ML20090M505
| ML20090M505 | |
| Person / Time | |
|---|---|
| Site: | Monticello  |
| Issue date: | 11/30/1973 |
| From: | NORTHERN STATES POWER CO. |
| To: | |
| Shared Package | |
| ML20090M503 | List: |
| References | |
| NUDOCS 9105030420 | |
| Download: ML20090M505 (128) | |
Text
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NORTilERN STATES POWER CCEPANY MONTICELLO NUCLEAR GENERATUiG PIM;T a
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r SEC0t:D RiiLG\D SUBMITTAL
.f N0VDiBER 1973
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Prepared By Infomation Supplied By Northern States General Electric
! Power Company Company i
9105030420 731119 i PLR ADDCK 05000263 P PDR
NORTilERN STATES POWER CQiPANY MONTICELLO NUCLEAR GENERATLNG PLANT 1
i SECOND RELMD SUBMINAL NOVEMBER 1973 I
Prepared By Infomation Supplied By Northern States General Electric Power Company Company
LIST OF ILLUSTRATIONS Titic Page Figure 2- 2 2-1 Reference Core Loading, Monticello Cycle 3 8x8 Reload Puel Assembly 3-2 3-1 3-3 3-2 Monticello R2 Reload Fuel Lattice Hot Average Void Infinite Lattice K, versus Exposure 5-3 -
5-1 5-2 2.63 7x7, 6.ft 8x8 Infinite Lattice K ,versus In-Channel Void Fraction Zero Exposure 5-6 AK Void Comparison 7x7 versus 8x8 from 0.40 Void to other Voids 5-7 5-3 200 mwd /t Doppler coefficients Uncontrolled 5-9 5-4 10,000 mwd /t Doppler coefficients Uncontrolled 5-10 5-5 5-6 6 versus Exposure Comparison 2.62 vt% 8x8, 2.63 vt% 7x7 Average Voids, Uncon'. rolled 5-12 Maximum Local Peaking versus Exposure Comparison 5-13 5-7 6-1 Doppler Reactivity Coefficient versus Average FL 1 6-13 Temperature as a function of Exposure and Moderator condition 6-2 Accident Reactivity Shape Functions for Cold Startup 6-14 8 = 0.0054 Scram Reactivity Function for Cold Startup 6-15 6-3 6-4 Cladding Temperature versus Time for the Recirculation Line $reak s'sh Failure of the LPCI Injection Valva (HPCI + 2CS + / 3S) AEC Assumptions 6-20 2
6-5 Perf ormance of ECCS with Failure of HPCI for Small (0.02 f t ) 6-22 Liquid Break (2CS + 4LPLI + ADS) AEC Assumptions 6-6 Cladding Temperature verlus Time f r a Small Break with Tsilure of HPCI (0.02 ft Break) (4LPCI + 2CS + ADS) 6-23 AEC Assumption 6-7 Peak Cladding Temperature Spectrum for Single Failure Conditions AEC Assumptions 6-24 Emergency Core Cooling System-Performance Capability 6-25 6-8 6-9 Perf ormance of ECCS for Main Steam Line Dreak Inside the .-
Dryvell with all ECCS Operating (HPCS + 2CS + 4LPCI + ADS)
AEC Assumptions 6-29 Core Flov and Pressure Following a Recirculation Line Break 6-30 6-10 6-11 Performance of ECCS with Failure of LPCI Injection Valve for the Design Basis Recirculation Line Break (KPCS + 2CS + ADS)
AEC Assumptions 6-32 6-33 6-12 Minimum Critical Heat Flux Ratio for DBA at Monticello 6-13 Cladding Temperature versus Time for an Intermediate Break with Failure of HPCI (0.1 ft 2 Break) (4LPCI + 2CS + ADS) 6-35 AEC Assumptions iv
! Table of Contents l ?.*K2
- 1. Introduction 1-1
'e
- 2. 2-1 Summary
'l 3. Mechanics! Design 3-1 3.1 Ceacrcl Design bescription 3-1 3.2 hechanical Design Bases 3-5 3.3 Results from Mechanical Design Evaluations 3-15 3.4 Fuel Operating and Developmental Experience 3-17 i References - Seccion 3 3-26
- 4. Thermal-llydraulic Characteristics 4-1 4.1 Fuel Assembly llydraulic Analysis 4-1
, 4.2 Fuel Assembly Thermal-Hydraulic Evaluction 4-4 4.3 Results of Thermal-Hydraulic Analysis 4-5 References - Section 4 4-8
- 5. Nuclear Characteristics 5-1 5.1 Introduction 5-1 5.2 Dundle Nuclear Description 5-1 5.3 Analytical Methods 5-11 5.4 Experience with GE Nuclear Models 5-15 5.5 Nuclear Characteristics of the Core 5-16 References - Section 3 5-19
- 6. Safety Analysis 6-1 6.1 Model Applicability to 8x8 Fuel 6-1 6.2 Results of Safety Analysis 6-6 References - Section 6 6- 6 f-
- 7. Technical Specifications 7-1 e
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- 1. INTRODUCTION This document provides the technical basis of the license submittal for the second reload of Northern States Power /Monticello Unit 1. Presented herein is a dcscription of the new fuel and the results of the evaluacion of the refueled core for the February,1974 outage.
The fuel et the site available for loading at the outage will be 116 Reload-2 fuel bundles, which are 8x8 bundles with an average enrichment of 2.62 wt% of U-235. There will also be available for reinsert 7 initial core bundles discharged at the end of Cycle 1 with an average exposure of about 7400 mwd /t.
The objective of this outage is to remove the 44 temporary coatrol cur-tains which remained in the ore during Cycle 2 and to load the core so as to assure the availability of the plant at high pour for approximately an annual cycle.
Sections 3, 4, 5 and 6 of this document, dealing with the subjects of reload fuel mechanical design and reloaded core thermal-hydraulic and nuclear charac-teristics, and safety analysis, present descriptions of design criteria, methods and results from design calculations and safety evaluations and represents com-plete information for the review of fuel assembly and core design.
4 1-1
LIST OF ILLUSTRATIONS (Continued)
Figure Title Page 6-14 Quality versus Time for DBA at Monticello 6-36 6-15 Heat Transfer coefficient for a Small Breah (0.02 ft2)
(4LPCI + 2CS + ADS) AEC Assumptions 6-37 6-16 !! eat Transfer Coefficient f or an Intermediate Break (0.1 f t )2 (ALPCI + 2CS + ADS) AEC Assumptions 6-38 6-17 Monticello Heat Transfer Coefficients for DBA with LPCI Injection Valve Pailure - AEC Analysis No. 1 (2CS + HPCI + ADS) 6-39 6-18 Performance of ECCS sith Failure of HPCI for Intermediate (0.1 ft2) Liquid Break (2CS + 4LPCI + ADS) AEC Assumptions 6-40 6-19 Power Ceneration Following a Design Basis Recirculation Lane Break Accident 6-41 6-20 Effects of Exposure on Maximum Cladding Temperature , 6-43 6-21 Fuel Rod Perforation Data 6-44 6-22 Distribution of Internal Pressure Within Rods 6-45 6-23 Percent Rod Perf oration versus Break Area (AEC Assumptions) 6-46 6-24 8x8 Reload Fuel Rod Identification 6-48 6-25 Fuel Type Locations 6-52 6-26 Base Rod Pattern for RWE 6-53 6-27 Monticello Composite MCHFR versus Rod Position Limiting Rod Withdrawal 6-54 6-28 MCHFR versus Rod Position Limiting Rod Withdrawal 7x7 6-55 6-29 MCHFR versus Rod Position Limiting Rod Withdrawal 8x8 6-56 6-30 Peak Heat Flux at.d Relative Power versus Rod Position Lindting Rod Withdrawal Monticello BOC 3 6-57 6-31 RBM Response - Channels A and C 6-58 6-32 RBM Response - Channels B and D 6-59 6-33 Scram Reactivity Curve 6-62 6-34 Control Rod Drive Scram Times 6-63 v
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- 3. MECRANICAL DESICN ,
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l 3.1 CENERAL DESIGN DESCRIPTION ,
\ The 8x8 fuel bundle contains 63 fueled rods and one spacer-capture water rod which are spaced and supported in a square (8x8) array by the upper and lower tie plates (see Figure }-1) . The lower tie plate has a nosepiece which ,
- has the function of supporting the fuel assembly in the reactor. The upper tie l plate has a handle for transferring the fuel bundle f rom one location to another. ,
e The identifying assembly number is engraved on the top of the handle, and a boss I projects from one side of the handle to aid in assuring proper fuel assembly ,
orientation. Both upper and lower tie plates are fabricated from Type-304 stain-less steel castings.
1
) Each fuel rod consists of high-density (95% TD) UO2 uel pellets stacked in a Zircaloy-2 cladding tube which is evacuated, backfilled wit *'. helium, and sealed by velding Zircaloy end plugs in each end. The fuel rod cladding thick- ,
ness is adequate to be " free-standing," 1.e., capabic of withstanding external reactor pressure without collapsing onto the pellets within. Although most-
. fission products are retained within the UO , a racti n the gaseous products 2 .
are released from the pellet and accumulate in a plenum at the top of the rod.
Suf ficient plenum volume is provided to prevent excessive internal pressure from i these fission gases or other gases liberated over the design life of the fuel.
- A plenum spring, or retainer, is provided in the plenum space to prevent move-t ment of the fuel column inside the fuel rod during fuel shipping and handling.
Three types of rods are employed in a fuel bundles tie rods, a water rod, l
l and standard rods. The eight tie rods in each bundle have threaded ena plugs .
which thread into the lower tie plate casting and extend through the upper tie '
plate casting. A stainless steel hexagonal nut and locking tab are installed g I
f on the upper ena plug to hold the assembly together. These tie rods support the
! weight of the assembly only during fuel handling operations when the assembly
, hangs by the handle; during operation, the fuel rods are supported by the lower
- tie plate. One rod in each fuel bundle (see Figure 3-2) is a hollow water tube .
i used to position seven Zircaloy-4 fuel rod spacers vertically in the bundle.
4 .
l l 3-1
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- 2.
SUMMARY
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i i
j The Monticello Unit I Reload-2 fuel will employ an 8x8 fuel assembly con-l figuration instead of the previously used 7x7. The pellet diameter, pellet 4
i length, cladding diameter, and rod pitch are changed from 7x7 design; how-
- ever, the_ assembly exterior. dimensions remain unchanged. The basic materials and fuel fabrication process used for the Reload-2 fuel assemblies are the same 1
1 as those used on the 7x7 design.
i The design reference core configuration for this license submittal consists
- of 116 Reload-2 new fuel bundles with an average enrichment of 2.62 wt% U-235, t
20 Reload-1 fuel bundles with an average exposure about 5550 MRd/t and 348 Ini-tial fuel bundles with an average exposure about 10,300 mwd /t. All temporary a
control curtains are removed from the core. The Reload-2 fuel bundle uses gado-linium for reactivity control augmentation. The relative location of the i
Reload-2 fuel bundles is shown in Figure 2-1. The design reference core was i
) developed from an extrapolation from expected cycle 2 operation to the end j of cycle and it contains extensive shuffling of irradiated fuel assemblies.
4 At the time of the outage the final core loading will be analyzed and compared i
l to the design reference case to insure that the core meets license requirements.
i Shutdown calculations.have been made on the Efull use of all Reload-2 bundles.
l Ample shutdown margin for the most reactive condition in the cycle has been i
i calculated for the design reference core.
These results indicate that the reactor will be able to operate safely at i 1670 NWt after the outage and satisfy all license requirements.
Table 2-1 1
FUEL TYPE AND NUMBER j Fuel Type Number Initial 348 4
Reload 1 20
! Reload 2 116
! Total 484 g
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, HM ENRICHMENT NUMSER TYPE wt % U-236 OF RODS 1 2.87 40 2 2,14 14 j 3 gg7 4 f
4 1,45 1 6 2A7 4 I W5 _ i i'
WS - SPACEH CAPTURE WATER ROD T - TIE RODS O - GADOLINIUM RODS Figure 3-2. Monticello R2 Reload Fuel Lattice ,
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Table 3-1 presents a summary of 8x8 design dimensions and a comparison to initial core and R1 fuel, and Figure 3-2 shows the location of the various
, fuel rod types within the Reload-2 assembly.
- 3.2 MECHANICAL DESIGN BASES J
In meeting the power generation objectives, the nuclear fuel shall be used '*
as the initial barrier to the release of fission products. The fission product retention capability of the nuclear fuel shall be substantial during normal modes '.
of reactor operation so that significant amounts of radioactivity are not released from the reactor fuel barrier.
The nuclear fuel shall be designed to assure (in conjunction with the core nuclear characteristics, the core thermal and hydraulic characteristics, the plant equipment characteristics, and the capability of the nuclear instrumenta-tion and reactor protection system) that fuel damage limits will not be exceeded during either planned operation or abnormal operational transients caused by any single equipment malfunction or single operator error.
3.2.1 Basis for Fuel Damage Analysis Fuel damage is defined as a perf oration of dhe fuel rod cladding which would permit the release of fission products to the reactor coolant.
The mechanisms which could cause fuci damage in reactor operational tran-sients ares (1) rupture of the fuel rod cladding due to strain caused by rela-tive expansion of the UO 2 Pellet; and (2) severe overheating of the fuel rod ,
cladding caused by inadequate cooling.
- A value of 1% plastic strain of the Zircaloy cladding has traditionally been defined as the limit below which fuel damage due to overstraining of the fuel cladding is not expected to occur. The 1% plastic strain value is based on General Electric data on the strain capability of irradiated Zircaloy cladding -
segments from fuel rods operated in several BWRs. None of the data obtained fall below the 1% plastic strain value; however, a statistical distribution fit -
I to the available data indicates the 1% plastic strain value to be approximately i
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The water rod is a hollow Zircaloy-2 rod equipped with a square bottom end plug to prevent rotation and assure proper location of the water rod within the fuel assembly. Several holes are drilled around the circumference of the water rod at each end to allow coolant water to flow through the rod. The spacers are equipped with Inconel-X springs and maintain rod-to-rod spacing. The remaining 55 rods in a bundle are standard rode the same active fuel length as the tie rods.
3~ The end plugs of the standard rods have pins which fit into anchor holes in the tie plates. An Inconel-X expansion spring located over the top end plug pin of each fuel rod keeps the fuel rods seated in the lower tie plate and allows them to expand axially and independently by sliding within the holes of the upper de plate.
The fuel pellets consist of high-density ceramic uranium dioxide manufactured by compacting and sintering uranium dioxide powder into cylindrical pellets with chamfered edges. The average UO2 pellet immersion density is approxiuately 95%
of theoretical density.
Four dif ferent U-235 enrichments are used in the fuel assemblies to reduce the local power peaking factor (see Figure 3-2). Fuel elema.nt design and manu-facturing procedures have been developed to prevent errors in enrichment location within a fuel assembly. The fuel rods are designed with characteristic mechani-cal and fittings, one for each enrichment. End fittings are designed so that it is not mechanically possible to completely put together a fuel assembly with any high enrichment rod.s in positions specified to receive a lower enrichment. As in the 7x7 assembly design, the 8x8 bundle incorporates the use of small amounts of gadolinium as a burnable poison in selected fuel rods. The gadolinia-urania fuel rods are designed with characteristic extended end plugs. These extended end plugs permit a positive, visual check on the location of each gadolinium-lg bearing rod af ter bundle assembly.
4 Most aspects of the 8x8 bundle design are similar to the current 7x7 design.
Specifically, the upper and lower tie plates, the fuel rod spacers,. the upper and lower end plugs, and other associated bundle hardware are geometrically similar to the 7x7 except for modeling down in size to be sv2patible with the increased j number of rods per bundle and the reduced rod diametral wall thickness. laut 8x8 fuel assembly outline dimensions are the same as the current 7x7 dimensions.
i
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i 3-4 2
l Table 3-1 INITIAL CORE AND RELOAD FUEL ASSEMBLY DESIGN SPECIFICATIONS Initial Reload Fuel Core Fuel R1 R2 Fuel Assembly Geome t ry 7x7 7x7 8x8 High Enrichment Rods 22 32 44 Medium High Enrichment Rode 19 10 14 Medium Low Enrichment Rods 8 6 4
- Low Enrichment Rods 0 1 1 Poison Rods 0 3 4 Water-Spacer Capture Rods 0 0 1 Rod Pitch (in.) 0.738 0.738 0.640 Water to Fuel Volume Ratio 2.47 2.53 2.60 Heat Transf er Area (f t2 ) 86.5 86.5 97.6 Fuci Rod
- Active Fuel Length (in.) 144.0 144.0 144.0 Gas Plenum Length (in.) 11.25 11.0 11.24 Fill Gas helium helium helium Getter no yes yes
, Fuel Material sintered UO 2 sintered UO sintered UO2 2
Initial Enrichment, wt/% U-235 Average for Bundle 2.25 2.30 2.62 High 2.95 2.56 2.87 Medium High 1.91 1.94 2.14 Medium Low 1.13 1.69 1.87 Low -
1.33 1.45 Pellet Diameter (in.) 0.487 0.477 0.416 Pellet Immersion Density (% TD) 95.0 95.0 95.0 Cladding
- Material Zr-2 Zr-2 Z r-2 Thickness 0.032 0.0 37 0.0 34
, Outside Diameter (in.) 0.563 0.563 0.493 Fuel Channel Material Zr-4 Zr-4 Z r-4 Outside Dimension (in.) 5.438 5.438 5.438
, Wall Thickness (in.) 0.080 0.080 0.080 Channel Length (in.) 162-1/8 162-1/8 162-1/8 3-6 l
Table 3-1 (Continued) i Initial Reload Fuel Core Fuel ~
R1
~
R2 Spacers Material Zr-4 with Zr-4 with Z r-4 with .
Inconel Inconel Inconel Springs Springs Springs Number per Bundle 7 7 7 .
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?I the 95% point in the total population. This distribution implies, therefore, a >
j small (<5%) probability that some cladding segments may have plastic elonga-l '
tion less than 1% at failure.
t 4
i
- For design purposes, critical heat flux (the onset of the transition from '
l nucleate boiling to film boiling) is conservatively defined as a design limit '
for fuel damage, although fuel damage is not expected to occur until well into the film boiling regime. Severe overheating of the fuel rod cladding is assumed .
~
i - to occur at a condition of minimum critical heat flux ratio (NCHFR - the minimus i
ratio of the critical heat flux correlation value at the corresponding fluid con-d
} ditions to the actual heat flux at a given point in the fuel assembly) less than ,
1 1.0. If MCHFR remains above 1.0 no fuel damage occurs as a result of inadequate j
cooling. The steady-state MCHFR and the resulting MCHFR during transients are discussed in more detail in Sections 4 and 6.
3.2.2 Ef fects of Radiation and Fuel Swelling
+
j Irradiation affects both fuel and cladding material properties. The effects include an increased cladding strength and a reduced cladding ductility. In addi-tion, irradiation in a thermal reactor environment resulta 1Ln the buildup of both j gaseous and solid fission products within the UO u*1 Pellet which tend to 2
increase the pellet diameter, i.e., fuel irradiation swelling. Pellet internal porosity and pellet-to-cladding gap have been specified in such a way that the thermal expansion and irradiation swelling are accommodated for the worst-case dimensional tolerances throughout life. The irradiation swelling model is based .
on data reported in References 1 and 2, as well as an evaluation of applicable l ,
high exposure data.3 l
l Observations and calculations based on this refined model for relative UO2 8
l fuel-cladding expansion indicate that the as-fabricated UO2 pellet porosity is l adequate (without pellet dishing) to accommodate the fission-product-induced UO l 2 swelling out to and beyond the peak exposures anticipated for this reload.
