ML18312A071
| ML18312A071 | |
| Person / Time | |
|---|---|
| Site: | Farley  |
| Issue date: | 10/30/2018 |
| From: | Southern Nuclear Operating Co |
| To: | Office of Nuclear Reactor Regulation |
| Shared Package | |
| ML18312A093 | List: |
| References | |
| NL-18-1299 | |
| Download: ML18312A071 (271) | |
Text
REV 21 5/08 HELB OUTSIDE CONTAINMENT 0.05 ft 2 BREAK AT 102-PERCENT POWER 30-min. OPERATOR ACTION TEMPERATURE VS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-1
REV 21 5/08 HELB OUTSIDE CONTAINMENT 0.2 ft 2 BREAK AT 70-PERCENT POWER TEMPERATURE VS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-2
REV 21 5/08 HELB OUTSIDE CONTAINMENT 0.2 ft 2 BREAK AT 102-PERCENT POWER TEMPERATURE VS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-3
REV 21 5/08 HELB OUTSIDE CONTAINMENT 0.4 ft 2 BREAK AT 102-PERCENT POWER TEMPERATURE VS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-4
REV 21 5/08 HELB OUTSIDE CONTAINMENT 0.6 ft 2 BREAK AT 102-PERCENT POWER TEMPERATURE VS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-5
REV 21 5/08 HELB OUTSIDE CONTAINMENT 0.8 ft 2 BREAK AT 102-PERCENT POWER TEMPERATURE VS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-6
REV 21 5/08 HELB OUTSIDE CONTAINMENT 1.2 ft 2 BREAK AT 102-PERCENT POWER TEMPERATURE VS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-7
REV 21 5/08 HELB OUTSIDE CONTAINMENT 1.1 ft 2 BREAK AT 102-PERCENT POWER TEMPERATURE VS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-8
REV 21 5/08 HELB OUTSIDE CONTAINMENT 4.6 ft 2 BREAK AT 102-PERCENT POWER TEMPERATURE VS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-9
REV 21 5/08 HELB OUTSIDE CONTAINMENT COMBINED TEMPERATURE PROFILE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-10
REV 21 5/08 HELB OUTSIDE CONTAINMENT 0.05 ft 2 BREAK AT 102-PERCENT POWER PRESSURE VS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-11
REV 21 5/08 HELB OUTSIDE CONTAINMENT 0.2 ft 2 BREAK AT 102-PERCENT POWER PRESSURE VS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-12
REV 21 5/08 HELB OUTSIDE CONTAINMENT 1.1 ft 2 BREAK AT 102-PERCENT POWER PRESSURE VS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-13
REV 21 5/08 HELB OUTSIDE CONTAINMENT 4.6 ft 2 BREAK AT 102-PERCENT POWER PRESSURE VS TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-14
REV 21 5/08 HELB OUTSIDE CONTAINMENT COMBINED PRESSURE PROFILE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-15
REV 21 5/08 HELB OUTSIDE CONTAINMENT COMBINED TEMPERATURE PROFILE FOR MODEL 54F CASES AT 102% POWER JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-16
REV 21 5/08 HELB OUTSIDE CONTAINMENT COMBINED TEMPERATURE PROFILE FOR MODEL 54F CASES AT 70% POWER JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3J-17
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.2-1 CRITERIA FOR INDENTIFICATION OF HIGH-ENERGY LINES AND EFFECTS CONSIDERED
Criteria Break Types and Effects Considered A Critical crack, jet impingement, flooding B Circumferential or longitudinal breaks, critical cracks, pipe whip, jet impingement, flooding, pressure and temperature effects on structural integrity of compartments, environmental effects
C Same as A D Piping systems with temperatures and pressures less than 200
° F and 275 psig not considered in this appendix.
Note: 1. Also includes temperature and steam/moisture effects. See section 3K.2.1.2.
Service 275 A B Pressure (psig) 0 D C 0 200 Service Temperature (F)
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.3-1 (SHEET 1 OF 2)
EQUIPMENT REQUIRED FOLLOWING A HIGH-ENERGY LINE BREAK - UNIT 1 (a)(b) (OUTSIDE CONTAINMENT)
Short-Term Long-Term Required for (<10 Min) (Hot Standby) Cooldown Reactor trip and safeguards act uation channels including sensors, Auxiliary feedwater system including pumps, water supply, Steam generator power operated relief valves (can be circuitry, and processing equipment (the prot ection circuits and system valves and piping (this system must be placed manually operated locally) used to trip the reactor on undervoltage, underfrequency, and in service to supply water to operable steam generators turbine trip may be excluded) within one minute after initiating signal)
Controls for defeating automatic safety injection actuation during Safety injection system, includi ng Reactor containment ventilation a cooldown and depressurization. pumps, the refueling wate r storage cooling units tank, and system valves and Residual heat removal system piping including pumps, heat exchanger and system valves and piping Diesel generators and emergency necessary to cool and maintain power distribution equipment the reactor coolant system in a cold shutdown condition Essential service water system including pumps and system valves and piping Capability for obtaining a reactor coolant system sample Essential component cooling water system including pumps, heat exchanger, and component cooling water surge tank
Main feedwater control valves(trip closed feature)(c)
Circuits and/or equipment required to trip the
) main feedwater pumps (c Main feedwater isolation valves (trip closed feature)(c)
Main steam line stop valv es (trip closed feature)(d)
Main steam line stop valve bypass valves (trip closed feature)(d)
Steam generator blowdown isolation valves (automatic closure feature)
Batteries
Control room ventilation
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.3-1 (SHEET 2 OF 2)
Short-Term Long-Term Required for (<10 Min) (Hot Standby) Cooldown Control room equipment must not be damaged to the extent that any equipment will be spuriously actuated or any of the equipment contained elsewhere in this list cannot be operated Emergency lighting In addition to the instrumentation required to operate the equipment on this list, indication of the following plant parameters should be available to the operator:
Wide range T hot or T cold (preferably T hot) for each reactor coolant loop Pressurizer water level Wide range reactor coolant system pressure Steam line pressure for each steam generator Wide range and narrow range steam generator level for each steam generator Containment pressure
- a. Instrumentation, interlocks, and power supplies required to operate the above equipment must be available.
- b. Support systems for the above equipment su ch as long-term diesel fuel storage, battery chargers, and a long-term water supply for the auxiliary feedwater system must be available.
- c. Required for steam line and steam generator blowdown line break only.
- d. Required for steam line, feed line, and steam generator blowdown line break only.
FNP-FSAR-3K 3K.A-i REV 21 5/08
ATTACHMENT A GENERAL INFORMATION REQUIRED FOR CONSIDERATION OF EFFECTS OF A PIPING SYSTEM BREAK OUTSIDE CONTAINMENT
FNP-FSAR-3K 3K.A-1 REV 21 5/08 ATTACHMENT A PART I GENERAL INFORMATION REQUIRED FOR CONSIDERATION OF THE EFFECTS OF A PIPING SYSTEM BREAK OUTSIDE CONTAINMENT
The following is a general list of information required for NRC review of the effects of a piping
system break outside containment, including the double-ended rupture of the largest pipe in the
main steam and feedwater systems, and for NRC review of any proposed design changes that
may be found necessary. Since piping layouts are substantially different from plant to plant, applicants and licensees should determine on an individual plant basis the applicability of each
of the following items for inclusion in their submittals.
A. The systems (or portions of systems) for which protection against pipe whip is required should be identified. Protection from pipe whip need not be provided if
any of the following conditions exist:
- 1. Both of the following piping system conditions are met:
- a. The service temperature is less than 200°F.
- b. The design pressure is 275 psig or less.
- 2. The piping is physically separated (o r isolated) from structures, systems, or components important to safety by protective barriers, or restrained
from whipping by plant design features, such as concrete encasement.
- 3. Following a single break, the unrestrained pipe movement of either end of the ruptured pipe in any possible direction about a plastic hinge formed at
the nearest pipe-whip restraint cannot impact any structure, system, or
component important to safety.
FNP-FSAR-3K 3K.A-2 REV 21 5/08 4. The internal energy level (a) associated with the whipping pipe can be demonstrated to be insufficient to impair the safety function of any
structure, system, or component to an unacceptable level.
B. Design basis break locations should be selected in accordance with the following pipe whip protection criteria; however, where pipes carrying high-energy fluid are routed in the vicinity of structures and systems necessary for safe shutdown of
the nuclear plant, supplemental protection of those structures and systems shall
be provided to cope with the environmental effects (including the effects of jet
impingement) of a single postulated open crack at the most adverse location(s)
with regard to those essential structures and systems; the length of the crack
size is taken to be one-half the pipe diameter in length and one-half the wall
thickness in width.
The criteria used to determine the design basis piping break locations in the piping systems should be equivalent to the following:
- 1. ASME Section III Code Class I piping (b) breaks should be postulated to occur at the following locations in each piping run (c) or branch run:
- a. The terminal ends.
- a. The internal fluid energy level associated with the pipe- break reaction may take into account
any line restrictions (e.g., flow limiter) between the pressure source and break location, and the
effects of either single-ended or double- ended flow conditions, as applicable. The energy level
in a whipping pipe may be considered as insufficient to rupture an impacted pipe of equal or
greater nominal pipe size and equal or heavier wall thickness.
