a slight effect upon MSLB mass and energy release data.

Since isolation of the main steam and main feedwater valves is initiated by the Steam and Feedwater Rupture Control System (SFRCS) on low secondary side pressure, the low RCS pressure setpoint has no impact upon the integrated mass release from the secondary side of the steam generators and the Main Steam System. However, a delayed RCS low pressure trip setpoint does have a slight impact upon the integrated energy release associated with an MSLB. For the design basis MSLB a 1.0 second delay in reactor trip would add approximately 2.7 MBTU of additional energy to the RCS. This additional energy would cause the RCS fluid temperature to be slightly higher than what is presently calculated for the design basis MSLB.

The increase in RCS temperature, in turn, would cause a slightly larger T between the RCS fluid and steam generator secondary side fluid. This increase would result in greater energy addition to the secondary side. If it is conservatively assumed that all the additional reactor energy is added to the integrated energy released by the MSLB, the increase in integrated energy is less than 2% of the total energy associated with the design MSLB. This increase in energy would have a negligible impact upon containment peak pressure and temperature results for the MSLB. The peak containment pressure associated with the MSLB with the increased energy addition is 22.6 psig. The MSLB would still be enveloped by the 2A hot leg break as the design basis accident for the containment.

15.4.4.3.3 Raising the Maximum Allowable Steam Generator Water Level Additional analyses were performed in order to support Technical Specification Amendment 192 (Reference 41 and 42) which allows the maximum steam generator (SG) water level to be 96%

of the Operate Range (OR). The primary concern associated with this change was the additional SG inventory available in MODE 3 which can be released during an accident condition. These analyses examined the consequences of this higher steam generator inventory on containment pressurization following a MSLB, environmental effects for line breaks outside containment, offsite radiological consequences and core criticality.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-40 UFSAR Rev 30 10/2014 Containment Pressurization:

During operation in Modes 1 and 2, revised operating limits have been implemented which are based on main steam superheat and indicated SG Operate Range. For plant operation up to 96% OR, these limits maintain the SG inventory at or below that used in the 100 percent Full Power MSLB containment analyses described in Sections 6.2.1.3.2 and 15.4.4.2.3.

While the plant is in Mode 3, if a Main Feedwater Pump is capable of supplying water to the SG and the SFRCS Low Pressure Trip is bypassed, the SG inventory is limited to 50 inches Startup Range. This limits the amount of energy available for release to the containment to less than that released during a MSLB at 100% full power.

While the plant is in Mode 3, with the SFRCS Low Pressure Trip active, the inventory in the SGs may be increased to a level of 96% OR. If the SFRCS Low Pressure Trip is bypassed, but the possible flow to the SG is limited to that available from the Motor Driven Feed Pump, the inventory may be increased to a level of 74% OR. For these conditions, the mass and energy released will remain bounded by the 100% Full Power MSLB containment analyses, including 10 minutes of feed to the SG.

Line Breaks Outside Containment:

Plant operating limits ensure that the SG inventory in Modes 1 and 2 at 96% OR is bounded by the SG initial inventory assumed for the existing HELB analyses.

Line breaks outside containment were examined to determine if the environmental effects remain bounding despite the larger steam generator inventory available in Mode 3 at 96% OR. The line breaks examined included the MSLB, main feedwater line break, main steam to auxiliary feed pump turbines line break and the Steam Generator Blowdown System line break.

In all cases the mass and energy released were bounded by the existing anal es in Section 3.6, except for one case. In Mode 3, with an initial water level of 96% OR with the MDFP supplying the SGs and with the SFRCS Low Pressure Trip active, the mass of water released is more than that assumed in Section 3.6.2.7.1.5. However, the energy content of the steam exiting the break is always lower at any given time in the transient because of the Mode 3 conditions. As a result.the environmental effects are bounded by the existing analysis.

Reactivity:

A MSLB with increased inventory in the SGs results in rapid overcooling of the RCS, thereby adding positive reactivity to the reactor. Administrative controls have been added to ensure adequate shutdown margin is present to prevent the reactor from attaining criticality during any postulated MSLB.

In Modes 1 and 2, with the maximum SG water level of 96% OR, plant operating limits maintain the SG inventory to that assumed by the existing 100% full power MSLB analyses. Therefore the cooldown is unaffected by the SG level change in Modes 1 and 2.

During Mode 3, the largest cooldown is caused by a MSLB with the steam generators at 74%

OR and the SFRCS Low Pressure Trip bypassed. The cooldown and positive reactivity insertion of this case bounds all other Mode 3 MSLB scenarios. To ensure the reactor remains subcritical in this event, administrative controls include the requirement to determine the boron concentration necessary to compensate for the calculated cooldown and procedural Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-41 UFSAR Rev 30 10/2014 requirements to establish the necessary boron concentration in the RCS prior to raising the SG level above low level limits.

When the plant is in MODE 4, the SGs can onl y induce a very limited cooldown of the RCS following any secondary side line breaks. Therefore, no additional reactivity requirements are needed. No specific feed pump requirements are needed for the same reason. The maximum

SG inventory limit is provided to ensure the SGs remain capable of decay heat removal by maintaining a steam flow path.

The assumptions in Section 15.4.4.2.3 related to Departure from Nucleate Boiling Ratio (DNBR) are consistent with the inventory in the SGs for operation in Modes 1 and 2. In Modes 3 and 4, departure from nucleate boiling can not occur due to the very low heat flux in the reactor.

Radiological Consequences:

The radiological consequences of a MSLB bounds all other steam line breaks. The radiological consequences of a HSLB in Mode 3 with a maximum initial steam generator water level was greater than the values given in Section 15.4.4.2.4., but remained below the NRC acceptance criteria. The results of the dose analysis are listed in Table 15.4.4-4a.

15.4.4.3.4 Lowering the SFAS RCS Low Pressure Trip Setpoint

The SFAS RCS Low Pressure Trip analytical setpoint, which initiates High Pressure Injection during a Steam Line Break (SLB), was revised to 1515 psia. All SLB analyses had used this value as an input assumption. Therefore, this change had no effect on the existing analyses.

15.4.4.3.5 Impact of Replacement Steam Generators

As part of the Steam Generator Replacement Project, AREVA performed an evaluation (reference 64) to determine the impact of the replacement Steam Generators on the analyses of record. That evaluation concluded that the reduction in the replacement Steam Generator steam nozzle area offsets the slight increase in the heat transfer capacity such that the MSLB analysis for core response, presented above, remains applicable with the replacement Steam Generators installed.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-42 UFSAR Rev 30 10/2014 TABLE 15.4.4-1 Steam Line Failure Parameters Steam generator inventory (I) as a function of fractional power level (P) for powers greater than 15% of nominal power (2772 MWt) I = 41,200P + 13,800 lb Maximum pipe size (ID), in. 33.9 Trip Variable Low Pressure Trip Delays Time, sec 0.6 Doppler coefficient (EOL), (k/k)/°F -1.77x10-5 (see note 1)

Moderator coefficient (EOL), (k/k)/°F -3.00x10-4 (see note 1)

Trip delay time (low pressure trip), sec 0.6 CRA movement time to 2/3 insertion during trip, sec 1.4 (1) An equivalent average temperature coefficient (combination of moderator and Doppler coefficients) of -3.1x10

-4 k/k/°F has also been used for temperatures at Hot Zero Power (HZP) and below (see Section 15.4.4.2.6.7), and this temperature coefficient is bounded by the values of moderator coefficient and Doppler coefficient shown in this table.

TABLE 15.4.4-2 Mass and Energy Releases for Building Pressure Analysis Mass, lb Energy, Btu Steam generator inventory (fouled) 62,500 35.9 x 10 6 Feedwater flow to affected steam generator (includes flow until trip and a 17 sec main feedwater control valve closing time) 18,550 8.2 x 10 6 Reactor coolant systems energy transferred --

65.7 x 10 6 Available mass in feedwater line between feedwater control valves and affected steam generator 37,800 16.7 x 10 6 Steam flow from unaffected steam generator (until isolation) 71,950 40.1 x 10 6 Total releases 190,800 166.6 x 10 6

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-43 UFSAR Rev 30 10/2014 TABLE 15.4.4-3 Summary of Steam Line Failure Analysis Minimum subcritical margin during transient, %k/k 0.69 Steam released to atmosphere from affected generator prior to feedwater isolation, lb 118,850 Steam released to atmosphere from unaffected steam generator prior to steam line isolation, lb 71,950Reactor coolant to secondary leakage during reactor coolant system depressurization, lb 1,788Minimum DNBR during transient 1.42 TABLE 15.4.4-4 Resultant Doses From a Steam Line Failure In Mode 1 (1) Exclusion area boundary 0-2 hr LPZ boundary 0-30 days Thyroid dose (Rem) 0.79 0.041 Whole body dose (Rem) 6.7 x 10-3 3.46 x 10-4 (1)See 15.4.4.2.4 referenced re-analyses for the resultant doses from a steam line failure.

TABLE 15.4.4-4a Resultant Doses From a Steam Line Failure In MODE 3 With SG Level at 96% Operate Range Exclusion Area boundary 0-2 Hr. LPZ boundary 0-30 Days Thyroid Dose (Rem) 0.951 0.063 Whole Body Dose (Rem) 3.0 x 10-3 2.0 x 10-4 Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-44 UFSAR Rev 30 10/2014 TABLE 15.4.4-5 Minimum Reactivity Margins for Various Main Steam Line Break Situations

Case Minimum Reactivity Margin, % k/k I. 102% of 2772 MWt. No offsite power. Break inside containment.

0.62 II. 102% of 2772 MWt. With offsite power. Break inside containment.

0.69 III. 102% of 2772 MWt. No offsite power. Break outside containment but upstream of main isolation valve.

0.62 IV. 102% of 2772 MWt. With offsite power. Break outside containment but downstream

of main steam isolation valve.

0.69 V. Hot standby or low power operation.

>0.69

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-45 UFSAR Rev 30 10/2014 TABLE 15.4.4-6 Steam Line Failure with Concurrent MSSV Failure Parameters

Initial Power 102% of 2772 MWt Steam generator inventory 55,000 lbm RPS high flux trip setpoint (with 0.4 second delay) 112% of 2772 MWt SFAS low pressure trip setpoint (with 30 second HPI delay) 1515 psia Doppler coefficient (EOL), (k/k)/°F -1.77x10-5 (see note 1)

Initial moderator coefficient (EOL), (k/k)/°F -3.00x10-4 (see note 1) Initial tripped rod worth, (k/k)/°F -3.5x10-2 Initial boron concentration 16 ppm SFRCS low steam line pressure setpoint 600 psia Turbine Stop Valve closure delay after SFRCS 1 second (.5 sec.

delay, .5 sec.

ramp) Main Steam Isolation Valve closure delay after SFRCS 6 seconds (1 sec.

delay, .5 sec.

ramp) Main Feedwater Stop Valve closure delay after SFRCS 18 seconds (1 sec.

delay, 17 sec. ramp) Auxiliary Feedwater delay after SFRCS 13 seconds

(1) An equivalent average temperature coefficient (combination of moderator and Doppler coefficients) of -3.1x10

-4 k/k/°F has also been used for temperatures at Hot Zero Power (HZP) and below (see Section 15.4.4.2.6.7), and this temperature coefficient is bounded by the values of moderator and Doppler coefficient shown in this table.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-46 UFSAR Rev 30 10/2014 15.4.5 Break in lnstrument Lines or Lines from Primary System That Penetrate Containment

15.4.5.1 Identification of Causes

A break in fluid-bearing lines which penetrate the Containment Vessel could result in the release of radioactivity to the environment. There are no instrument lines connected to the Reactor Coolant System which penetrate the Containment Vessel. There are, however, other piping lines from the Reactor Coolant System to the Makeup and Purification System and the Decay Heat Removal System which do penetrate the Containment Vessel. Leakage through fluid penetrations not serving accident-consequence-limi ting systems is minimized by a double barrier design so that no single, credible failure or malfunction of an active component will result in loss of isolation or intolerable leakage. The installed double barriers take the form of closed piping, both inside and outside the Containment Vessel, and various types of isolation valves.

The most severe pipe rupture with regard to radioactivity release during nonnal station operation occurs in the Makeup and Purification System. This would be a rupture of the letdown line just outside the Containment Vessel but upstream of the letdown control valves. The occurrence of a rupture at this point results in a loss of reactor coolant until the temperature switches on the outlet of letdown coolers close the redundant valves on the inlet to the coolers.

Additionally, the rupture outside containment could be isolated by an SFAS level 2 signal on low RCS pressure. An SFAS level 2 signal would close the letdown line containment isolation valves thereby terminating release of reactor coolant system fluid outside contaimnent. Although both the redundant temperature switches and the SFAS isolation signal are available for isolation of the letdown system, the design analyzed in Subsection 15.4.5.2.3 uses the SFAS actuation on low reactor coolant system pressure to terminate release to the environment. This assumption represents a bounding analysis for total integrated doses by maximizing the blowdown duration.