The primary purpose of the gap between the UO 2 fuel pellet and Zircaloy cladding is to accommodate differential diametral expansion of fuel pellet and 3-8 e- - ,-----.-..e-, --,--,-,w--,--- .w.- , - . ..m t-re w w - ., -w,,..------.,,;,,w-,-, ,.e-~w+.m-.es-...-,w-.s . - . .--ep ,.nz,-,,c,i-*-w w, c. - ww , m p --
cladding and, thus, preclude the occurrence of excessive gross diametral cladding strain. A short time af ter reactor startup, the fuel cracks radially and redis-tributes out to the cladding. Experience has shown that this gap volume remains available ia the form of radial cracks to accommodate gross diametral fuel expansion.4 The thermal conductance across the pellet / clad gap, in theory, depends upon ,
the gas conductivity and the distance of the pellet from the cladding when pellet '
and clad are not in contact, and upon the pressure of the fuel on the cladding if they are in contact. Initially, the gap is filled with helium. As the fuel accumulates exposure, a namber of phenomena which can influence the pellet-clad thermal conductance can become important. Fission gases are released from the fuel and dilute the helium gas to form a mixture of He, Kr. Xe and UO I"E"#I*Y 2
volatiles with lower thermal conductivity than pure helium in the free volume within the fuel rods. In addition, it has been postulated that the phenomenon of fuel densification may tend to cause an increase in the pellet-to-cladding gap with an attendant feedback on pellet-clad thermal conductance. The important observation in this regard is that there is a phenomenon which tends to counter-act the adverse effects of fission gas dilution and fuel densification. Spe-cifically, it has been observed that for high power BVR fuel rods, the fuel pellet-to-cladding gap closes progressively with exposure in spite of any effect of densification on pellet diameter, with the result that the pellets and clad-ding achieve intimate contact with increasing exposure, thus reducing the impor-tance of the gas conductivity to good thermal conductance.4 This qualitative discussion serves merely to describe the phenomena influ- .
encing pellet-clad thermal conductance with increasing exposure. In the integral
- models employed in the detailed mechanical design analysis of BWR fuel, the value of pellet-clad thermal conductance is held constant for convenience. The constant '
value employed is 1000 Btu /h-f t 2 ,.F. The use of this constant value has been found to be a conservative assumption when applied in conjunction with the inte-gral fuel design models employed by General Electric. Specifically, the design .
fission gas release model employed in the determination of fuel rod plenum size and cladding wall thickness has been shown to overpredict available data on fie- -
sion gas release when applied with a pellet-clad thermal conductance value of 3-9 i
1 J
1000 Stu/h-ft 'F. Similarly, the design model for relative fuel-cladding expansion (pellet-to-cladding interaction) also has been shown to be very con-
, servative relative to available data when a value of 1000 Btu /h-f t 2 ,.F is used i
for pellet-cladding thermal conductance. The basis for these integral fuel i
design models is described in more detail in Reference 3.
l 1
.. Fission-product buildup also tends to cause a slight reduction in fuel melt-ing temperature. The melting point of UO 2 is e naidered to reduce with irradia-tion at the rate of 32('C)/10,000 (mwd /Te).
- 7 j In the temperature range of interest (>500'C) the fuel thermal conductivity
- is not considered to be significantly affected by irradiation.
A small fraction of the gaseous fission products (approximately 20%) are released from the fuel pellets to produce an increase in fuel rod internal gas pressure. In vneral, such irradiation effects on fuel performance have been characterized by available data and are considered in determining the design features and performance. Thus, the irradiation effects on fuel performance are inherently considered when determining whether or not the stress intensity limits and temperature limits are satisfied.
l 3.2.3 Maximum Allowable Stresses The strength theory, terminology, and stress categories presented in the ASME Boiler and Pressure Vessel Code,Section III, are used as a guide in the mechanical design and stress analysis of the reactor fuel rods. The mechanical l
design is based on the maximum shear stress theory for combined stresses. The
}
equivalent stress intensities used are defined as the difference between the most
- positive and least positive principal strasses in a triarial field. Thus, stress t ..
, intensities are directly comparable to strength values found from tensile tests. *
- Table 3-2 presents a summary of the basic stress intensity limits that are i
applied for Zircaloy-2 claddingt l
I
- 3-10 l-i-
. . _ . _ . - ~ - - - - - . . ._.. - . - . - - . - . . . - . . - . _ . . - _. .
1 1
i j Table 3-2 q STRESS INTENSITY LIMITS '
Yield Strength Ultimate Tensile Categories (Sy) Strength (Su) l Primary Membrane Stress 2/3 1/2 i Primary Membrane Plus Bending Stress ,
l Intens i ty 1 1/2 to 3/4 Primary Plus Secondary Stress Intensity 2 1.0 to 1.5
, In the design of BWR Zircaloy-clad UO 2 Pellet fuel, no continuous func-3 tional variations of mechanical properties with exposure are employed since the s. .
I irradiation effects become saturated at very low exposure. At beginning of life, !
I the cladding mechanical properties employed are the unirradiated values. At subsequent times in life, the cladding mechanical properties employed are the l saturated irradiated values. The only exception to this is that unirradiated
{ mechanical properties are employed above the temperatures for which irradiation I
ef fects on cladding nachanical properties are assumed to be annealed out. It-is
)
] significant that the values of cladding yield strength and ultimate tensile 2
strength employed represent the approximate lower bound to data on cladding f ab-ricated by General Electric, i.e., approximately two standard deviations below f
l the mean value.
Design analyses have been performed for the 8x8 -reload fuel which show that the stress intensity limits given in the above table are not exceeded during continuous operation with linear heat generation rates up to the operating limit ,
of 13.4 kW/f t, nor for short-term transient operation up. to 16% above the peak operating limit of 13.4 kW/ft, i.e., 15.6 kW/ft. Stresses due to external cool- .
4 ant pressure, internal gas pressure, thermal effects, spacer contact, flow-induced vibration, and manufacturing tolerances were considered. Cladding mechanical properties used in stress analyses are based on test data of fuel . ,
rod cladding for the applicable temperature.
}-11
l 3.2.4 Capacity for Fission Gas Inventory A plenum is provided at the top of each fuel rod to accommodate the fission 4
gas released from the fuel during operation. The design basis is to provide suf-ficient volume to limit the fuel rod internal pressure so that cladding stresses do not exceed the limits given in Table 3-2 during normal opetation and for short-term transients of 16% or less above the peak normal operating conditions.
3.2.5 Maximum Inte rnal Gas Pressure Fuel rod internal pressure is due to the helium which is backfilled at one j
atmosphere pressure during rod f abrication, the volatile content of the UO '
2 and the fraction of gaseous fission products which are released from the UO '
2 The most limiting combination of dimensional tolerances in assumed in defining the hot plenum volume used to compute fuel rod internal gas pressure. A quantity of 1.35 x 10' gram moles of fission gas are produced per HWd of power produc-tion.
In fuel rod pressure and stress calculations, 4.0% of the fission gas pro-duced is calculated to be released from any UO2 v lume at a temperature less than 3000*F and 100% from any UO above 3000*F. The above basis has been demon-2 strated by experiment to be conservative over the complete range of design ten-perature and exposure conditions. The calculated maximum fission gas release fraction in the highest design power density rod is <20%. This calculation is conservative because it assumes the most limiting peaking factors applied to this rod. The percentage of total fuel rod radioactivity released to the rod plenum is less than 20% because of radioactive decay during diffusion from the UO '
2 3.2.6 Internal Pressure and Cladding Stresses During Normal Conditions
~ The maximum internal pressure is applied coincident with the minimum appli-cable coolant pressure to compute the resulting cladding stresses which, com-bined with cladding stresses from other sources, must satisfy the stress limits described in Table 3-2. The maximum internal pressure generally does not exceed 1800 psia.
3-12 e
- - _ - - . - - = . - . . . . - . - - - . .-. -- - - _ - . - - _ - _ _
1 I
3.2.7 Cyclina and Fatigue Limits .
1 The design basis for fuel fatigue limits consists of the linear cumulative -
j damage rule (Miner's hypothesis)$ and the Zircaloy f atigue design basis of i Reference 6. The fatigue life analysis is based on the estimated number of tem-perature, pressure, and power cycles. During fuel life, less than 5% of the allowable fatigue life is consumed.
i Cyclic condition Estimated Cycles '.i Room temperature to 100% power . . . . .. . . . . .. . . . ...... N4/yr j Hot standby to 100% powe r . . . . . . . . . . . . . . .. . . . . , . . . . N12/yr 50% power to 100% power . . .. . . . . . . . . . . ... ... . . . . . N60/yr 75% power to 100% power . . . . . . . . .. . . . .... .. ..... N250/yr 100% power to 116% power . .. .. . ... . . . . . .. . .. . . . . N1/2 yr 3.2.8 Deflection i
The operational fuel rod deflections considered are the deflections due tos
- 1. Manuf acturing tolerances
- 2. Flow-induced vibration
- 3. Thermal effects
- 4. Axial load
- There are two criteria that limit the magnitude of these deflections. One criterion is that the cladding stress limits must be satisfied; the other is that the fuel rod-to-rod and rod-to-channel clearances must be sufficient to allow .
i free passage of coolant water to all heat transfer surf aces. Thermal hydraulic .
j testing has demonstrated that allowing a statistical minimum clearance of 0.060 inch at two standard deviations away from the nominal clearance is sufficient to a I
assure a very low probability of local rod overheating due to occurrence of critical heat flux.
I i
i t
d 3.2.9 Flow Induced Fuel Rod Vibrations m
Flow-induced fuel rod vibrations depend primarily on flow velocity and fuel
- rod geometry. For the range of flow rates and geometrical variations for the plant, vibrational amplitude does not exceed 0.002 inch. The maximum vibrational amplitude occurs midway between spacers due to the constraint of the spacer. The
. stress levels resulting from the vibrations are negligibly low and well below the endurance limit of all affected components.
3.2.10 Fretting Corrosion 1
Fretting wear and corrosion have been considered in establishing the fuel mechanical design basis. Individual rods in the fuel assembly are held in post-tion by spacers located at intervals along the length of the fuel rod. Springs are provided in each spacer cell so that the fuel rod is restrained to avoid excessive vibration. Tests of this design have been conducted both out of reactor as well as in reactor prior to application in a complete reactor core basis. All tests and post-irradiation examinations have indicated that fretting corrosion does not occur. Post-irradiation examination of many fuel rods indi-cates only minor fretting wear. Excessive wear at spacer contact points has never been observed with the current spacer configuration.
(
l 3.2.11 Potential for Hydriding The design basis for fuel in regard to the cladding hydriding mechanism is to assure, through a combination of engineering specifications and strict manu-l f acturing controls, that production fuel will not contain excessive quantities of moisture or hydrogenous impurities. An engineering specification limit on
, moisture content in a loaded fuel rod is defined which is well below the thresh-
- old of fuel failure. Procedural controls are utilized in manufacturing to pre-
. vent introduction of hydrogenous impurities such as oils, plastics, etc., to the f fuel rod. Hot vacuum outgassing (drying) of each loaded fuel rod just prior to final end-plug welding is employed to assure that the level of moisture is well below the specification limit. As a further assurance against possible fuel rod perforation resulting from inadvertent admission of moisture or hydrogeneous a
e 3-14 l
e9
- ,, --,e -
., . - - - - . _ - . - - - - ~. _ -.. . . - - -. ..- - - - . -. -.
l i
i l
l
.i Impurities into a fuel rod, General Electric is now using a sirecnium alloy l hydrogen getter material in all fuel rods. This getter material has been proven ef fective by both in-pile and out-of-pile tests.
1
- 3.2.12 Dimensional Stability l
j The fuel assembly and fuel components have been designed to assure dimen- -.-
l sional stability in-service. The fuel cladding and channel specifications 4
include provisions to preclude dimensional changea due to residual stresses. '
In addition, the fuel assembly has been designed to arr.osmodate dimensional
] changes that occur in-service due to thermal diff erential expansion and irradia-
, tion effects: for example, the fuel rods are free to expand lengthwise inde-i pendent of each other, and the channel is free to expand relative to the fuel bundle.
3.3 RESULTS FROM MECilANICAL DESIGN EVALUATIONS l 3.3.1 Steady-State Mechanical Performance Reload fuel is designed to operate at core rated power with suf ficient design margin to acconriodate reactor operations and satisfy the mechanical design bases discussed in detail in Section 3.2. In order to accomplish this objective, the 8x8 reload fuel is designed under the most limiting conditions l at 100% of rated power, to operate at a maximum steady-state linear heat genera-i tion rate of <13.4 kW/ft.
l Thermal and mechanical analyses have been performed which demonstrate that the mechanical design bases are met for the maximum operating power and exposure .
combination throughout fuel life.
l 3.3.2 Fuel Damage Analysis
- For fresh UO 2 fuel the calculated linear heat generation rate (LHGR) cor-responding to 1% diametral plastic strain of the cladding is approximately 25.4 kW/ft. Later in life the calculated linear heat generation rate correspondiag ,
i i
3-15 i
l 1
T
- - - . - - . n.e -, - , . , , - , , , ,. , . . - . , , . . . , .,,,n,....... ,,..w,.- .,. -, - , - m.n..,,
to 1% diametral plastic strain decreases to approximately 23.8 kW/f t at 25,000 mwd /t and approxicately 21.1 kW/ft at 39,500 H4d/t. However, due to a depletion of fissionable material, the high exposure fuel has less nuclear capability and will operate at correspondingly lower powers; therefore, a wide margin is main-tained throughout life between the operating LHGR ani the LHGR calculated to cause 1% cladding diametral strain.
~
The addition of small amounts of gadolinia to UO 2 results in a reduction in the fuel thermal conductivity and ??lting temperature. The result is a reduc-tion in the LHGRs calculated to cause 1% plastic diametral strain for gado11nia-urania fuel rods. However, the gadolinia-urania fuel rods are designed to oper-ate at lower power to compensate for this and provide margins similar to standard UO 2 r de.
For the 8xB reload fuel design analysis has shown that the power required to produce 1% plastic strain throughout life for all rod types in the assembly is equal to or greater than 180% of the maximum steady-state power.
3.3.3 Incipient U02 Center Helting For the 8x8 reload fuel, incipient center melting is expected to occur in fresh UO 2 fuel r da at a linear heat generation rate of approximately 20.4 kW/ft. This condition corresponds to the integral:
melt kdT = 93 w/cm t .
32'F
~
8 where k=
692 61+T + 6.02366 x 10-12 (T + 460)3 B tu/h-f t *F, and T is in *F.
3-16
V The value of the above integral decreases slightly with burnup, as a result of the decrease in fuel melting temperature with increasing exposure.
3.4 FUEL OPERATING AND DEVELOPMENTAL EXPERIENCE 3.4.1 Fuel Operating Experience -
The peak linear heat generation rate design limit for steady-state opera- ,
tion in 13.4 kW/f t which corresponds to a heat flux of 354,250 Btu /h-ft2 . This condition is well within the bounds of available production and developmental fuel experience.
The fuel operating limit and the fuel damage limit have been established based on operating experience and experimental teste covering the complete range 4
of design power and exposure levels. Tables 3-3 and 3-4 present a summary of power reactor production fuel experience. Tables 3-5 and 3-6 show the ranges of development fuel irradiations which have already been completed or are in progress. This experience has been used in establishing design features and in the analysis of performance characteristics. A large volume of experience has been obtained over the past 10 to 15 years with production fuel in commercial power BWRs and numerous developmental irrt.diations.
i The large volume of production experience, starting with the f$rst load of fuel in Dresden 1 Nuclear Power Station in 1960, has provided feedback on the adequacy of the design for, and the effects of, operation in a commercial power reactor environment. Production fuel experience has also provided feedback on ,
the incidence and effect of flaws and impurities which occur statistically in
- large volume production processes.
1' The production Zircaloy-clad UO2 Pellet fuel experience is supplemented i
by a large amount of in-pile and out-of-pile developmental work. The develop-mental work to date has been employed to test a wide range of design charac- ,
teristics, to investigate various mechanisms affecting the performance of the j fuel rod, and to extend irradiation experience to higher local combinations of fuel rod power and exposure than covered by production fuel.
3-17
g . .
Table 3-3 StM OF LEADING EIFERIEN G ON CCREENT1.Y OPERATING FROtRICTION ZIRCAIDY-CIAD U0 3 F11.1ZT Ft EI AS OF OCIORE.R 1 1971 Exposure Exposure Design mad er of Design Fuel Fellet-to-Feek Average Time Acties F1selon Gaa Segments Man Seat Peak Rod Clad Clad cap Feilet Assembly 1'acore Flum(a,b) IJECR( a,b)
Fuel Flenssa (Vol or Rede Die Thicknese (Nominal) Length For Unit 5tt11 in Reactor (Wd/Te) (W d/Te) (Years) (8tu/h-ft .)2 (kW/ft) (afle)
(in.) .( elle) (ie.) Fuel Tol) Core Dresden I Type 111 8 25.800 16.450 5.45 360000 15.4 0.555 Dresden I Type 111 F 35 7.5 109.0 0.040 3.780 28.200 19.480 4.45 360000 15.5 0.5625 35 Dresden I Type V to 108.25 0.048 2.592 21.320 13.250 2.45 360000 15.5 0.5625 ".5 Carigilano Type A to 108.25 0.C48 3.492 26.120 15.180 7.35 252000 10.3 0.534 30 5 105.7 Carigliano Type SA 0.0 31 6.804
& SB 15.160 7.270 2.95 320000 14.6 0.59 3 37 18 107.0 0.030 Consumers (BRF) 7.936 Type 8(e) 35.380 23.430 4.95 434000 15.0 0.449 34 8 70.0 0.048 242 Consumers (8RF)
Type E(e) 15.500 8.730 2.75 410000 17.7 0.5625 40 Consumers (BEF) 11 70.0 ".048 1.386 Type EC(d.e) 15.559 7.930 2.00 410000 17.7 0.5625 40 11 70.0 0.048 2.079 Consumers (BRF)
Type F 0.75 41*X)00 17.7 0.5625 40 11.5 70.0 0.048 1.771
, Humboldt Type 11 21.598 13.230 4.00 325000 12.1 0.486 33 Hue oldt Type 111 to 79.0 0.043 3.724 i 14.332 6.615 2.00 389000 16.8 0.563 32 11 79.0 as KR8 0.062 3.384 22.409 14.634 4.45 367000 15.8 0.5625 35 10 130.0 0.058 KRS-KD 11.124 7.050 1.15 367000 15.8 5.784 0.563 32 11 130.0 0.058 648 Tarapur I 13.738 8.337 1.80 365000 15.8 0.5625 35 10.5 144.0 0.059 10.224 Tarapur 11 13.407 7.818 2.00 365000 15.8 0.5625 Oyster Creek I 35 10.5 144.0 0.059 10.224 11.976 8.307 2.35 400000 17.5 0.570 35.5 11 144.0 0.078 Nine Mile Point 8.412 5.106 2.05 400000 27.440 17.5 0.570 35.5 11 544.0 0.078 Dresden II 4.825 2.900 2.00 405000 26.068 17.5 0.563 32 12 144.0 0.078 Dresden II (reload) 1.600 970 0.5P 405000 54.941 17.5 0.563 32 12 144.0 0.078 Dresden III 575 290 0.25 405000 10.535 17.5 0.563 32 12 144.0 0.075 Tsuruga 12.037 7.107 1.75 400000 35.476 17.5 0.570 35.5 12 144.0 0.078 Millstone 5.293 0.85 *4.700 3.065 400000 17.5 0.570 35.5 12 144.0 0.078 Fukushima-1 4.410 0.85 28.420 3.300 400000 17.5 0.570 35.5 12 144.0 0.078
&mttcello 2.945 0.75 19.600 s 1.582 405000 17.5 0.563 32 12 144.0 0.078 Nucienor 2.070 0.60 23.716 1.290 400000 17.5 0.570 35.5 12 144.0 0.078 KEM 500 500 0.10 19.600 428000 18.5 0.563 32 12 144.0 0.11 SVR/4(c) 45.000 5.00 11.172 27.500 428000 18.5 0.563 37 12 144.0 0.11 8x8 Raiond(c) 45.000 28.000 5.00 354000 13.4 0.493 34 9 144.0 0.08 a = at rated power b = license limit c = typical design as opposed to proven performance in preceding entries d = includes 15 assemblies with 2 rods per bundle of plutonitan e = values as of February it. 1971 J
Table 3-4 51hMARY OF PRODUCTIC1: .w(L f.XPES.!ENCE
!!8CAIET-CLAD 00 2 FELLrt FLTL AS Or OCf08t.R 1 1971 Mumber of tiesf3n or Averste hatteus Toers of Weight Number Fuel Bode Warranted Assembly Assembly Operation of Fuel of Fuel or Segments trpos ur e Exposure taposur e in (5.Segmente) (Wd/Tel (W d/Te) (Wd/Te) Ite ac t or_
Identificetton 1Q Amst e lles Dresdesi 1 Type 1 132,400 534 77.184(S) 7.400 9,100 23,100 1960-1969 Type 111-8 43.500 192 6.912 14.900 16.750 20,650 1964-1971 Type til-F 21.400 104 3.7 44 16,500 17.900 25.950 1965-1971 Type V 24,800 106 3.816 16,500 13.050 19,600 1967-1971 100 7,200(S) 8.800 11.000 21.000 1960 7 RVE-KAHL 14.100 Ger!811ono 19,750 1963-1971 Type A 111,100 229 16,848 12,100 13.600 . '
Type SA 29,380 66 4,224 19.300 15.200 17.900 1968-1971 .