- b. Piping is a pressure-retaining component consisting of straight or curved pipe and pipe
fittings (e.g., elbows, tees, and reducers).
- c. A piping run interconnects components, such as pressure vessels, pumps, and rigidly fixed
valves, that may act to restrain pipe movement beyond that required for design thermal
displacement. A branch run differs from a piping run only in that it originates at a piping
intersection as a branch of the main pipe run.
FNP-FSAR-3K 3K.A-3 REV 21 5/08 b. Any intermediate locations between terminal ends where the primary plus secondary stress intensities S n (circumferential or longitudinal) derived on an elastically calculated basis under the
loadings associated with 1/2 safe shutdown earthquake and
operational plant conditions (a) exceed 2.0 S m (b) for ferritic steel and 2.4 S m for austenitic steel.
- c. Any intermediate locations between terminal ends where the cumulative usage factor (U)(c) derived from the piping fatigue analysis and based on all normal, upset, and testing plant
conditions exceeds 0.1.
- 2. ASME Section III Code Class 2 and 3 piping breaks should be postulated to occur at the following locations in each piping run or branch run:
- a. The terminal ends.
- b. Any intermediate locations between terminal ends where either the circumferential or longitudinal stresses derived on an
elastically calculated basis under the loadings associated with
seismic events and operational plant conditions exceed 0.8.
- a. Operational plant conditions include normal reactor operation, upset conditions (e.g.,
anticipated operational occurrences) and testing conditions.
- b. S m is the design stress intensity as specified in Section III of the ASME Boiler and Pressure Vessel Code, "Nuclear Plant Components."
- c. U is the cumulative usage factor as specified in Section III of the ASME Boiler and Pressure
Vessel Code, "Nuclear Power Plant Components."
FNP-FSAR-3K 3K.A-4 REV 21 5/08 (S h + S A)(a) or the expansion stresses exceed 0.8 S A.
The requirement to postulate arbitrary intermediate breaks has been eliminated from the structural design basis (including resultant dynamic
and environmental effects) as allowed by NRC Generic Letter 87-11, "Relaxation in Arbitrary Intermediate Pipe Rupture Requirement".
C. The criteria used to determine the pipe break orientation at the break locations as specified under (b) above should be equivalent to the following:
- 1. Longitudinal (b) breaks in piping runs and branch runs, 4-inches nominal pipe size and larger.
- 2. Circumferential (c) breaks in piping runs and branch runs exceeding 1-inch nominal pipe size.
D. A summary should be provided of the dynamic analyses applicable to the design of Category I piping and associated supports which determine the resulting
loadings as a result of a postulated pipe break including:
- a. S h is the stress calculated by the rules of NC-3600 and ND-3600 for Class 2 and 3 components, respectively, of the ASME Code Section III Winter 1972 Addenda.
S A is the allowable stress range for expansion stress calculated by the rules of NC-3600 of the ASME Code,Section III, or the USA Standard Code for Pressure Piping, ANSI B31.1.0-1967.
- b. Longitudinal breaks are parallel to the pipe axis and oriented at any point around the pipe
circumference. The break area is equal to the e ffective cross-sectional flow area upstream of the break location. Dynamic forces resulting from such breaks are assumed to cause lateral
pipe movements in the direction normal to the pipe axis.
- c. Circumferential breaks are perpendicular to the pipe axis, and the break area is equivalent to
the internal cross-sectional area of the ruptured pipe. Dynamic forces resulting from such
breaks are assumed to separate the piping axially and cause whipping in any direction normal
to the pipe axis.
FNP-FSAR-3K 3K.A-5 REV 21 5/08 1. The locations and number of design basis breaks on which the dynamic analyses are based.
- 2. The postulated rupture orientation, such as a circumferential and/or longitudinal break(s), for each postulated design basis break location.
- 3. A description of the forcing functions used for the pipe-whip dynamic analyses, including the direction, rise time, magnitude, duration, and initial
conditions that adequately represent the jet-stream dynamics and the
system-pressure difference.
- 4. Diagrams of mathematical models used for the dynamic analysis.
- 5. A summary of the analyses which demonstrates that unrestrained motion of ruptured lines will not damage, to an unacceptable degree, structures, systems, or components important to safety, such as the control room.
E. A description should be provided of the measures, as applicable, to protect against pipe whip, blowdown jet, and reactive forces, including:
- 1. Pipe restraint design to prevent pipe whip impact.
- 2. Protective provisions for structures, systems, and components required for safety against pipe whip, blowdown jet, and reactive forces.
- 3. Separation of redundant features.
- 4. Provisions to separate physically piping and other components of redundant features.
- 5. A description of the typical pipe-whip restraints and a summary of number and location of all restraints in each system.
F. The procedures that will be used to evaluate the structural adequacy of Category I structures and to design new seismic Category I structures should be
provided, including:
- 1. The method of evaluating stresses, e.g., the working stress method and/or the ultimate strength method that will be used.
- 2. The allowable design stresses and/or strains.
- 3. The load factors and the load combinations.
G. The structural design loads should be provided. They include the pressure and temperature transients; the dead, live, and equipment loads; and the pipe and
equipment static, thermal, and dynamic reactions.
FNP-FSAR-3K 3K.A-6 REV 21 5/08 H. Seismic Category I structural elements, such as floors, interior and exterior walls, building penetrations, and the buildings as a whole, should be analyzed for
eventual reversal of loads due to the postulated accident.
I. If new openings are to be provided in existing structures, the capabilities of the modified structures to carry the design loads should be demonstrated.
J. Verification that failure of any structure, including nonseismic Category I structures, caused by the accident, will not cause failure of any other structure in
a manner to adversely affect:
- 1. Mitigation of the consequences of the accidents.
- 2. Capability to bring the unit(s) to a cold shutdown condition.
K. Verification that rupture of a pipe carrying high-energy fluid will not directly or indirectly result in:
- 1. Loss of required redundancy in any portion of the protection system (as defined in IEEE-279), Class IE electric system (as defined in IEEE-308),
engineered safety feature equipment, cable penetrations, or their
interconnecting cables required to mitigate the consequences of that
accident and place the reactor(s) in a cold-shutdown condition.
- 2. Environmentally induced failures caused by a leak or rupture of the pipe, which would not of itself result in protective action but does disable
protection functions. In this regard, a loss of redundancy is permitted but a
loss of function is not permitted. For such situations plant shutdown is
required.
L. Assurance should be provided that the control room will be habitable and its equipment functional after a steam-line or feedwater-line break or that the
capability for shutdown and cooldown of the unit(s) will be available in another
habitable area.
M. Environmental qualification should be demonstrated by test for that electrical equipment required to function in the st eam-air environment resulting from a
high-energy-fluid-line break. The information required for our review should
include the following:
- 1. Identification of all electrical equipment necessary to meet requirements of K above. The time after the accident in which they are required to operate
should be given.
- 2. The test conditions and the results of test data showing that the systems will perform their intended function in the environment resulting from the
postulated accident and time interval of the accident. Environmental
conditions used for the tests should be selected from a conservative
evaluation of accident conditions.
FNP-FSAR-3K 3K.A-7 REV 21 5/08 3. The results of a study of steam systems identifying locations at which barriers will be required to prevent st eam jet impingement from disabling a protection system. The design criteria for the barriers should be stated and
the capability of the equipment to survive within the protected environment
should be described.
- 4. An evaluation of the capability for safety-related electrical equipment in the control room to function in the environment that may exist following a
pipe-break accident. Environmental conditions used for the evaluation
should be selected from conservative calculations of accident conditions.
- 5. An evaluation to ensure that the onsite power distribution system and onsite sources (diesels and batteries) will remain operable throughout the event.
N. Design diagrams and drawings of the steam and feedwater lines, including branch lines, showing the routing from containment to the turbine building should
be provided. The drawings should show elevations and include the location
relative to the piping runs of safety-related equipment, including ventilation
equipment, intakes, and ducts.
O. A discussion should be provided of the potential for flooding of safety-related equipment in the event of failure of a feedwater line or any other line carrying
high-energy fluid.
P. A description should be provided of the quality control and inspection programs that will be required or have been utilized for piping systems outside
containment.
Q. If leak-detection equipment is to be used in the proposed modifications, a discussion of its capabilities should be provided.
R. A summary should be provided of the emergency procedures that would be followed after a pipe-break accident, including the automatic and manual
operations required to place the reactor unit(s) in a cold-shutdown condition. The
estimated times following the accident for all equipment and personnel
operational actions should be included in the procedure summary.
S. A description should be provided of the seismic and quality classification of the high-energy-fluid piping systems, includi ng the steam and feedwater piping that runs near structures, systems, or components important to safety.
T. A description should be provided of the assumptions, methods, and results of analyses, including steam- generator blowdown, used to calculate the pressure
and temperature transients in compartments, pipe tunnels, intermediate
buildings, and the turbine building following a pipe rupture in these areas. The
equipment assumed to function in the analyses should be identified, and the
capability of systems required to function to meet a single active component
failure should be described.
FNP-FSAR-3K 3K.A-8 REV 21 5/08 U. A description should be provided of the methods or analyses performed to demonstrate that there will be no adverse effects on the primary and/or
secondary containment structures due to a pipe rupture outside these structures.