15.4.5.2 Analysis of Effects and Consequences

15.4.5.2.1 Safety Evaluation Criteria

The safety evaluation criterion for this accident is that resultant doses shall not exceed 10CFR100 limits.

Beginning with cycle 5, the fuel cycle length was extended to 18 months. The plant Technical Specifications limit the RCS activity to a value which is significantly less than the iodine activity associated with 1% failed fuel, assumed in the evaluation of this accident presented in section 15.4.5.2.3, 15.4.5.3.1 and 15.4.5.3.2. Therefore the resultant doses presented in 15.4.5.2.3, 15.4.5.3.1 and 15.4.5.3.2 bound the radiation doses for any fuel cycle. The plant Technical Specification limits are based on the NRC eval uation as documented in the NRC SER. For this reason no additional evaluation was performed for this accident to support extended 24 month

fuel cycles.

15.4.5.2.2 Methods of Analysis

A digital computer program was used to determine loss of coolant characteristics of this accident. The multinode model included a detailed description of the Reactor Coolant System.

The model provides mass, energy, and momentum balances for the Reactor Coolant system nodal arrangement.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-47 UFSAR Rev 30 10/2014 This analysis assumed a complete severance of the 2-1/2 inch letdown line. No operator action was assumed. Coolant was assumed to flow out until the isolation valve was fully closed. Credit was not taken for the reduction in flow during the last seconds while the valve was closing. The normal makeup system was assumed to function which results in a slightly longer time to reach the low reactor coolant pressure setpoint and a correspondingly higher mass

release.

15.4.5.2.3 Environmental Consequences

Figure 15.4.5-1 shows the Reactor Coolant System pressure as a function of time. As can be seen from this figure, the time to 1600 psig is approximately 95 seconds. Adding on the 10 seconds for the valve closure time, coolant escapes for a total period of 105 seconds. The total mass of reactor coolant released is 7,955 pounds.

Assuming the reactor operated with 1 percent defective fuel cladding, the individual isotopic activities released to the Auxiliary Building are listed in Table 15.4.5-1. Since the building ventilation exhaust passes through activated charcoal adsorbers, a reduction factor of 20 is assumed for iodine, based on an adsorber efficiency of 95%.

Atmospheric dilution is calculated using the 2-hour dispersion factor developed in Section 2.3.

The total integrated doses at the exclusion distance as shown in Table 15.4.5-2 are well below the limits of the 10CFR100 guideline.

15.4.5.3 Effects of Plant Changes

15.4.5.3.1 Lowering the RPS Low Pressure Reactor Trip Setpoint

Technical Specification Amendment 149 approved a reduction of the Reactor Protection System (RPS) Reactor Coolant System (RCS) low pressure trip to 1900 psig. This accident was reanalyzed (Reference 32) in support of a reduced RPS RCS Low Pressure trip. For the re-analysis the reactor trip is assumed to occur due to an RCS low pressure trip. As with the present USAR analysis, no credit was taken for isolation from the existing temperature switches. Coolant flow out of the break was assumed until closure of the letdown line isolation valve following an SFAS low RCS pressure actuation.

Following the break, the RCS depressurizes due to the loss of inventory to atmosphere. Based on the B&W analysis, the low RCS pressure condition was reached at approximately 130 seconds following break initiation. Following the reactor trip and assumed coincident loss of offsite power, the RCS continues to depressurize until an SFAS low RCS pressure condition occurred at approximately 250 seconds after break initiation. The SFAS trip at this time caused the letdown line isolation valve to close, thereby terminating the release of RCS inventory.

Assuming a 10 second valve closure time, RCS coolant release occurred for approximately 260 seconds. The total mass of coolant released to the atmosphere by the break is 44,460 lb with an average enthalpy of 430 BTU/lb. The diesel generator and sequence delay times were assumed to be included in the 250 second duration of the transient (i.e., coincident reactor trip and loss of offsite power).

The radiation doses associated with the above mass release are calculated in Reference 33 using the following assumptions.

- All the noble gas activity in the reactor coolant discharged through the break is released to the environment.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-48 UFSAR Rev 30 10/2014

- The iodine activity contained in the portion of reactor coolant that flashes into steam is assumed to become airborne and is released to the environment.

- Although the EVS will be actuated by SFAS, the analysis does not take credit for EVS filters, or plateout of iodine on surfaces.

- The reactor coolant activity is based on activities given in USAR Section 15A for 1% failed fuel, which are considerably higher than the Technical Specification limit of 1 Ci/gm dose equivalent I-131.

The resultant doses due to the revised RPS RCS Low Pressure Trip are:

Exclusion Area Boundary Low Population Zone Thyroid (Rem) 3.52 0.18 Whole Body (Rem) 0.03 0.002 The above doses are higher than those presently reported in the USAR Table 15.4.5-2. This is primarily attributed to the assumptions made in the re-analysis of not taking any credit for any iodine removal due to operation of EVS. Regardless, these results satisfy the Standard Review Plan (NUREG 0800) Section 15.6.2 acceptance criteria that doses be well below 10% of 10CFR100 guideline values.

15.4.5.3.2 Lowering the SFAS Low Pressure Trip Setpoint

In support of twenty four (24) month operating cycles, it was necessary to revise the SFAS RCS Low Pressure Trip Setpoint. The analytical value used in analyses is 1515 psia. Reducing this value from the previous 1600 psig (1585 psia) delays isolation of the letdown line break to 45.51 seconds. The additional mass and energy releases associated with the time delay were determined to be 7,104 lb and 3,054,720 BTU, respectively.

The values were then used in a radiation dose calculation to determine the increase in doses at the exclusion area boundary and the low population zone. The resultant doses due to a RPS RCS Low Pressure Setpoint of 1900 psig and an SFAS RCS Low Pressure setpoint of 1515 psia are:

Exclusion Area Boundary Low Population Zone Thyroid (Rem) 4.83 0.25 Whole Body (Rem) 0.04 0.002 These does are higher than previously reported in Table 15.4.5-2. However, the results still satisfy the acceptance criteria of NRC Standard Review Plan (SRP) Section 15.6.2, "Radiological Consequences of the Failure of Small Lines Carrying Primary Coolant Outside Containment," which requires that doses do not exceed a small fraction of 10CFR100 guideline values, that is, 2.5 Rem and 30 Rem for the whole-body and thyroid doses, respectively.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-49 UFSAR Rev 30 10/2014 15.4.5.3.3 Impact of Replacement Steam Generators As part of the Steam Generator Replacement Project, AREVA performed an evaluation (reference 64) to determine the impact of the replacement Steam Generators on the analyses of record. That evaluation concluded that the replacement Steam Generator design differences do not affect the inputs or integrated leak flow used in the analysis of a letdown line break.

Therefore, the existing analyses remain applicable with the replacement Steam Generators installed.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-50 UFSAR Rev 30 10/2014 TABLE 15.4.5-1 Activity Released to Auxiliary Building (1) From Letdown Line Rupture

Isotope Activity (Ci) Kr-83m 1.58 Kr-85m 8.37 Kr-85 47.2 Kr-87 4.58 Kr-88 14.7 Xe-131m 11.8 Xe-133m 15.4 Xe-133 1330 Xe-135m 5.07 Xe-135 27.9 Xe-138 2.8 I-131 17.6 I-132 26.4 I-133 20.8 I-134 2.72 I-135 10.38 (1)See 15.4.5.2.1 and 15.4.5.3 referenced re-analyses for the activity released from letdown line rupture. TABLE 15.4.5-2 (1) Resultant Doses From Letdown Line Rupture Exclusion area boundary 0-2 hrs LPZ boundary 0-30 days Thyroid dose (Rem) 0.123 6.37 x 10

-3 Whole body dose (Rem) 0.015 7.67 x 10

-4 (1)See 15.4.5.2.1 and 15.4.5.3 referenced re-analyses for the resultant doses from the letdown line rupture.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-51 UFSAR Rev 30 10/2014 15.4.6 Major Rupture of Pipes Containing Reactor Coolant Up To and Including Double-Ended Rupture of the Largest Pipe in the Reactor Coolant System (Loss-of-Coolant Accident)

The computer model CRAFT (Model for Equilibrium LOCA Analysis) was originally used to describe this accident. All other methods and assumptions that were used are described in BAW-10034, Rev. 3 (May, 1972). Subsequent large break loss-of-coolant accident analyses performed in accordance with AEC ECCS "final acceptance criteria" were performed with methods and assumptions as described in BAW-10104, Rev. 5 (November 1988) and

BAW-10105, Rev. 1 (July, 1975).

For Cycle 13 and onward, the large and small break LOCA spectrum was reanalyzed using the RELAP5/MOD2-B&W-based evaluation model (BAW-10192PA, July 1998). The analysis results are summarized in Reference 51 and demonstrate that the acceptance criteria of 10CFR50.46 are met. A discussion of the non-radiological aspects of this accident is provided

in Chapter 6.

15.4.6.1 Accident Analysis

The nonradiological aspects of this accident are discussed in Chapter 6.

15.4.6.2 Safety Evaluation Criterion

The safety evaluation criterion for this accident is that resultant doses shall not exceed 10CFR100 guideline values.

15.4.6.3 Environmental Analysis of Loss-of-Coolant Accident

Safety injection is designed to prevent significant cladding melting in the event of a loss-of-coolant accident. The analysis in Chapter 6 demonstrates that safety injection will prevent cladding melting for a loss-of-coolant accident resulting from Reactor Coolant System ruptures ranging in size from small leaks to the complete severance of a hot-leg coolant pipe.

Without cladding melting, only the radioactive material in the coolant at the time of the accident plus some gap activity is released to the Containment Vessel. The environmental consequences from a loss-of-reactor coolant accident are analyzed by assuming that 1 percent of the fuel rods are defective before the release of reactor coolant to the Containment Vessel.

In addition to the coolant activity, the activity associated with the gap of all fuel rods is also assumed to be released. While perforation of fuel cladding will require some time, it is conservatively assumed that all gap activity in the fuel rods is released directly to the Containment Vessel at the time of the accident. Appendix 15A lists the maximum activity in the coolant in Table 15A-4 and the total fuel rod gap activity in Table 15A-2. The resultant doses due to the maximum break size LOCA evaluated using the guidelines of Safety Guide 4 are given in Table 15.4.6-1.

15.4.6.4 Maximum Hypothetical Accident The analysis in Chapter 6 has demonstrated that, even in the event of a loss-of-coolant accident, no significant core melting vill occur. Hovever, to demonstrate in a still more conservative manner that the operation of the station does not present any undue hazard to the general public, a hypothetical accident involving a gross release of fission products is evaluated. No mechanism whereby such a release occurs is postulated, since this would require a Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-52 UFSAR Rev 30 10/2014 multitude of failures in the engineered safety features, which are provided to prevent such an occurrence.

Fission products are assumed to be released from the core as stated in TID-14844; namely, 100 percent of the noble gases, 50 percent of the halogens, and 1 percent of the solids. Of the iodine released, 50 percent is assumed to plate out and the other half is assumed to remain in

the Containment Vessel atmosphere where it is available for leakage. Although the Containment Vessel leakage rate will decrease as the pressure decays, the leakage is assumed to remain constant at the design leak rate for the first 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />. Thereafter, the leak rate is assumed to be reduced to one-half the design leak rate and to remain at this value for the duration of the accident. These assumptions are consistent with Regulatory Guide 1.4.

The 2-hour thyroid and whole-body doses at the exclusion distance, the 30-day doses at the low-population-zone distance (evaluated using Safety Guide 4), and a Dose Reduction Factor (DRF) for containment sprays are used for boric acid spray credit for the 0-2 hours dose at the site boundary and at the LPZ. The dose to the whole body from the passing cloud has been calculated using the same meteorological conditions used in determining the thyroid dose. The dose values shown in Table 15.4.6-2 are less than 10CFR100 limits.

A complete description of the Containment Vessel Spray System is given in Subsections 6.2.2.2.2 and 6.2.3.

Spray removal rates used in the analysis were evaluated using the model described by Parsly in ORNL-TM-2412 Part VII.

For elemental iodine, =0.24 hr-1 for injection of water from the Borated Water Storage Tank (BWST), and I=0.163 hr

-1 for water from the Containment Vessel Emergency Sump. Spray credit during the recirculation phase was conservatively taken from the time that recirculation starts until a DF of 100 was attained. The pH of the recirculation water is rising from the boric acid pH of 4-5 to the neutral pH of 7 prior to the time of recirculation (Ref 59).

For particulate iodine, =0.2 hr-1 for injection of water from the Borated Water Storage Tank, and =0.134 after the start of recirculation. No credit is taken for spray removal of organic iodine.

The injection phase is assumed to terminate at 40 minutes, the minimum time at which the water from the BWST could be exhausted. After this time reduced iodine removal credit is taken for the recirculation phase. Spray removal credit is terminated at a DF of 100 for both elemental and particulate iodine.