Type 58 78,770 64 4.096 2.900 5,300 1970-1971 JDPR 9,800 76 5,472(5) 8.800 3.800 N/A 1963-7 Haboldt Type 11 28,600 169 8,281 15.400 13,600 17.250 1965-1971 Type 111 23.702 140 5,040 20.700 4.950 13.150 1971-Coneunero Type B 30 3,630 16.500 19.800 24.600 1966-1971 8.700 Type E 12,640 42 3,234 16.500 8,700 10,800 1968-1971 Type EG 11.617 38 2.926 16.500 7.700 10,800 1969-1971 Type F 6.977 23 1.771 22,000 - - 1971-KR3 Type A 104,100 37 1 13.248 16,500 15,300 18.900 1966-1971 Type KD 5.110 18 648 16.500 7.050 8,450 1970-1971 Tarepur 1 18 5,'a 20 284 13.916 16.500 8.300 10.450 1969-1971 Terepur 2 185.120 284 13.916 16.500 7,800 9,900 1969-1971 Oyster Creek 242.900 560 27.440 16,500 8,300 1,550 1969-1971 Nine Mile Pt. 230,760 532 26,068 16.500 5,100 6,250 1969-1971 Tsuruga 136,200 314 15,092 16.500 7.100 9,300 1969-1971 Dresden 2 314.050 753 35.476 20.900 2.900 3.450 1970-1971 Dresden 2 84 toad 87.420 215 10.535 20.900 1.000 1.300 1971-Fukushtaa 1 172,484 400 19,600 20.900 3.300 3,400 1971-Monticello 206,813 484 23.716 20.900 1.300 1.600 1971-580 28.420 20.900 3.100 3.700 1971-M111 stone 250.625 uucienor 172,818 400 19.600 20.900 1.300 1.550 1971-724 20.900 300 350 1971- i Dresden 3 314.050 35.476 228 11.172 20,900 <500 <500 1971-KIM 97.017 3-19 O
1 e
1
. Table 5-9 CENERA1. ELECTRIC DEVI24PMDITAL 1RAAD1ATICNS
!!RCALOY-Ct.AD 952 TD U0 PELLET FUEL RODS 2
4 Puel No. Rod Clad Wall Pellet-to- Peak Heat Peak Peak d
of Dio. Thicknese Clad Cap Plua LHCR Exposure Reactor 2 kame Rode itM (in.) (mile) (Stu/h-ftj (kW/ft) (wd/Te) Status Dr. eden Prototype VBh'R 9 0.565 0.030 3.0-16.0 460.000 19.94 12.000 completed Puel C 3 -le (R 6 D). VBWR 144 0.424 0.022 2.0-8.0 509.000 16.6 13.800 Completed j Dresden Prototypes VBWR 52 0.565 0.028 5.0- 8. 0 407.000 17.64 10.000 Completed High Per-formance CETR 12 0.565 0.030 4 0-6.0 630.000 27.0 1.500 completed h UO 1.126.000 49.0 2
High Per-formance CETR 2 0.565 0.030 4.0-11.0 1.355.000 58.0 14.000 Completed' UO g Sa-1* Dresden 1 98 0.424 0.022 4.0-8.0 400.000 13.0 40.000 completed d
D-1.2.3 Consumere 363 0.424 0.030 7.0 434.000 14.2 30.000 Completed I
t>.50 consumers 36 0.570 0.035 12.0 507.000 22.0 15.400 s.1 D-52.53 Consumers 58 0.700 0.040 13.0 525,000 27.0 4.600 t l CE-Malden Halden 21 0.563 0.032-0.060 7.0-14.0 510.000 22.0 (.300 Continuing a e USAEC Contract AT(04-3) - 189 Project Agreement 11 b = USAEC Contract AT(04-3) - 189 Project Agreement 17 e = USAEC contract AT(04-3) - 189 Project Agreement 41 d
- USAEC Contract AT(04-3) - 361 e a Rollow Pellet f a USArc Contract AT(04-3) - 189 Project Agreement 50
=
g tight fuel rods failed during second operating cycle due to-2 normal crud and scale deposition
, h = One rod fatture 6 49 kW/f t 1
- Puel assemblies presently out of reactor pending approval for reinsertion 3-20 1
=
, ,. ,. - . _ r,, - - = ,- " #
't Table 3-6 GENERAL ELECTRIC DEVELOP! ENTAL IRRADIATIONS 'i ZIRCAIDY-CLAD 95% TD UO PELLET CAPSULES l j GENERAL ELECTRIC TbST REACTOR i e
A Number Fuel Rod Clad Wall Pellet-to- Peak Heat Peak Peak 1 of Diameter Thickness Clad C-ap Flux LHCR Exposure Capsule Rods (mils) (kW/ft) f- (in.) (in.)
(Stu/h-ftf) _ (Mid/Te) Status .
- A 3 0.425 0.024-0.032 1.4-10.2 750,000 24.5 88,000 Complete 1 0.488 0.032 11.2 785,000 29.4 34,000 Complete B 6 0.489 0.034 7.8-11.6 504.000 18.9 65,000 Complete '
C 5 0.557 0.036 2.0-15.0 475,000 20.3 59,000 Complete l
D 5 0.557 0.036 2.0-14.0 540,000 23.0 36,500 Complete
- y i j g E 5 0.250 0.015 6.5 735,000 14.1 100,000 Complete l r
i j F 3 0.443 0.030 3.0-13.0 480 000 , 16.3 29,000 Complete
,1 .
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3
- a s
1 1
- I 3 i l
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i i r 4
i 4
e a s , * *
.- _ . . . . - . . _-_ L
More than 25 production fuel types have been designed, manuf actured, and operated in more than 19 BWRs. When all production fuel types are considered, a total of more than 440,000 Zircaloy-2-clad UO2 fuel r de have been operated in CE-designed BWks. Out of this number of rods. N180,000 of which went into opera-tion during 1970 and 1971, enly %0.2% have been detected to have failure due to wall perforation, and this includes fuel which failed after having exceeded design performance conditions.
i Peak linear heat generation rates (LHGR) from approximately 10 to 17 kW/ft have been experienced with the production fuel. Individual fuel assemblies have achieved average exposures greater than 23,500 Wd/Te and have operated more than 9 years in-co;e residence. In comparison, the 8x8 reload fuel has the following proposed operating characteristics :
13.4 kW/ft maximum LHGR (Operating Limits),
45,000 Wd/Te maximum local exposure, and 4-6 years in-core residence time.
Fuel rod diameters in the range of 0.425 to 0.570 inch o.d. with cladding wall thickness from 30 to 40 mils and pellet-to-cladding gaps from 3 to 11 mils have been used in production fuel. Rod-to-rod pitch has varied from 0.533 to 0.874 inch, with rod-to-rod spacing varying from 0.128 to 0.213 inch. Active fuel column lengths have varied from 59.8 to 144.0 inches with fission gas plenum volume per unit of fuel volume from 0.013 to 0.100. Such fuel rods have been licensed and operated in 6x6, 7x7, 8x8, 9x9,11x11 and 12x12 fuel bundle con-figurations. In comparison, the design for this 8x8 reload fuel has the follow-ing physical characteristics:
, . Bundle geometry = 8x8 Active fuel length = 144 in.
Fission gas plenum volume (volume per unit fuel volume) = 0.08 Fuel rod o.d. - 0.493 in.
Pellet-to-cladding gap = 0.009 in.
Rod pitch = 0.640 in.
Rod spacing = 0.147 in.
3-22 elb ar - -
I I 3.4.2 Fuel Developmental Erperience I P The production Zitcaloy-clad U02 Pellet fuel experience described in the previous section is supplemented by a large amount of in-pile and out-of-pile developmental work. The developmental work to date hes been employed to test a wide range of design characteristics, to investigate various mechanisms affect- ,
ing the performance of the fuel rod, and to extend irradiation experience to higher local combinations of fuel rod power and exposure than covered by produc- ,
tion fuel. The following presents a discussion of the pertinent developmental fuel experience which, in combination with the production fuel experience, pro-vides the basis for the current BWR fuel design and operacing limits.
Tables 3-5 and 3-6 present a summary of design details and performance con-ditions for Zircaloy-clad UO2 pellet fuel rods and capsules
- irradiated under General Electric or USAEC-General Electric development test programs. These data
, complement the BWR production fuel experience by providing edditional data et higher local combinations of fuel rod power and exposure. Ovarall, more '.han 800 fuel pins with design characteristics similar to the current BWR fuel have been irradiated under General Electric or USAEC-General Electric programs. The irradiations have been performed with BWR environment in both test reactors and in commercial power BWRs. Test reactors employed in General Electric develop-mental irradiations summarized in Tables 3-5 and 3-6 are the Vallecitos Boiling Water Reactor (VBWR), the General 21ectric Test Reactor (GETR) and more recently
]
' the Halden Reactor. Developmental fuel irradiations have also been performed in the Consumers Big Rock Point and Dresden Unit 1 commercial power BWRs.
i
! The range of peak performance conditions covered by the varioue development irradiations goes beyond the design performance conditions for fuel in this class
, of reactor. The development performance conditions includes ,
13.0 - 58.0 kW/ft maximum LHGR, and 1500 - 100,000 mwd /Te maximum local exposure.
I *
- A capsule, as used herein, refers to a test fuel rod. or group of rods combined, with all features similar to production fuel rods except for having reduced -
active fuel length (as low as approximately 3 in.).
l l
3-23
- - - . . - . , _ . - -. ., , , . . - , , , _ , . , - , - ~ . , ,.,.n , - - - - , , , .
g ,. ,
The corresponding operating conditions for this reload fuel are:
4 1
13.4 kW/ft maximum LHGR, and s45,000 mwd /Te maximum local exposure, a .
The range of design characteristics and dimensions covered by the various developmental irradiations also encompasses the characteristics and dimensions
~
employed in the current BWR fuel design. The range of design characteristics and dimensions covered by the various developmental irradiations include the following:
Fuel rod o.d. - 0.250 to 0. 700 in. ,
Clad wall thickness - 0.025 to 0.060 in.,
Pellet-clad gap - 0.0014 to 0.016 in., and Pellet length - 0.3 to 0.95 in.
The corresponding fuel design characteristics for this reload fuel are l
Fuel rod o.d. - 0.493 in.,
Clad wall thickness - 0.034 in. ,
Pellet-clad gap - 0.009 in. , and Pellet length - 0.420 in.
Considering the range of power levels and peak fuel burnups attained in the broad base of operating and developmental fuel experience, it has been concluded that the current 8x8 fuel design is a conservative application of this experience.
I A more complete review of GE BWR fuel experience is provided in Reference 3.
3.4.3 Fuel Damage Experience g
Although the incidence of f ailure in General Electric Zircaloy-clad UO 2 fuel has been quite low (so.2% out of more than 440,000 fuel rods), fuel has been operated at Dresden Unit 1 and elsewhere with perforated cladding. Dresden Unit I has operated with some failures in the Type 1 Zircaloy-clad UO 2 f""1' i" the Type 11 stainless-steel-clad UO uel, and more recently in fuel Typea 2
4 3-24
111-B, IV-y, and V. The Humboldt Bay and big Rock Point reactors have also oper-ated with f ailures in stainless steel fuel (Humboldt Type 1 and Big Rock Type A).
The Big Rock reactor has operated with some fuel f ailures in both the Type b and Type E Zircaley-clad UO uel designs as wel as a number of f ailed high power 2
(22 to 27 kW/f t) fuel rods in the center-melt developmental fuel assenblies. KnB, Dresden 2. Tsurugs, and Fukushima have operated with some perforated fusi rods ..
during initial operation. Failures have resulted from manufac'turing defects, incompatibility of ,ainless steel as cladding material in the BWR core steaar -
water environment, inadequate volume for accommodation of fuel expansion and/or fission ges pressure for fuel operated beyond design exposares, cladding over- E temperature caused by excessive deposits of crud on fuel rod surf aces resulting from materials in the feedwater system, f retting wear caused by foreign debris trapped in fuel rod spacers, local internal hydridias of the cladding, and local clad strains due to pelle./c1 riding interaction. In essentially all cases, the mechanisms causing the fuel rode to fail in service have been carefully identi-fled. Appropriate corrections have beer, made to the manuf acturing process and to the fuel or system design and operation to reduce the probability of future recurrence of such failures.
Operation with f ailed fuel rods has shown that the fission product release rate f rom defective fuel rods can be controlled by regulating power level. The rate of increase in released activity appc'ently associated with progressive deterioration of failed rods has been deduced from chronological plots of the offgas activity measurerents in operating plants. These data indicate that the activity release level can be lowered by lowering the local power lansity in the vicinity of the fuel rod f ailure. These measured data also indicate that sudden or catastrophic failure of the fuel assembly does not occur with continued opera- ,
tion and that the presence of a failed rod in a fuel asseebly does not result in propagation of failure to neighboring rods. Shutdown can be scheduled, as t required, for repairing or replacing fuel assemblics that have large defects.
Eva\uating the fission product release rate for f ailed fuel rods shows a wide variation in the activity release levels. Designers have attempted to relate the release rates to defect type, size, and specific power level. These ,
data support the qualitative observations that fission -roduct release rater are functions of power density and that progressive deterioration is a function of time.
~
3-25
t t i A more detailed summary of General Electric experience with BWR Zircaloy-clad U02 pellet fuel, including recent production and development data, has
! been documented (see Reference 3).
i 1
3.4.4 Fuel Dennification i
The amount of in-pile fuel densification in BWR Zitcaloy-clad UO2 pellet 4
J ,
fuel has been observed to be small and is not considered to have any significant i effects on fuel. performance. Detailed consideration of the occurrence and poten-
} tial ef fects of in-pfle fuel densification in Ceneral Electric BWRs is reported in Reference 4. The AEC staf f has recently issued a model for analysis of densi-fication effects in BWRs. This model is considered by Ceneral Electric to be overly conservative in light of observations on DVR fuel. A separate submittal will be provided to present the results of analysis employing the AEC staf f model.
REFERENCES - SECTION 3
, 1. WAPD-TH-263, "Ef fects of High Burnup on Zircaloy-Clad. Bulk U0 Plate Fuel 2
Element Samples," September 1962.
- 2. WAPD-TH-629 "Irradiatior. Behavior of Zircaloy-Clad Fuel Rods Containing Dished End 00 2 Pellets," July 1967.
1
- 3. Williamson, H. H. , and Ditmore, D. C. , " Experience with BWR Fuel Through
! September 1971," May 1972 (NEDO-10505).
. 4. Ditmore, D. C., and Elkins, R. B., "Densification Considerations in BWR Fuel Design and Performance," December 1972 (NEDH-10735).
- 5. Miner, M. A., " Cumulative Damage in Fatigue," Journal of Applied Mechanics, 2
,1,2,, Transactions of the ASME, g , 1945.
- 6. O'Donnel, W. J. , and Langer, B. F., " Fatigue Design Basis for Zircaloy Com-ponents," Nuelear Science and Engineering, jp,1964.
3-26
-e - - , - . -- ,-- ..-,.. - _ _ , - - - 4
l i
i.
1 j
1 4
- 4. THERMAL-HYDRAULIC CHARACTERISTICS ~
1 i 4.1 FUEL ASSD(BLY HYDRAULIC ANALYSIS ,
6 l
j 4.1.1 Cors Pressure Drop, Hydraulic Loads, and Correlations 1
j ne flow distribution to the fuel assemblies is calculated on the assump-l tion that the pressure drop across all fuel assemblies is the same. This assump- ,,
i tion has been confirmed by measurements of the flow distribution in modern boil-a j ing water reactor as reported in References 1 and 2. The components of bundle ,
pressure drop considered are friction, local, elevation, and acceleration. Pres-sura drop measurements made in operating reactors confirm that the total measured core eressure drop and calculated cure pressure drop are in good agreement. ,
) nere is reasonable assurance, therefore, that the calculated flow distribution I throughout the core is in close agreement with the actual flow distribution of I an operating reactor.
9 i
- 4.1.1.1 Friction Pressure Drop a
{ Friction pressure drop is calculated using the model relation ,
d i
- 2 v ft AP l g = 2gp 2 'TPF 8
DA Hd 3
where AP g = friction pressure drop, poi, .
w = mass flow rate, g = acceleration of gravity, .,
p = water density, D
H
= channel hydraulic diameter, Ag = channel flow area, '
{ L = length, l f = friction factor, and Ph ase frictioc multiplier.
f (TPF " tw 4-1 y--y..,,,,,+,.n.,_- .c m , ,,y,,,,.,p, y ,.n,.,,... ,,,,,,7 .~,,,,,_.,,,.__,,,.._,.,,,_4 --m_r,,-y~,., . . . - ,. - - - _ -,, _.,y.- , - . . - ,.
This basic model is similar to that used throughout the nuclear power industry.
The formation for the two-phase multiplier is based on data which compare closely to those found in the open 11terature.3 General Electric Company has taken significant amounts of friction pressure drop data in multirod geometries representative of modern BWR plant fuel bundles
, and correlated both the friction factor and two-phase cultipliers on a best-fit basis using the above pressute drop formulation. Checks against more recent data are being made on a continuing basis to ensure that the best models are used over the full range of interest to boiling water reactors.
4.1.1.2 Local Pressure Drop The local pressure drop is defined as the irreversible pressure loss asso-cisted with an area change such as the orifice, tie plates, and spacers of a fuel assembly.
The general local pressure drop model is similar to the friction pressure drop and is 2
v K AP g = I 2gp
} 'TPL where AP g = local pressure drop, psi, K = local pressure drop loss coefficient, A = reference area for local loss coefficient, (TPL
- bio-Phase local multiplier.
, and v g, and p are defined the same as for friction. This basic model is similar to that used throughout the nuclear power industry. The formulation for the two-phase multiplier is similar to that reported in the open literature' with the addition of empirical constants to adjust the results to fit data taken ,
at General Electric Company for the specific designs of the BWR fuel assembly.
Tests are performed in single-phase water to calibrate the orifice and lower tie 4-2 m
w
l I
plate, and in both single- and two-phase flow to arrive at best-fit design values for spacer and upper tio plats pressure drop. The range of twst variables is
]
- specified to include the range of interest to boiling ** ster reactors. Full scale
- 8x8 tests have been performed to determine the local loss coefficients for upper and lower tie plates and fuel rod spacers. These loss coefficients are in turn used in hydraulic analyses of the core for determination of local pressure losses. .
'l 4.1.1.3 Elevation Pressure Drop ,
I The elevation pressure drop is based on the well-known relation t
APg = pL pa p g (1 - o) + p gos where AP = elevati n pressure drop, psi, E
L = length,
- 0 = average water density, o = void fraction, and pg.pg = saturated water and vapor density, resp.
1 The void fraction correlation is similar to models used throughout the 4
nuclear power industry and includes effects of pressure, flow direction, mass velocity, quality, and subcooled boiling. Checks against new data are made on a continuing basis to ensure that the best models are used over the full range of interest to boiling water reactors.
4.1.1.4 Acceleration Pressure Drop 4
i .
The pressure drop component due to acceleration includes the pressure change experienced by the fluid at an area change and the pressure change resulting from density change, such as that which occurs in steam formation. The formulation for the acceleration pressure drop is as follows:
. 4-3
._. - - _ . _ - _ _ - - - - . ._. _ - _ - . - _ - . - _ - _ _ - _ _. ~ _ _ _ _ - __ _ - .
Acceleration Pressure Change due to riow Area Change 4
i v2 A"
- (' ~ '2 2go A 2I ' ' 'A I AP 8 ACC 2
1
, - where APACC " accel ***Li n Pressure drop,
. A = final flow area,
. 2 4
Ag = initial flow area, i
and other terms are as previously defined.
I
.! Acceleration Pressure Change due to Jensity Changes 0 * ~ 8 i ACC ch . Olff IN . ,
where
_1 , x 42 _ (1-x) 2 '
og pa p g (1-a) g pg= momentum density, x = steam quality, and other terms are as previously defined. The total acceleration pressure drop in boiling water reactors is on the order of less than 5 percent of the total pressure drop.