FNP-FSAR-3K 3K.A-9 REV 21 5/08 PART II POSTULATED BREAK AND LEAKAGE LOCATIONS IN THE MAIN STEAM LINE High-Energy Fluid System Piping A. Fluid Systems Separated from Essential Systems and Components
For the purpose of satisfying the separation provisions of plant arrangement as specified in B.1.a of the Branch Technical Position APCSB 3-1, a review of the
piping layout and plant arrangement drawings should clearly show that the
effects of postulated piping breaks at any location are isolated or physically
remote from essential systems and components. At the designer's option, break
locations as determined from B.1.c and B.1.d of Branch Technical Position MEB
3-1 may be assumed for this purpose.
B. Fluid System Piping in Containment Penetration Area
Breaks need not be postulated in those portions of piping identified in B.2.c of the Regulatory Position APCSB 3-1, provided they meet the requirements of the
ASME Code,Section III, Subarticle NE-1120, and the following additional design
requirements:
- 1. The following design stress and fatigue limits should not be exceeded:
For ASME Code,Section III, Class 2 Piping
- a. The maximum stress ranges as calculated by equation 9 and 10 in Paragraph NC-3652, ASME Code Section III, considering upset
plant conditions (i.e., sustained loads, occasional loads, and thermal
expansion) and an OBE event should not exceed 0.8 (S h + S A).
- b. The maximum stress as calculated by equation 9 in Paragraph NC-3652, under the loadings resulting from a postulated piping
failure of fluid system piping beyond these portions of piping, should
not exceed 1.8 S
- h. The deflection-limited stresses are included in equation 9.
- 2. Welded attachments, for pipe supports or other purposes, to these portions of piping should be avoided except where detailed stress
analyses, or tests, are performed to demonstrate compliance with the
limits of B.1.b(1) of MEB 3-1.
- 3. The number of circumferential and longitudinal piping welds and branch connections should be minimized.
- 4. The length of these portions of piping should be reduced to the minimum length practical.
FNP-FSAR-3K 3K.A-10 REV 21 5/08 5. The design of pipe anchors or restraints (e.g., connections to containment penetrations and pipe whip restraints) should not require welding directly
to the outer surface of the piping (e.g., flued, integrally forged pipe fittings
may be used) except where such welds are 100-percent volumetrically
examinable in service and a detailed stress analysis is performed to
demonstrate compliance with the limits of B.1.b(1) of MEB 3-1.1.
C. Fluid Systems Enclosed Within Protective Structures
- 1. With the exception of those portions of piping identified Part B above, breaks in Class 2 and 3 piping (ASME Code,Section III) should be
postulated at the following locations in those portions of each piping and
branch run within a protective structure or compartment designed to
satisfy the plant arrangement provision of B.1.b or B.1.c of Branch
Technical Position APCSB 3-1:
- a. At terminal ends of the run if located within the protective structure.
Terminal ends include those locations identified in APCSB 3-1, paragraph B.2.c(3).
- b. At intermediate locations selected by one of the following criteria:
- i. At each pipe fitting (e.g., elbow, tee, cross, flange, and nonstandard fitting), welded attachment, and valve.
Where piping contains no fittings, welded attachments, or valves, at one location at each extreme of the piping within
the protective structure. (A terminal end, as determined by
B.1.c(1)(a) of MEB 3-1, may be considered as one of these
extremes.)
ii. At each location where the stresses exceed 0.8 (S h + S A).
NOTES:
- 1. Tees and junctions having comparable sizes and fixtures need not be considered as terminal
ends for purposes of break locations when so justified in the stress analysis.
- 2. Stresses under normal and upset plant conditions, and an OBE event as calculated by
equations 9 and 10, Paragraph NC-3652 of the ASME Code,Section III.
- 3. Select two locations with at least 10-percent difference in stress or, if stresses differ by less
than 10 percent, two locations separated by a change of direction of the pipe run.
FNP-FSAR-3K 3K.A-11 REV 21 5/08 The requirement to postulate arbitrary intermediate breaks at locations where the stresses do not exceed 0.8 (S h + S A) has been eliminated from the structural design basis (including resultant dynamic and environmental effects) as
allowed by NRC Generic Letter 87-11, "Relaxation in
Arbitrary Intermediate Pipe Rupture Requirements".
- 2. The main steam piping downstream of the MSIVs was designed, fabricated, and constructed to the requirements of ANSI B31.1.0-1967
through 1971 addenda, including Code Cases 74 and 95. The main
steam piping upstream of the MSIVs was designed, fabricated, and
constructed to the requirements of ASME Section III, Class 2, through
Summer 1971 addenda. The stress analyses performed on both portions
of piping were carried out using identical methods of analysis. Breaks, and the use of no break criteria, in the ANSI B31.1 portions of this piping
were postulated using the same criteria that were applied to the ASME
Section III portions. The following discussion provides justification for
using this approach.
- c. A comparison of materials, quality assurance, welding heat treatment, and nondestructive examination for piping and fittings for
the main steam system was made between the ASME Section III, Class 2, portion of the system as installed and the ANSI B31.1.0
portion of the system as installed. The results are as summarized
below.
The materials including weld filler metal for both portions have the same physical and chemical properties. The same quality
assurance provisions for welding apply throughout both portions.
The design material specifications require 100-percent radiography
of all longitudinal and circumferential butt welds in both portions of
the system. All welding is post-weld heat treated in both portions;
i.e., stress relief at 1100-1200°F.
- d. The only differences between the as-fabricated piping systems are as follows:
- i. ASME Section III portion requires Code Data Forms; ANSI portion does not. ii. ASME Section III portion requires third party inspection; ANSI portion does not. iii. All piping and associated welding filler metal that is part of the containment penetration is Charpy impact tested.
Impact testing is required by ASME Section III and may be
advisable for cold hydrostatic testing, but is not needed for
system operation as brittle fracture would not occur at
main steam operating temperature.
FNP-FSAR-3K 3K.A-12 REV 21 5/08 In summary the ANSI portion of the main steam piping is equivalent to the ASME Section III portion. However, the "N" stamp cannot be applied to
ANSI piping because of items i and ii above. This difference does not
affect the quality of material or workmanship.
D. Augmented Inservice Inspection
Inservice examination and related des ign provisions in the containment penetration area and throughout the no break region should be in accordance
with the following:
- 1. The protective measures, structures, and guard pipes should not prevent the access required to conduct the inservice examinations specified in the ASME Boiler and Pressure Vessel Code,Section XI, Division 1, "Rules
for Inspection and Testing of Components in Light-Water Cooled Plants."
- 2. For those portions of fluid system piping identified in B.2.c of APCSB 3-1, the extent of inservice examinations completed during each inspection
interval (IWA-2400, ASME Code, 1974 Edition with Addenda through Summer 1975,Section XI) should provide 100-percent volumetric
examination of circumferential and longitudinal pipe welds within the
boundary of these portions of piping to the extent practical.
- 3. The areas subject to examination should be defined in accordance with Examination Categories C-F and C-G for Class 2 piping welds in Table
FNP-FSAR-3K 3K.B-i REV 21 5/08 ATTACHMENT B PIPE WHIP RESTRAINT DESIGN
FNP-FSAR-3K 3K.C-i REV 21 5/08 ATTACHMENT C METHODS USED TO CALCULATE BLOWDOWN RATES FOR HIGH ENERGY FLUID LINE RUPTURES
REV 21 5/08 CRITICAL MASS VELOCITY vs RESERVOIR PRESSURE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.C-1
REV 21 5/08 CRITICAL MASS VELOCITY vs RESERVOIR QUALITY AND RESERVOIR PRESSURE (VIA MOODY CORRELATION)*
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.C-2
REV 21 5/08 REDUCTION IN MASS FLOWRATE DUE TO PIPING FRACTION LOSSES JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.C-3
REV 21 5/08 UNIT 1 SIMPLIFIED SCHEMATIC OF STEAM LINE HEADER AND BREAK LOCATION THREE LOOP PLANTS JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.C-4
REV 21 5/08 MASS RELEASE RATE vs TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.C-5
REV 21 5/08 ENERGY RELEASE RATE vs TIME JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.C-6
REV 21 5/08 CRITICAL MASS VELOCITY AS A FUNCTION AT PRESSURE AND QUALITY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.C-7
REV 21 5/08 ENTHALPY OR FLUID AS A FUNCTION AT PRESSURE AND QUALITY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.C-8
REV 21 5/08 STEAM GENERATOR STEAM MASS AS A FUNCTION AT INITIAL PRESSURE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.C-9
FNP-FSAR-3K 3K.D-i REV 21 5/08 ATTACHMENT D COMPARTMENT PRESSURE TEMPERATURE ANALYSIS COMPUTER PROGRAM DESCRIPTION FNP-FSAR-3K
3K.D-1 REV 21 5/08 ATTACHMENT D COMPARTMENT PRESSURE TEMPERATURE ANALYSIS COMPUTER PROGRAM DESCRIPTION (COPDA)
3K.D.1 INTRODUCTION This appendix describes the analytical techniques used to evaluate high-energy pipe rupture.