This analysis assumed the TID 14844 release of 25 percent of the core iodines available for release from the containment. Beginning with cycle 5, the fuel cycle has been extended. The total activity in the fuel and fuel rod gaps is given in Table 15A-6. The 24 month fuel cycle is used in the evaluation of MHA related doses. Of this, in accordance with Regulatory Guide 1.4, 4 percent was conservatively assumed to be organic. Iodine removal from the Containment air is accomplished by spray removal, decay, or leakage to the annulus. The iodines removed by the spray are deposited in the sump solution. Thus, when the sump is used as a source of spray water, the partition coefficient is smaller than for clean spray, resulting in the smaller spray removal constant. Partition coefficients can be found in ORNL-TM-2412 Part IV. Only one of the two containment spray pumps was assumed to be working at any one time, although this would not be the case unless there was an equipment failure.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-53 UFSAR Rev 30 10/2014 The assumptions used in this analysis were from L. F. Parsly's analysis in ORNL-TM-2412 Part Vll and are considered conservative. The r eport BNP-100 by Postma, Coleman, and Hillard reported experimental results much greater for boric acid sprays than those calculated here.

The Containment Spray System (Subsection 6.2.2.2.2) is completely independent of the Emergency Ventilation System (EVS) (Subsection 6.2.3.1). The spray system works inside the containment. The EVS collects leakage into the annular space between the Containment Vessel and the Shield Building.

The dose analysis incorporated the following additional assumptions:

a. The annular region between the Containment Vessel and the Shield Building is at atmospheric pressure upon accident initiation; 13 minutes are required to obtain a negative pressure in that region. It is assumed that all activity escaping the Containment Vessel during that time is released directly to the atmosphere without benefit of filtration or mixing. After the negative pressure has been obtained, 3 percent of the leakage is direct to the environment and the other 97 of the leakage is collected by the Emergency Ventilation System and exhausted through 95 percent efficient HEPA filters and charcoal adsorbers. Ref (35)

b. Control room personnel on duty at accident initiation remain on duty for 96 hours0.00111 days <br />0.0267 hours <br />1.587302e-4 weeks <br />3.6528e-5 months <br /> (4 days).

c. The operator absorbing the highest dose is assumed to be on duty for the initial 4-day period and 40 percent of the time for the period of 4 to 30 days.

d. The Control Room Ventilation System is isolated by closure of the normal ventilation system upon receipt of a high-pressure (4 psig) or high-radiation signal (per Amendment 221, SFAS no longer actuates on high containment radiation) from the Containment Vessel, a low reactor pressure, or a high-radiation signal in the station vent. This isolation requires a maximum of eleven seconds execution time (assuming 6 seconds maximum for instrument response). The radioactive concentration entering the control room will be nonexistent during this time since it takes in excess of 15 seconds for the radioactive cloud to travel from the containment to the control room intake.

e. Ten minutes after the start of the accident, the Emergency Control Room Ventilation System is started in the intake and recirculation mode.

f. Inleakage to the control room is 63 cfm for the first 10 minutes.

g. A fresh air (filtered) intake of 300 cfm is taken at 10 minutes and continues to 30 days. This intake slightly pressurizes the control room. In addition, it is assumed that 10 cfm unfiltered air reaches the control room due to door openings during the 10 minute to 30 day period.

h. A direct gamma dose is received through the walls from the Containment Vessel/Shield Building structure.

i. The dispersion factors were calculated by use of

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-54 UFSAR Rev 30 10/2014 aC 1 Q X for the control room dose analysis.

where C = shape factor = 0.5 A = vertical cross-sectional area of the containment

= average wind speed.

The meteorological conditions were chosen to correspond with those stipulated in Safety Guide 4, Section C, Paragraph 3, with the wind speed adjusted to correspond to actual site conditions, and the area calculated at 2597.4 m

2. The dispersion factors used in the analysis are:

0 to 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> 5.85 x 10

-4 sec/m 3 1 to 4 days 2.51 x 10

-4 sec/m 3 4 to 30 days 5.7 x to 10

-5 sec/m 3 The control room recirculation flow is 3300 cfm in the isolation mode and 3000 cfm in the emergency mode. In both cases, the air flows through charcoal filters that are 95 percent efficient for elemental, particulate, and organic materials.

j. Deleted

Table 15A-6 gives the isotopic inventory in curies at the containment at the beginning of the accident. These inventories are defined as A 0 in the equation that follows.

Mathematical Model:

A o = initial activity released to containment at time zero (Ci)

A 1(t) = initial activity with decay (Ci)

- r t = A o e .

A 2(t) = activity in primary containment at any time (Ci)

- 2 t = A 2 (t o)e A 4(t) = activity in Shield Building at any time (Ci)

.eteetA t 404 t 4 t 224021 Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-55 UFSAR Rev 30 10/2014 R(t) = release rate of activity to the atmosphere (Ci/sec)

= (F 1 3) (A 4 (t)), where F 1 is the filter nonremoval fraction.

IA R(t) = total activity released to atmosphere over time interval (Ci) 0 0 4 t 40431 4 t 4 2 t 22402231 t R Ae1tAFe1e1tAF R CR(t) = intake rate of activity for the control room (Ci/sec)

= F 2V X/Q R(t)

F 2 = filter nonremoval fraction; V = intake rate m 3/sec. A CR(t) = activity in the control room at any time 47 t 7 t 427 t 7 t 22402131 2eeeetAFQ/XVF t 70CR47 t 7 t 40431 2etAeetAFQ/XVF IA CR(t) = integrated control room activity (Ci-sec) 7 t 7 4 t 447 7 t 7 2 t 2272402131 2e1e11e1e11tAFQ/XVF 7 t 70CR 7 t 7 4 t 4470431 2e1tAe1e11tAFQ/XVF Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-56 UFSAR Rev 30 10/2014 Symbol Definition A 0 = initial activity released to the containment (Ci)

A 1(t) = initial activity with decay (Ci)

A 2(t) = activity in primary containment (Ci)

A 4(t) = activity in secondary containment (Ci) R (t) = release rate of activity from secondary (Ci/sec)

IA R(t) = total activity released over time (Ci)

R cr (t) = control room activity intake rate (Ci/sec)

A cr (t) = activity in the control room (Ci)

IA cr (t) = Integrated activity in the control room (Ci - sec) r = primary decay rate 1 = primary leak rate R 1 = primary cleanup rate - recirculation + spray 2 = total primary loss rate = r + 1 + R 1 R 2 = secondary cleanup rate 3 = secondary leak or purge rate 4 = total secondary loss rate =

R 2+ 3 + g 6 = control room cleanup rate CR = control room leak or purge rate = intake rate 7 = total control room loss rate = 6 + CR + r Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-57 UFSAR Rev 30 10/2014 Symbol Definition DS-TH = site boundary thyroid dose (rem)

D S- = site boundary beta skin dose (rem)

DCR-TH = control room thyroid dose (rem)

D CR- = control room beta skin dose (rem)

DCF TH = dose conversion factor thyroid Ci rem DCF WB = dose conversion factor - whole body total Ci rem . sec m 3 BR = breathing rate (m 3/sec) X/Q = meteorological dispersion factor (sec/m

3) D S- = site boundary whole body gamma dose (rem)

D CR- = control room whole body gamma dose (rem) = concentration time interval for the cloud 3 msecCi DCF-SKIN dose conversion factor-beta skinCi rem sec m 3 Site Boundary and LPZ Boundary

Thyroid Dose DS-TH = IA R . DCFTH .B.R. .

Q X rem = Ci Ci rem . sec m 3 . 3 m sec Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-58 UFSAR Rev 30 10/2014 Whole Body Dose R S IA Q X DCFE25.0D rem Ci m sec sec m Ci rem msecCi sec m Ci rem 3 3 3 3 R skin S IA Q X DCFE23.0D Control Room Doses

Thyroid Dose vol 1.R.BDCF IA D TH CRTHCR rem 3 3 m 1 sec m Ci remsecCi Whole Body Dose vol 1IADCF vol 1IAE25.0D CR CR CR rem 3 3 3 3 m 1secCi sec m Ci rem m 1secCi sec m Ci rem vol = control room volume (m

3) vol 1 IA DCF vol 1IAE23.0D CR skin CR CR Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-59 UFSAR Rev 30 10/2014 To calculate the activity released to the atmosphere over any time period, IA R (t) or the integrated control room activity, IA CR(t) over any time period, equations for IA R (t) and IA CR(t) may be evaluated over incremental time periods within which all parameters are constant, and the activities from these increments summed. Time = t o refers to the start of each time increment, and time = t is the length of the time increment.

For example, for a 10-second release where the filter nonremoval fraction, F 1 = 1 for the first three seconds and F 1 = 0.5 for the next seven seconds.

IA R (t) = IA R (0 to 3 sec) + IA R (3 sec to 10 sec) 4 3 4 2 3 2240213e1e10tA1 4 7 4 2 7 2240213e1313tA5.0 15.4.6.5 Effects of Engineered Safety Features Leakage During the Maximum Hypothetical Accident An additional source of fission product leakage during the maximum hypothetical accident can occur from leakage of the engineered safety features external to the containment vessel during the recirculation phase for long-term core cooling. A single failure analysis on engineered safety features is given in Tables 6.2-21 and 6.3-6.

It is assumed that the water being recirculated from the Containment Vessel Emergency Sump through the external system piping contains 50 percent of the core saturation inventory. This is the entire amount of iodine released from the Reactor Coolant System. The assumption that all the iodine escaping from the Reactor Coolant System is absorbed by the water in the Containment Vessel is conservative since much of the iodine released from the fuel will be plated out on the vessel's walls. It is assumed that all the iodine contained in water that flashes is released to the Auxiliary Building atmosphere. The peak sump temperature during sump recirculation is 248°F. As the cooldown continues, the sump temperature decreases with time. This results in a corresponding reduction in flashing fraction. Although the temperature drops below 212°F within 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />, it is conservatively assumed that the flashing fraction is 0.041 (using the peak sump temperature of 251°F) for the first 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />. After 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />, the sump temperature is less than 212°F, and the partition factor for Iodine is 0.01. The activity is assumed to be released through the 95 percent efficient HEPA filters and charcoal adsorbers of the Emergency Ventilation System to the station vent. Atmospheric dilution is calculated using the dispersion factors developed in Section 2.3.

The peak sump temperature determined by the Containment Vessel analysis described by Section 6.2.1.3 is slightly different than the evaluated peak temperature of 251°F. This Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-60 UFSAR Rev 30 10/2014 difference was evaluated to have a negligible impact on the dose analysis (see Reference 6.5-42).

The total leakage outside containment was determined. Plant procedures limit total combined measured leakage during normal plant operation from both trains of ECCS to 40 gph. 40 gph was used in the calculation of doses from ECCS leakage (Reference 60). In Table 15.4.6-2, the "Resultant Doses from ESF Leakage" are included in the "Resultant Doses from MHA."

The leakage and the resultant thyroid dose at the exclusion distance and Low Population Zone (LPZ) are shown in Table 15.4.6-2. Based on the assumptions above, the leakage outside the containment gives a negligible contribution to the exclusion area boundary and LPZ boundary

doses. Technical Specifications requires that the low pressure injection system and containment spray system be included in a program to reduce leakage as low as practical. The program includes periodic visual inspection requirements and integrated leak test requirements. This ensures that the exclusion area boundary and LPZ doses are below the 10CFR100 guideline values.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-61 UFSAR Rev 30 10/2014 TABLE 15.4.6-1 (1) Resultant Doses From Maximum Break Size LOCA Exclusion Area Boundary 0 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> LPZ Boundary 0 to 30 days Thyroid dose (Rem) 41.0 3.25 Whole-body dose (Rem) 1.03 0.128 (1)The environmental consequences for a LOCA, discussed in USAR Section 15.4.6.3, are bounded by the analyses provided in USAR Section 15.4.6.4 for a MHA.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-62 UFSAR Rev 30 10/2014 TABLE 15.4.6-2 Resultant Doses From MHA

Exclusion

Area Boundary 0 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> LPZ Boundary 0 to 30 days Thyroid dose (rem) 234 25.2 Whole-body dose (rem) 3.5 0.4

Control Room Operator Doses (rem)

Thyroid - Skin -Whole Body Inside the control room (1) 18.9 11.9 0.55 Direct gamma dose (Ref 61) 0 0 0.57 TOTALS 18.9 11.9 1.12 Resultant Doses from ESF Leakage Assumptions: (Reference 60)

Liquid leakage, gph 40 Leakage which flashes, gph 1.6 Two-hour thyroid dose at exclusion area boundary, rem 2 Thirty-day thyroid dose at LPZ, rem 0.8 (1) The operator dose "inside the control room" includes an ECCS leakage term (References 58 and 60).

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-63 UFSAR Rev 30 10/2014 15.4.6.6 Control Room Habitability An analysis of release of noxious gas (chlorine) within the station and its effects on the control room is presented in Subsection 15.4.8. (Note: Chlorine Tank cars are no longer on site.)