- 4.2 IVEL ASSEMBLY THERMAL-HYDRAULIC EVALUATION 4.2.1 Critical Heat Flux and Minimum Critical Heat Flux Ratio The critical heat flux (CHF) condition (the onset of the transition from nucleate boiling to film boiling) is one of the important denign considerations in boiling water reactors. It occurs whenever excessive heat is being trar.sferred 4-4
\
w y - - - - - .- _ ., . , - . ,-_,
- . ~ -- -
to boiling or evaporating water and is usually accompanie/ ,y a rapid ,
deterioration of the heat transfer process. The critical heat flux is a func~
tion of the 1cesi steam quality, mass flow rate, press,ree, and flow area geometry.
Analyses of CHF are based on the concept of the minimum critical heat flux ,
ratio (MCHFR) . The steen quality distribution, calculated by sear.a of energy balances between the fue; and coolant, is used with the CHF correlation to cal- ,
culate the spatial distribution of CHF values. Dividing these values by actual '[
design reactor heat fluxes yields the design MCHFR.
4.2.2 Steady-State Thermal-Hydrait,e Licensina Criteria ;
For purposes of maintaining adequate thermal margin during normal steady-state operation, the puviously establie'ad license limits of HCHFR t 1.9 and MLHGR 1 17.5 kW/ft were applied to the 7x7 initial core and reload fuel. For ;
the 8x8 reload fuel, the limits of HCHFR 11.9 and MLHGR 113.4 kW/f t were employed. Results from safety analyses usinis these steady-state operating limits as initial conditions are discussed in Section 6. Results of full scale 8x8 CHF testing will be made available to the US,.rC upon completion of this ongoing test 6
program.
4.3 RESULIS OF THERMAL-HYDRAULIC ANALYSIS i Analyses were performed for a variety of core loadings to fully assess the I
i effect of the 8x8 reload assembly on core thermal-hydraulic characteristics.
i "
i The fiw core configurations considered are described as fol. lows:
l i
i
- i 1. Core loaded with 7x7 fuel (representative of initial etere or Reload-1 ,
l core loading).
t i
! 2. Core loaded with 7x7 fuel and a single 8x8 reload fuel assembly. ;
I !
- 3. One-quarter of the core loaded with 8x8 reload asseselies ad the '
i remainder loaded with 7x7 fuel. ,
i !
l l
l 4-5 l.
. . _ . _ . _ _ _ ,,._ .,_. - _ _ _ ._ _ . _ _ .. , _ ~ _ . _ . _ .. _ ,_ . ,,~. _ _
I 4
I J
l !
i
-I 4 I One half of the core loaded with Bx0 reload assemblies and the l l remainder loaded with 7x7 fuel.
]
[
i .
i 5. Tull core loaded with 8x8 reload assemblies.
i L
{ The thermal-hydraulic analyses were performed for the follwing reactor -
l l ,, conditions l {
t Reactor Powers
. i 1670 WL ;
f* Reactor Pressure: 1040 psia (steam dome) '
! Recirculation f1w rate 6 i 57.6 X 10 lb/h i i Inlet enthalpy 523 Stu/lb i
{ Bypass flow 10% of total core f1w
- 1. !
The same design basis power distribution as was previously employed was used in j the analyais. The power peaking factors are as follws:
)
i Power Peaking Factor 7x7 8x8 r
) Radial 1.47 4 1.47 '
A Axial 1.57 1.57 i
' Local 1.24 1.22
. L Table 4-1 presents a tabulation of significant thermal-hydraulic character-l istics calculated for the identified cases. The results shw that. Arrespective I of the number of 8x8 fuel assemblies loaded in the core, both the 7x7 and 8x8 l
] fuel assemblies receive adequate coolant f1w. The margin to CHF for the limit- !
i
- i. ing assembly in an 8x8 core, or in a mixed 7x7-8x8 core, is always equal to or j
greater than the margin to CHF for the limiting assenbly in a 7x7 core. Further-
, l f more, due to the increased heat transfer area and correspondingly lwer operating i *. f heat flux of the 8x8 assembly relative to the 7x7 assembly, the 8x8 fuel has 1
greater margin to CHF than does the 7x7 fuel. ';
4 i .
I i
! 4-6 1 1
. - - - , -ms,,,. ,,, , , , . - . ,-..,.,v,-..r.,_, .w-w. ,m._.www--~-~~ m ,,v-,,4-.,,,w-,,w.m--,,-w.,- , , , - , ~
i t
f .
Table 4-1 l AssVLTS OF 11tIRMAIAffDBAULIC ANALYSB8 I !
Case Nue er i 2 3 4 5 I
- Core Average Void '
i Fraction, 2 27.4 27.4 27.4 27.5 27.5 ,
J Core Frassure Drop, poi 17.9 17.9 18.2 18.4 18.9 J
Water Rod Flow, 1 of
} Total Core Flow W/A 0.0008 0.09 0.17 0.35
! Asse e ly Type 7x7 7x7 8x8 7m7 8x8 7x7 8m8 8x8 l Number 484 483 1 363 121 242 242 484 Ret Channel Coolant Flow,103 lb/h 109 109 102 111 103 112 104 106 i
Bot Channel MCKFR 2.01 2.01 2.27 2.03 2.30 2.05 2.31 2.36
! Case Descriptions Case 1 Full core loading (484 assemblies) of 7m7 fuel.
Case 2: Same se case 1 with one 7x7 assembly replaced with an 8s8 aseeely.
Case 3: one-quarter core load of 8x8 reload assemblies with the remainder 7x7 assemblies.
Case 4: One-half core load of Sat reload assemblies with the remainder 7x7
, aseambitas. ,
l Case 5: Full core loading of Sm8 reload fuel.
I l
r l
e.
I l
l 1
i I
i l
J _ REFERENCES - SECTION 4
- 1. " Core F1w Distribution in a Modern Boiling Water Reactor as Measured in
]
Monticello," Licensing Topical Report, January 1971 (NEDO-10299).
- 2. Kim,11. T. and Smith, H. Sc , " Core Flow Distribution in a General Electric Boiling Water Reactor as Measured in Quad Citier Unit 1," Licensing Topical l
J Report, December 1972 (NEDo-10722) .
k
- 3. Martinelli, R. C. and Nelson, D. E., " Prediction of Pressure Drops during i Forced Convection Boiling of Water," ASME Trans., 70, pp. 695-702,1948.
- 4. Barooty, C. J., "A Systematic Correlation for Two-Phase Pressure Drop," Heat Transfer Conference (Los Angeles) AICilE, Preprint No. 37, 1966, i
- 5. Monti;ello Nuclear Generating Plant Safety Analysis Report, Docket No. 50-263.
- 6. Hinila, J. A. (General Electric Co.) letter to J. M. llendrie (USAEC), March 30, i 1973.
i 1
4 9 l
f 1
i 4-8 r-- -
. _ _ - - _ _ . _ _ . _ _ _ _ - = _- _-- _ _ __-_ - _ _. -. .. - - _ _
j 5. HUCLF.AR CHARACTERISTICS i
i
5.1 INTRODUCTION
The nuclear design of the 8x8 reload bundles described in this section has ,
been performed with the same analytical models and design methods used for General Electric 7x7 reload cores licensed by the USAEC over the past several 1
years. ,
No changes have been made in the analytical models or in the design methods. The 8x8 reload bundles will be loaded into the cores that have been closely followed by CE using these same analytical models and design methods.
A high degree of confidence can be expressed regarding the verification of GE nuclear models and methods for these plants. In addition, these same models and tethods have been routinely used for cores having lattices in the range from 6x6 to 11x11, and in coros with mixtures of either 6x6 and 7x7 or 8x8 and 9x9.
The 8x8 fuel beina licensed is well within the range of physical parameters of previous General Electric fuel designs, and no decrease in accuracy can be expected because of the change to an 8x8 fuel design.
l
- 5.2 BUNDI.E NUCLEAR DESCRIPTIONS The mechanical description and physical parameters of the 8x8 reload fuel have been given in Section 3.
This section describes the calculated nuclear parameters of the 8x8 reload bundles and makes comparisons to previously licensed 7x7 fuel designs.
4 There are few real " limits" on the bundle design itself. The real limits are generally expressed in terms of core parameters (e.g., shutdown margin or l maximum neat flux). The results of analyses involving core nuclear character-1stics are discussed in Section 5.5. The intent herein is to describe the nuclear parameters and to show that for 7x7 and 8x8 bundles of the same average enrichment, the calculated nuclear parameters are either not remarkab1;' differ-ent or are different in a manner that would be expected. The choice of some parameter (say hot reactivity) for reload fuel is dependent on the environment in which the reload fuel bundle will be used: that is, the reload bundle 5-1
i l i
t ;
1 I i< l.
l 4
f requirements would be slightly different for a very early hutdown than for an j outage following a period of operation beyond full power exposure capability. j j Generally, for reload fuel, the enrichments and rear +Ntles of the bundles will '
l be higher than for initial cores. Specifically, a much higher value of reac- !
! tivity is allwable for the lw exposure reload bundle than is allwable for the ,
initial core bundle (at the same low exposure) because the reload fuel bundle is I loaded into an environment of highly exposed bundles of generally 2Wer average reactivity. ,
l 5.2.1 2.62 wtt U-235 8x8 Bundle Desian I
i j 5.2.1.1 Reactivity d
I i
j Pigure 5-1 shows the hot average void reactivity of the 2.62 wt% U-235 bundle ;
) versus exposure and compares this bundle to a 7x7 bundle of 2.63 wt% U-235 enrich- !
j ment used at Oyster Creek as the first reload batch. The Oyster Creek R-1 bundle l
} has the same number of gadolinium-containing rods as the 8x8 bundle, but the r
i 3
average gadolinium concentration is about half that .of the 8x8. The result of !
this difference is that the gadolinium is worth less and burns out faster in the !
t j 7x7 bundle than in the 8x8, and thus the 7x7 imdle has a higher initial reac- ;
l tivity than the 8x8. Following the gadolinium burnout, hwever, the reactivities
}
1 of the bundles are essentially identical.
i
) Table 5-1 compares some physical parameters for these two bundles, while Table 5-2 comparea the aero exposure cold reactivities of the two bundles. As ;
) ,
can be seen in Figure 5-1, while the initial reactivities of the bundles differ !
l . because of dif ferent gadolinium worths, af ter the gadolinium is gone the reac-f tivities of the bundles are essentially identical.
e 1
{ 5.2.1.2 Void Reactivity l The variation of reactivity with void is of importance in the stability of-the reactor core while at normal power operation. Thre is no design criteria l , placed on the void coefficient except that the overall. void coefficient be nega-l tive at every point in the operating cycle. Overall void coefficients refer to ,
um 5-2 c -,- . ,=e--, --.y.m4-- m., --.,_...,m.34 _-_,-..--.-,,~,,,.w,,,,_.,.o.,y,..--w,,-.-,-,-.,,,,y,,44,c,--,,,,,,,,,m-,,, ,,y.,,-, - - - .,, ,-,.,n..-,.
. ~ - ~ -w a .e -----e aa g m.-- -n--g-sa.a - -,,-s.-+. -..-uus.~.u.sg~m1-.-a-,.-ea-e . -.---._----...-ns -
..z_-s s.s- ----as. . . - . --s-_ s i
1 4
1 4 1
i
)
R 1
1
$ i
~,
I ~
w '
4 E u O
e:
=
Ie 3 a
s a
- g I
i 1
{t E3 o5 -
o
- M E
a t
er2 .#
4 e
w U 4 a y
, 0 e:l a 4
1 30 i * >
i e
- ~
?
u
^;
G W n >
E a
> - . e
/ 3 t
(
A n
4
/ -
g ,
g i 1
( E
\
_ ~
s N * % * *m=
l l _ D5 ' l I I q e z 3 4 2 8 8 8
. a a - - - - -
0311GW ANO3NA "w 5-3
_ _ . . . - . _ _ _ - , . _ _ ,. . . . . _ - _ _ . . . - _ _ . - ~ . , _ _ . . _ . ~ _ - , _ , - . _ . _ . - - - . _ _ ,
Table 5-1 PHYSICAL PARAWTERS OF 2.63 7x7 MID 2.62 8x8 4
7x7 8x8 Pellet Outside Dianeter (in.) 0.487 0.416 Rod Outside Diameter (in.) 0.563 0.493
.- Rod-to-Rod Pitch (in.) 0.738 0.640 Water-Tuel Ratio (cold) 2.43 2.60 U Bundle Weight (pounds) 427.8 404.6 Cladding Thickness (mils) 32 34 Table 5-2 COLD REACTIVITY WMPARISON Zero Exposure Condition Controlled 7x7* 8x8**
k-infinity Cold No 1.163 1.166 l k-infinity Cold Yes 0.988 0.981 ok/k Control Strength 0.150 0.157 i
, , *2.63 wt% 7x7
- 2.62 wt% 8x8 i
i 5-4
the core response. It will be sufficient here to give results of infinite lattice
- calculation of reactivity versus in-channel void fraction rusd to show that the same behavior is aten for both 8x8 and 7x7 fuel. Figure 5-2 compares the void reactivity of the asse two bundles de',cribed above, and, se can t;e seen, the variation in reactivity with void is very close for both the controlled and
]
uncontrolled states at zero expreure. Again, note that the value of the 7x7 bundle is lower than that of th e 8x8 because an initially larger volume fraction
, of the bundle contains gado11niue. Figure 5-3 compares the lattice ak= going f from 0.40 void to other voids as a function of exposure. As can be seen, the -
l void reactivity characteristics are very similar, s
5.2.1.3 Doppler Reactivity The Doppler coefficient is of prime importance in reactor safety. The Doppler coefficient is a measure of the reactivity changa, associated with an increase in the absorption-of-resonance-energy neutrons caused by a change in the temperature of the material in question. The Doppler reactivity coef ficient provides instantaneous negative reactivity feedback to any rise in fuel tempera-ture, on either a gross or local basis. The magnitude of the Doppler coef fi-cient is inherent in the fuel design and does not vary significantly among Bk'R reactor designs having low fuel enrichment. For most structural and moderator materials this ef f ect is not significant, but in U-238 and Pu-240 an increase in temperature produces a comparatively large increase in the absorption crcss sec-
. tion. The resulting nonfission absorption of neutrons causes a significant loss in reactivity. In BWR fuel, in which approximately 98% of the uranium in the UO 2 is U-238, the Doppler coefficient provides an innediate reactivity response that opposes fuel fission rate changes.
Although the reactivity change caused by the Doppler effect is small com- g pared to other power-related reactivity changes during normal operation, it becomes very important during postulated rapid power excursions in which large fuel temperature changes occur. The most severe power excursions are those i associated with rapid removal of control rods. A local Doppler feedback asso-ciated with a 3000 to 5000'F temperature rise is available for terminating the initial burst.
5-5
l l
0 92
, 1.13
- 0 91 1.17 - l 1
1 0 to 1.16
?XF UNCONTROLLE D 1
i - 0.99 i t 16 -
i 0 to 1.14
,j 1r -
0 87 1 13 ,
o i 8 Y
$ O" O
E U 1.12 -
0 96 6 ex8 g
" UNCONTHOLLED 0 t
w -
1.11 - 0 86 7X7 CONTROLLE D 1
1,10 -
0M 4x8 CONT ROLL E D 1.00 -
0 83 i
l '
l os - c.s2 i
1 07 -
0.81 e
I I I I I I I , , ,
040 0 50 0 60 0.70 0 80 0 0.10 0 70 0.30 2
IN CHANNEL VOt0 FR ACTION Figure 5-2. 2.63 7x7, 2.62 8x8 Infinite Lattice K.
versus In-Channel Void Fraction Zero Exposure 5-6 e9
,a,.-.-- - n. , - ,
1 op i
i c as -
.f I os v ia i j o ca -
7 20% V
.l _
I c ot -
'1 A_
sT
.i O l h' w
~
4 0 o1 e 7x7 A ext o or 4
{
0 03 , ,
_ 70s v a
4 e
I I o e to ts EXPOSURE (GWdhi Figure 5-3. AK Void Comparison 7x7 versus 8x8 from 0.40 Void to other Voids l 5-7 l
i
-n. - _ . , , _ . ,, .. . . . _ , . - _ _ . . _ . _ . - , --
-- . . _ ~ _- . ._ . . . _ - . . - . . - - - - ._ __ _ . _ - - . _ - - - - _ - . . _ - --
1 1
The Doppler reactivity decrement is derived directly from the lattice cal-culations which are performed to generate the nuclear constants. The lattice methods currently being employed in the f ast and resonance-neutron-energy regions I
are based on the method of Adler, Hinman and Nordheim with the inclusion of
) the intermediate resonance approximation. This provides an adequate calculation i of both the spatial and energy self-shielding for the resonance absorbers that j .. explicitly includes temperature, moderator density, and geometry ef fects. A fine i group B-1 slowing-down calculation of the f ast and epithermal neutron spectrum provides the proper weighting of the resonance absorptf or. to yield eff ective i resonance integrals or cross sections that accurately represent the BWR I
environment.
1
-i The Doppler decrement is determined by doing the lattice calculations at I
several fuel temperatures holding all other input parameters constant. This 1
results in a change in the neutron multiplication factor which is solely due to a change in the fuel temperature, which is the Doppler effect. From these analy-
] ses it has been determined that the Dopple r defect, Akg ,p can be reprc cnted j very accurately by the followi.ng expressions Ak
- DOP = CDOP (( - %) .
l Therefore, the toppler reactivity decrement increases proportionally with the square root of fuel temperature. T, and CDOP is the constant of proportion-ality. The Doppler reactivity coefficient is derived using the same techniques described above. The following equation is used to calculate the Doppler reac-tivity coefficient
,1_ dk ,
CDOP k dT kg + CDOP
- 2 2
Figures 5-4 and 5-5 compare the Doppler coef ficients for two 7x7 fuel designs
. o the Doppler coef ficients for the 2.50 wt% U-235 8x8 bundle at 200 HJd/t and at 10,000 mwd /t, respectively. The Doppler coef ficient of the 8x8 2.62 is essen-tially identical to that of the 8x8 2.50 bundle.
i 5-8 i
I
_ _ _ _ __.- _ _ _ ._... _ _.- _.,__, . . .._y,_ .
I i l
1 i
i r
1 1
i l
i 1
- O 4 voso a
1 ""
i s
/ ,,, f
! 40%
i p
,1 06 -
./
f /p# ..
70%
/ #
/ /
3
/
/ /
i
/
/ /
i / /
/
1
' o8 -
/
/ /
- / /
i
%. /
1 /
/ / ,
1 2
.i o - /
3
/ - - - O 7x7iwitAL I / A 7x7 attoAD
/ O exe l / /
- / /
! c' /
- .a -
4 j
)
- i .4 l l 0 M 2000 3000 4000-1 T E h4PE R ATURE (*F) l l
Figure 5-4.- 200 mwd /t Doppler Coefficients Uncontrolled l
l 5-9
l I
1 I
i 1
I
. VOIO 0%
,/
i 06 -
- l 1 / 40%
/ /
. 70%
A
] # #
/
l -c e - #
1
/
/ /
l / /
/
~, C
- . ,0 _ /'
{-
/ /
1 5 /
^
/
7 /
! = = -0 7x7 :Niis AL
- 1.2 -
[ O 7X7 RELOAD
/ O SX8
. /
14 -
1 i .
4 ,
'THE 8x0 VALUES ARE AT 11.200 mwd /t
, D
.i. l l l l 0 1000 2000 3000 4000 TEMPE RATURE ton Figure 5-5. 10,000 Wd/t Doppler Coef ficients Uncontrolled 5-10
,vr 7 V - ' " " ' ' " '
l l
l i
4 It should be understood that the data presented in these figures are for an infinite lattice. In a finite reactor system the power J1stribution, and hence fuel temperature distribution, will vary spatially. This in turn results in a spatial variation in the Doppler feedback with larger Doppler reactivity decre-ments occurring in the high temperature and thus in high neutron flux regions of the reactor core. Therefore, high Doppler reactivity feedback can occur for ,
) relatively low core average power increases since the larger Doppler reactivity oecrements will occur in the high flux, or importance weighting, regions of the I core. Results of core calculations are reported in Table 5-3.
1 5.2.1.4 Delayed Neutron Fraction Given in Figure 5-6 is a comparison of the delayed neutror fraction for the 2.62 8x8 bundle and the 2.63 7x7 bundle at hot average void conlitions. As can be seen the dif ferences are negligible.