3K.D.2 INITIAL COMPARTMENT CONDITIONS The masses of air and water as steam in the compartments are determined using the initial
input conditions of temperature, pressure, relative humidity, and compartment volumes. The
specific humidity of saturated air at the compartment temperature is read from a correlation
table of temperature and water vapor in saturated air. The compartment specific humidity is
obtained by:
SH = (RH) x (SSH)
where:
SH = specific humidity of compartment air, lb steam/lb air
RH = relative humidity of compartment air
SSH = specific humidity of saturated air at compartment temperature, lb steam/lb air.
The vapor pressure of the water is determined by:
PW = SH 0.623 (SH)(PT)+ where:
PW = vapor pressure of water at compartment temperature, psia
PT = total compartment pressure, psia.
The air pressure in the compartment is determined by:
FNP-FSAR-3K
3K.D-2 REV 21 5/08 The mass of air in the compartment is evaluated using the perfect gas law equation:
MA = (T)n R(V)(PA)(144) where:
V = volume of compartment, ft 3
R = gas constant, 1545.3
T = compartment temperature, °R
n = molecular weight of air, 28.97 lb/lb mole
PA = partial pressure of the air lb/in 2
The mass of water in the compartment, MS, is:
MS = (MA)(SH)
The masses of air and water in the remaining compartments are determined in the same
manner.
The internal energy of the air, UA(I), in each compartment is calculated using 0°F as a base:
UA(I) = [CV][MA(I)][TP]
where:
CV = specific heat of air at constant volume, 0.171 Btu/lb-°F
TP = compartment temperature, °F
The internal energy of the water vapor in each compartment is calculated by the equation:
US(I) = [MS(I)][UG]
where:
UG = internal energy of the steam evaluated from the saturated steam tables at the compartment temperature.
FNP-FSAR-3K
3K.D-3 REV 21 5/08 3K.D.3 CONSERVATION OF MASS AND ENERGY IN COMPARTMENTS
The inventory of the total mass and energy in the compartments is maintained from the inlet and
exit flows during the time increment:
()()()()UA(I)UV(I)UV(I)US(I)UW(I)UV(I)MSO HGOMSIHGI(I)SUUS(I)MWOHOMWIHI(I)WUUW(I)
UAO UAI(I)AUUA(I)
MA(I)MV(I)MT(I)MS(I)MW(I)MV(I)MSO MSI(I)SMMS(I)
MWO MWI MW(I)MW(I)MAO MAI(I)AMMA(I)
N N N N N N N N N N N N+=+=+=+=+=+=+=+=+=+= where: Primed () values refer to end of previous time step; all other values refer to current time step.
MA(I) = mass of air in compartment (I), lb
MW(I) = mass of water in compartment (I), lb
MS(I) = mass of steam in compartment (I), lb
MV(I) = mass of water and steam in compartment (I), lb
MT(I) = total mass in compartment (I), lb
MAI = mass of air entering compartment, lb
MAO = mass of air leaving compartment, lb
MWI = mass of water entering compartment, lb
MWO = mass of water leaving compartment, lb
FNP-FSAR-3K
3K.D-4 REV 21 5/08 MSI = mass of steam entering compartment, lb
MSO = mass of steam leaving compartment, lb
UAI = total energy of air entering compartment, Btu
UAO = total energy of air leaving compartment, Btu
HI = enthalpy of water entering compartment (I), Btu/lb
HO = enthalpy of water leaving compartment (I), Btu/lb
HGI = enthalpy of steam entering compartment (I), Btu/lb
HGO = enthalpy of steam leaving compartment (I), Btu/lb
UA(I) = energy in air in compartment (I), Btu
UW(I) = energy in water in compartment (I), Btu
US(I) = energy in steam in compartment (I), Btu
UV(I) = energy in two-phase mixture in compartment (I), Btu
UT(I) = total energy in compartment (I), Btu.
3K.D.4 COMPARTMENT PRESSURE CALCULATIONS
The compartment pressure is calculated using the total mass and energy in the compartment after the flow from the upstream compartments and/or the blowdown has been
added to the compartment inventory of mass and energy. A convergence procedure is used to arrive at the equilibrium thermodynamics conditions in the compartment using temperature as
the trial argument. The equilibrium thermody namic state is considered determined when the trial temperature provides properties such that the ratio of the difference between the trial
energy balance and the energy inventory is less than 0.001. The state properties of the steam
and water mixture at the trial temperature are obtained from the saturation tables. The mass of
steam is then determined by:
MS = VG(VL) (MW) - (V) 1 where:
V = volume of compartment, ft 3
VL = specific volume of water, ft 3/lb FNP-FSAR-3K
3K.D-5 REV 21 5/08 VG = specific volume of steam, ft 3/lb MW 1 = mass of water from previous iteration, lb
The mass of water (MW) is determined by:
MW = MV - MS
A trial energy balance is calculated:
ETRIAL = (MS)(UG) + (MW)(UL) + 0.171(MA)(TP)
The procedure is repeated varying the value of TP until the relation:
001.0 UTETRIAL -UT is satisfied.
If, after establishing the thermodynamic equilibrium conditions, MW 0, the compartment is considered to be superheated. The equilibrium conditions are recalculated by setting the steam mass equal to the vapor mass and calculating the steam pressure at the search temperature by:
PS = 0.5961(MS)
V T PS = pressure of steam, psia
T = compartment search temperature, °R
V = compartment volume, ft 3 0.5961 = )144(183.1545)144( Weight Mole R= The internal energy of the steam at the pressure and temperature is obtained from the
superheat tables and a trial energy balance calculated by:
ETRIAL = (MS)(UG) + 0.171(MA)(TP)
The procedure is repeated varying the value of TP until the relation:
001.0 UTETRIAL -UT is satisfied.
FNP-FSAR-3K
3K.D-6 REV 21 5/08 where:
UL = internal energy of water at the compartment temperature, Btu/lbm
UG = internal energy of steam, Btu/lbm
The total pressure in the compartment is the sum of the steam pressure and the air pressure
with the latter being calculated by:
PA = V688.459TPMA37.0+ where:
0.37 = ()()14497.283.1545 144 Weight Mole R= 3K.D.5 FLOW CALCULATION
Two-flow equations are provided for calculati ng the flow between compartments. The Moody Equation is used for the analysis of reactor cavity pressures resulting from the decompression
of the primary coolant system and for other co mpartments where the blowdown results in single
component two-phase flow fairly early in the tr ansient. A compressible fluid flow equation is used for the analysis of steam generator compartment pressures for the main steam line breaks
and for other compartments where the blowdown re sults in two component two-phase flow for all of the transient or that portion of the transient through the maximum peak pressure.
In the application of the Moody Equation for ca lculating the flow from compartment 1 to
component 2, the flow is assumed to be critical if the pressure in compartment 2 is less than 0.55 times the pressure in compartment 1. If the flow is critical, the throat pressure is set equal
to 0.55 times compartment 1 pressure.
For subcritical flow the form of the Moody equation is:
()()()2 1 2 22X1K2X2VF2X1 K2VG2X2H1HOJGc2 G+xx+xxx= 1 - Corresponds to upstream compartment 2 - Corresponds to downstream compartment
FNP-FSAR-3K
3K.D-7 REV 21 5/08 All the terms of the formula are evaluated using the following equations:
HO1 = MV1 HV1 H2 = HF2 + X2 x HFG2 X2 = SF2- SG2SF2 - SO2 SO2 = SO1 (Since Moody's Model assumes isentropic flow)
SO1 = SF1 + X1 x SFG1 X1 = HF1 - HG1HF1 - H01 K = 2P2P1P224.12VL2VG 3 1x where:
HO1 = stagnation enthalpy of the fluid in compartment 1, Btu/lb
HV1 = internal energy of the vapor in compartment 1, Btu
MV1 = mass of vapor in compartment 1, lb
SO1 = specific stagnation entropy of fluid in compartment 1, Btu/lb - °R
SF1 = specific entropy of water in compartment 1, Btu/lb - °R
SG1 = specific entropy of steam in compartment 1, Btu/lb - °R
HF1 = specific enthalpy of water in compartment 1, Btu/lb
HG1 = specific enthalpy of steam in compartment 1, Btu/lb
SFG1 = specific entropy of vaporization in compartment 1, Btu/lb - °R
P1 = total pressure in compartment 1, psia
X1 = quality (vapor mass flow fraction) in compartment 1
H2 = specific enthalpy of fluid in compartment 2, Btu/lb
FNP-FSAR-3K
3K.D-8 REV 21 5/08 HF2 = specific enthalpy of water in compartment 2, Btu/lb
HFG2 = specific enthalpy of vaporization in compartment 2, Btu/lb
SO2 = specific stagnation entropy of compartment 2, Btu/lb - °R
SF2 = specific entropy of water in compartment 2, Btu/lb - °R
SG2 = specific entropy of steam in compartment 2, Btu/lb - °R
VG2 = specific volume of steam in compartment 2, ft 3/lb VL2 = specific volume of water in compartment 2, ft 3/lb P2 = total pressure in compartment 2, psia
K = slip ratio dimensionless
The state properties for compartments 1 and 2 are obtained from the saturation tables at the
pressures in the compartment.