Control Room habitability following a main steam line break accident is discussed in Section

15.4.4.2.5.

Area radiation monitors are provided for the control room which continuously give the background radiation level. In case of any abnormal increase in the background level, the operator can manually isolate the normal ventilation system and start the Control Room Emergency Ventilation System if needed. The control room doses following a loss-of-coolant accident are presented in Table 15.4.6-2.

15.4.6.7 Partial Loop Flow LOCA

Power operation with less than four reactor coolant pumps in operation is allowed by plant technical specifications. However, with less than four pumps in operation, the core power level must be reduced. Large break LOCA analyses are performed to determine allowable rod position limits in this configuration. The full-power LOCA linear heat rate limit is typically preserved during partial pump operation. Although the core power level is reduced, the initial stored energy in the fuel pellet (fuel average temperature) remains the same as with the full power case. At reduced power, a positive moderator temperature coefficient is also allowed.

The purpose of the analysis is to demonstrate that the full power and full flow case bounds partial pump operation with reduced core power but with the same stored energy in the fuel and a positive MTC.

There are 5 possible break configurations at the pump discharge for partial loop operation:

1. 3-pump operation

a. break in idle pump discharge

b. break in active pump discharge of loop with the idle pump

c. break in pump discharge of loop with two active pumps

2. 2-pump operation, one idle pump in each loop

a. break in active pump discharge

b. break in idle pump discharge

Power operation with only two reactor coolant pumps running is not allowed by Davis-Besse License Condition C.3.a.

Partial pump (3-pump) LOCA analyses were performed with the RELAP5/MOD2-B&W-based LOCA evaluation model that is described in BAW-10192PA. Various sensitivity studies are performed in the evaluation model that generically demonstrates that the model provides an appropriate and conservative converged solution. In accordance with the SER on BAW-10192PA, plant-specific analyses for partial pump operation are required to be performed to demonstrate that partial pump operation is bounded. The partial pump analyses are performed at the maximum allowed linear heat rate limit from Reference 51 and a conservative Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-64 UFSAR Rev 30 10/2014 (positive) MTC of +1.0 pcm/°F. The LOCA analyses are based on a nominal core power level of 2966 MWt and includes a 2% heat balance error, e.g., the full power case was run at 102% and the partial pump case at 77% power. The results of the partial pump cases are summarized in Reference 51. The partial pump cases show that the peak clad temperature for the full power full flow case is higher and thus bounds the results for partial pump operation provided a linear heat rate penalty is applied to the full-power limit. The penalty is a function of power level and a maximum of 0.2 kW/ft.

15.4.6.8 Additional Analyses and Related Plant Modifications

15.4.6.8.1 DELETED

15.4.6.8.2 Positive Moderator Temperature Coefficient (MTC) Analysis

The large break LOCA analyses documented in Reference 51 contain sensitivity studies to determine the allowable moderator temperature coefficient (MTC) as a function of the core power level with all of the reactor coolant pumps operating. The analysis was based on a nominal core power level of 2966 MWt. The results concluded that the MTC must be negative for core power levels above 80%,(2372.8 MWt) i.e., the allowable MTC is +9.0 pcm/°F at hot zero power and 0 pcm/°F at 80% power and above. The results are reported based on nominal core power, but the analysis included a heat balance error of 2% of full power.

15.4.6.8.3 DELETED 15.4.6.8.4 DELETED

15.4.6.8.5 Impact of Replacement Steam Generators

As part of the Steam Generator Replacement Project, AREVA performed an evaluation (reference 64) to determine the impact of the replacement Steam Generators on the analyses of record. That evaluation documented that this event addresses only the doses resulting from a maximum hypothetical event. It also documented that the dose calculations do not directly depend on transient nuclear steam supply conditions. This hypothetical accident involves a gross release of fission products. No mechanism whereby such a release occurs is postulated. The design and performance differences of the replacement steam generators have no impact on this postulated gross release of fission products. Therefore, the replacement steam generators do not affect the analysis of record for the maximum hypothetical event.

15.4.7 Fuel-Handling Accident

15.4.7.1 Identification of Causes

Spent fuel assemblies are handled entirely under water. Before refueling, the reactor coolant and the refueling canal water above the reactor are increased in boron concentration so that, with all control rods removed, the keff of a core is no greater than 0.99. In the spent fuel storage pool, the fuel assemblies are stored under water in storage racks that have Boral neutron absorbing material to prevent criticality. Under these conditions, a criticality accident during refueling is not considered credible. Mechanical damage to the fuel assemblies during transfer operations is possible but improbable. The mechanical damage type of accident is considered the maximum potential source of activity release during refueling operations.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-65 UFSAR Rev 30 10/2014 Radiological consequences due to a postulated dry fuel storage cask drop accident have been evaluated in Section 15.4.7.2.5.3. Additional evaluations of potential accidents related to onsite dry fuel storage facility operations are discussed in the Davis-Besse Site Certified Safety Analysis Report (CSAR).

15.4.7.2 Accident Analysis - Accident Outside Containment

15.4.7.2.1 Safety Evaluation Criterion

The safety evaluation criterion for this accident is that resultant doses shall not exceed

10CFR100 guideline values.

In support of 18 month cycle operation, re-analyses of the Fuel Handling Accident (FHA) outside containment were performed (reference 30) based on three 450 EFPD cycles. Dose calculations were performed using more conservative source terms. Results of the evaluations showed that the offsite radiological dose for this accident was below the acceptance criterion value in the Standard Review Plan (NUREG 0800).

This evaluation crit erion was approved by the NRC via approval of the Cycle 6 Reload Report. The fuel handling accidents have been re-evaluated in USAR sections 15.4.7.2.5.1 and 15.4.7.3.1 for a maximum fuel assembly burnup of 60,000 MWD/MTU. That burnup limit is not changed for the extended fuel cycle.

15.4.7.2.2 Methods of Analysis

The assumptions and guidelines of Safety Guide 25 were used in this analysis. For convenience, the major assumptions made for this analysis are shown in Table 15.4.7-1. The reactor is assumed to have been shut down for 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br />, which is the minimum time for Reactor Coolant System cooldown, reactor closure head removal, and removal of the first fuel assembly.

It is further assumed that the entire outer row of fuel rods in the assembly, 56 of 208, suffers mechanical damage to the cladding. Since the fuel pellets are cold, only the gap activity is released. The fuel rod gap activity is calculated using the escape rate coefficients and calculational methods discussed in Section 11.1.

See 15.4.7.2.1 referenced re-analyses for a description of the methods of analysis performed to support extended fuel cycles.

15.4.7.2.3 Results of Analysis

The gases released from the fuel assembly pass upward through the spent fuel storage pool water prior to reaching the fuel-handling-area atmosphere. Although there is experimental evidence that a portion of the noble gases will remain in the water, no retention of noble gases is assumed. In experiments whereby air/steam mixtures were bubbled through a water pond, Diffey, et al. (reference 3) demonstrated decontamination factors of about 1000 for Iodine.

Similar results for iodine were demonstrated by Barthoux, et al. (reference 4) and predicted by Eggleton (reference 5). Based conservatively on these references, 99 percent of the iodine released from the fuel assembly is assumed to remain in the water. The fuel-handling area ventilated and is discharged through Emergency Ventilation Systems 95 percent efficient HEPA filters and charcoal adsorbers to the station vent.

See 15.4.7.2.1 referenced re-analyses for results of calculations performed to support 18 month

operation.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-66 UFSAR Rev 30 10/2014 See 15.4.7.2.5 for analyses to support fuel enrichments to 5.0 wt% U-235 and fuel burnups to 60,000 megawatt days per metric ton (MWD/MTU).

The Spent Fuel Pool has been re-racked to provide 1624 storage locations. Based on the physical and analytical design of the racks, it was determined no changes to the Fuel Handling Accident were required.

15.4.7.2.4 Environmental Consequences

The activity is assumed to be released over two hours from the station vent. Atmospheric dilution (for site and LPZ boundary) is calculated using the two-hour atmospheric dispersion coefficient developed in Section 2.3. Table 15.4.7-2 gives the total integrated dose at the exclusion distance for the whole body and the thyroid gland. These results are less than 10CFR100 guideline values.

See 15.4.7.2.5.1 for analyses for environmental consequences involving fuel enrichments up to 5.0 wt% U-235 and fuel burnups up to 60,000 megawatt days per metric ton (MWD/MTU) and applicability to extended fuel cycles.

15.4.7.2.5 Additional Analyses

The following analyses supersede all previously documented USAR fuel handling accident analyses for an outside containment fuel handling accident.

15.4.7.2.5.1 Effects of Extended Fuel Cycles, Fuel Burnup and Increased Fuel Enrichments

ln order to evaluate the effects of extended fuel burnup and increased fuel enrichment on the consequences of the fuel handling accident outside containment, the source term and offsite dose calculations were reperformed assuming that fuel assembly-average burnups could be as great as 60,000 MWD/MTU and that initial fuel enrichments could be as great as 5.0 weight percent uranium-235 (wt% U-235). Analysis assumptions are summarized in Table 15.4.7-1a.

All assumptions were made in accordance with Nuclear Regulatory Commission Regulatory Guide 1.25 ("Assumptions Used for Evaluating the Potential Radiological Consequences of a Fuel Handling Accident in the Fuel Handling and Storage Facility for Boiling and Pressurized Water Reactors (Safety Guide 25)," dated March 25, 1972).

For the extended fuel cycles, additional source term analyses were performed as described in Section 15.A.7.0 of Appendix 15A. Those analyses maintained the maximum fuel assembly burnup limit of 60,000 MWD/MTU and did not affect previously analyzed environmental consequences of the fuel handling accident.

15.4.7.2.5.2 Results

References 38 and 48 contain a detailed description of these fuel handling accident analyses. Source terms were determined to be extremely weak functions of fuel enrichment, with lower enrichments in general producing larger offsite doses than higher enrichments. Table 15.4.7-6 contains the source term results for fuel assemblies having initial enrichments of 3.0 and 5.0 wt% U-235. The values in Table 15.4.7-6 are fuel assembly-average fission product activities, and do not represent atmospheric release activities, as is the case in Table 15.4.7-3.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-67 UFSAR Rev 30 10/2014 The offsite doses for these fuel handling accident outside containment analyses are shown in Table 15.4.7-2a. In summary, these results are applicable to the extended fuel cycle as well, since the pertinent maximum fuel assembly burnup limit is maintained. They meet the acceptance criteria provided in NUREG-0800 and, therefore, are well within the dose guidelines set forth in 10CFR100.

15.4.7.2.5.3 Dry Fuel Storage Cask Drop As a part of the evaluation of the onsite dry fuel storage facility, additional evaluations related to a postulated cask drop accident have been performed.

USAR Section 9.1.4.3 implies that the maximum number of fuel assemblies in a shipping cask is ten. Since the dry fuel storage canister transfer cask contain 24 assemblies, the offsite radiological evaluations were performed as a part of the onsite dry fuel storage facility evaluation using the guidance contained in Standard Review Plan (SRP) 15.7.5; Spent Fuel Cask Drop Accidents. SRP 15.7.5 states that evaluation of radiological consequences need not be performed if the cask drop distances are less than 30 feet and appropriate impact limiting devices are employed during cask movement, radiological consequences are evaluated below.

The following assumptions were considered in performing the analysis:

a. The top shield plug will be secured to prevent displacement of assemblies out of the cask in the postulated cask drop accident.

b. The gap activity time for Spent Fuel Assemblies that will be stored in the dry fuel storage canister is five years. During this decay period all the gaseous activity except for Kr-85 and I-129 will decay to negligible levels.

c. The gap activity for these isotopes is assumed to be 30 percent of the total activity of these isotopes in the assembly.

d. All the gap activity is released. No credit is considered for plateaus or any other passive removal mechanisms.

e. Although the station emergency ventilation system will be activated if the radiation levels exceed the radiation monitor setpoints, no credit is taken for this system for removal of iodine.

f. The activity is released over a two hour period. The X/Q value applicable for this time period is 1.9x10

-4 sec/m 3 The Resultant Site Boundary Dose are:

WHOLE BODY: 0.008 REM

THYROID: 0.07 REM

These dose are less than the offsite consequences given for a Fuel Handling Accident Outside of Containment, and they are significantly less than the SRP 15.7.5 acceptance criteria values (i.e., 75 REM to the Thyroid and 6 REM to the Whole Body).

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-68 UFSAR Rev 30 10/2014 15.4.7.2.6 Impact of Replacement Steam Generators As part of the Steam Generator Replacement Project, AREVA performed an evaluation (reference 64) to determine the impact of the replacement Steam Generators on the analyses of record. That evaluation concluded that steam generator performance does not impact this accident in any way. Therefore, the existing analyses remain applicable with the replacement Steam Generators installed.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-69 UFSAR Rev 30 10/2014 TABLE 15.4.7-1 (1) Fuel-Handling Accident Parameters Outside Containment Fuel burnup, full power days 1017 Power level for the assembly during operation, MWt 27.9 Filter efficiency for iodine removal, %

95 Atmospheric dispersion at exclusion distance, s/m 3 , assuming ground release 1.9 x 10-4 The gap activity of the highest power fuel assembly is given in Table 15A-3, and the release by isotope is given in Table 15.4.7-3.