5.1.1.5 Peaking Factors i
The calculated waximum local peaking factors at average void for the 7x7 2.63 and 8x8 2.62 wt1 U-235 bundles are compared in Figure 5-7. Of more impor-tance to the reactor operator is the increased heat transfer area of the 8x8 bundle, leading to much lower peak kW/f t, as noted in Sectior '
5.3 ANALYTICAL METHODS The analytical methods and nuclear data used to determine the nuclear chat-acteristics are similar to those used throughout the industry for water-moderated -
sys tems.2 The Lattice Physics Model is used to generate few-group-neutron cross sec-tions for use in calculating lattice reactivities, relative fuel rod powers within assemblies, and averaged few-group cross sections. These cross sections and reactivities are calculated at various void and exposure conditions and are used for calculatir.g two- and three-dimensional reactor power distributions.
Local fuel rod powers are calculated for an extensive combination of parameters 5-11
---_m ._. m__.- a 4*.aw.m4 4--%4u,4u sm__s- - - - ~ &a m.- . - _ -
4 G
S N
8 a
8
=
4
- o
- to s
G te k
) 0
/ <
/ ~
N k
u
/ n e
a
/
f Y
$ E O
CD NW f U E
5 4
a
/. o
~ u e &
O G
.I N w i = D N
8 p
w a
y
. l l I . @
e h l
a 8
e
- G ks N U 60 Ene 5-12
n
-a~-- ->.+ --..n-- ----_ a.aa,---.sm+ s -_ m ,2 __ _ . _ _ . - - ---..._.-+__a-s . -m - aw ,
1 i
l 1 l t
4 1 A 1
/
/
/
r y
1 e
Y h O
t N .
g 4
/,
/ o 4 O 1
n 4
' 3
- e J
lr l' f
~
/
/ S*/ 4 ~
J i
/
A
- N i 40
- 4 2
4 .
, C l .
. .. / / / / /
N k- $ N E N k- k-MV 3d 1VD01 WOWlXYW I
l l 5-13
including fuel and moderator temperatures, burnup, steam voids, and the presence or absence of adjacent control rode. These few-group calculations are performed over either single-bundle cells or groups of four bundles characteristic of repeating arrays in the loaded reactor core. The fast and resonance-energy cross sections are computed by GAM-type program. ' The fast energieb ?** 'reated
.- by multigroup, interral collision probabilities to account for geometrical effeuts in fast fission. Resonance cross sections are computed using the intermediate resonance approximation, and the epithernal spectrum is obtained from a B-1 multigroup solution. Account is taken of position and energy-dependent Dancof f factors. Changes due to concentration self-shielding and spectral effects of isotopic composition are recomputed as a function of fuel exposure. THERHOS-type calculations are used to determine the spatially varying thornal spec-trum throughout the fuel bundle. The ef fects of control blades on the cross sec-tions of adjacent materials are calculated and accounted for. Power and flux distributions, infinite multiplication f actors, and material and flux-weighted cross sections are calculated using twc-dimensional, few-group dif fusion theory on fuel assemblies and arrays of fuel assemblies. Burnup calculations are per-formed by integrating the secular equations describing the fuel depletion process with spatial neutron flux and energy distributions typical of reactor operat.ing conditions. At selected burnup intervals, the nuclide concentrations are used to recalculate revised cross sections with the lattice model, and these are again recycled through two-dimensional dif fusion theory.
A large three-dimensional boiling water reactor simulation code providing for representation and calculation of spatially varying voids, control rods, burnable poisons, and other variables is used to compute power distributions, exposure, and reactor thermal-hydra"lic characteristics at the beginning of core life and as burnup progresses. Gado11nia is distributed in a few rods within each fuel assembly for supplementary control. This feature makes it necessary to compute the radial space-time dependence of the cd-155 and Gd-157 concentra-tions within the fuel rods. Experimentally, verification of the calculated
. reactivity ef fect as well as the calculated removal rate of the high cross-section isotopes has been accomplished. Observation of the operating control rod pattern during full power operation has shown the renoval of the gadolinia control to be well matched to the fissionable isotope removal. The effective rate of depletion can be monitored by observing the operating reactivity status.
~
5-14
= ' ' = = m - . - . - _ _ _ ~ _ _ _ _ _ _ _ _ _ _
Thus, any trer.d toward an unacceptably small shutdown margin caused by faster-then-anticipated absorber removal could be detected and remedial action applied before any unsafe condition cuuld be created. Any tendency toward slower removal rates would affect only cycle length and would be an economic problem unrelated to esfety.
.i Operating reactor and critical experiments compared to theoretical data pro-vide the precision necessary for reactor design. 'O The reactivity calculation of these analytical methods is frequently compared to the actual performance of operating .eactors. Specific comparisons have been made for the Oyster Creek and Dresden 2 plante. The results of these comparisons show that the calculated and actual results agree within experimental and manufacturing tolerances. The design methods have been shown to be able to compute local powers to within 13%,
fuel assembly segment powers to within 210%, Pu-U ratios versus exposure to within 13%, and core reactivities and cold shutdown margin to within 0.005 ok.
Experimental tests have also been used to verify the analytical calculations of both reactivity and isotopic composition for lattices in the range from 6x6 to 8x8. These tests give results nearly identical to the comparisons with the operating plants. The most recent experimental comparison is documented in Reference 9.
5.4 EXPERIENCE WITH CE NUC1. EAR MODELS The analytical methods described in Section 5.3 have been used by General Electric to design and follow cores having lattices in the range from 6x6 to 11x11 aside from the normal 7x7 reload cores. Of special interest in this ,
regard are the Humboldt Bay and the Garigliano reactors. These cores are oper-ating with nixed lattices and have operated successfully for some time. In the .'
case of Humboldt, the core has operated since July of 1969 with a mixture of 6x6 and 7x7 reload fuel bundles in the core. This mixed lattice reload core has been licensed by the USAEC following General Electric analysis using the same analytical methods described above. Also of note in this regard is the Gariga 11ano reactor which has operated since October of 19As with a mixture of 8x0 and 9x9 fuel bundles. This reload core has been licensed by a regulatory agency 5-15
- - _ . . . - - . _ ~ - -
_ _ _= -- -- . - _ _ _ . --
comparable to the USAEC following uvneral Elaetric maalysis. All nucleat license submittal information supplied by General Electric for the pant neveral years has been developed using these same well proven analytical methods. There has been adequate experimental and operational verification of these methods to lattice designs of other than 7x7 fuel. No decrease in accuracy can be expected because of the change to an 8x8 fuel design.
5.5 NUCLFAR CHARACTERISTICS OF TliE CORE Earlier sections have discussed the infinite lattice steady-state reactivi-ties and teactivity coef ficients of the new 8x8 reload fuel bundles and have made comparisons to previously used 7x7 reload bundles. This section discusses the results of core calculations on shutdown margin (including the liquid poison system) and core average reactivity coef ficients.
5.5.1 Core Ef fective Multiplication. Control System Worth and Reactivity Coefficients A tabulation of the typical nuclear characteristics of the pre- and post-outage cores is provided in Table 5-3. Because the nuclect characteristics of the reload fuel are close to those of the initial fuel, the temperature and void dependent characteristics of the reload core will not differ significantly from the values previously reported.
5.5.2 Reactor Shutdown Margin TI.e refueled core fully meets criteria established for the initial core in that it may be maintained suberitical in the most reactive condition throughout the subsequent operating cycle with the most reactive control rod in its full-out position and all other rods fully inserted. The shutdown margin at BOC3 is 0.010, anC with an R value of 0.00 in Cycle 3, this shutdown margin is the mini-mum Cycle 3 value.
d i
1 5 *6 w
. . . - - _ . , + - -
b Table 5-3 r I
NUCLEAR QlARACIERISITCS OF THE DESIGN REFERENCE CORE g Core-Effective Multiplication e and Control System Worth Pre-Outage Reload (0% Voids. 20'C) Core Core ;
K,gg Uncontrolled 1.093 1.122 Ak Poison Curtains 0.008 --
E
! Ak Control Rods 0.153 0.163 K,gg Fully controlled 0.932 0.959 K,gg Strongest Rod Out 0.946 0.990 Increase in Core Reactivity with Exposure into Cycle <0.0 <0.0 Reactivity Coefficients, Range During Operating Cycle for Reload Core Steam Void Coefficient at 40% Voids (1/k)(Ak/AV), 1/% Void -10.1 x 10 -10.3 x 10 to
-10.0 x 10
Power Coef ficient at 16'O )Mt and 524 Btu /lb Inlet Enthallyg
( Ai/k) /( AP/P) -0.051* -0.062 to
-0.053 Fuel Temperature Coef ficient at 650'C3 (1/k)( Ak/AT),1/'F Fue) -1.17 x 10
-5 -1.06 x 10 -5 t,
-1.17 x 10 -0
~
- Reference power level about 1400 )Mt at EOC2.
5-17
3 3
5.5.3 Liquid Poison System 1
1 j The liquid poison system is designed to provide the capability to bring the
} :cactor from full power (1670) to a cold xenon-free shutdown condition (K,gg
, <0.97) assuming none of the cratrol rods can be inserted. 'the requirements of j ,
, this system are dependent pri:narily on the reactor power level and on the reac-l tivity ef fects of voids and temperature between full-power and the cold. xenon-
{ free condition. The liquid poison system has been examined and has been found j to be adequate since the reference power level of 1670 Wt has not changed and j the core reactivity effects of voids and temperature have not been significantly
] altered by the introduction of reload fuel.
I f 5.5.4 Reactivity >l Fuel in Storage 1
J There is no new safety implication with the spent fuel storage pool and the j new fuel storage rack configuration, because the K= of t' e reload fuel is less j than the K= of the initial fuel asseablios without temporary control curtains, i
I i
i
) .
i e
l 1
l i
I
- 5-111 1
n E p i+ w A ..
s REFERENCES - SECTION 5 1
- 1. Carter, J. L. , Jr. , " Computer Code Abstracts, Computer Code-HRG " Reactor Physics Dept., Technical Activities Quarterly Report: .uly. August, Septem-ber 196G October 15, 1966 (BNWL-340).
- 2. Chernick, J., " Status of Reactor-Physics Calculations for U.S. Pruer Reac-tors " % actor Technology, 3
_1_3,, 4, (Winter 1970-1971).
- 3. Wilcox, T. P. and Perkins, S. T. , "AGN-CAM, an IBM 7090 Codo to Calculate Spectra and Multigroup Constants," April 1965 (AGN-TH-407).
- 4. Carter, J. L., "HRG3 A Code for Ca?.culating the Slowing Down Spectrum in the P3 or B Approximatior.," October 15, 1966 (BNWL-340).
2
- 5. Honeck, H. C., " THERMOS - A Thermalization Transport Theory Ccde for Reactor Design," June 1961 (BNL-5826) .
- 6. Crowther, R. L. , Petrick, W. J. , and Weits. berg, G. A. , "Three Dimensional BWR Simulation," ANS National Topical Meeting, April 1969.
- 7. Fuller, E. D., " Physics of Operating Boiling Water Reactors," Nuclear Appli-cations and Technology. H. November 1969.
- 8. Aline, P. G., et al. , "The Physics of Non-Uniform BWR Lattice," BNES Inter-National Conference on the Physics Problems in Thermal Reactor Design, June 1967. -
- 9. "Contain=d Burnable Neutron Absorber as Supplementary Control," Quad Cities . ;
Units 1 and 2 FSAR, Ammendment 9.
W 5-19
. - _ _ _ - - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ l
- 6. SAFETY ANALYSES 6.1 HDDEL APPLICABILITY TO 8x8 FUEL This section provides information on the applicability (to the 8x8 design) of existing models used f or safety analysis. Where changes in fuel design affect model applicability, the capacity of the models to accommodate these changes is discussed.
6.1.1 Control Rod Drop Accident (RDA)
The postulated sequence of events for this the worst case accident involves an abnormally high worth rod becoming disconnected from its drive, being stuck in the fully inserted position, the drive being withdrawn and the control rod falling out of the core to the rod drive position. Analysis of this accident is performed at various reactor operating states; the key reactivity feedback mechanism af f ecting the shutdown of the initial prompt power burst is the Doppler coefficient. Final shutdown is achieved by scramming all but the dropped rod. The methods utilized to evaluate the rod drop accident have been updated on a continuing basis to reflect improvements in analytical capability. '
The change from a 7x7 to an 8x8 fuel lattice has no effect on the excursion model used in the analysis of the RDA or on the reactivity feedback ef fect due to Doppler which is used in the analysis. The number of fuel pins failed due to the RDA is dependent on the fuel pin (local) power peaking factors in the bundle and final peak fuel entbalpy in the core. The local peaking factors and the peak fuel enthalpy are inherently known for an 8x8 lattice, the local peaking f actors from the lattice design calculations,I and the peak fuel enthalpy from the RDA 1
analysis.
Homogenized bundle cross sections and nuclear constants are calculated using etandard lattice design techniques as noted in Section S. Since the bundle cross sections, which are produced from the lattice calculations and which are used in the RDA excursion model, are homogenized, the RDA excursion model does not recog-nize the lattice type used to produce the bundle cross sections.
6-1
___ . _ _ ~ _ _ _ _ . _ _ _ -
I l
1 4
i 5 l 1
) I j A mixture of 7x7 and 8x8 fuel bundles in a reloaded core present no analyti- l 1
cal problem. The homogeniced cross sections and nuclear constants used to repre-sent each fuel bundle in the RDA analysis are calculated using methods which have previously been used for lattice designs from 6x6 to 11x11 geometry and in cores j with mixtures of either 6x6 and 7x7 or 8x8 and 9x9 -(refer to Section 5) . Local j power peaking at RDA conditions is explicitly calculated, j -. :
] 6.1.2 Loss-of-Coolant Accident (LOCA)
] The Emergency Core Cooling System models which are used for the LOCA analy-
! sia for 8x8 fuel are essentially those which have been prcviously used for the
}
7x7 fuel designs. They are described and exemplified in Reference 5. The spe-
- cific models as applied to the 8xS fuel design will be discussed in the follow-i ing paragraphs in their order of presentation in Referenca 5.
j i
j 6.1.2.1 Short-Term Thermal-Hydraulic Model I '
j The significant parameters used by the short-term thermal-hydraulic model l will remain essentially the same in changing fuel designs. The exceptions arer ;
i i
l 1. Core pressure drop - the total core pre-transient pressure drop for a 1
1 1
full 8x8 core is %1 pai higher than for a full 7x7 core. Since maxi-7 d
mization of the core pressure drop is conservative, a partial 7x7/ par-4 tial 8x8 core is assumed to be fully 8x8.
i l
j 2. Core heat flux - the core heat flux versus time is consistent with 8x8 fuel operating LHGR and stored energy as well as 8x8 geometry. As in '
l the case of core pressure , drop the effect is small. The most siF nifi-cant parameters, the core thernal power, the maximum steam flow, and ,
the recirculation flow, remain unchanged with this change in fuel design.
1 The changes listed aboie result in only a small change in core flow and l pressure responses.
l 1
4 h
l 6- 2
, , , , . - ,, ,.w. , , , , , . -y~y--, _r.-,,21--.-,, , ., , -. , ... s., , . , ,,- - - ,,p.. .w,e,., , , - - r i-s- .,y w ,p,,-.,me.y%.,.,. , - -.,y-, a#wc--, ,ee-
i j
l 1
)
l I
j 6.1.2.2 Long-Term Thermal-Hydraulic Model 1
The only significant chanEe to the long-term thermal-hydraulic model is the change in bundle geometry and therefore a small change in the core total hydraulic s
diameter of the core. The long-term thermal-hydraulic model has the capacity to l model various geometries; therefore, such smail changes resulting from the change ,
j in fuel design do not represent an " extrapolation" in the model. The important parameters, e.g. , core power, steam flow, recirculation flow and basic reactor
~
. geometry, remain unchanged.
- 1 l
b.1.2.3 Transient critical Heat Flux Model 4
The transient critical heat flux model will change only in that the bundle ,
I geometry and IRGR will change. Test data taken in the new ATLAS loop with full
~
power Ax8 bundles is being provided to the AEC by General Elcetric to verify the applicability of the existing model. If modification of the model is
~
required, it will be made based on the results of these extensive tests.
4 6.1.2.4 Core Heatup Model The core heatup model used for 8x8 analyses is essentially that described in Reference 5 with the incorporation of the modifications described in Refer-ence 6 and the obvious change to the 8x8 bundle geometry and LHGR. The model l has been used to predict the results of a number of ECCS transient tests of a full scale stainless steel clad 8x8 heater red bundle. These tests fully con-firm the applicability of the Core Heatup Model as modified for 8x8 fuel. Full l- scale ECCS tests with pressurized Zircaloy heaters were conducted in October of 1973 for further demonstration of the applicability of the Core Heatup Model.
l I
i 6.1.2.5 Total LOCA Analysis
- The total LOCA analysis which includes the four above models will not change in procedure. The only changes in the results will be due to changes in fuel geometry and linear heat generation ra te, which are handled by the existing J
i i.
6-3 d
. . , . . - - . - - . - - ,.r--- -., , . - , , - . , - . -,n- , , , - - - , . ,- ~ .- - ~ , e - - - , , n--*-
l models without modification with the possible exceptions noted in 6.1.2.3. Cen-I eral Electric is presently discussing the applicability of current LOCA models for licensing 8x8 fuel with the AEC, and confirmation is expected shortly.
1 6.1.3 Transient Anelysis and Core Dynamics A complete range of single f ailure caused events which are abnormal but
- reasonably expected during the life of the plant were analyzed for 7x7 fuel as r.
l part of the original plant licensing. Results from these analyses were included in the FSAR and subsequently reviewed for 7x7 reload fuel. A reanalysis of l these events has been carried out incorporating 8x8 reload fuel and planned i
modifications to the reactor pressure relief system. The purpose of this section is to demonstrate the applicability of the current analytical models to 8x8 fuel and mixed core analysis.
6.1.3.1 Transient Analysis Model Applicability to 8x8 Fuel The documentation of transient analysis methods for General Electric BWRs is provided in Reference 7. This document includes not only the equations of a the transient model, but aleo 6 parameter study and comparison of rafety analy-
. ses applying the model to plant startup data. The mathematical model described in Reference 7 is applied to both new and reloaded cores. The model as presently constituted is a " lumped" thermodynamic model with single bundle representations
- for average and hot channels. The neutron kinetics representation is a poin' reactor using the point reactor kinetics equations. This brief model review serves as a basis to point out that the model, which 1 very generally defined,
- does not change or lose validity due to a mixture of 7x7 and 8x8 or a total core of 8x8 fuel. Parts of the model lump or average system components for computa-i tion purposes. These model parts, such as thermodynamic regions or neutron kinetics, are affected by simple input parameter changes due to fuel changes. ,-
The most affected part of the model is the actual fuel heat transfer model.
There are several objectives in transient analysis which affect the fuel model.
Briefly these can be broken down as: (1) computation of fuel thermal margins, i
6-4 i
. . _ _ _ _ _ _ _ _ _ _ _ _ _ _ . ._ _ _ _ _ _ _ _ _ _ = _ . _
t i
e a
(2) conservative heat flux computation for the system transients, and (3) compu-
- tation of everage fuel temperature for Doppler. If the systou contains a com-plate load of either 7x7 or 8x8 fuel, the fuel model input is straightforward because the entire core is represented by the same fuel parameters. In the case
- of mixed 7x7 and 8x8 core fuel loading, the average core thermal calculations i are not completely characterized by either the 7x7 or 8x8 fuel type. The mixed '
j fuel loading can be adapted conservatively to the model however. This is achieved I by doing three things for input to the dynamic models (1) using the fuel type conservative for the fuel thermal margin as the hot channel fuel type in the j transient analyses; (2) since mixed load core dynamic performance is the aver-age of 7x7 and 8x8 fuel, choosing the conservative fuel type for plant tran-sients to yield overall system conservative results in the dynamic analysis; and
- (3) using Doppler coefficient input data which is conservative to the overall
} core design when coup),d with the conservative fuel design of (2) above. The i use of the dynamic model as outlined above vill allow a totally conservative dynamic analysis for any fuel looding.
1 J
- 6.1.4 Rod Withdrawal Error (RWE) i j The rod withdrawal error reactivity insertion event is normally included in j the Transient Analysis portion of reload fuel safety analysis submittals. How-i ever, since the event is analyzed by methods other dian the transient mathemati-cal models referred to in 6.1.3, the model applicability of analysis of this event to 8x8 fuel or mixed cores is discussed separately.