For critical flow the form of the Moody Equation is:
[]2 3 3 2 3 2 2 1VGTXTVFTXT(1 HT)J(H01Gc2 Gx+xxx= All the terms of the formula are evaluated using the following equations:
HO1 = UV1/MV1
HT = HFT + XT x HFGT XT = SFT SGT SFT SOT SOT = SO1 since isentropic
SO1 = SF1 + X1 x SFG1 X1 = HF1 - HG1HF1 - HO1 FNP-FSAR-3K
3K.D-9 REV 21 5/08 where:
XT = quality at the throat
HT = specific enthalpy of the fluid at the throat
HFT = specific enthalpy of the water at the throat
HFGT = specific enthalpy of the vaporization at the throat
SOT = specific stagnation entropy of the fluid at the throat
SGT = specific entropy of the steam at the throat
SFT = specific entropy of the water at the throat
The other variables were defined previously. The state properties for compartment 1 and the
throat are obtained from the saturation tables at the respective pressures in the
compartment and throat.
The throat pressure is calculated as follows:
PT = P1 x 0.55 With Moody flow for both the subcritical and critic al flow conditions, the calculated value of the flow is decreased to sixty percent of the flow (Moody Multiplier = 0.6).
In the application of the compressible fluid equation, if the ratio of the pressure in
compartment 2 to the pressure in com partment 1 is less than RC as obtained by:
RC = 1K KK1 2+ the flow is considered to be critical.
The form of the flow equation is:
G = 2 11K1K1K 2RH01P1KGc+xxxx+ The isentropic exponent K for the air, steam, and water mixture is calculated by:
K = P1 PA1KA + P1 PS1KGF xx FNP-FSAR-3K
3K.D-10 REV 21 5/08 where:
KA = isentropic value of K for air (= 1.4)
KGF = isentropic value of K for steam-water mixture RHO1 = specific density of the mixture in compartment 1, lb/ft 3
P1 = total pressure of compartment 1, psia
PS1 = pressure of steam in compartment 1
PA1 = pressure of air in compartment 1
RH01 is calculated using the equation:
RH01 = MT1/VOL1
where:
MT1 = total mass of fluid in compartment 1, lb
VOL1 = volume of compartment 1
If the flow is subcritical, the form of the flow equation is:
G = 2 1 K1K K 2)R(R1K KRH01P1Gc2xxxx+ where the terms are as previously defined and R =
1 2 P P The mass flow for both the compressible fluid flow equation and the Moody equation is
calculated by:
total MF = CAGxx air MAF = 1MT1MA MF water MWF = 1MT1MWMF steam MSF = 1MT1MS MF FNP-FSAR-3K
3K.D-11 REV 21 5/08 The energy transferred by the flow is:
air UAF = T1CPMAFxx water UWF = HL MWFx steam USF = HG MSFx where:
A = area of flow path, ft 2
G = mass flow, lb/ft 2-s C = flow coefficient calculated external to code
CP = specific heat of air at constant pressure
HL = enthalpy of water at compartment temperature
HG = enthalpy of steam at compartment temperature
MA1, MW1, MS1, and MT1 are the same as the one previously defined.
The flow coefficient "C" was calculated using the same method as outlined in the COPRA
computer program which has been previously submitted for NRC review in NS-731-TN, "Containment Pressure Analysis," Power and Industrial Division, Bechtel Corporation, San Francisco, California, December 1968.
FNP-FSAR-3K
3K.E-i REV 21 5/08
ATTACHMENT E CALCULATION METHODS FOR COMPARTMENT PRESSURIZATION
FNP-FSAR-3K
3K.E-1 REV 21 5/08 3K.E.1 PRESSURE AND TEMPERATURE ANALYSIS The results presented below have been superseded with regard to the pressure and
temperature transient for the main steam line break in the main steam valve room. The details
and results of the new analysis are presented in appendix 3J. The discussion below is retained for completeness.
3K.E.1.1 Compartment Model The worst break within the main steam room was determined by analysis. The main steam
room is modeled as one large room. Venting to the atmosphere from the main steam room is possible through either the pipe chase or the penthouse.
The worst break within the pipe chase was determined by analysis. Venting to the atmosphere
is either directly from this chase or through the main steam room.
3K.E.1.2 Flow Model Flow coefficients for expansions and contractions were calculated by the methods outlined in
reference 1. When flow was through highly restricted vents, such as through grating, a
conservative flow coefficient was applied. A flow model for the main steam room is given in
figure 3K.E-3. A flow model for the pipe chase is given in figures 3K.E-6 and 3K.E-6A.
3K.E.1.3 Results A double-ended guillotine break in the 36-in. O.D. line in the main steam room results in the
most severe localized pressure response. The pressure reaches a peak of 20.5 psig at 0.123 s.
As the transient continues, heat absorption by t he walls, which is conservatively neglected, in addition to the decrease of the blowdown, will cause the pressure and temperature to decrease.
The temperature and pressure responses for a break in the main steam room are plotted in
figures 3K.E-1, 3K.E-1A, 3K.E-2, and 3K.E-2A. The maximum pressure in the pipe chase is
28.8 psig. The pressure and temperature are plotted in figures 3K.E-4 and 3K.E-5.
The penthouse, which was added to provide both additional volume and venting to the
atmosphere, was optimized by varying both its volume and vent area until construction and
pressurization limitations were satisfied. The final design provides adequate venting so that
overpressurization does not occur for any break in the main steam room.
A pressure-temperature response in the turbine-driven auxiliary feedwater pump room, resulting
from the severance of a 4-in. auxiliary turbine pump steam line, is also analyzed. At the low
flowrate through the 4-in. O.D. line, the valve closure time, approximately 10 seconds, is short
compared to the time to reach 4-percent quality flow -- approximately 60 seconds. The
pressure response curve for this room is shown in figure 3K.E-8.
FNP-FSAR-3K
3K.E-2 REV 21 5/08 In addition to pressure and temperature response curves for the main steam system, additional curves and flow models for the following com partments containing high energy lines have been
calculated using the computer model described in attachment D:
A. Turbine-driven auxiliary feedwater pump room.
B. CVCS letdown heat exchanger room (upper and lower levels).
C. Piping tunnel from the letdown line penetration room, elevation 100 to the CVCS heat exchanger room.
D. Letdown line penetration room elevation 100.
E. Recycle holdup tank compartments (3).
F. BTRS alternate letdown line valve compartment elevation 121.
FNP-FSAR-3K
3K.E-3 REV 21 5/08 REFERENCES
- 1. "Containment Pressure Analysis," NS-731-TN , Bechtel Corporation, Power and Industrial Division, San Francisco, California, December 1968.
- 2. "Subcompartment Pressure and Temperature Transient Analysis," BN-TOP-4, Rev. 1 , Bechtel Power Corporation, October 1977.
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.E-1 BLOWDOWN-AUXILIARY STEAM LINE
time m h (s) (lb/s) (Btu/lb) 0.0 274.9 1191.4 0.156 274.9 1191.4 0.156 91.56 1191.4 10.0 91.56 1191.4
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.E-2 CVCS LETDOWN LINE RUPTURE:
BLOWDOWN-PENETRATION ROOM (el 100 ft)
time m h (s) (lb/s) (Btu/lb) 0.0000 876.00 353.74 0.0201 520.89 353.69 0.0436 389.88 353.60 1.0670 389.88 353.60 1.0670 272.04 353.60 3.2581 272.04 353.60 8.2581 194.94 353.60 8.7609 194.94 353.60 8.7609 0.00 353.60 1.0E+6 0.00 353.60
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.E-3 UNIT 1: CVCS LETDOWN LINE RUPTURE:
BLOWDOWN-LETDOWN HEAT EXCHANGER ROOM
time m h (s) (lb/s) (Btu/lb) 0.00000 632.94 353.7 0.09006 389.88 353.6 2.01000 389.88 353.6 2.01000 194.94 353.6 4.46420 194.94 353.6 9.32370 194.94 353.6 9.32370 45.93 353.6 12.3026 0.00 353.6 1.0E+06 0.00 353.6
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.E-4 UNIT 2: CVCS LETDOWN LINE RUPTURE:
BLOWDOWN-LETDOWN HEAT EXCHANGER ROOM
time m h (s) (lb/s) (Btu/lb) 0.00000 632.94 353.7 0.09006 389.88 353.6 2.01000 389.88 353.6 2.01000 194.94 353.6 8.33410 194.94 353.6 8.33410 0.00 353.6 1.0E+06 0.00 353.6
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.E-5 CVCS LETDOWN LINE RUPTURE:
BLOWDOWN-SEAL WATER HEAT EXCHANGER ROOM
time m h (s) (lb/s) (Btu/lb) 0.0000 876.00 353.74 0.0263 654.10 353.72 0.0637 419.60 353.64 0.0758 389.90 353.60 1.5593 389.90 353.60 1.5593 272.00 353.60 8.2279 272.00 353.60 8.2279 0.00 353.60 1.0E6 0.00 353.60
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.E-6 CVCS LETDOWN LINE RUPTURE:
BLOWDOWN-PIPE TUNNEL
time m h (s) (lb/s) (Btu/lb) 0.0000 9.82 353.74 3.9 9.82 353.74 5.5171 4.37 353.60 445.5171 4.37 353.60 445.5171 0.00 353.60 1.00E+06 0.00 353.60
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.E-7 CVCS LETDOWN LINE RUPTURE:
PEAK TEMPERATURES AND PRESSURES
Temperature Pressure Room (°F) (psig) el 100-ft penetration room 171 2.7 Letdown heat exchanger room 216 2.5 Seal water heat exchanger room 219 2.6 Piping tunnel 216 2.5 el 100-ft rooms 162, 155 175 el 100-ft rooms 160, 161, 163 142 el 121-ft hallway areas Bound by the BTRS line break
el 139-ft hallway areas Bound by the BTRS line break
el 155-ft hallway areas Bound by the BTRS line break