Accident duration, hr 2 Bases:

a. 56 fuel pins fail, releasing gap activity.

b. 99 percent iodine remains in water.

c. 72-hour decay

d. Hottest assembly

e. The control room ventilation system is isolated upon receipt of a high-radiation signal from the station vent. The isolation requires a maximum of ten seconds from attainment of the setpoint to closure of damper, during which time the intake rate of outside air is 10,960 cfm.

f. Control room inleakage of outside air is 1 cfm.

g. The release point (station vent) is 160 feet horizontal distance and 180 feet vertical distance.

(1)See 15.4.7.2.5 referenced re-analyses for the analysis assumptions.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-70 UFSAR Rev 30 10/2014 Table 15.4.7-1a Fuel Handling Accident Assumptions - Outside Containment

The following assumptions were used in the analysis that determined the doses listed in Table 15.4.7-2a:

1. The accident occurs at 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br /> following reactor shutdown.

2. Noble gas and iodine gap activities are based on Regulatory Guide 1.25.

3. The failure of 56 fuel pins.

4. A time-averaged radial peaking factor on an assembly basis of 1.4 is utilized.

5. The gap activity in the damaged fuel assembly is assumed to be released to the pool. All the noble gas activities that are released to the pool are assumed to escape from the pool; one percent of the iodine activities that are released to the pool are assumed to escape

from the pool.

6. An instantaneous release (very high escape rate) from the Spent Fuel Pool is assumed to ensure that all the activity coming out of the pool is released to the environment in a short time. 7. Atmospheric dispersion factor (X/Q) at site boundary is 1.9x10-4 sec/m 3 and at LPZ it is 9.9x10-6 sec/m

3. 8 Emergency Ventilation System charcoal filter efficiency for iodine removal is 95 percent.

9. The air intake of the Control Room's normal HVAC is automatically isolated upon receipt of a high radiation signal from the Station Vent radiation monitors.

Note: The assumptions utilized for the FHA outside the containment are identical to those for the FHA inside the containment with the exception of Items (3) and (8).

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-71 UFSAR Rev 30 10/2014 TABLE 15.4.7-2 (1) Resultant Doses From Fuel-Handling Accident Outside Containment

Exclusion Area Boundary 0 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> LPZ Boundary 0 to 30 daysThyroid dose (rem) 0.106 5.58 x 10-3 Whole body dose (rem) 0.106 5.59 x 10-3 Control Room 0 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> Thyroid dose (rem) 0.116 -skin dose (rem) 0.024 Total body gamma dose (rem) 5.53 x 10-3 (1)See 15.4.7.2.5 referenced re-analyses for resultant doses.

TABLE 15.4.7-2a Resultant Doses From Fuel Handling Accident Outside Containment - Extended Fuel Burnup (60,000 MWD/MTU)

Exclusion Area Boundary 0 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> LPZ Boundary 0 to 30 days Thyroid Dose (rem) 0.85 4.4 x 10-2 Whole Body Dose (rem) 0.15 8.0 x 10-3 Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-72 UFSAR Rev 30 10/2014 TABLE 15.4.7-3 (1) Activity Released to the Atmosphere Due to the Postulated Fuel-Handling Accident Outside Containment (Ci)

I-131 2.18 x 10 1 I-132 1.47 x 10-9 I-133 5.52 x 10-1 I-134 3.66 x 10-26 I-135 1.09 x 10-3 Xe-131m 1.40 x 10 2 Xe-133m 7.91 x 10 1 Xe-133 1.19 x 10 4 Xe-135m 0

Xe-135 1.52 x 10 2 Xe-137 0 Xe-138 1.82 x 10-73 Kr-83m 5.00 x 10-11 Kr-85m 1.27 x 10-3 Kr-85 1.23 x 10 3 Kr-87 4.60 x 10-16 Kr-88 3.45 x 10-6 Kr-89 0

(1)See 15.4.7.2.5 referenced re-analyses for activity released to atmosphere.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-73 UFSAR Rev 30 10/2014 15.4.7.3 Accident Analysis - Accident Inside Containment 15.4.7.3.1 Safety Evaluation Criterion

The safety evaluation criterion for this accident is that resultant doses shall not exceed

10CFR100 guideline values.

In support of 18 month cycle operation, re-analyses of the Fuel Handling Accident (FHA) inside containment were performed (reference 30) based on three 450 EFPD cycles. Dose calculations were performed using more conservative source terms. Results of the evaluations showed that offsite radiological dose for this accident was below the acceptance criterion value in the current NRC Standard Review Plan (NUREG 0800). This evaluation criterion was approved by the NRC via approval of the Cycle 6 Reload Report.

The fuel handling accidents have been evaluated in USAR sections 15.4.7.2.5.1 and 15.4.7.3.4.1 for a maximum burnup of 60,000 MWD/MTU. That burnup limit is not changed for the extended fuel cycle.

15.4.7.3.2 Analysis

The following assumptions were used in the analysis:

a. The accident occurs at 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br /> following reactor shutdown.

b. Noble gas and iodine gap activities are based on Regulatory Guide 1.25.

c. One entire assembly is considered damaged.

d . A time-averaged radial peaking factor on an assembly basis of 1.4 is utilized.

e. The gap activity in the damaged fuel assembly is assumed to be released to the pool. All the noble gas activities that are released to the pool are assumed to escape from the pool; one percent of the iodine activities that are released to the pool are assumed to escape from the pool.

f. Containment isolation is not assumed.

g. An instantaneous release (very high escape rate) from the containment is assumed to ensure that all the activity coming out of the pool is released to the environment in a short time (see Table 15.4.7-5).

h. Atmospheric dispersion factor (X/Q) at site boundary is 1.9x10

-4 sec/m 3 and at LPZ it is 9.9x10

-6 sec/m 3.

i. No filtration is assumed.

See 15.4.7.3.4.1 referenced re-analyses for a description of extended fuel cycle re-analysis and assumptions used.

See 15.4.7.3.4.1 for analyses supporting fuel enrichments up to 5.0 wt% U-235 and fuel burnups up to 60,000 megawatt days per metric ton (MWD/MTU).

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-74 UFSAR Rev 30 10/2014 15.4.7.3.3 Environmental Consequences The activity is assumed to be released over 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />. Atmospheric dilution for site boundary and LPZ is calculated using the 2-hour atmospheric dispersion coefficient developed in Section 2.3.

Table 15.4.7-4 gives the caiculated doses. These results are well within the 10CFR100

guideline values.

See 15.4.7.3.4.1 referenced re-analyses for the extended fuel cycle re-analysis of environmental consequences.

See 15.4.7.3.4.1 for analyses for environmental consequences involving fuel enrichments up to 5.0 wt% U-235 and fuel burnups up to 60,000 megawatt days per metric ton (MWD/MTU).

15.4.7.3.4 Additional Analyses

The following analyses supersede all previously documented USAR fuel handling accident analyses for an inside containment fuel handling accident.

15.4.7.3.4.1 Effects of Extended Fuel Cycle, Fuel Burnup and Increased Fuel Enrichments

In order to evaluate the effects of extended fuel burnup and increased fuel enrichment on the consequences of the fuel handling accident inside containment, the source term and offsite dose calculations were reperformed, assuming that fuel assembly-average burnups could be as great as 60,000 MWD/MTU and that initial fuel enrichments could be as great as 5.0 weight percent uranium-235 (wt% U-235).

All assumptions were made in accordance with Nuclear Regulatory Commission Regulatory Guide 1.25 ("Assumptions Used for Evaluating the Potential Radiological Consequences of a Fuel Handling Accident in the Fuel Handling and Storage Facility for Boiling and Pressurized Water Reactors (Safety Guide 25)," dated March 25, 1972).

For the extended fuel cycles, additional source term analyses were performed as described in Section 15.A.7.0 of Appendix 15A. Those analyses maintained the maximum fuel assembly burnup limit of 60,000 MWD/MTU and did not affect previously analyzed environmental consequences of the fuel handling accident.

15.4.7.3.4.2 Results

Reference 38 contains a detailed description of these fuel handling accident analyses. Source terms were determined to be extremely weak functions of fuel enrichment, with lower enrichments in general producing larger offsite doses than higher enrichments. Table 15.4.7-6 contains the source term results for fuel assemblies having initial enrichments of 3.0 and 5.0 wt% U-235. The values in Table 15.4.7-6 are fuel assembly-average fission product activities, and do not represent atmospheric release activities, as is the case in Table 15.4.7-5.

The offsite doses for these fuel handling accident inside containment analyses are shown in Table 15.4.7-4a. In summary, these results meet the acceptance criteria provided in NUREG-0800 and, therefore, are well within the dose guidelines set forth in 10CFR100.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-75 UFSAR Rev 30 10/2014 15.4.7.3.4.3 Control Room Dose Analysis

Additional analysis of the fuel handling accident inside containment was completed (Reference 44). This analysis calculated the control room dose for a fuel handling accident inside containment using the following assumptions:

1. The accident occurs 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br /> after reactor shutdown.

2. The noble gas and iodine fuel activities are based on the 3 wt% enrichment U-235 for an extended burnup of 60,000 MWD/MTU and are consistent with Table

15.4.7-6.

3. Noble gas and iodine release fractions are consistent with RG 1.25: 10% of fuel assembly Xe, Kr and I except for Kr-85 which is 30%.

4. All fuel pins (208) of one assembly are assumed to release their activity instantaneously to the pool.

5. All the noble gases and 1% of the iodine are released from the pool to the containment. The activity is assumed to be released from containment to the atmosphere over 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />. No credit is taken for containment isolation. The 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> atmospheric dispersion coefficient for the control room is 5.85x10

-4 sec/m 3. No filtration is assumed prior to atmospheric release.

6. The control room normal HVAC air intake, which is more than 160 feet from the release point, is isolated prior to the release from containment reaching it. The HVAC air intake is automatically isolated upon receipt of a high radiation signal from the station vent.

7. Conservative values of Control Room volume (50,000 ft

3) and Control Room unfiltered inleakage (59 cfm) were utilized.

8. No credit is taken for the control room emergency ventilation system to remove iodine activity leaked into the control room.

9. Normal control room HVAC is established 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> following initiation of the accident.

Table 15.4.7-4a gives the results of the control room dose calculation. The doses are well within the control room dose acceptance criteria as given in GDC 19 of 10CFR50 Appendix A.

15.4.7.3.5 Impact of Replacement Steam Generators

As part of the Steam Generator Replacement Project, AREVA performed an evaluation (reference 64) to determine the impact of the replacement Steam Generators on the analyses of record. That evaluation concluded that steam generator performance does not impact this accident in any way. Therefore, the existing analyses remain applicable with the replacement Steam Generators installed.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-76 UFSAR Rev 30 10/2014 TABLE 15.4.7-4 (1) Resultant Doses From Fuel-Handling Accident Inside Containment

Exclusion Area Boundary LPZ Boundary Thyroid dose (rem) 44.7 2.33 Whole body dose (rem) 0.17 8.86x10-3 (1)See 15.4.7.3.4 referenced re-analyses for resultant doses from the FHA inside containment.

TABLE 15.4.7-4a Resultant Doses From Fuel Handling Accident Inside Containment - Extended Fuel Burnup (60,000 MWD/MTU)

Exclusion Area Boundary 0 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> LPZ Boundary 0 to 30 days Thyroid Dose (rem) 62.6 3.26 Whole Body Dose (rem) 0.55 3.0x10-2 Control Room (1) 0 to 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> Thyroid Dose (rem) 17.6 Whole Body Dose (rem) 0.10 (1)See 15.4.7.3.4.3 referenced reanalyses from the FHA inside containment.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-77 UFSAR Rev 30 10/2014 TABLE 15.4.7-5 (1) Activity Released to the Atmosphere Due to the Postulated Fuel-Handling Accident Inside Containment (Ci)

I-131 4.26 x 10 2 I-132 3.11 x 10-7 I-133 1.18 x 10 2 I-134 1.47 x 10-22 I-135 6.86 x 10-1 Xe-131m 3.50 x 10 2 Xe-133m 1.28 x 10 3 Xe-133 8.55 x 10 4 Xe-135m 0

Xe-135 5.07 x 10 2 Xe-137 0 Xe-138 5.67 x 10-70 Kr-83m 2.43 x 10-8 Kr-85m 2.93 x 10-1 Kr-85 2.48 x 10 3 Kr-87 3.70 x 10-13 Kr-88 1.23 x 10-3 Kr-89 0 (1)See 15.4.7.3.4 referenced re-analyses for the activity released from the FHA inside containment.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-78 UFSAR Rev 30 10/2014 TABLE 15.4.7-6 Fuel Assembly Fission Product Activities (Curies) For Extended Fuel Burnup (60,000 MWD/HTU) (1)

Isotope Enrichment = 3 wt% U-235 Enrichment = 5 wt% U-235 I-131 5.86E+05 5.81E+05

I-132 5.58E+05 5.56E+05 I-133 1.29E+05 1.30E+05 I-134 1.05E-18 1.07E-18

I-135 7.02E+02 7.03E+02 Xe-131m 8.32E+03 8.18E+03

Xe-133m 2.56E+04 2.55E+04

Xe-133 1.12E+06 1.12E+06 Xe-135m 1.13E+02 1.13E+02

Xe-135 1.35E+04 1.35E+04 Xe-138 0 0 Kr-83m 2.49E-04 2.62E-04

Kr-85m 1.71E+00 1.85E+00 Kr-85 8.18E+03 9.53E+03 Kr-87 1.87E-12 2.06E-12

Kr-88 6.59E-03 7.29E-03

(1) Corrected to include Regulatory Guide 1.25 Radial Peaking Factor and Radioactive Decay for 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br />.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-79 UFSAR Rev 30 10/2014 15.4.8 Effects of Toxic Material Release on the Control Room

In accordance with Sections 2.2 and 6.4, toxic materials are not stored in volumes which would affect control room habitability.