I i
t Analysis of the rod withdrawal error is performed on the assumption that the maximum worth rod is fully inserted and adjacent rods are withdrawn in a j ,
manner which will allow full design reactor power with operating limits attained
, near the inserted rod. This is an abnormal rod pattern which is not normally
{' employed, but it maximizes the rod worth of the inserted rod for purposes of the l conservative analysis. The maximum worth rod is than inadvertently withdrawn until rod block occurs, initially assuming the worst allowable LPEH bypass condi-tions. The results depend primarily on the capability of the flux monitors to
- detect the local change in the fuel around the control rod as it is withdrawn j ,
and to stop the control rod before damage limit conditions occur.
i t
6-5 i
,__m .. . . ..-_ , _ , _ _ . _ , _ , _ _ , _ _ _ , , . . , . , , . . ,,,_,.,,,,,.m. ,,,,._e 7 , .. ~ .s , ,.vy,,-
1 It should be noted that there are two rod block systems currently in use in GE BWRs. The first is described in Reference 8 and is employed in Oyster Creek Unit 1 and Nine Mile Point Unit 1. All other GE BWRs utilize the rod block sys-i ten described in Reference 9.
The Oyster Creek 1/Nine Mile Point 1 system uses the APRMs on a quadrant b.cis 0and the other system uscs the LPRM strings surrounding the control rod .
boing withdrawn. In both cases the sensors in the system are reading neutron j
flux. Also in both cases, analysis of the transient is performed assuming worst ,
! case allowable LPRM bypass conditions.
1 The total analysis of the RWE transient utilizes the three-dimensional coupled nuclear-thermal-hydraulie representation of the core as described in Ssetion 5.2 for determination of neutron flux levels at instrumented locations end for determination of fuel assembly flow rate. The responses by instruments .
to changes in flux levels is independent of the fuel type.
6.2 RESULTS OF SAFETY ANALYSIS 6.2.1 Core Safety Analyses Use of the Hench-Levy correlation to determine the safety limit and to estab-lish margina from the normal operating points to the safety limit was established in previous licensing submittals. The same considerations, margins, and damage limits described in detail before, have been applied in evaluating the reloaded core. The operating limit on LHGR for the reload fuel is lower than previously loaded fuel. A further discussion of these controlling factors in the core .
ccfety analyses is presented below.
6.2.1.1 Fuel Damage Limits Fuel damage from perforation of the cladding and a subsequent release of fission products can result from overheating or excessive strain of the claddf us. ~
The former is assumed to occur when MCHFR reaches 1.0 based on the Hench-Levy correlation and the latter is assumed to occur when MLHGR reaches 25.4 kW/f t l
I 6-6 l
l i
i (see Section 3). The mechanical design of the reload fuel is to the same design i criteria and bases as the initial core fuel and the same damage limits are applicable.
i 6.2.1.2 Operating Limits b .-
l The R-2 bundlco are designed to operate with the same MCHFR limit as the iro radiamd fuel in the same environment and with a lower MLHCR. That is, MCHFRs
- will be greater than 1.9 and MLHCRs will not be greater than 13.4 kW/ft. The I limiting values of HCHFR for the reload fuel during normal operation are the same as for the initial core fuel based on the similar design conditions a
established for the fuels. The limiting value of MLHGR is lower for the reload fuel since it has 63 rods instead of 49 rods in the initial core fuel with
! approximate'ly the same bundle power.
l 6.2.1.3 Operating Margins With the previously given damage limits and design limits for the reload i
{
3 fuel, operating margins between the two limits for the reload fuel are expected to be greater than the previously loaded' fuel. However, this is based on the .
maximum design condition. Actual reload fuel operating conditions of MCHFR and MLHGR are expected to be well below the design limits as has been the experience
! for previously loaded fuel. Thus, actual operating margins will continue to be greater than the minimum allowable 'ralues used in the analyses discussed below.
(!
6.2.1.4 Abnormal Conditions j The minimum allowable operating margins described above are consetvatively l used in analyses of events such as abnormal operational transients and uncer-tainties concerning steady-state fuel operating conditions. Since these margins are not reduced with the reload fuel, the results of these analyses are not expected to change appreciably with the insertion of reload fuel except where dynamic changes are occurring on the reactor and its characteristics have been changed by the reload fuel in such a way as to significancly affect the tran-I sient results. These considerations involve the transient analyses which are
[ ~~~
covered separately below.
l l 6-7 I .
_ , , , . _ _ . - , . _ . _ . _ _ _ , . . . ~ , _ ._,mm...m, . , _-- . . _ - , - - , . _ _ ,
6,2.2 Accident Analysen 6.2.2.1 Main Steam Line Break Accident i The analysis of the main steam line break accident depends on the operating th2rms1-hydraulic parameters of the overall reactor: such as the pressure, and tho overall factors affecting the consequences, such as primary coolant activity. ,
Insertion of 8x0 reload fuel will not change any of these parameters so the
- previously reviewed results of this analysis will not change.
6.2.2.2 Refueling Accident The analysis of the refueling accident depends on mechanical damage caused by a fuel bundle falling back onto the top of the core while it is being removed, which will not change with the use of the reload fuel. The consequences depend on the fission product inventory in the fuel and various factors affecting the caount and kind of releases to the atmosphere. The fission product inventory in not expected to increase even with the large number of rods in the 8x8 reload fu]3 bundle. Thus, even if more rods were damaged, the total fission product inventory is not increased, but there will be slight changes in the relative caounts of different constituents because of the slight differences in enrichment and gadolinia concentration. The effects of these small differences will be inconsequential in terms of the releases caused, and undetectable when the various reduction factors are applied to determine offsite consequences. Therefore, the l provously reviewed results of this accident analysis will not change.
6.2.2.3 Control Red Drop Accident l
6.2.2.3.1 Identification of Causes There are many ways of inserting reactivity into a boiling water reactor.
However, most of them result in a relatively slow rate of reactivity insertion cud therefore pose no threat to the system. It is possible, however, that l
6-8
\
I l
i i
i rapid removal of a high worth control rod could result in a potentially sig- ,
! nificant excursion. Therefore, the accident which has been chosen to encom-a d
pass the consequences of a reactivity excursion is the control Rei Drop Acci-i
, dent (RDA).
6.2.2.3.2 Starting Conditions and Assumptions t
Before the control rod drop accident la poasible, the following sequence of events must occurt
- 1. The complete rupture, breakage or disconnection of a fully inserted control rod drive from its cruciform control blade at or near the i coupling.
- 2. The sticktrg of the blade in the fully inserted position as thu rod 3 drive is withdrawn (worst case) .
4 3. The falling of the blade after the rod drive is fully withdrawn (worst case).
This unlikely set of circumstances makes possible the rapid removal of a centrol rod. The dropping of the rod results in a high local k ,in a small region of the core. For large, loosely coupled cores, this would result in a highly peaked power distribution end subsequent shutdown mechanisms. Significant shif ts in the spatial power generation would occur during the course of the excursion. Therefore, the method of analysis must be capable of accounting for i any possible ef fects of the power dis tribution shif ts.
I In order to limit the worth of the rod which could be dropped, the rod worth minimizer system or a second operator controls the sequence of rod withdrawal.
This assures no movement of an out of sequence rod before the 50% rod density configuration is achieved and limits movement of rods to in-sequence segnents beyond the 50% rod density configuration during startups. The 50% rod density s
configuretion occurs during each reactor startup and corresponds to the condi-l tion in which 50% of the rods are fully inserted in dhe core and 50% are fully withdrawn, ss 6-9
I
> 6.2.2.3.3 Accident Description 1
4 The accident is defined ass i
- 1. Tha highest worth rod that can be developed at any time in core life
> under any operating condition drops from fully inserted position to fully withdrawn position (rod increments only beyond 50% rod density).
i .
l 2. The rod drops.
- 3. The scram is that defined in the technical specifications.
The detailed analysis of this accident is discussed in References 1, 2, 3 and 4. A continuing effort is being made in the area of analytical methods to assure that nuclear excursion calculations reflect the lates t "s tate-of-the-art."
The sequence of events and the approximate times of occurrence are as follows:
Approximate Event Elapsed Time (1) Reactor is at a control rod density pattern corresponding to maximum in-sequence rod worths (2) Rod worth minimizer or operators are functioning to restrict rod withdrawalm to in-sequence rods or rod increments. Maximum worth in-sequence control blade ,
becomes decoupled.
(3) Operator selects and withdrews the control rod drive of the decoupled maximum worth in-sequence rod to its fully with-drawn positica (rod increments only beyond 50% rod density).
(4) Blade sticks in the full inserted position.
6-10
.=_ ~ - - - . . . - .. - . - - . ~ . - - . - - - . - - . - _. - - -
< t r
i i Approximate
! Event Elapsed Time ,
i j (5) Blade becomes unstuck and drops at the maximum vtlocity j determined from experimental data (3.11 fps). O d
(6) Reactor goes prompt critical and initial power burst is i .
. terminated by the Doppler Reactivity Feedback. <1 sec
], (7) APRM 120% power signal scrams reactor.
i j (8) Scram terminates accident. <5 sec i 6.2.2.3.4 Identification of Operator Actions 1
The termination of this excursion is accomplished by automatic safety fea-1 tures or inherent shutdown mechanisms. Therefore, no operator action during the f excursion is required.
i
! 6.2.2.3.5 Analysis of Effects and Consequences 4
) 6.2.2.3.5.1 Methods, Assumptions and Conditions The methods, assumptions, and conditions for evaluating the excursion aspects of the control rod drop accident are described in detail in References 1, 2, 3, and 4.
. Reference (1) is the topical report on rod drop and is applicable to begin-
! - ning of life conditions for curtained cores. Reference (2) is the first supple-ment to Reference (1) and is applicable to beginning of life conditions for i -
Cado11nia cores. Ref erence (3) is the second supplement to Reference (1) and is applicable to exposed cores. Reference (4) is not a supplement to Reference (1),
however, the information contained therein is supplemental since it is a direct
. expansion of the described methods applied to a parametric study of worst cased l
variables resulting in a boundary approach to rod drop accident evaluation.
l 6-11 F
i
,_ _ . ._.. . _ _ _ = _ _ _ _ _ _ . . _ . _ . - . - . - - . - - , _ _. ,-__ _ . . . . m . . , , -- . . _ , , - . . , . . . . . . ~ , - . -
l 1
The technical bases which are presented in Reference (4) were used to verify that the result of a rod drop excursion in the reloaded core would not exceed the design criteria, as described below.
J Although there are many input parameters to the Rod Drop Accident Analysis, the resultant peak fuel enthalpy is most sensitive to three basic conditions.
These are: 1) Doppler reactivity feedback, 2) accident reactivity characteris-4 tics, and 3) scram reactivity feedback.
! If all other parameters remained unchanged, the rod drop excursion for exposed cores would be less severe than for initial cold clean cores under the same set of conditions since the Doppler reactivity feedback will be more nega-tive. This is due to the fact that Pu-240, which has a large negative Doppler a
ef fect, builds up with exposure. iigure 6-1 shows the comparison between the actual Doppler coef ficient and *,ne T-chnical Bases Doppler (Reference 4) coef ficient.
The accident reactivity characteristics have varying effects on the rod drop excursian results. These characteristics are accident reactivity shape, total control rod worth, local peaking factor, and the delayed neutron fraction. The total control rod reactivity worth (worth of the dropping rod) has a major eff ect on the accident results; this will not change substantially with the insertion of reload fuel. The local peaking factor and the delayed neutron frcction were inputs to the evaluation and were in the same range as those shown for actual plant experience in Ref erence(4) . A comparison of the calculated accident i
i reactivity shape function with the Technical Bases shape function is shown in Figure 6-2. Figures depicting hot startup conditions are not shown since l the maximum in-sequence rod for that condition is not as limiting as for cold -
startup conditions.
The scram reactivity feedback function has a significant ef fect on the results of a rod drop excursion. The scram reactivity feedbuck shape was evaluated based on Technical Specification scram times and shown to be above that used in establishing the boundary in Reference (4). (See Figure 6-3).
The evaluation of each of these parameters by comparison to the boundary values presented in Reference (4) shows that the maximum rod worth is not as great as the i
6-12 l
~ _ ._. -
l 4
q 4
e A
) -as -
1
) 8 1
. - 1.0 -
C i A l- /
~
/
j - 1.2 -
~
/ /
)
1
{ - 1.4 -
/
< n f f
2 1 8 l
)? - 1.6 -
TYPICAL OPE RATING PLANT DATA MONTICELLO B0C
~
MONTICELLO E0C
- 1.8 -
, A 0 GNIT, o% VOIDS, COLD
~
B o GM/T, os VOIDS. HOT C o GM/T,40% VOIDS. HOT l D 6 GM/T,40% VOIDS, HOT l E 16 GM/T,40% VOID 6, HOT i - 2.0 -
i
// .
-2.2 -
l
. l . l . I i I,i I , I , I ,
- o 400 ano 1200 1eco 2000 2400 2000 3200 l 1
, AVERAGE FUEL TEMPERATURE,(C'l i
Figure 6-1.
i Doppler Reactivity Coefficient versus Average Fuel Temperature as a Function of Exposure and Moderator Condition for R2 Fuel l
l l
6-13 l i
o.ois l
I
- note -
P g. i .2 i .s
'9 o.oi -
290 CAL /GM PEAK FUFL ENERGY CONTENT o.olo -
1 2
E
_ HONTICELLO = RELOAD 2 lo 4008 -
IN-SEQUENCE ROD E
8 5
5 aoos -
t o.ood -
\ -
aco2 -
o I I I I l .
0 2 4 g a to 12 CONTROL ROD POSITION (ft WITHDR AWN)
Figure 6-2. Accident Reactivity Shape Functions for Cold Startup S = 0.0054 6-14
4
-am j
I t
4 5
- -o.o7 -
4 i
1 .-
r i
t -o.os -
4 .
I i
.I I
_aos -
c t.
E -o.oe -
I k E
i 5
-nos -
4 280 CAL /GM PE AK FUEL MONT! CELLO - E= sagy contaur PERFORMANCE
-o.02 -
i
\
-not -
l o 1 , I g o 1 2 1 4 5 6 ELAPSED TIME AFTER SCRAM SIGNAL (esel I
Figure 6-3. Scram Reactivity Function for Cold Startup l
- - 6-15 i
f
- __ _ , _ - - , - , . , ,,--m.. _ ,, ,_..-.,_4
l i
i 1
} .
j 1.3% Ak derived. This verified that che consequences of a rod drop excursion i from any in-sequence control rod would be below the 280 cal / gram design limit, cince maximum in-sequence rod worths af ter this reload will be well below the ,
1.3% Ak allowable.
{
j 6.2.2.3.4.2 Fuel Damage l
1 The fuel damage thresholds are based on both experimental and theoretical data. This information is discussed in Section-14 of Reference 10, and Section f ,
i VI of Reference 11. 4 l
The rod drop accident analysis is sensitive to spatial variations in the core i design such as fuel loading patterns, gadolinium distribution, etc. An estimate t
] of the radiological exposures has been made and is based on the failure of all
) fuel rods above an energy content of 170 cal /gm assuming the maximum enthalpy l reaches 280 cal /gm during the accident. This is consistent with the boundary l cpproach established in Reference (4) . 'Ihe number of _ failed fuel rods and >
l the released fission products are therefore approximately the same as those
! discussed in the boundary approach document. The resulting doses are well
! within the 10 CFR Part 100 buidelines.
4 l 6.2.2.4 Loss-of-Coolant Accident j The following evaluation is based on-the 8x8 reload fuel. The results of l the 7x7 fuel evaluation has not changed and can be found in previous submittals.
I 6.2.2.4.1 Design Bases ,
l j The objective of the emergency core cooling systems (ECCS), in conjunction "
! with the containment, is to limit the release of radioactive materials following -
! a loss-of-coolant accident so that resulting radiation exposuras are within the values provided in published regulations.
Safety design bases and functional requirements for the emergency core i
cooling systems are given in the FSAR and have not changed.
I l
6-16 s
I
, ,_ ,-. ~...- .- _ --_.-. ._. --._ . -
- . - _ - ~ _ . - _ - _ _ . . _ . . _ , _ _ _ _ _ _ . _ _ . . _ . , - _ . _ _ , , _ . _ _ . - . . - - . , _ _
6.2.2.4.2 System Design 4
- The ECCS, comprising four separate subsystems, is designed to satisfy the 2 following performance objectives
1 a
- 1. To prevent fuel cladding melting as the result of any mechant .
i .' failure of the nuclear boiler system up to, and including, a break i equivalent to the largest coolant recirculation system pipe.
a
- 2. To provide this protection by at least two independent, automatically j actuated cooling systems.
1
- 3. To function with or without external (off-site) power sources.
- 4. To permit testing of all ECCS by acceptable methods including, where-l ever practical, testing during power plant operationa.
i
- 5. To function under assumed seismic conditions dascribed in Section 12 of Reference 10.
5 i The operational capability of the various emergency core cooling systems to l meet functional requirements and performance objectives is as follows.
i
! During the first ten minutes following the initiation of operation of the
. ECCS, the functional requirements are satisfied for all combinations of single
- active component failure and single pipe breaks, including pipe breaks in any ECCS subsystem which might partially or completely disable that subsystem.
1 Af ter the first ten minutes, and in the event of an active or passive failure I
in the ECCS or its essential support system, long term core and containment cool-in; in provided by any one LPCI or core spray pump delivering water .o the reac-tor vessel and by one RHR pump supported by one RHR heat exchanger with 100%
j service water flow.
i 4
The description and detailed design information on specific parts of the l emergency core cooling system is presented in the FSAR (Reference 10).
I 6-17
d 6.2.2.4.3 Performance Evaluation Summa ry. To achieve reliability, each energency . ore cooling subsystem uses the minimum feasible number of components tha cre required to actuate.
All equipment is testable during operation. Twc different cooling methods -
spraying and flooding - provide diversity.
Evaluation of ECCS controle and instrumentation for reliability and i redundancy shows that a failure of any single initiating sensor cannot prevent ,
i or falsely start these cooling systems. No single control f ailure can prevent dhe combined cooling systems from adequately cooling the core. The controls and instrumentation are calibrated and tested to assure adequate response to conditions representative of accident situations.
The emergency core cooling systems are provided to remove the residual and decay heat from the reactor core so that fuel cladding temperature is kept below 2300'F. The intent of the ECCS temperature criterion is to prevent gross core meltdown and fuel cladding fragmentation. Under extreme conditions highly oxidized Zircaloy could fracture on cooling. Based on the AEC's model in the Interim Acceptance Criteria for ECCS, cladding fragmentation on cooldown is prevented (for the time scale of interest here) if the maximum cicading tempera-ture is limited to less than 2300'F. This is therefore the design temperature criterion for ECCS system performance. The actual performance of the core cool-ing systems is such that peak temperatures much lower than 2300'F will be maintained throughout the complete break spectrum.
l A summary of peak cladding temperatures calculated to occur in R-2 fuel .
for the worst intermediate break and the design-basis break is in Table 6-1.
l
\
l l
l 6 '8 l
I i
l 3
Table 6-1 FEAK CLADDING TEMPEltATURES Large Intermediate Intermediate Break Break Break Temperature Temperature Size Single Failure Assumed (*F) (*F) (Ft2) 4 AEC Index of Worst Single Failure 2300 2300 X i Acceptability *
.C.a_s e
' AEC Assump- LPCI injection valve,
- 1. 2030 X X tions
. tions
- Calculated metal-water reaction is less than 0.2% of cladding for all cases above.
AEC acceptability index is 1%.
t Four LPCI pumps, two CS pumps, and ADS remaining.
X Does not apply.
f Two CS pumps, one HPCI pump, and ADS remaining.
Evaluation Model. The performance analysis of the ECCS is based upon analy-tical models used to conservatively predict reactor vessel pressure, liquid inven-tory, and fuel cladding temperature variations with time af ter a break. These models are identified, exemplified, and fully explained in Reference 5. There have been no deviations f am the evaluation model described in Appendix A, Fart II of AEC ' Interim Folicy Statement.
Fuel Clad Effects. Figure 6-4 shows peak cladding temperatures as a func-tion of time for the worst single failure case which leaves 2CS + HPCI and the
~
ADS operable. As shown, the maximum cladding temperature for this break, the most severe design-baris accident, is substantially limited by the emergency core cooling systems.
s 6-19
, 4 j
d
==
i "1
HOT M1 nePLO0000 1
i 1f
~
ROS B 0 .,
i i aoo i.
.I 1
1 I
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Y ao* "
f RATED MAY
! If i
l i
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j Mee -
i f
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i j
,
- ME PMMJRE M POR ROO IDENTIFICA110N
, 000 -
l r .-
i l .