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.E-8 BTRS ALTERNATE LETDOWN LINE RUPTURE:
BLOWDOWN-HOLDUP TANK ROOM (NO. 156)
time m h (s) (lb/s) (Btu/lb) 0.0000 820.8 357.7 0.0843 472.1 357.7 0.1654 358.8 357.7 0.1654 358.8 353.6 1.5704 358.8 353.6 1.5704 265.5 353.6 6.6102 265.5 353.6 6.6102 71.0 353.6 7.30 0.0 353.6 1.0E+06 0.0 353.6
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.E-9 BTRS ALTERNATE LETDOWN LINE RUPTURE:
BLOWDOWN-HOLDUP TANK ROOM (NO. 157)
time m h (s) (lb/s) (Btu/lb) 0.0000 820.8 357.7 0.0952 453.6 357.7 0.1614 358.8 357.7 0.1614 358.8 353.6 1.3773 358.8 353.6 1.3773 265.5 353.6 6.6529 265.5 353.6 6.6529 71.0 353.6 7.45 0.0 353.6 1.0E+06 0.0 353.6
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.E-10 BTRS ALTERNATE LETDOWN LINE RUPTURE: BLOWDOWN UNIT 1 HEAT EXCHANGER ROOM/VALVE COMPARTMENT
time m h (s) (lb/s) (Btu/lb) 0.0000 820.8 357.7 0.1079 412.6 357.7 0.1407 358.8 357.7 0.1407 358.8 353.6 0.7108 358.8 353.6 0.7108 265.5 353.6 6.8906 265.5 353.6 6.8906 71.0 353.6 17.75 71.0 353.6 22.75 71.0 353.6 1.0E+06 0.0 353.6
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.E-11 BTRS ALTERNATE LETDOWN LINE RUPTURE: BLOWDOWN UNIT 2 HEAT EXCHANGER ROOM/VALVE COMPARTMENT
time m h (s) (lb/s) (Btu/lb) 0.0000 820.8 357.7 0.1079 412.6 357.7 0.1407 358.8 357.7 0.1407 358.8 353.6 0.7108 358.8 353.6 0.7108 265.5 353.6 6.8906 265.5 353.6 6.8906 71.0 353.6 8.4 0.0 353.6 1.0E+06 0.0 353.6
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.E-12 BTRS ALTERNATE LETDOWN LINE CRITICAL CRACK: BLOWDOWN UNIT 1 HEAT EXCHANGER ROOM/VALVE COMPARTMENT
time m h (s) (lb/s) (Btu/lb) 0.00000 9.20 357.7 200.00 9.20 357.7 208.64 4.02 357.7 208.64 4.02 353.6 673.16 4.02 353.6 673.16 0.00 353.6 1.0E+06 0.00 353.6
FNP-FSAR-3K
REV 21 5/08 TABLE 3K.E-13 BTRS ALTERNATE LETDOWN LINE RUPTURE PEAK TEMPERATURES AND PRESSURES
Temperature Pressure Room (°F) (psig) Recycle holdup tank room (No. 156) 208 2.3 Recycle holdup tank room (No. 157) 207 2.3 Reheat heat exchanger/valve room 211 2.0 el 121-ft room 207 (hatch area) 170 el 121-ft other hallway areas 139 (rooms 205, 208, 209, 218, 222, and 237) el 100-ft hallway area Bound by the CVCS line break el 139-ft hallway areas 123 el 155-ft hallway areas 107
REV 21 5/08 DIFFERENTIAL PRESSURE ACROSS STEAM ROOM AND PIPE CHASE WALL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-1
REV 21 5/08 TEMPERATURE IN MAIN STEAM ROOM RESULTING FROM A DOUBLE-ENDED MAIN STEAM BREAK IN MAIN STEAM ROOM JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-1A
REV 21 5/08 PRESSURE IN MAIN STEAM ROOM RESULTING FROM A DOUBLE-ENDED BREAK IN MAIN STEAM ROOM JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-2
REV 21 5/08 TEMPERATURE DISTRIBUTION IN PIPE CHASE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-2A
REV 21 5/08 MAIN STEAM ROOM FLOW MODEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-3
REV 21 5/08 UNIT 1 PIPE CHASE PRESSURE (COMPARTMENTS 1, 4)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-4
REV 21 5/08 UNIT 1 PIPE CHASE TEMPERATURES (COMPARTMENT 1)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-5
REV 21 5/08 PIPE CHASE, MAIN STEAM ROOM FLOW MODEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-6
REV 21 5/08 UNIT 1 PIPE CHASE, MAIN STEAM ROOM FLOW MODEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-6A (SHEET 1 OF 2)
REV 21 5/08 UNIT 1 PIPE CHASE, MAIN STEAM ROOM FLOW MODEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-6A (SHEET 2 OF 2)
REV 21 5/08 TURBINE-DRIVEN AUXILIARY FEEDWATER PUMP ROOM JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-7
REV 21 5/08 TURBINE-DRIVEN AUXILIARY FEEDWATER PUMP ROOM JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-7A
REV 21 5/08 AUXILIARY TURBINE ROOM FLOW MODEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-8
REV 21 5/08 el 100-ft PENETRATION ROOM PRESSURES (LETDOWN LINE BREAK IN el 100-ft PENETRATION ROOM)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-9
REV 21 5/08 el 100-ft PENETRATION ROOM TEMPERATURES (LETDOWN LINE BREAK IN el 100-ft PENETRATION ROOM)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-10
REV 21 5/08 UNIT 1 LETDOWN HEAT EXCHANGER ROOM PRESSURES (LETDOWN LINE BREAK IN LETDOWN HEAT EXCHANGER ROOM)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-11
REV 21 5/08 UNIT 1 LETDOWN HEAT EXCHANGER ROOM TEMPERATURES (LETDOWN LINE BREAK IN LETDOWN HEAT EXCHANGER ROOM)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-12
REV 21 5/08 UNIT 2 LETDOWN HEAT EXCHANGER ROOM PRESSURES (LETDOWN LINE BREAK IN LETDOWN HEAT EXCHANGER ROOM)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-13
REV 21 5/08 UNIT 2 LETDOWN HEAT EXCHANGER ROOM TEMPERATURES (LETDOWN LINE BREAK IN LETDOWN HEAT EXCHANGER ROOM)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-14
REV 21 5/08 SEAL WATER HEAT EXCHANGER ROOM PRESSURES (LETDOWN LINE BREAK IN SEAL WATER HEAT EXCHANGER ROOM)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-15
REV 21 5/08 SEAL WATER HEAT EXCHANGER ROOM TEMPERATURES (LETDOWN LINE BREAK IN SEAL WATER HEAT EXCHANGER ROOM)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-16
REV 21 5/08 PIPING TUNNEL PRESSURES (LETDOWN LINE CRITICAL CRACK IN PIPING TUNNEL)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-17
REV 21 5/08 PIPING TUNNEL TEMPERATURES (LETDOWN LINE CRITICAL CRACK IN PIPING TUNNEL)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-18
REV 21 5/08 CVCS LETDOWN LINE RUPTURE FLOW MODEL: A CRITICAL CRACK IN THE PIPING TUNNEL IN UNIT 2 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-19 (SHEET 1 OF 2)
REV 21 5/08 CVCS LETDOWN LINE RUPTURE FLOW MODEL:
LINE BREAK IN el 100-ft PENETRATION ROOM JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-19 (SHEET 2 OF 2)
REV 21 5/08 CVCS LETDOWN LINE RUPTURE FLOW MODEL: UNIT 1 - LINE BREAK IN LETDOWN HEAT EXCHANGER ROOM JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-20 (SHEET 1 OF 2)
REV 21 5/08 CVCS LETDOWN LINE RUPTURE FLOW MODEL: UNIT 2 - LINE BREAK IN LETDOWN HEAT EXCHANGER ROOM JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-20 (SHEET 2 OF 2)
REV 21 5/08 RECYCLE HOLDUP TANK ROOM (NO. 156) PRESSURES (BTRS ALTERNATE LETDOWN LINE BREAK IN THE TANK ROOM - NO. 156)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-21
REV 21 5/08 RECYCLE HOLDUP TANK ROOM (NO. 156) TEMPERATURES (BTRS ALTERNATE LETDOWN LINE BREAK IN THE TANK ROOM - NO. 156)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-22
REV 21 5/08 RECYCLE HOLDUP TANK ROOM (NO. 157) PRESSURES (BTRS ALTERNATE LETDOWN LINE BREAK IN THE TANK ROOM - NO. 157)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-23
REV 21 5/08 RECYCLE HOLDUP TANK ROOM (NO. 157) TEMPERATURES (BTRS ALTERNATE LETDOWN LINE BREAK IN THE TANK ROOM - NO. 157)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-24
REV 21 5/08 UNIT 1 REHEAT HEAT EXCHANGER/VALVE ROOM PRESSURES (BTRS ALTERNATE LETDOWN LINE BREAK IN THE HX/VALVE ROOM IN UNIT 1)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-25
REV 21 5/08 UNIT 1 REHEAT HEAT EXCHANGER/VALVE ROOM TEMPERATURES (BTRS ALTERNATE LETDOWN LINE BREAK IN THE HX/VALVE ROOM IN UNIT 1)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-26
REV 21 5/08 UNIT 2 REHEAT HEAT EXCHANGER/VALVE ROOM PRESSURES (BTRS ALTERNATE LETDOWN LINE BREAK IN THE HX/VALVE ROOM IN UNIT 2)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-27
REV 21 5/08 UNIT 2 REHEAT HEAT EXCHANGER/VALVE ROOM TEMPERATURESS (BTRS ALTERNATE LETDOWN LINE BREAK IN THE HX/VALVE ROOM IN UNIT 2)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGIURE 3K.E-28
REV 21 5/08 UNIT 1 REHEAT HEAT EXCHANGER/VALVE ROOM PRESSURES (BTRS ALTERNATE LETDOWN LINE CRITICAL CRACK IN THE HX/VALVE ROOM)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-29
REV 21 5/08 UNIT 1 REHEAT HEAT EXCHANGER/VALVE ROOM TEMPERATURES (BTRS ALTERNATE LETDOWN LINE CRITICAL CRACK IN THE HX/VALVE ROOM)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-30
REV 21 5/08 UNIT 2 FLOW MODEL OF el-121 HALLWAY AREA JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-31
REV 21 5/08 LINE BREAK: FLOW MODEL OF UNIT 1 HALLWAYS IN el 100, 121, 139, AND 155 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-32 (SHEET 1 OF 2)
REV 21 5/08 LINE BREAK: FLOW MODEL OF UNIT 1 HALLWAYS IN el 100, 121, 139, AND 155 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-32 (SHEET 2 OF 2)
REV 21 5/08 CRITICAL CRACK: FLOW MODEL OF UNIT 1 HALLWAYS IN el 100, 121, 139, AND 155 JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.E-33
FNP-FSAR-3K
3K.F-i REV 21 5/08 ATTACHMENT F METHODS USED TO CALCULATE PIPE WHIP THRUST LOADS AND JET IMPINGEMENT FORCES FNP-FSAR-3K
3K.F-1 REV 21 5/08 ATTACHMENT F ANALYSIS OF PIPE RUPTURE THRUST AND JET FORCES 3K.F.1 GENERAL
Methods for calculating pipe rupture thrust and jet impingement forces are given in this attachment. Single- and two-phase blowdowns are analyzed to evaluate the nature and
magnitude of these forces which result in pipe whip and impingement loads on structures and
preventive barriers.