15.4.8.1 Identification of Causes

(Note: Chlorine Tank Cars are no longer on site).

Chlorine is the only toxic material stored on site which could affect control room operation.

Chlorine from a 30-ton railroad car is used to supply the station chlorination system. The major chlorination equipment, which includes the evaporators and dispensers, is located in the Water Treatment Building. One-and-one-half-inch supply piping from the car to this building is in a covered trench, except for passage over the cooling tower channel, where the pipe is encapsulated in a larger pipe.

15.4.8.2 Accident Analysis

The supply car (top of rail El. 585 ft) is approximately 560 feet from the air intake to the control room (El. 654 ft). The closest distance to this intake from the supply pipe is about 380 feet.

Major assumptions relating to an accident are as follows:

a. Meteorology:

Pasquill F, wind one meter per second (same as for short-term accident diffusion estimates, Subsection 2.3.4).

b. Tank Car Accident:

A complete rupture is assumed in this case.

c. Supply Line Accident:

A maximum flow rate of 3.89 pounds per second is equivalent to the maximum flow permitted by the two excess-flow valves in the tank car. At this flow rate, the excess-flow valves are designed to close. (

Reference:

Chlorine Manual, The Chlorine Institute, Inc.)

d. All doors in the control room are closed. Free volume of control room is 7618 cubic feet. e. Two chlorine detectors with independent essential power supplies are located in the control room fresh air intake vent and two detectors with independent essential power supplies, located at the chlorine tank car, actuate the normal ventilation system intake and exhaust dampers to close so that supply air volumetric flow rate linearly decreases with time.

f. Initial control room chlorine concentration is zero when chlorine detectors are actuated.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-80 UFSAR Rev 30 10/2014 g. Effective height of intake above grade is one-half the actual intake height. This assumption was made because the size of the station structures may cause the wind to carry a ground-level chlorine release up and over the building, and this assumption attempts to include funneling effects.

h. Full-face, self-contained breathing apparatuses will be provided at readily accessible locations in the control room.

i. Flow into and out of the control room is that supplied by ventilation blowers.

j. System initial design flow for control and other rooms is 21,920 cfm. Initial air flow to and from the control room is 4,100 cfm.

k. Control room atmosphere is homogeneous.

l. The inleakage rate of contaminated air to the control room after isolation is 25 cfm.

m. The net free volume of the rooms surrounding the control room is 104,234 ft

3. n. All these rooms are isolated during the accident.

o. In the tank car rupture case, it was assumed that 25 percent of the total chlorine initially flashes to vapor.

p. In the pipeline break case, it was assumed that 25 percent of the flow flashes as it flows out.

q. A spill area of 1650 ft 2 was used (reference 16).

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-81 UFSAR Rev 30 10/2014 Control Room Concentration Model:

a. Atmospheric Diffusion:

1. The diffusion equation for a continuous ground-level release is: (reference 14) 2 z 2 yzy c z2/1exp y2/1exp u Q X (1) where X = the short-term concentration, g/m 3 Q c = amount of chlorine as continuous release, g/sec u = wind speed, m/sec

y = horizontal standard deviation of the plume, m Z = vertical standard deviation of the plume, m

2. The diffusion equation for an instantaneous (puff) ground-level release with a finite initial volume (reference 15) is:

2 I 2 z 2 2 I 2 y 2 2 I 2 x 22/12 I 2 z 2 I 2y,x 1 z y x2/1exp87.7 Q X (2) where X = concentration at coordinates x, y, z from the center of the puff, g/m 3 Q 1 = puff release quantity, g

, , , = standard deviations of the gas x y z concentrate in the horizontal

alongwind, horizontal crosswind, and vertical crosswind directions, respectively (assume x = y), m 7.87 = 2 1/2 3/2 Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-82 UFSAR Rev 30 10/2014 I = initial standard deviation of the puff, m

= o 1X87.7 Q where X o is the density of chlorine at standard conditions, g/m 3

The variation of unit concentration at a specific stationary receptor is determined by evaluating X in the exponential term in Equation (2) as follows:

X = D= ut

where D is the source - receptor distance, m

u is the wind speed, m/sec t is the time after release, sec

b. Dilution Inside Control Room:

The differential equation used to describe chlorine concentration, C, inside the control room is dt dC= Q (X - C) where Q = flow rate of air, cfm

i. before damper closed - normal intake rate

ii. while damper is closing - linearly decreasing

iii. after damper is closed - inleakage rate V = net free volume of control room, ft 3

X = outside concentration of both puff and continuous release, ppm

i. before damper is closed - concentration at damper

iii. after damper is closed - concentration at subject rooms

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.4-83 UFSAR Rev 30 10/2014 15.4.8.3 Results of the Analysis The following assumptions were made in the analysis of the chlorine-release accident. No credit was taken for the channeling of the cold and dense chlorine gas cloud around buildings and along ditches. The chlorine detectors were located at the chlorine tank car for the tank car accident and at the air intake for the pipe break and were assumed to have a response time of

five seconds at a set point of 5 ppm. The transport time from the air intake to the isolation damper was determined to be five seconds, and a five-second closing time of the isolation dampers was used.

A 25-cfm inleakage of contaminated air was considered after the isolation dampers were closed.

For the tank car accident, the chlorine concentration inside the control room when the isolation dampers closed was 0.0 ppm, and the maximum chlorine concentration reached was 16.0 ppm after approximately 14 hours1.62037e-4 days <br />0.00389 hours <br />2.314815e-5 weeks <br />5.327e-6 months <br />. This was assuming that the control room ventilation was not restarted after the chlorine cloud had passed. The operators in the control room have well over two minutes to put on self-contained breathing apparatus. Five ppm is not reached until approximately 1.9 hours1.041667e-4 days <br />0.0025 hours <br />1.488095e-5 weeks <br />3.4245e-6 months <br /> after the chlorine spill. This is in accordance with Regulatory Guide 1.95 (ref. 17.)

For the pipeline break case, the maximum chlorine concentration reached inside the control room is 4.02 ppm. Plots of the chlorine concentration in the control room versus time are presented in Figures 15.4.8-1 and 15.4.8-2.

15.4.8.3.1 Impact of Replacement Steam Generators

As part of the Steam Generator Replacement Project, AREVA performed an evaluation (reference 64) to determine the impact of the replacement Steam Generators on the analyses of record. That evaluation concluded that steam generator performance does not impact this accident in any way. Therefore, the existing analyses remain applicable with the replacement Steam Generators installed.

C loo u z 1: 0. 1 0.2 0.3 0. Y 0.5 0.6 0.7 0.8 Control Ron ASS~DIY Ltth, C,r~k/k OAVl S-BESSE NUCLEAR POWER STIT l ON PEAK NEUTRON POWER AS A FUNCTION OF EJECTED CRA WORTH FIGURE 15.4.3-1 REVISION 0 JULY 1982 I 0.2 0.3 0. Y 0.5 0,,6 0.7 0.8 Control Rod Assm~Iy Wortn, C ak/k b OAV l S-BESSE NUCt EAR POWER STlrT l ON PEAK THERMAL POWER AS A FUNCTION OF EJECTED CRA WORTH FIGURE 15.4.3-2 Rrtd Powor, 801 REVISION 0 JULY 1982 High ?reamre High Flux R8t.d Pow?, EOL loo3 Rated Pour, BOL 6 CRA Wo r tn, : a k/k DAVl S-BESSE NUCLEAR POWER STAlT l ON PEAK ENTHALPY OF HOTTEST FUEL ROO AS A FUHCTION OF EJECTED CRA WOElTH FIGURE 15.4.3-3 Rated Power, 8OL 7 Rated Power, EOL L High Pressure REVISION 0 JULY 1982 Rated Power, EOL Rated Power, BOL - Fl ux Trip i loo3 Rated Power,BOL 4 Doppler Coefficient.

(&/k)/F x lo5 DAVI S-BESSE NUCLEAR POWER STAT l ON PEAK NEUTRON POWER AS A FUNCTi ION OF DOPPLER COEFFlCl EN1 FOR AN EJECTED CRA WORTH OF 0.65% ~k/k AT BOTH loo3 RATED POWER AN0 RATED POWER FIGURE 15.4.3-4 REVISION 0 JULY 1982 b I I I > Rated Power, BOL H~gh tlux High Pressure -0.6 - 1.0 -I. Y - 1.8 -2. 2 Doppler Coefficeiant,(&/k)/~

x lo5 DAV I S-BESSE NUCLEAR POWER STAT I ON PEAK THERMAL POWER AS A FUNCTION OF DOPPLER COEFFICIENT FOR AN EJECTED CRA. WORTH OF 0.65% Ak/k AT BOTH 1 oo3 RATED POWER AND RAT ED POWER FIGURE 15.4.3-5 REVISION 0 JULY 1982 lo' -5.0 -3.0 - 1.0 +l.O 43.0 Moderator Coefficient. (h1k)I~ x 10' DAVI S-BESSE NUCLEAR POWER STAT l ON PEAK NEUTRON POWER AS A FUNCTION OF MODERATOR COEFFICIENT FOR AN EJECTED CRA WORTH OF 0.652 ~k/k AT BOTH 10.~' RATED POWER AND RATED POWE,R FIGURE 15.4.3-6 REVISION 0 JULY 1982 22!0 t I I 1 Rated Powcr. SOL 180 , Rated Power, EOL 160 --

• 4) s 0 e 120 , Nominal ,BOL L (0

• rO 4) Moderator coefficient. (dkIk)/F x 10' DAVI S-BESSE NUCLEAR POI* R STAT l ON PEAK THERMAL POWER AS A FUNCTION OF MODERATOR COEFFICIENT FOR AN EJECTED CRA WORTH OF 0.65% ~k/n AT BOTH 10-3 RATED PWER AND RATED POWER FIGURE 15.4.3-7 REVISION 0 JULY 1982 Trip Delay Time, s J L Rated Power, BOL Nom i n a1 - Rated Power, EOL 5 I Nomi ~ial 10-3 ~atcd Power. BOL (tli gh Pressure Tri p) C loo3 Rated Power, EOL DAY1 S-BESSE NUCLEAR POWER STAT l ON PEAK THERMAL POWER AS A FUNCTION OF TRIP DELAY TIME FOR AN EJECTED CRA WORTH OF 0.651 ak/k AT BOTH 1 Oo3 RATED POWER AND RATED F'OWER FIGURE 15.4.3-8 REVISION 0 JULY 1982 CRA Warth, %bk/k DAVI S-BESSE NUCLEAR POWER STAT l ON PERCENT PINS EXPERIENCING ONE AS A FUNCTI ON OF EJECTED CRA WORTH A'r RATED POWER, BOL FIGURE 15.4.3-9 REVISIOIY 0 JULY 1982 Time After Break, s DAVIS-BESSE NUCLEAR POWER STATION DOUBLE-ENDED RUPTURE OF 36- INCH STEAM L llSE BETWEEN STEAY GENERATOR AND MAIN STEM ISOLATION VALVE FIGURE 15.4.4-1 REVISION 0 JULY 1982 LO. 000 m 3 - 3 'S - rn 20,000 ' r Y y. r -8 10 2 0 30 4 0 50 6 0 Time After Break, s OAVI S-BESSE NUCLEAR POWER STAT1 ON OOUB1,E-ENOED RUPTURE OF 36- INCH STEW LINE BETWEEN STEAM GENERATOR AN0 MAIN STEM ISOLATION VALVE FIGURE 15.4.4-2 REVISION 0 JULY 1982 Timv After Break, s DAVI S-BESSE NUCLEAR POWER STATION DOUBLE-ENDED RUPTURE OF 36- INCH STEAM L INE BETWEEN STEM GENERATOR AND MAIN STEAM ISOLATION VALVE FIGURE 15.4.4-3 REVISION' 0 JULY 1982 Z o.v e =*b Li --_li 3-?n ;d CE X1; *lH= E= =O cEF9 ll- tE v, +Gur ul -EEF d;PH J lgITtEO H E HE H =F 2.o I U FII u5 li CG!lll3l-H ac t'O g 3 I I I l\\\\\\t t ll rl , I 5 St'l.U lio.j xa<Gl tt I t I I t I fr'.l i,, Ii3 !IF f ro I t;7\\'a E ll gs el.:;Itd.r!u F a Steam Line Break Size, ft2 DAVI S-BESSE NUlr EAR POWER STATION TIWE FROM flUPNRE TO TRIP ( INCLUDING DELAYS) VERSUS STEM LINE BREAK SIZE FIGURE 15.4.4-5 REVISION 0 JULY 1982 Rcaclor Power, Fraction of 2828 MWT RCS Pressure (Hot Leg), pe Total Reactivity, m Note:

• MSLB occurs at 2 seconds DAVIS-BESSE NUCLEAR POWER STATION Double-Ended Rupture of 36-inch Sbtam Line Between Steam Generator and Main Steam Isohion Valve with Sakry Valve Stuck Open at Rupm on Unaffected Steam Gcncmor REVISION :14 JULY 199:L RCS Cold Leg Ttmpe-, 'F Stram Generator Secondary Prwsurts, P* Note:

• MSLB occurs at 2 seconds DAVIS-BESSE NUCLEAR POWER STATION D~uble-Edd Rupame of 36inch Stam Line Between Stcam Gtnaatonr and Main Steam Won Valve with Safety Valve Stuck Opcn at Rqme on UnaBF& Socam Gaaamr Figure 15.4.4-7 REVISION 14 JULY 1991 Prtssuriztr Liquid Volumt, ff Note:

• MSLB occurs at 2 seconds 0 10 20 30 40 50 60 70 80 90 100 110 120 Transient Ti, S~C* Double-Endcd Rupture of 36inch Stcam LiDe Between Sttam Generator and Main Steam Isolation Valve with Safety Valve Stuck Opcn at Rupture on Unaffected Steam Generator REVISION Is JULY 1991 TIHE ISECONDS)

UNAFFECTED SG TIME OF BREAK DAVIS-BESSF NUCLE?? PONE2 STATION P8ESSLqE TRANSIENT AT SFRCS PRES52RE Tb? LOCATION TlME (SECOhTDS)

AFFECTED SC TIME (SECONDS)

UNAFFECTED SG TIME OF BREAK DAVIS-BESSE NUCLEAR POWER STATION PRESSURE TRANSIENT AT SFRCS PRESSURE TAP LOCATION 'FIME (SECONDS)

AFFECTED SG Revision 3 July, 1987 DAV I S-BESSE NUCLEAR POWER STAT l ON REACTOR COOLANT PRESSURE AS A f:UNCT ION OF TIME FOR THE COMPLETE SEVERANCE OF A LETDOWN L l NE FIGURE 15.4.5-1 REVISION 0 JULY 1 982 DAVIS-BESSE NUCLEAR POWER STATION CHLORINE CONCENTRATION AFTER TANK CAR RUPTURE FI WRE 15.4.8-1 REVISION 0 JULY 1982

-ta o z o g ul-tr F'ct lrlY o p orv I s- BEssE lfucl[rn P0IER sTrTl0ll$LoRI)l E CoxCEllIRrT l0ll rffEnI PIPE LI}IE BRETT tl EURE l!. l. N'2 REVISION O JULY 1982 Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.5-1 UFSAR Rev 30 10/2014

15.5 REFERENCES

1. L.C. Watson, A.R. Bancroft, and C.W. Hoelke, "Iodine Containment by Dousing in NPD-11," AEC-1130.

2. M.A. Styrikovich, et al., "Transfer of Iodine from Aqueous Solutions to Saturated Vapor," Soviet Journal of Atomic Energy 17, July 1964.

3. H.R. Diffey, et al., "Iodine Cleanup in a Steam Suppression System," International Symposium on Fission Product Release and Transport Under Accident Conditions, Oak Ridge, Tennessee, CONF-65047, Vol 2, pp 776-804 (1965).

4. A.J. Barthoux, et al., "Diffusion of Active Iodine Through Water, With the Iodine Being Liberated in CO 2 Bubbles at High Temperature," AEC-TR-6144, June 1962.

5. A.E.J. Eggleton, "A Theoretical Examination of Iodine-Water Partition Coefficients," AERE-R-4887, February 1967.

6. B. Weidenbaun and S. Naymark, "Potentialities of Molten UO 2 as a Reactor Fuel," Transactions of the ANS, 7, No. 2, pp 527-528 (1964).

7. R.C. Liimatainen and F.J. Testa, Studies in TREAT of Zircaloy-2 Clad, UO 2 -Core Simulated Fuel Elements," Argonne National Laboratory Chemical 2 Engineering Division Semi-Annual Report, ANL-7225, January-June 1966.

8. J.F. Whited, GE-NMPO, direct communication of experimental data to be published.

9. L.N. Grossman, "High-Temperature Themal Analysis of Ceramic Systems," Paper presented at the American Ceramic Society 68th Annual Meeting, May 1966.

10. J.A. Redfield, "CHIC-KIN - A Fortran Program for Intermediate and Fast Transients in a Water-Moderated Reactor, WAPD-TM-479, January 1965.

11. A.F. Hery and A.V. Vota, "WIGL2 - a Program for the Solution of the One-Dimensional, Two-Group, Space-Time Diffusion Equations Accounting for Temperature, Xenon, and Control Feedback," WAPD-TM-532, October 1965.

12. W.R. Wise, Jr., and J.F. Proctor, "Explosion Containment Laws for Nuclear Reactor Vessels," NOLTR-63-140, August 1965.

13. W.R. Wise, Jr., An Investigation of Strain Energy Absorption Potential as the Criterion for Determining Optimum Reactor Vessel Containment Design, NAVORD Report 5748, June 1958.

14. D.H. Slade, Ed., Meteorology and Atomic Energy, USAEC, 1968.

15. USNRC, Regulatory Guide 1.78, "Evaluating the Habitability of a Nuclear Power Plant Control Room During a Postulated Hazardous Chemical Release," Revision 1, December

2001.

16. A.E. Howerton, "Estimating Area Affected by a Chlorine Release," The Chlorine Institute, Inc., New York, 1967.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.5-2 UFSAR Rev 30 10/2014

17. USNRC, Regulatory Guide 1.95, "Protection of Nuclear Power Plant Control Room Operators Against an Accidental Chlorine Release," 1975.

18. J. W. Pegram, C. W. Tally, and J. A. Weimer, "Justification for Raising Setpoint for Reactor Trip on High Pressure," BAW-1890, Babcock & Wilcox, September 1985.

19. C. L. Ritchey, "DB-1 Startup Event Reanalysis," B&W Document 32-1167957-00, Babcock & Wilcox, July 17, 1987.

20. C. L. Ritchey, "DB-1 Startup Event Reanalysis Summary," B&W Document 86-1167960-02, Babcock & Wilcox, August 31, 1987.

21. D. J. Skulina, "MSLB Analysis with Failed Open MSSV" B&W Document 32-1178648-01, Babcock & Wilcox, August 10, 1990.

22. R. L. Harne and J. H. Jones, Thermal-Hydraulic Crossflow Applications, BAW-1829, Rev 0, Babcock & Wilcox, Lynchburg, VA, April, 1984.

23. BAW-10179P-A, Latest rev. per App.4-B "Safety Criteria and Methodology For Acceptable Cycle Reload Analysis," (See App.4-B).

24. DELETED

25. DELETED

26. DELETED

27. DELETED

28. DELETED

29. DELETED

30. S. A. Skidmore, "DB-1 FHA Thyroid Doses", B&W Document 51-1164134-00, Babcock & Wilcox, dated May 2, 1986.

31. J. R. Worsham III, "Safety Analysis to Change MTC" at Davis-Besse, B&W Document 51-1201479-00, Babcock & Wilcox, dated December 21, 1990.

32. B&W Calculation 32-1171668-00, "DB-1 Letdown Line Break", dated May, 1988.

33. TED Calculation C-NSA-065.01-010, Rev. 0.

34. Docket No. 50-346, Amendment No. 149 to Facility Operating License No. NPF-3 (TAC No. 69569), June 4, 1990.

35. Docket No. 50-346, Amendment No. 160 to Facility Operating License No. NPF-3 (TAC No. 80347), August 23, 1991.

36. TED Calculation C-NSA-60.00-001, Rev. 1.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.5-3 UFSAR Rev 30 10/2014 37. TED Calculation C-NSA-059.01-010, Impact of 1900 PSIG RPS Trip on Containment Response to MSLB.

38. M. A. Rutherford, "DB FHA Extd BUP Calculation," B&W Document 32-1224894-00, B&W Nuclear Technologies, dated September 24, 1993.

39. Log No. 3077 (EXT-89-07488), Evaluation of the Davis-Besse Nuclear Power Station Compliance with 10 CFR 50.62 Requirements for Reduction of Risk from Anticipated Transients without Scram (ATWS) (TAC 59086)," September 29, 1989.

40. B&W Calculation 32-1224901-01, "DB-1 Rod Ejection Analysis," dated August, 1994.

41. TED Calculation C-NSA-63.01-005, Line Breaks in Mode 3, Rev. 0.

42. TED Calculation C-NSA-63.01-007, MSLB In Mode 3 with the MFWP, Rev. 0.

43. Docket No. 50-346, Amendment 202 to Facility Operating License No. 20 NPF-3 (TAC No. M92533), November 17, 1995.

44. TED Calculation C-NSA-028.01-002, Control Room Radiation Dose due to Fuel Handling Accident Inside Containment, Rev. 2.

45. B&W Calculation 32-1159090-01, Davis-Besse Unit 1 SFRCS Accident Analysis with Licensing Assumptions, March 2, 1987.

46. B&W Calculation 32-1171148-00, Davis-Besse Loss of Feedwater 220 inch Pressurizer Setpoint, March 14, 1988.

47. BWNT Document 51-1245290-00 "D-B Cy10 EOC T ave Reduct Man," 2/15/96.

48. Docket No. 50-346, Serial Number 2180, "Supplemental Information for License Amendment 181," October 5, 1993.

49. TED calculation C-NRE-062.02-072, "Two Year Cycle Source Term Analysis, Rev. 1."

50. DELETED

51. TED Calculation, C-NSA-064.02-036, "DB-1 LOCA Summary Report."

52. DELETED

53. FTI Document 51-5006137-01, "DB Cy13 PSC 15-74 Statept.," 2/2/00.

54. Bechtel Calculation 61.01, FSAR Chapter 15 Accident Analysis.

55. Bechtel Calculation 60.16 Accident Analysis.

56. BAW 10193PA, RELAP5/MOD2-B&W for Safety Analysis of B&W Designed Pressurized Water Reactors, Rev 0, January 2000.

57. C-NSA-060.00-014 Revision 0, Loss of Feedwater (LOFW) Analysis (Framatome ANP document 86-5027672-00, "DB LOFW Analysis at 2827.44 MWT," September 26, 03).

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15.5-4 UFSAR Rev 30 10/2014

58. Calculation, C-NSA-028.01-005, Revision 3, "Control Room Radiation Doses Following a Maximum Hypothetical Accident."

59. Calculation, C-NSA-049.01-007, Revision 0, "DB1 Post-LOCA pH Analysis."

60. Calculation, C-NSA-028.01-007, Revision 0, "Control Room Doses due to ECCS Leakage to the BWST and Auxiliary Building."

61. Calculation, C-NSA-028.01-008, Revision 0, "Dose in the Control Room due to EVS and CREVS filter activity."

62. C-NSA-060.00-016, Revision 0, "Reanalysis of Startup Event" (Areva Document 86-5027875-00, "DB-1 Startup Event Analysis Results," June 24, 2003).