, l i l I l l l l l l l l l l l l l 5 3 4 4 W M 40 M 100 300 Tuna lamel plyse M Ondeng Terapermoso n Tirrw fe the Recarculation bras Break weeh pedure ed abe LPCI j ispeeden Yaive (HPG + 3CE + ADS) ABC Aasenepeone
, 6-20 l
r m . _. . - . . . . _ . _ , _ . , _ . . . , . . , _ . . _ . , . _ , . . _ . , . , , . . _ _ . , , . . . _ . _ . . . . . _ _ _ , _ . . . . _ _ , . . _ . , . . . . . _ . _ _ . -
~
An example of the integrated system performance is shown in Figure 6-5 for a typical small size break with failure of RPCI. Peak cladding temperature for this case is shown in Figure 6-6.
Figure 6-7 is a break area spectrum analysis of the peak cladding tempera-ture and percent metal-water reaction for the worst single failures. The single
,. failures are the loss of the RPCI or the loss of the LPCI injection valve resulting in ECCS degradation to 4 LPCI + 2CS + ADS and RPCI + 2CS + ADS, r
respectively. Adequate cooling is maintained.
ECCS Performance Individual System Performance. The capability of the individual subsystems of the ECCS is shown on the bar chart (Figure 6-8) . A whole bar represents the capability of an individual system to protect the core without assistance from another subsystem. A half bar represents the range of break sizes for which a low pressure system must rely upon a high pressure system for additional inven-tory makeup and/or more rapid vessel depressurization. The ADS provides no inventory makeup and therefore cannot protect the core individually. The bar chart reveals subsystem characteristics but should not te applied to ECCS per-formance evaluations. No single failure could be hypothesized that would result in only one subsystem of the ECCS being available.
Integrated Operation of Emergency Core Cooling Systens. Two different nethods and at least two independent core cooling systerm are provided to limit fuel cladding temperature, over the entire spectrum of postulated reactor pri-
, mary system breaks, as required by the design bases.
The following discussion is directed toward the integrated performance of the ECCS; that is, how the ECCS will actually operate to provide core cooling for the entire spectrum of loss-of-coolant accidents. The discussion is sub-divided based on the two types of loss-of-coolant accidents; a break of a liquid
, line and a break of a steam line.
6-21
1200 60 4
1 1
! 1000 -
50 l
\
WATER LEVEL INSIDE SHROUD goo _ _
40 l
1 1 w
000 -
g _
TAF -
30 W
\ t s d - 5 g a f
400 - - 20 BAF i -
4 i
REACTOR VESSEL PRESSURE 200 - - 10 i
t
, 1 I l i 0 0 e 200 400 000 300 TIME (ses) i Mgure 6 5. Performance of ECCS with Failure of IIPCI for Small (0.02 ft2) Liquid Break (2CS + 4LPCI + ADS) AEC Awumptions 6-22
3000 1600 -
E L
c E
N - HOT SPOT REFLOODED l
l1000 w -
>=
ROD 18' ROO 04 5 MOD =
u RDO 37 600 -
'SEE FIGURE 6-24 FOR ROD IDENTIFICATION I I I O
O 200 400 e00 300 TIME leec)
Figure 64. Cladding Temperature vs. Time for a Small Break with Failure of IIPCI (0.02 ft2 Break)
(4LPCI + 2CS + ADS) AEC Assumption 6-23
_. . s_. . .. . . . . . . - . -. _. . - . -. _ . _ _ ._ _- - . . _. .
i l
2soo j
2
.I i.
t I 2000 - .
t i
! s WI FAILURE I ^
(4LPCI + 2G + AD$1 LPCI INJECTION VALVE FA1 LURE m (HPCI + 2CS + ADS)
L
" 1000 -
1 I
E l
4 1 u g 1000 -
l e.2 t
i y
- 5 9
m E
' too 8 c.1
' E E .
E 1
I i .
i
) 0 ' ' ' ' ' 'l ' ' ' ' ' ' i ' ' '
O O.01 402 0J15 0.1 OJ 0.5 1.0 2.0 5.0 BREAK AREA (ft 2g 4
1 i
Figun 6 7 Peak Cladding Tempeture Spectrum for Sinpe Failure Conditions AEC Assumptions i 6-24 8
4 e --- - ~, ,w , < , , , - . . , ,,e ,,m, --t , - , - - - - . - - - - - - * ,n
. - - . _ . , ..-n....~,.. . . _ . - - . - - - . . . . . . - . . . . . . - _ - - - . - . - ~ . ... _. ~. . . - -
l l 2
1 i
1 e
1 3
t STEAM SREAK$
l AUTO-RE LIE F
1 1
I k f_J
?.
4 8 LPCI I
l 0 I CORE SPRAY l
a i
2 i
4 4 1 1
1 4
1-f AUTO-RE LIE F 1
l
! LloulO SREAKS n El
} MIXIMUM REClRCULATION LINE SREAK SIZE I E ,
1 mi l 9 l t t
I l
C t l -
CORE SPRAY f _f I
l A l
eIlltttttlttiti11#1! I t lt I t l l' I f l t I i lID 1 l l t I f ilI l' G,0 0.1 0.2 1.0 2.0 10 4.0 5.0 5.0 7.0 SREAK AREA (ft ty i . SCALE CHANGES I
Figure 6-8. Emergency Core Cooling System-Performance Capability 6-25 r
{
. - . . . - - - . , _ , , . . , . _ . . _ . . . . _ . , , , . - , . . . . _ , - , . , , . _ .. ~ . _ .-._.,_,_ -. _ .. _,.. _..._ ,-. ....._. .,_ - -. , . ~ . . _ . . _ _ . - . . _ . . - _ _ .
For convenience, the breaks are classified according to the location of the penetration on the reactor vessel. The break types will fall into one of three categories. These, along with the lines that fall into these categories, are described below:
- 1. Steam Type Breaks. These are breaks in which the reactor vessel .
penetration is exposed to the steam regions inside the vessel.
- a. Steam Lines
- b. Some Instrument Lines
- 2. Steam / Liquid Type Breaks. These are breaks in which the reactor ves-sel penetration is either exposed to the two-phase regions inside the vessel or to regions which are exposed to liquid, but are near the water level and would therefore turn into steam breaks very shortly af ter the break occurred. These are located above the core.
- a. Feedwater Lines
- b. Core Spray Lines
- c. Some Instrument Lines
- 3. Liquid Type Breaks. These are breaks in which the reactor vessel pene-tration is well below the vessel water level and below the top of the core,
- a. Recirculation Pump Suction Lines
- b. Recirculation Pump Discharge Line
! c. Drain Line 4
j d. CRD Housir,3 -
1
- c. Incore Housing
- f. Jet Pump Instrument Line For a given size break, the lower the line penetration is located on the vessel, the higher the prak clad temperatures; 1.e. , the peak clad temperature i
for a given size break will be higher for those lines in liquid type breaks than 6-26
- . ~ - . . - . . - _ . - . . , - . - . . - , --
1 l
in steam / liquid type breaks and those in steam / liquid type breaks will be higher than those in steam type breaks. In demonstrating the performance and capability of 'the ECCS, recirculation line breaks are analyzed since these will result in
]
the highest ECCS peak clad temperatures for a given break size. The rupture and consequences of a main steam line break have also been analyzed since this is the most severe case with regard to containment performance.
) .-
.l For purposes of core performance and cladding integrity the most severe l .
accident (design basis accident) is the loss-of-coolant accident (recirculation line break). By analyzing breaks in the main steam line, the ef fects of all other steam type breaks are covered. For liquid type breaks, the spectrum analysis performed on the recirculation line breaks covers the ef fects of all other type liquid breaks such as the RHR suction and return lines, and recircu-
! lation riser lines.
The peak clad temperatures for the steam / liquid type breaks will be less than for the comparable size liquid breaks. This was shown in part in Millstone 12 Unit 1. AEC Docket No. 50-245 Amendment 14 in which the effects of various size feedwater breaks were analyzed.
Steam Line Breaks. The most severe steam pipe break is one that occurs inside the drywell, upstream of the flow limiters. Although Chc isolation valves i close within 10.5 seconds (10-second valve action time plus 0.5-second instru-ment response), such a break permits the pressure vessel to continue to depres-surize. For putposes of analysis, pre-accident conditions assumed are the reac-tor operating at design power, steam dome at maximum design operating pressure.
l
. ssram low wate.r level in the pressure vessel, and loss of auxiliary power
~
coincident with the steam pipe break.
The accident sequence starts with an instantaneous, guillotine severance of
~
the steam pipe upstream of the steam flow restrictors. The steam flow accelerates to its limiting critical flow value in the break at the pressure vesrel end and at the flow-limiter end. Steam loss exceeds the generation rate and results in rapid depressurization of the pressure vessel and steam pipes. The first 10 seconds of this accident are similar to the break outside the drywell. However, 6-27 i
a n- -
, , m - ,
l I
i i
1 i
a j for the break inside the drywell, closure of the isolation valves reduces the blowdown rate but does not prevent the vessel f rom depressuriting The vessel l
I continues to depressurite caasing suf ficient voids to immediately shut down the !
reactor.
1 4
g A scram is initiated by a position switch in each isolation valve (at approxi-1 J I mately 10% closure) so control rod insertion begins within 1.5 seconds after the j break. Low water level or high drywell pressure also initiate a scram.
1 Loss of reactor coolant through blowdown f rom the double-ended break con-sists of three intervals: first steam blowdown, then mixture blowdown, and fin-
) ally steam blowdown again. As the reactor vessel depressurites, flashing causes the water level to rise. When the level reaches the steam pipes, the break flow changes from a steam blowdown to a steam-vater mixture blevdown. Mass flow rate I through break increases sharply. At 10.5 seconds the isolation valves are j closed, which reduces the blowdown rate. As coolant is expelled and pressure '
j decreases, the wat2r level eutside the shroud drops below tre steam pipe eleva-ticu and eteam blowdown begirm again. The long term pressure transient and level olevation transient are shown in Figure 6-9.
j Approximately 40 seconds after tbs break occurs, both core spray systens and j the LPCI dystem start to inject coolant into the vessel. For this analysis the i aornaA situction where all ECCS pumps are operating is assumed. Analyses of f degraded situationt, in which only a portion of ECCS operates also show that the
! core remains covered and cooled throughout the entire blowdown transient, with l cladding integrity maintained.
I 1 . .
! Liquid __Line B eaks. The double-ended recirculat$on line break is the design i
basis accident for the emergency Lore cooling systems. The reactor is assumed
- I to be operating at design power when a complete circumferential rupture instantly j occurs in one of the two recirculation system suction lines. Normal a-c power l supply to the recirculation pumps is assumed to fail at the time of the accident.
i Core inlet flow and vessel pressure follouing the accident are shown in Figure 6-10.
i l
l l 6-28 I
i
', m - . . _ . . _ . _ ._ _ . ~ . _ - . . . - - - ._.. -._ _ __,__.____ _ . _
i 4
4 s
t i l
,,oo .o l
i i
a i
I 5000
- so )
' l 4 ,
1 WATER LtvLL
]
l
/ WATER LEVEL IN64DE SHROUD J
4 1
soo -
4 i i I
1 d
l-i 4
1
- w
- Ono .-
3 ,
E s :
.J t 2
d W !
> . b i d l +
< s 1 w p
a,i a
(
R eso -
~ 30 ,
j RAF 1
. ,co _ /- Rf Act0R visstL PREtSURE _ ,o f
I J
l I l n n
, o 90 too too 300 Tlut (seel l
Figure H. Performance of ECCS fur Main Steam Line Break inside the Drywell with all FCCS Operating
) (HPCS + 2CS + 4LPCI + ADS) AFC Assumpt ons i
- 6-29 t t
I i
4 ee -~ -n. .- , , .- , , _ , . . - , . , - - . . .-. .+_,,,--ne,,-,-,we-e.,.,w..- -.~n , ,mm,..--,,y-men _,
1 I \
i l
1 1J 1300 l
1 l
i
, i j 1.0 -
tuas
- 1 1
1-I h
1 i
u -
[ cont entasuRt -
soo 1
4 I
- g 0.4 300g i
i E 2
)
LOWER PLEDWUM j
PLOW P MN j COASTDOWN o - _
t 1
i l OJ -
- 300
- ~4-i
?
. l L l l l l l l 0 0 l 0 3 4 8 8 10 13 14 1e TIME (anal l
Figure 610 Core Flow and Pressure Fouowing a Recirruhtion une Break 6-30 -
l l
1 Initially, the rotating energy stored in the pump and motor of the unbroken recirculation system line provides continuing flow into the lower plenum, main-tatning a relatively high level of core flow. The flov is assumed to coaca when the falling level in the downconer reaches the jet pump suction level, i I I
! When the break flow in the severed tecirculation line changes to steam, the
, associated high vessel depressurization rate causes the water in the lower plenum to vigorously and immediately fle.sh to steam. This will force a two-phase f1w l j up through the core and through the jet pump diffusers. As the lower plenum inventory is depleted, the mass flow rate into the core diminishes.
l Calculations indicate that the reactor vessel depressurizes in approximately 3
40 seconds. The ECCS is initiated by either the low water level sensors in the i reactor vessel or high drywell pressure sensors. The ECCS begins delivering flow '
1 to the vessel at s30 seconds of ter the accident. Figure 6-11 shows the vessel 4
pressure and watcr inventory transient following the accident.
J i
The transient minimum critical heat flux ratio (MCHTE) for the highest pwered fuel bundle during the blowdown is shown in Figure 6-12. The axial power shape q was chosen to ass're that the fuel bundle was initially operating at thermal ,
limits.
I j
As la evident from the figure, the MCHFR decreases initially af ter the acci- "
dent occurs, increases slightly, and then decreases to less than 1.0 when core flow stagnates due to the uncovering of the jet pumps. Steam then blankets the reactor core and film boiling is established. However, this heat transfer is conservatively neglected for this analysis (i.e., the heat transfer coefficient is set to zero).
+
, MCliFR becomes greate than unity when the high core flow rates caused by water in the lower plenum flas; to steam. This flashing forces large quantities of water through the core and jet pumps. (See Figure 6-10.) With MCl!FR greater 4
than unity, reestablished nucicate boiling would quickly cool the cladding to near saturation temperature. However, no credit for rewetting is taken. The Groeneveld f11m boiling correlation (AECL-3281) is used to determine the convec- ,
tion coefficient as instructed by the AEC Interim Acceptance Criteria (IAC).
6-31
I I
1200 60 1000 - -
90 s
= m 300 - ~
40 t
1 J
000 -
- 3 E
.a 9A,
. g 1
.A l> 1 -
WATER LEvil INSIDE SHROUD 400 -
- 30 "E /
4 200 -
to REACTOR VitstL PRES $URE I I f O 0 0 to toc 40 200 TIME lasci Figure 611. Perforrr.snee of ECCS with Fadure of LPCI Injection Valve for the Design Basis Recireviation Line Break (llPCS + 2CS + ADS) AEC Assumptions 6+32
-.~~~..-.,n-.~.._.. . _ . _ _ , . . . . . . . . , - . - . . - . . - - . .... .. . . - . _ - - - - - - - . . _ _ - - . _
i 4
a l
1, j
I i
, 7 1
i 1
1 6 -
f i .
1 i
- 1 4 . 6 - -
1 1 i,
g o
E E
$, b a 4 -
4 E
1 k I
J h
- r l 52 -
4 -
J.
1 s
I d
i 2
3 4
4 i
t
] 1 i
i 1
1
. f I ! !
l . .
O 0 2 4 6 8 to 12 TIME (ses) 4 l
t I i I Figure 612. Knimum Critical liest Flux Ratio for Di,A at Monticello
!._ 6-33 ;
1 i
- c. ,-.._,,._n-- - _ , - __ _ ,.,,_.__--,-,_.,,,.--,,7-.,7,.,_,y. _ _ . _ _
,,.r_,_,w. r,%__. m. ,,--,,,y,, , . . _ _ _ ~ .m._ ...,,,--m_,. .-y,. . , , , , .-
When the core uncovers, it is assumed to be insulated. A drywell high pres-sure or reactor vessel low water level signal starts the HPCI and LPC1, the LPCS, and the standby a-c power supply. When the core spray flow reaches rated value or when the core is reflooded, the appropriate coef ficients are applied.
Figures 6-6 and 6-13 show the peak cladding temperatures for four rod groups for a small and intermediate break, respectively. The intermediate break size shown is one that results in high peak cladding temperature in the smaller break sire range.
Figure 6-14 shows core inlet and outlet quality versus time for the DBA. The curve is shown for only the DBA, because quality af fects the film boiling heat transfer coefficient. For small and intern.ediate size breaks, nucleate boiling is assured as Ic>ng as the core is covered. Nucleate boiling heat transfer coefficients are independent of fluid quality. When the core is uncovered, the heat transfer coefficient as assumed to be zero, even though a significant steam cooling coef ficient would exist.
Figures 6-15, 6-16, and 6-17 show the heat transfer ccefficient versus time for the small break, intermediate break, and design basis accident, respcetivaly.
Figures 6-5 and 6-18 show the reactor vessel (RPV) water level versus time f or the small break and intermediate break, respectively.
Figure 6-12 shows the minimum critical heat flux ratio (MCllFR) versus time for the design basis LOCA. Because the flow trsnsient for the small and inter-mediate break sizes is mild compared to the LOCA, it is not shown. As long as the core is covered, the MCHFR for smaller breaks is always greater than unity, and nucleate boiling is always assured.
Figure 6-19 shows the assumed power generation following a design basis accident.
System Capacity. System capacity description and detailed infornation is contained in the FSAR and has been shown to be adequate to keep peak cladding temperatures <2300'F.
6-34
- - - - . ~ . _ - . ~ . . _ . . . ~ . . - . . . _ . . . - - - . ~ . . - . ~ . - . . - . . . . . _ _ . - . . . . _ _ . . . _ . - . - - . . . - _ _ - - . _ . . . - - . _
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l 17:gure 613.
Caddang Temperature n. Time for an Intermediate Brtak with Failure of IIPCI 2(0.1 ft Break) i (4LPCI + 2CS + ADS) AEC Assumptions i
6-35
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i I Figure 614. Quality vs. Time for DBA st MonticeUo l
f 6-36 l
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l Figure 613. Ileat Transfer Codficierit for a Small Break (0.02 f t2 ) ,
l (4LPCI + 2CS + ADS) AEC Assumptions i 6-37 I
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l (4LPCI + 2CS + ADS) AEC Amamptbns i
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- 6-39 i
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, Liquid Break (2CS + 41.PCI + ADS) AEC Assumptions ,
6-40 ,
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Long Term Core Cooling. Long term cooling is defined as cooling after the initial thermal transient has been terminated until the fuel can be safely removed. Long term cooling conditions have not changed with insertion of the reload fuel.
n Peaking Factorn Figure 6-20 fu a plot showing the typical behavior of cladding temperature versus exposure for the DBA. peaking factors giving the highest peak cladding .
temperatures occur at N15,000 mwd /T exrosure. These peakin6 factors were then used in conjunction with the IAC calculational models to determine the stated pedk cladding temperatures for Monticello.
Fuel Rod perforations. The mechanism of fuel rod perforation during the LOCA has been studied extensively and le well understood. A fuel rod will perforate if the cladding hoop stress exceeds the ultimate strength of Zitcaloy at the peak cladding temperature experienced during the LOCA. The number of fuel rods perforated is therefore a functic of the predicted peak cladding temperatures as well as the experimentally determined internal gas pressure distribution and perforation st.ess data. A plot of stress at perforation and ultimate strength of Zircaloy at various temper 2 cures is presented in Figure 6-21.
The calculated fission gas internal pressure distribttion within the core is shown in Figure 6-22 for a typical 7x7 fuel design. The distribution is obtained by integrating the expected fission gas release rate for normal opera-tion to the end of an equilibrium cycle when the accunulated fission products are maximum. In addition, the partial pressure of volatile materials and initial gas pressure is included. Because the 8x8 fuel design has lower opera-ting fuel temperatures, a smaller diameter and thicker cladding relative te the
- 7x7, the stress distribution produced by Figure 6-22 for 7x7 fuel can be con-servatively applied to 8x8 fuel. These data are used with the heatup analysis to determine the maximum percentage of fuel rod perforation for any size break .
and the worst single failure assumption (see Figure 6-23).