3K.F.2 JET THRUST FORCES In the event of a high energy pipe break, the fluid blowdown and the propagation of pressure disturbance produce jet loads that may result in pipe whip and jet impingement forces.
Immediately after the break, while the pressure disturbance propagation is settling down and
the blowdown rate is building up, the resulting jet forces are a function of time, asymptotically
acquiring steady-state value provided the system stagnation pressure P o remains constant.
Methods of calculating the steady-state values of these forces are given in the following:
3K.F.2.1 Steady-State Thrust Calculation The generalized steady-state thrust equation as developed by Shapiro (1) is ()eae c eAPP gVm F+= (1) where:
m = fluid mass flowrate (lb m/s) V e = fluid exit velocity (ft/s)
g c = gravitational constant (32.2 lb m ft/lb f s 2)
P e = fluid exit pressure (psf)
P a = ambient pressure (psf)
A e = exit area (ft
- 2)
FNP-FSAR-3K
3K.F-2 REV 21 5/08 A convenient nondimensional thrust can be defined by dividing through by P o and A e obtaining P P - P + P g GV = AP F oae o c eeo (2) One-dimensional continuity, m = VA and the definition G = m/A can be used with equation (2) to obtain the alternate expressions P P-P +P g GV = AP F oae o c eeo (3) and P P-P+p gG= AP F oae o c e 2eo (4) where e is exit mass density (lb m/ft 3). Two blowdown situations are considered for rupture of steam and water lines. They are as
follows:
- 1. Blowdown of steam from superheated or saturated conditions.
- 2. Blowdown of a steam-water mixture or subcooled water.
3K.F.2.2 Saturated Steam
Thrust forces associated with the blowdown of saturated steam are obtained from figure 3K.F-1 (when fL/D effects are considered) or from figure 3K.F-2 (when effects of a flow restrictor only
are considered).
As can be seen from figure 3K.F-2, thrust forces associated with the critical flow of two-phase
mixtures through upstream restrictions are lower than those associated with the blowdown of
saturated steam through the same restriction. Therefore, the thrust forces associated with the
blowdown of saturated steam were used to evaluate the effects of a main steam line rupture.
FNP-FSAR-3K
3K.F-3 REV 21 5/08 3K.F.2.3 Saturated Steam-Water Mixture or Subcooled Water
Although fluid escaping from a rupture in a subcooled system involves a two-phase mixture, the
subcooled forces only were conservatively us ed for the analysis applying the following Moody equation from reference 2.
()c m 2 m m B t gV)(GPP A F+= (5) where
P m = Maximum pressure at the break
G m = Maximum flowrate at the break
V m = Specific volume (V f) at P m P = Atmospheric pressure A B = Break Area
G m and P m were obtained from figures 3K.F-3 and 3K.F-4, respectively, using a stagnation enthalpy (h f) for the system temperatures and the system source pressure P o given in table 3K.F-1. When the corresponding data points did not fall within the envelopes in figures 3K.F-3
and 3K.F-4, a point on the saturated liquid boundary at the system pressure was used to obtain
P m and G m. For conservatism, no D Lf effects were considered for subcooled forces.
3K.F.3 FLUID JET IMPINGEMENT FORCES
In the event of a pipe break, the fluid flowing through the pipe emerges out as a jet impinging at
nearby structures or equipment. Various blowdown situations considered here are described in
subsection 3K.F.2. On emerging from the breakpoint, the jet undergoes free rapid expansion to
the ambient pressure at relatively short distance -- a few diameters of break area. For this asymptotic distance, momentum and shear interactions with jet environment can reasonably be
neglected. As such, applying forward momentum conservation, the total jet force, F j , is constant throughout its travel, and theref ore, as assumed by Moody; (2)
F j = F (6) where F is the total thrust force defined in subsection 3K.F.2. Methods of calculating F are also
given there.
FNP-FSAR-3K
3K.F-4 REV 21 5/08 For the purpose of this attachment, it is further assumed that F j remains constant for all distances beyond the asymptotic area. This assumption is conservative. Therefore, the jet
pressure at any location along the axis of the jet is given by:
(x) A F (x)P j j j= (7) where
A j (x) is the expanded jet area at location x along the jet axis. See figure 3K.F-5 for system geometry.
Moody (2) has developed a simple analytical model for estimating the asymptotic jet area for steam, saturated water, and steam/water bl owdown situations. Evaluations of LOFT (5) experimental results tend to indicate that, for s ubcooled water and steam blowdown situations, the jet area expands uniformly at half angle of about 15 degrees, whereas steam/water
blowdown expands much more rapidly because of large-scale water flashing. Results of
Moody's analytical analysis agree, at least qualitati vely, with LOFT results. In addition, Moody's analytical analysis predicts results of other experiments, as discussed in reference 2.
In this attachment, an empirical approach has been adopted combining Moody's analytical
model with the uniform half angle approach, as shown in figure 3K.F-5. The half angle is conservatively assumed to be = 10 degrees.
According to this empirical model, the distance of jet travel is divided into three regions.
Region 1 extends to the asymptotic area, at which point the jet expansion area is calculated
according to Moody's method; in Region 2, jet area remains constant; then in Region 3, the jet expands at half angle = 10 degrees.
For subcooled water blowdown, this model assumes half angle approach, = 10 degrees, uniformly in all the three regions, since Moody's model is not truly applicable for this case.
To follow Moody, the extent of Region 1 is taken as
x 1 = 5D e (8) and the jet area at location x 1 is given by the equation:
A j (x 1) = R 2 j 1 (9) = jc 1 2 eFg V G)(A where:
D e = Equivalent diameter of pipe break area FNP-FSAR-3K
3K.F-5 REV 21 5/08
A e = Pipe break area
R j1 = Radius of the expanded jet at location x
- 1. R j1 is constant in region 2.
F j = F, thrust force (equation 6)
v 1 = Specific volume. v 1 is calculated as described in reference 2
For two-phase blowdown, mass flowrate G is taken from reference 3. Region 2 extends to the
location x 2 given by:
A j (x 1) = A j (x), x = x 2
where
A j (x) is the jet area in Region 3 and is calculated by any one of the following equations. (See figure 3K.F-6 for jet geometrical configurations):
- 1. Guillotine break:
A j (x) = 2 e) tan D2x (1 Ae+ where = 10 degrees is the half angle of jet expansion
- 2. Longitudinal (slot) break:
A j (x) = ) tan w 2x )(1 tan 2x (1 Ae++ where = 2D e and w =
e D 8 and and w are slot length and width, respectively
- 3. Circumferential crack:
A j (x) = )) tan2(1 x )(1 tan w2x (1 Ae+++
where = 2 1 D e and w =
2 1 wall thickness and and w are slot length and width, respectively
FNP-FSAR-3K
3K.F-6 REV 21 5/08 In Region 1, the additional conservative assumption is made that the jet area increases uniformly from A j at x = 0, to A j (x 1) at x = x 1 , or A j (x) = 1 2 e j 1 exxofor,1 R R x x1A 1+ where j1 e e eRand, A 2 D R== is given by equation 9.