63. Areva NP Document 51-9004090-005, "Davis-Besse MUR Summary Report."

64. Areva NP Document, 51-9209222-001, "DB-2 ROTSG Disposition of Events."

65. Areva NP Document, 51-9211669-003, "Disposition of ROTSG for DB-1 LOCA Analysis."

66. Calculation, C-NSA-081.05-001, "DB Caldon Dropped Rod Analysis."

67. Areva NP Document 86-5007079-00, "Davis-Besse SG Overpressure Protection."

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15A-i UFSAR Rev 30 10/2014

APPENDIX 15A

RADIATION SOURCES

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15A-ii UFSAR Rev 30 10/2014 APPENDIX 15A RADIATION SOURCES TABLE OF CONTENTS

Section Title Page 15A.1.0 GENERAL 15A-1 15A.2.0 ACTIVITY IN CORE 15A-1 15A.3.0 ACTIVITY IN HIGHEST POWER FUEL ASSEMBLY 15A-1 15A.4.0 FISSION PRODUCT ACTIVITY IN REACTOR COOLANT 15A-2 15A.5.0 FISSION PRODUCT ACTIVITY IN SECONDARY COOLANT 15A-2 15A.6.0 DELETED

15A.7.0 SOURCE TERM FOR EXTENDED CYCLE EVALUATION 15A-7

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15A-iii UFSAR Rev 30 10/2014 LIST OF TABLES

Table Title Page 15A-1 Escape Rate Coefficients 15A-4 15A-2 Total Core Fission Product Inventory in Fuel and Fuel Rod Gaps 15A-4 15A-3 Fission Product Inventory in Fuel and Fuel Rod Gap of Highest Power Fuel Assembly 15A-5 15A-4 Maximum Fission Product Activity in Reactor Coolant 15A-6 15A-5 Activities Expected in Secondary Coolant 15A-6 15A-6 Comparison of Core Fission Product Inventory for a 24 Month Fuel Cycle with Source Terms from Table 15A-2 and TID 14844 15A-8 Davis-Besse Unit 1 Updated Final Safety Analysis Report 15A-1 UFSAR Rev 30 10/2014 APPENDIX 15A RADIATION SOURCES

15A.1.0 GENERAL

This appendix lists the quantities of fission products in the fuel, fuel rod gap, reactor coolant and secondary system that are used to evaluate the environmental consequences of the Chapter 15 accidents. The calculational methods discussed in Section 11.1 were used, but more conservative assumptions have been used to calculate these accident sources. This results in higher fission product activities and conservative dose values.

See Section 15A.7.0, Source Term for Extended Cycle Evaluation.

15A.2.0 ACTIVITY IN CORE

The fission product activity in the fuel and the fuel rod gap is calculated by the method discussed in Subsection 11.1.1, using the following assumptions:

a. Full power operation at 2772 MWt for a 433 day initial cycle and two 277 day equilibrium cycles.

b. Operation with no defective fuel rods so the maximum core fission product inventory is calculated. The calculation was performed with the escape rate coefficients shown in Table 15A-1.

The resulting fuel and fuel rod gap fission product inventories are given in Table 15A-2.

See 15A.1.0 referenced re-analyses for the assumptions used to determine the activity in core.

15A.3.0 ACTIVITY IN HIGHEST POWER FUEL ASSEMBLY

The fission product activity in the fuel and gap of the highest power fuel assembly is calculated by the same method as was used for the core activity (Section 15A.2), except that:

a. The activity inventory is based on the highest power fuel assembly removed from the core at the end of any core cycle.

b. The assembly contains no defective fuel rods so that the maximum fission product inventory is calculated.

The resulting fuel and gap fission product inventories in the highest power fuel assembly are

given in Table 15A-3.

See 15A.1.0 referenced re-analyses for the calculated activity in the highest power fuel assembly.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15A-2 UFSAR Rev 30 10/2014 15A.4.0 FISSION PRODUCT ACTIVITY IN REACTOR COOLANT The fission product activity in the Reactor Coolant System is calculated by the method discussed in Subsection 11.1.2.1, using the following assumptions:

a. Full power operation at 2772 MWt for a 433 day initial cycle and two 277 day equilibrium cycles.

b. One percent of the fuel rods in the core develop cladding defects at the beginning of the third operating cycle. All of the defects are assumed to occur in the high burnup fuel rods which have been irradiated for the 433 day initial cycle and one 277 day equilibrium cycle. The defective fuel rods are assumed to leak continuously during the entire 277 day irradiation cycle.

c. Base loaded operation.

d. Continuous reactor coolant purification through the purification demineralizer at an average flow rate of one Reactor Coolant System volume per day with a zero removal efficiency for krypton, xenon, molybdenum, cesium, yttrium, tellurium and tritium, and a 99% removal efficiency for all other nuclides.

e. All isotopes, except krypton, xenon and tritium are removed with a 99.9% efficiency during the processing of the reactor coolant bleed. Krypton and xenon are removed with a 99.9% efficiency during the first 267 days of an equilibrium cycle, but after that time, when the deborating demineralizers are used for bleed stream processing, the removal of krypton and xenon ceases.

The maximum reactor coolant activities are given in Table 15A-4.

See 15A.1.0 referenced re-analyses for the fission product activity in secondary coolant.

15A.5.0 FISSION PRODUCT ACTIVITY IN SECONDARY COOLANT

The equilibrium iodine activity in the secondary coolant is given by the ratio of production terms to loss terms.

Production is from a 1 gpm steam generator leak. Losses are:

a. 90% removal in the 67% of the flow which sees the demineralizers.

b. Secondary leakage of 10 gpm.

c. A DF of 10 4 in the steam jet air ejector.

d. Decay.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15A-3 UFSAR Rev 30 10/2014 The flow rate of the secondary system is 70 systems/day.

The noble gas activity will be negligible due to gas removal in the steam generator and the condenser. Table 15A-5 lists the activities expected in the secondary coolant.

See 15A.1.0 referenced re-analyses for the fission product activity in reactor coolant.

15A.6.0 DELETED

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15A-4 UFSAR Rev 30 10/2014 TABLE 15A-1 (1) Escape Rate Coefficients

Element Escape rate coefficient, sec-1 Krypton 1.0 x 10-7 Xenon 1.0 x 10-7 Iodine 2.0 x 10-8 (1)See 15A.1.0 referenced re-analyses for the escape rate coefficients used.

TABLE 15A-2 (1) Total Core Fission Product Inventory in Fuel and Fuel Rod Gaps

Activity at end of third cycle Isotope Fuel activity, Ci Fuel rod gap activity, Ci Kr-83m 8.26(+6) 1.06(+4) Kr-85m 2.46(+7) 5.58(+4) Kr-85 5.32(+5) 4.33(+5) Kr-87 4.54(+7) 3.06(+4) Kr-88 6.85(+7) 9.84(+4)

Xe-131m 5.9(+5) 8.72(+4) Xe-133m 3.45(+6) 1.03(+5) Xe-133 1.43(+8) 9.21(+6)

Xe-135m 3.53(+7) 3.40(+4) Xe-135 1.76(+7) 1.86(+5) Xe-138 1.2 7(+8) 1.87(+4)

I-131 7.46(+7) 1.46(+6)

I-132 1.09(+8) 2.01(+5)

I-133 1.44(+8) 3.09(+5)

I-134 1.83(+8) 1.92(+4)

I-135 1. 4 (+8) 9.73(+4) (1)See 15A.1.0 referenced re-analyses for the total fission product inventory in fuel and fuel rod ga Davis-Besse Unit 1 Updated Final Safety Analysis Report 15A-5 UFSAR Rev 30 10/2014 TABLE 15A-3 (1) Fission Product Inventory in Fuel and Fuel Rod Gap of Highest Power Fuel Assembly

Isotope Fuel activity, Ci Fuel rod gap activity, Ci Kr-83m 6.15(+4) 7.56(+1)

Kr-85m 1.76(+5) 3.99(+2) Kr-85 5.14(+3) 4.59(+3) Kr-87 3.23(+5) 2.19(+2) Kr-88 4.89(+5) 7.03(+2)

Xe-131m 4.22(+3) 6.23(+2)

Xe-133m 2.46(+4) 7.38(+2)

Xe-133 1.02(+6) 6.58(+4)

Xe-135m 2.52(+5) 2.42(+2) Xe-135 1.26(+5) 1.33(+3) Xe-138 9.07(+5) 1.34(+2)

I-131 5.33(+5) 1.05(+4)

I-132 7.79(+5) 1.44(+3)

I-133 1.03(+6) 2.21(+3)

I-134 1.31(+6) 1.37(+2)

I-135 1.0(+6) 6.95(+2)

(1)See 15A.1.0 referenced re-analyses for the fission product inventory of highest power fuel assembly.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15A-6 UFSAR Rev 30 10/2014 TABLE 15A-4 (1) Maximum Fission Product Activity in Reactor Coolant 1

Isotope Activity, Ci/cc Kr-83m 3.12(-1) Kr-85m 1.65 Kr-85 9.30 Kr-87 9.04(-1)

Kr-88 2.90 Xe-131m 2.32 Xe-133m 3.03 Xe-133 2.63(+2) Xe-135m 1.00 Xe-135 5.50 Xe-138 5.52(-1)

I-131 3.48 I-132 5.20 I-133 4.11 I-134 5.39(-1) I-135 2.05 1 Coolant density = 44.5 lbm/ft 3 (1)See 15A.1.0 referenced re-analysis for the maximum fission product activity in reactor coolant.

TABLE 15A-5 (1) Activities Expected in Secondary Coolant Isotope Activity, Ci/gm I-131 3.44(-4) I-132 4.39(-4) I-133 3.99(-4)

I-134 3.66(-5)

I-135 1.92(-4)

(1)See 15A.1.0 referenced re-analysis for the activities expected in secondary coolant.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15A-7 UFSAR Rev 30 10/2014 15.A.7.0 Source Term for Extended Cycle Evaluation The evaluation below discusses the impact, from a source term standpoint (i.e., for dose consequences), of extending the fuel cycle to 24 months.

Among the accidents analyzed in Chapter 15, only the control rod ejection accident (15.4.3), the Loss of Coolant Accident (15.4.6), and the fuel handling accident (15.4.7) utilize the activity in the fuel and/or in the fuel rod gap in the calculation of off site doses. For the other accidents, the dose consequences were evaluated based on the fission product activity in the reactor coolant. Since the coolant activity used in the accident analysis is significantly higher than the Technical Specifications limits, the USAR dose analysis is bounding for accidents that do not result in a release from the fuel rods. The fuel handling accident has been evaluated in USAR sections 15.4.7.2.5.1 and 15.4.7.3.4.1 for a maximum burnup of 60,000 megawatt days per metric ton. This burnup limit is not changed by the 24 month fuel cycle. The remaining two accidents are evaluated below.

The total fission product inventory in the core (i.e., in the fuel matrix as well as in the fuel rod gap) was analyzed in Reference 49. Also, implicit in the analysis is the core's operation for 690 Effective Full Power Days (EFPDs) at a power level of 2827 MWt (i.e., 102% of 2772 MWt).

Based on the expected fuel batch size and 60,000 megawatt days per metric ton (MWD/MTU) burnup limit, fuel enrichments of approximately 4.5 to 5 percent are needed to support full power operation for two years. However, in order to maximize the calculated iodine inventory, the 24 month fuel cycle source term was analyzed by assuming a fuel enrichment of 4% by weight.

The core inventory for a 277 day equilibrium cycle was given in USAR Table 15A-2. However, when Davis-Besse was licensed this was not used in the evaluation of dose consequences due to a postulated Maximum Hypothetical Accident (MHA) in USAR section 15.4.6. Instead, the core inventory used in the evaluation of the MHA was based on TID-14844. TlD-14844 is the historical document that describes a source term and methodology for calculation of offsite dose consequences. A comparison between the source terms from the original USAR Table 15A-2, those for the 24 month cycle and TID-14844 (used in the analysis of the MHA) is in Table 15A-6. This table shows that, although there are minor variations in the calculated core inventories, the source term for the 24 month cycle is not significantly different than the one

given in the original USAR Table 15A-2.

The original USAR source term was not used in the MHA analyses and assumptions used in the rod ejection accident for the original USAR analysis (with respect to availability of the condenser) were not as conservative as the NRC's independent analysis in the Safety Evaluation Report (SER) for Davis- Besse. Therefore, additional evaluations using the 24 month cycle source term for these two accidents were performed as described in sections 15.4.6.4 and 15.4.3.2.7.

Davis-Besse Unit 1 Updated Final Safety Analysis Report 15A-8 UFSAR Rev 30 10/2014 TABLE 15A-6 Comparison of Core Fission Product Inventory for a 24 Month Fuel Cycle with Source Terms from Table 15A-2 and TID 14844*

Total Activity in the Fuel and Fuel Rod Gaps, Ci Isotope From Table 15A-2 24 Month Fuel Cycle TID 14844 Kr-83m 8.27E6 9.24E6 1.17E7 Kr-85m 2.47E7 1.93E7 3.67E7 Kr-85 9.65E5 9.32E5 1.16E6 Kr-87 4.54E7 3.69E7 6.60E7 Kr-88 6.86E7 5.18E7 9.05E7 Xe-131m 6.77E5 8.59E5 7.34E5 Xe-133m 3.55E6 4.68E6 3.91E6 Xe-133 1.52E8 1.51E8 1.59E8 Xe-135m 3.53E7 3.08E7 4.40E7 Xe-135 1.78E7 3.31E7 1.52E8 Xe-138 1.27E8 1.28E8 1.35E8 I-131 7.61E7 7.71E7 7.08E7 I-132 1.09E8 1.12E8 1.08E8 I-133 1.44E8 1.56E8 1.59E8 I-134 1.83E8 1.71E8 1.86E8 I-135 1.40E8 1.46E8 1.44E8

• Note: These values are calculated values and do not reflect limits.