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l 6-43 e
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70.000 6 APED D AT A IN AIR ($lNGLt RODI 6 Y APED PREOxlDelED tslNott ROOD g g IN Sit AM 10.000 - O APIo oATA 'N AIR to ROD f t st 11 Q APE D DAT A IN AIR 19 ROD TEST 11) 6 O NMPO DA1 A 0 075 cm = 0 63 em 2 64 O.eGE Seit iT $PECIME WS. t 0 06 men 1 IN ARGON O NMPO IRR ADI Af t D Fr TUBING IN ARGON Q NMPO 1RR ADI Af tD Cr TUBING IN STE AM 5,000 6
$t i NE DO 10329 F OR FULL EXPLANATION
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', 0 100 900 1000 iL 1500 i INTE RN AL PRESSURE (poiel
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i Fpre 6 23. Percent Rod Perforation n Break Area (AEC Amumptions) 6-46 i
__. _ - - _ _ __ .~. , , _ _ _ _ , - - _ , _ . _ . . _ . . . _ . - _ _ _ _ _ _.- . - _ . . . _ . . .
Conformance With Interim Acceptance C;i$eria. In the analyses discussed above there have been no deviationa f rom the evaluation model described in Appendix A. Part 2 of the AEC Interim Policy Statement.
Effects of ECCS Oparation on the Core. The mechanical effects of ECCS opera-I tion on the core, react.? coolant system and ECCS are those associated with the
,. thermal ef fect of injecting water into these systena which is cooler than j these systems and components. These thermal stresses have been considered in the design of the core, reactor coolant system and ECCS.
l There are no nuclear effects resulting from ECCS operation, since all con-trol rods are inserted and the reactor remains suberitical during the injection of the cooler ECCS water.
There are no chemical additives in the ECCS vater and thetefore no chemical ef f ects on the core, reactor coolant system or ECCS.
Lag Times. The system time delays assumed in the LOCA accident are as follows.
Maximum Allowable Time From Maximum Time Delay Af ter Signal Receipt Until the Pttps Receipt of Signal Until All j Have Reached Rated Speed Valve Motion is Complete Syatem (sec) (sec) kWC1 30 30 CS 30 30 j LPCI 43 43 ADS -
120 Fuel Dennification ,
r Actus1 fuel density data and a supplemental page to NEDH-10735 (Reference
- 13) will be provided so that Maximum Average Planar Linear Heat Cencration Rate (KAPLHGR) and Power Spiking Penalties (AP/P) for the 8x8 fuel can be deter-mined. This information will be available prior to startup of the reloaded
, core. Appropriate changes to the densification models occurring in the interim will be similarly addressed.
6-47
'--"a WIDE WlDL CORNin l
1 I 2 3 4 6 6 7 11 9 10 11 12 I? 14 16 16 17 18 19 20 21 22 23 24 25 ?6 27 78 29 30 31 32 33 34 36 36 37 38 39 40 41 42 43 44 45 46 47 ed 49 60 61 62 63 64 66 2 67 68 69 60 61 61 63 64 e
4 p'llI0' I' . k O' I Isl{ ll5 8 B0 $ 8'lI l $$'f{ k lt lll 6-48
l 4 6.2.3 Rod Withdrawal Error l 6.2.3.1 Identification of Causes
}l Startine Conditions and Assumptione. The reactor is operating at a power level above hot standby at the time the control rod withdrawal error occurs.
I The reactor operator has followed procedures and up to the point of the with-1 i
drawal error is in a normal mode of operation (i.e., the control rod pattern,
, flow set point, etc. , are all within normal operating limits). For these con-i i
ditions it is assumed that the withdrawal error occurs with the maximum worth control rod. Therefore, the maximum positive reactivity insertion will occur.
j Event Description, k'hile operating in the power range in a normal mode of operation the reactor operator makes a procedural error and withdraws the maxi-mum worth control rod to its fully withdrawn position. Due to this positive 4
reactivity insertion, the core average power will increase. More importantly, j the local power in the vicinity of the withdrawn control rod will increase and potentially could cause localized fuel failures due to either achieving critical h3at flux (CHF) or by exceeding the 1% plastic strain limit imposed on the cladding as the transient failure threshold. The following list depicts the sequence of 4
events for this transient.
Approximate Event Elapsed Time 4
(1) INent boghs; operator selects and withdraws at maximum 0 rod speed the maximum worth control rod (2) Core average and local power increases
. (3) LPRM's alarm <5 sec
~
(4) Event ende-rod block by RBM <30 sec 4
Identification of Operator Actions. Under most normal operating conditions no operator action will be required since the transient *:hich will occur will be
- very mild. If the peak linear power design limits e.te exceeded, the nearest local povar range monitors (LPRM's) will detect this phenomenon and sound an alarm. The operator must acknowledge this alarm and take appropriate action to rectify the situation.
6-49
?
if the rod withdrawal error is severe enough, the rod block monitor (REM) system will sound alarms at which tir- the operator must acknowledge the alarm and take corrective action. Even for extremely severe conditions (i.e., for highly abnormal control rod patterns, operating conditions, and assuming that the operator ignores all alarms and warnings and continues to withdraw the control rod) the RBM system will block further withdrawal of the control rod bef ore fuel damage occurs.
6.2.3.2 Analysas of Effects and Consequences Methods, Assumptions, and Conditions. The analysis considers the con-tinuous withdrawal of the maximum worth control rod at its maximum drive speed from the reactor which is operating at rated power with a control rod pattern which results in the core being placed on thermal design limits (i.e., MCHFR =
1.9 and a peak linear power of 13.4 kW/f t for the B X 8 fuel or 17.5 kW/f t for the 7 X 7 fuel). A worst case condition is analyzed to ensure that the results obtained are conservative. Also, this approach serves to demonstrate the func-tion of the RBM system.
The worst case situation la established for the most reactive reactor state and assumes that no xenon is present. This ensures that the maximum amount of excess reactivity which must be controlled with the movable controi rods is present. During a normal startup sufficient time would be available to achieve some xenon and samarium buildup, and after some short period of operation camarium will always be present. This assumption makes it possible to obtain a worst case situation in which the maximum worth control rod is fully inserted and the remaining control rod pattern is selected in such a way as to achieve design thermal limits in the fuel bundles directly adjacent to or diagonally adj acent to the inserted naximum worth control rod which is to be withdrawn.
It should be pointed out that this control rod configuration would be highly abnormal and could only be achieved by deliberate operator action or by numerous operator errors during rod pattern manipulation prior to the selection and com-plete withdrawal of the maximum worth rod.
6-50
_ -_ ___ _ _ l
Figure 6-25 shows the locations of the initial, first reload and 8 X 8 fuel. Figure 6-26 indicates the rod withdrawn in tha limiting case. Figure 6-27 is a composite MCHFR curve where the worst 7 X 7 and 8 X 8 bundles have been combined, thereby showing the rod position where further withdrawal of the con-trol rod of interest could result in !!CHFR <1.0 in one of the fuel bundles in-fluenced by that rod. Figures 6-28 and 6-29 show the worst case MCHFR for 7 X 7 and 8 X 8 bundles, respectively. The most limiting case of each is used to
.' establish the curve in Figure 6-27. Figure 6-30 shows the peak heat flux and relative bundle power for the limiting case considering both 7 X 7 and 8 X 8 fuel. The analyzed peaks occur at 22.4 kW/f t and 16.8 kW/f t well below the damage limits of 28.0 kW/ft and 25.4 kW/ft for 7 X 3 and 8 X 8, respectively.
From the composite MCHFR curve (Figure 6-27), the control rod maximum withdrawal point is shown to occur at 6.5 ft. From the diagrams in Figures 6-31 and 6-32 which reflect the various system responses to combinations of allowable LPRM f ailures, the RBM reading corresponding to withdrawal of the limiting rod is shown. In the case of Channels A and C (Figure 6-31), a rod block occurs at 5 f t., well below the limiting 6.5 f t. point.
For channels B and D, (Figure 6-32) the rod block occurs at 6.2 f t.; with the RBM setting of 108%, the RBM satisfies the analytical requirements.
The effects of continuous withdrawal of an in-sequence rod were alsu analyzed; as expected, no in-sequence event produces results more severe than the limiting cases described above, e
6.2.4 Transient Analysis and Core Dynamics This section describes the transient analysis for Monticello, Cycle 3, based on the analyses of the Final Safety Analysis Report (FSAF.) . This analysis considers the large reactor distrubances which serve as a basis for a comprehensive evaluation of the plant dynamic safeguards.
6-51 1
- . . - - ,-- . .. . . . . . . . .;.-'. : ; .'m e...~ia'.7;.'.;.~.". C.T.'.T; .C ;
1 1
l 1
I slTE NOutNCLATURE Jr 01 03 06 07 08 11 13 15 17 19 21 23 26 s2 1 H R 2 se 3 4 R R R 4 ,
M R R S 42 R R R R e ao R e 38 R R R R R S 36 R B B 9 34 A R R R R 10 32 R 8 5 11 30 R R R R R R 12 d
28 13 1 2 3 4 5 6 7 8 9 10 11 12 13 R = 8 X 0 (RELOADJ)
B e (RE LDAD 11 DESIGN NOMthCLATURE 4
Figure 6J*. l'u. l T 3pc len ali in.
6-52
l l
1 81 0 0 0 of 24 0 0 24 43 0 24 14 24 0 38 to M M 0 34 34 24 at M E 14 0 24 21 0 0 F M 42 M 34 0 0 27 0 14 0 18 42 0 42 18 0 14 0 23 0 0 M M 42 34 M 0 0 18 24 0 14 0 24 13 24 b M 0 34 34 24 11 0 24 to 24 0 7 24 0 24 0
- 3 0 0 0 1
0 2 6 10 14 18 22 26 30 34 38 42 44 60 F= TRANSIENT ROD BLANK = ROD OUT XX = NOTCHES ROO WITHDRAWN FULL CORE e
Figurr 6 26. l(a.c l(ini l'onirrn hit 1(% l'.
, 6-53
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ROD PO&lTION Pigure 6 27. Monticello Composite MCliFR vs. Rod Poution L.imiting Rod Withdrawal 6-$4
m 42 4.0 -
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Figure 6-29. MCHFR vs. Rod Position Limitirig Red Withdrawt] 8 X 8 6-56
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r Figure 6 30. '
- H-st Flux and Relative Power vs. Rod Position Limitini Roo Withdrawal Monticebo BOC3
!. 6-57 1
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i 100 0 2 4 5 8 10 12 i4 l MTROL ROD POSITION (FEET WITHDRAWN) i 2 Figure 6-31 RBM Raponse to Control Rod Modon - Monticello Channel A + C Umiting Rod withdrawal 4
6-58 a
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1 0 2 4 0 3 10 12 14 CONTROL ROO PostTION (FEET WITHDRAWW) l
- Figure 6-32. RBM Response to Control Rod Motion in Monticello
- {. Channel B + D Limiting Rod Withdrawal 6-59
, , _ - . . . . , _ _ _ . , . . . - . , , , , , . _ _ . , , _ . . . . . . . _ - . . . . . . 2,. -,.r,_.,,,,,. . . . . . . . . . . . _ , , . . . . , , _ . ~ . . , , . . _ , . . , , .
Two msjor design changes, as well as the introduction of 8x8 fuel and '
modified contrM rod scram times, are being planned for Cycle 3 and have been incorpor. in the digital model used for the analysis. The replacement of the four s, _, safety valves with combination safety / relief valves of the the same type as the existing RV's is a proposed change dt.igned to improve pressure relief margins and eliminate the interaction of relief and safety valves. Six Target Rock combination safety / relief valves are considered in this analysis. A second major design change, a Prompt Relief Trip (PRT)
- system, is planned to increase the effectiveness of the safety / relief valves i in maintaining pressure and thermal margins when the plant is subjected to j severe steam flow disturbances.
}
f The control rod scram time used was the same as the present Technical Specification scram time up to 507, insertion; a straight extension to the 907.
insertion was used, changing that point from 5.0 seconds to 3.5 seconds. The analyses used the Design Basia Scram Reactivity Curve (D Curve) d? fining the generic outer bound scram reactivity function. The scram reactivity curve and j scram time curve are shcwn in Figures 6-33 and 6-34.
The PRT system provi'aes an immediate trip of six s&fety/ relief valves in response to a turbine trip (stop valve closure) or a generator load rejection (fast closure of the control valves) from high reactor power levels. This
- system is a major improvement for the scram reactivity considerations dis-I cussed at length in earlier submittals. 11,15,16 The PRT system will be discussed in detail in a separate document to be submitted later; included j will be a functional descriptior of the system, its objectives, operating .
l characteristics and other considerations. The reload analyses were based i
on the followitag operating conditions (except for special cases requiring initial conditions at less than rated power and flow):
{
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Thermal Pow n 1670 Wt (T-G Design)
T-G Lasign Steam d ov 6.77 x 106 lb/hr i
i 6-60 1
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I Turbine inlet Pressure 980 psig Jet Pump M Ratio 1.59 i
Bypass capacity 15% - Design Flow j
Safety / Relief Valve Capacity 71.1% - Design Flow (6 valves at 1080 psig +1%)
Safety / Relief Valve Time Delay 0.4_sec Safety / Relief Valve Stroke Time 0.1 see i,
I Scram Rod Drive (Figure 6-33) 67S (3.5 seconds at 90% stroke)
Scram Curve (Figure 6-34) Design Basis "D" i
i Feeawater Capacity 105% - Design Flow Feedwater Temperature 3760F 1
6.2.4.1 Identification of Abnormal Operational Transients ,
i A cpmplete range of single-failure-caused events which are abnormal but ra_asonably expected to occur during the life of the plant were analyzed ar part of the original licensing of the plant. These analyses I
' were described in the FSAR.
Subsequent submittals have, where appropria'.e, included additional consideration of those events of significance to the j
concept being reported (i.e. ,
reloads, changes to transient analysis parameters, or plant modifications). 11,14,15,16 A complete evaluation of all transient events (abnormal operational transients) was performed in support of this reload to ensure all pre -
1 viously established requirements were met.
This extensive reanalysit was deemed necefsary in view of the significant plant changes bet.g concurrently applied to Monticello during the forthcoming teraeling outage.
1 6-61 i
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The transient reapslyses included the following nuclear system parameter
! variations:
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- 1. nucicar system pressure increases;
- 2. reactor vessel water (modera?or) temperature Jecreases;
- 3. positive reactivity insertions; '
- 4. reactor vessel coolant inventory decreases;
- 5. reactor core coolant flow tacreases; and .
- 6. reactor core ecolant flow decreases.
A The complete reanalysis of the affected FSAR abnormal operational n
transients shows that the reload core satisfies MCHFR, heat flux and over
- pressure requirements as described here and in Reference 10. Only those transients resulting in system pressure increases are significantly affected by the planned reactor pressure relief system modifications, the use of the D acram reactivity curve, and the new scram times. The detailed results of all Abnormal Operational Transients will be presented in the farthcoming analysis based on these permanent solutions to changing scram reactivity
- conditions. While specific details of the modification may change prior to the installation, the basic operational aspects should not; the analyses are expected to remain essentially the same.
6.2.5 Loading Error l
l The worst case loading error for the c 2arence core configuration occurs i
when a reload bundle is rotated 180 degrt.s in a location near the center of .
i the core.
Proper orientation of fuel assemblies in the reactor is readily verified by isual observation and is assured by verification procedures during core loading. Five separate visual indications of proper fuel assembly orientation -
exist:
4 6-64
. _ _ _ . - - ~. __ _ _ _ - _ ,. , ., _, - . _ . - . _ _ _
NEDS-20016
- 1. The channel fastener assemblies, including the spring and guard used to maintain clearances between channels, are located at one corner of each fuc1 assembly adjacent to the center of the control rod.
- 2. The identification boss on the fuel cosecbly handle pointo toward the adjacent control rod.
- 3. The channel spacing buttons are adjacent to the control rod passage
. area.
- 4. The assembly identification numbers on the fuel assembly handles are all readable f rom the direction of the center of the cell.
- 5. There is cell-to-cell replication.
4 Experience has demonetrated that these design features are clearly visible so that any miroriented fuel assembly would be readily distinguished during core loading verification.
If, however, through an error, a fuel assembly were installed rotated 180 0 from the proper orientation, which is the norst case rotational error for any expecure level, no fuel damage would be incu. ted during the subsequent power operation, even if the misoriented assembly were operating at the maximum permitted power. Analysis shows that this error would result in a MLHGR 2516.3 kW/ft and a MCHFR 2el.51 for a rotated R2 bundle. These are less than the damage limit established for this fuel. Should the loading error involve
, , one of the irrad .ted assemblies, the analysis in Reference 11 (reporting that no fuel damage would be incurred) is applicable for the Cycle 3 core.
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6-65
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l NEDS-20016 REFERENCES - SECTICN 6
- 1. Paone, C. J., and Woolley, J. A., " Rod Drop Accident Analysis for Large l Boiling Water Reactors," Licensing Topical Report, March 1972 (NEDO-10527).
1
- 2. Stirn, R. C., Paone, C. J and Young, R. M. , " Pad Drop Accident Analysis for Large BWRs," Licensing 'Ispical Report, July 1972 (NEDO-10527, Supple- -
ment 1). ,
- 3. Stirn, R. C. , Paone, C. J. , and Haun, J. M. , " Rod Drop Accident Analysis for Large Boiling Water Reacters Addendum No. 2 Exposed Cores," Licensing Topical Report, January 1973 (NEDO-10527, Supplement 2).
- 4. " Technical Basis for Changes to Allevable Rod Worth Specified in Technical Specification 3.3.8.3 (a)," submitted to AFC October 4,1973. Dkt. 50-263.
- 5. Slifer, B. C., and Rogers, A. E., " Loss-of-Coolant Accident and Emergency Core Cooling Models for General Electric Boiling Water Reactors," Licens-ing Topical Report, April 1911 (NEDO-10329 and NEDO-10329 Supplement 1) .
- 6. Duncan, J. D. , and Leonard, J. E. , "Modeling the BWR/6 Loss-of-Coolant Accident: Core Spray and Bottom Flocling Heat Transfer Effectiveness,"
, March 1973 (NEDE-10801).
I
- 7. Linford, R. B., " Analytical Methods of Plant Transient Evaluations for the General Electric Boiling Water Reactor," February 1973 (NEDO-10802).
1
- 8. "In-Core Nuclear Instrumentation Systecs for Oyster Creek Unit 1 and Nine
- Mile Point Unit 1 Reactors," August 1968 (APED-5456). .
- 9. Morgan, W. R., "In-Core Neutron Monitoring System for General Electric Boiling Water Reactors," November 1968, revised April 1969 (APED-5706).
- 10. Monticello Nuclear Generating Plant, FSAR, Dkt. 5h 263.
4 6-66
i 1
- NEDS-20016
- 11. Monticello Muc1ser Generating Flent, First Reloed License Submittal, l February 1973.
i
- 12. Millstone Unit 1. FSAR Amendment 14, Dkt. 50-245.
a l
l 13. " Fuel Densification Effects on General Electric Boiling Water Reactor Fuel, Supplement 6, 7, and 8. Composite," August 1973 (NEDM-10735).
,* 14. "Results of Transient Resnalysis for Monticello Nuclear Generating Plant l
i with End-of-Cycle Core Dynamit. Chat acteristics," February,1973.
l
[ 15. "Monticello - Safety Valve Setpoint Increase Analysis," Change request i dated September 3, 1973.
d l 16. "Monticello Cycle 2 Scram Reactivity Considerations, Analyses and Modifi-
} cations " October 1973.
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a j 7. TECHNICAL SPECIFICATIONS
- There are four areas of the Technical Specifications affected by the preceeding information. Changes made necessary by the reactor pressure relief i
! system modifications discussed in Section 6.2.4 will be outlined in the forth-coming submittal on that subject. The formal request for Technical Specification f change = will be a separate, subsequent submittal. Specifications affected by this submittal include the following:
Section 2 - The heat flux of a 7x7 fuel assembly operating up to 17.5 kw/f t results in a 3.-08 total peaking factor. Changes should reflect the use of 8x8 l fuel operating up to 13.4 kw/f t resulting in a 3,04 total peaking factor.
Section 3.3.C - The transient analysis (Section 6.2.4) was done based on a control rod scram time to 90% insertion of 3.5 seconds rather that 5.0 seconds as presently allowed. The Specification will be changed accordingly.
{ Section 3.5.K - The bx8, R-2 fuel will have unique properties for consideration of postulated fuel densification phenomena. Since the AEC staff model requires the use of measured pellet theoretical density, this information can not be i finalized until the fuel is fabricated.
l Section 5.2 - The facility description states that fuel assemblies have 49 fuel rods each. This must be changed to allow the use of 8x8, 63 fuel rod assemblies.
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