3K.F.3.1 Impingement Loads on Targets
Once the jet area A j is calculated by the method described above, the jet pressure is readily calculated according to equation 7, i.e.,
P j = j j A F and the jet impingement load on the target is given by
F T = P j A te where A te is the effective target area. Calculation of A te for various geometries is outlined below:
- 1. Flat Surface
If the target with physical area A t cancels all the fluid momentum in the jet, then:
A te = A t For the case where target is oriented at angle with respect to the jet axis and there is no flow reversal:
A te = A t sin 2. Pipe Surface
Let
D p = Diameter of pipe, and
FNP-FSAR-3K
3K.F-7 REV 21 5/08 D j = Diameter of jet impinging on pipe
= A 4 j then, for D p > D j A te = CA j where C is pipe curvature factor and =2 C For D p < D j A te = tAC where jptDDA=(conservative approximation)
FNP-FSAR-3K
3K.F-8 REV 21 5/08 REFERENCES
- 1. Shapiro, A. H., The Dynamics and Thermodynamics of Compressible Fluid Flow , Ronald Press Co., New York, 1953.
- 2. Moody, F. J., "Prediction of Blowdown and Jet Thrust Forces," ASME Paper 69 HT-31 , August 6, 1969.
- 3. Moody, F. J., "Max imum Two-Phase Vessel Blowdown from Pipes," APED-4827 (65APE4), General Electric Co., April 20, 1965.
- 4. Moody, F. J. "M aximum Flow Rate of a Singl e Component, Two-Phase Mixture," APED-4378 , General Electric Co., October 25, 1963.
- 5. Dietz, K. A., Editor, Quarterly Technical Report , Engineering and Test Branch, October 1 through December 31, 1967, Phillips Petroleum Company, IDO-17242, May 1968.
- 6. Moody, F. J. "Fluid Reaction and Impi ngement Loads," presented at the Specialty Conference, Structural Design of Nuclear Plant Facilities at Chicago, Illinois, December 17-18, 1973. (Published in Volume 1 of the conference notes.)
FNP-FSAR-3K TABLE 3K.F-1 THRUST LOADS DUE TO A FULL AREA PIPE RUPTURE
REV 21 5/08 System Line Size Temperature
(°F) Pressure (P o) (psig) Thrust Force (lb f) Main steam 32 in. 547 1005 285,000 36 in. 547 1005 278,100 Main feedwater 14 in. 442 1055 122,700 Auxiliary steam 3 in. 547 1005 7,300 4 in. 547 1005 5,400 Auxiliary feedwater 4 in. 442 1055 10,500 8 in. 442 1055 39,400 10 in. 442 1055 62,000 CVCS and BTRS 3 in. 380 550 6,900 Steam generator blowdown 2 in.
547 1055 4,730
REV 21 5/08 FRICTION EFFECT ON STEADY BLOWDOWN FORCE (REF. 6)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.F-1
REV 21 5/08 STEADY BLOWDOWN FORCE WITH RESTRICTION (REF. 6)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.F-2
REV 21 5/08 MAXIMUM STEAM WATER FLOWRATE AND LOCAL STAGNATION PROPERTIES (REF. 4)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.F-3
REV 21 5/08 LOCAL STATIC PRESSURE AND STAGNATION PROPERTIES AT MAXIMUM STEAM/WATER FLOWRATE (REF. 4)
JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.F-4
REV 21 5/08 JET GEOMETRY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.F-5
REV 21 5/08 FLUID JET GEOMETRY JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.F-6
FNP-FSAR-3K
3K.G-i REV 21 5/08 ATTACHMENT G MAIN STEAM ROOM AND PIPE CHASE STRUCTURAL STRESS ANALYSIS FNP-FSAR-3K
3K.G-1 REV 21 5/08 3K.G.1 INTRODUCTION The purpose of this attachment is to describe the analysis performed on the main steam room of
the Farley Nuclear Plant, Unit 1. Since the Unit 2 main steam room is a mirror image of the
Unit 1 main steam room, the results and conclusions of these analyses are applicable to both
units. A computer finite element analysis was undertaken for the walls and slab of the main
steam room because of the relatively complex geometry of the room, the number of possible loading cases involved, and the numerous points of application of loads. A conventional
analysis was performed on the portion of the containment wall adjacent to the main steam room.
These analyses are described in more detail in subsection 3K.G.2 of this attachment.
Subsection 3K.G.3 contains the summary of results and conclusions.
3K.G.2 DESCRIPTION OF ANALYSIS 3K.G.2.1 Finite Element Model The structure under investigation was modeled using a finite element method. The finite
element mesh is shown in figures 3K.G.2-3 and 3K.G.2-4. The model includes the north and
west wall of the main steam room, the slab at elevation 127 ft 0 in., including beams and the
cable chase, and two partition walls in the main steam room between the three main steam
lines. A total of 758 nodal points and 693 quadrilateral and triangular elements were utilized in
this model. Forty-seven beam elements were also used to represent the two beams and the cable chase. The boundary conditions which were input to the program are shown on figures
3K.G.2-3 and 3K.G.2-4. These boundary conditions include partial or complete fixity against
rotation, combined with partial or complete fixity against displacement.
3K.G.2.2 Containment Wall Analysis The method of analysis of the portion of the containment wall adjacent to the main steam room
is a method described by P. P. Bijlaard in a paper titled "Stresses from Radial Loads in
Cylindrical Pressure Vessels." (1) The three worst combinations of pressure loads from a postulated pipe rupture in the main steam room were applied to the containment wall. The
method described in the aforementioned paper was then used to determine the maximum
forces and moments. A conventional working stress design method was then used to evaluate
the stresses in the containment wall and, finally, the margin of safety percentage.
3K.G.2.3 Input Loads
A. Dead Load (D) -
The concrete deadweight of 150 lb/ft 3 was used for the walls and slab.
FNP-FSAR-3K
3K.G-2 REV 21 5/08 B. Live Load (L) -
A uniformly distributed live load of 300 lb/ft 2 was applied to the entire floor slab at elevation 127 ft 0 in. Of this
300 lb/ft 2 , 200 lb/ft 2 was for miscellaneous live loads and 100 lb/ft 2 was for piping and conduit live loads.
C. Thermal Load -
(T a) Operating thermal effect was incorporated according to
the following data and table.
Initial concrete temperature = 70°F
Main steam room operating temperature = 120°F
Operating temperature in other auxiliary
building area = 80°F
Inside (°F) Outside(°F)
Wall A 120 120 Wall B 120 80 Wall C 120 120 Wall D 120 120 Slab 120 80 The resulting thermal stresses from the subsequent higher
temperature in the main steam room following a pipe break
were not incorporated. This is because these stresses
would not occur simultaneously with those from pressure
and jet-impingement forces.
D. Thermal Pipe -
Reaction (R a) Forces to the structures from pipe reactions under thermal
conditions generated by a postulated break were not
included for the same reason as item C above. At the
normal operating condition the effect was examined and
found to be negligible. Hence, the load (R a) was not incorporated in the load combinations of the finite element
analysis.
FNP-FSAR-3K
3K.G-3 REV 21 5/08 E. Pressure (P a) - The calculated pressure values in each main steam room compartment following a pipe break were all multiplied by
a factor, 1.4 x 1.2 = 1.68, to obtain the equivalent static
pressure. The value of 1.4 is a safety factor for the
pressurization calculation, and the value of 1.2 accounts
for the dynamic load factor. In the actual load input, these
equivalent static pressures were further multiplied by 1.25
or 1.50 to comply with the load combinations.
F. Pipe -
Restraint Force (Yr)
The structural steel pipe restraints in the main steam room
are framed together so that the load taken is shared by
each of the concrete separating walls. Any load which is
taken by one restraint is transferred to each of the walls. It
was found that because of the interaction between
restraints and the stiffness of the walls, the displacement
and thus the stresses in each wall are negligible.
Therefore, the finite element model which analyzes the
walls and slab considers the restraint point as a point of
support.
G. Jet Force (Yj) - The jet forces were calculated in accordance with attachment F. These forces, as in the pressure case, were
multiplied by a factor, 1.2 x 1.2 = 1.44, to account for the
safety and dynamic load factors, prior to being applied to
H. Missile Impact -
Load (y m) Pipe restraints are spaced such that no pipe missile will be
generated by, or during, a postulated break.
Consequently, this load was not included in the load
combinations.
I. Seismic Force -
(Feqo, Feqs) 1. Vertical components of the seismic force were superimposed on the dead load as follows:
7 percent g for OBE (1/2 SSE) 1.25 x 7 percent =
8.75 percent g for 1.25 OBE, and 9 percent g for
SSE.
FNP-FSAR-3K
3K.G-4 REV 21 5/08
- 2. Lateral components of the seismic force were applied to the walls as uniformly distributed lateral
pressures. For walls A and B (figure 3K.G.2-4)
which have supports on four sides and, thus, relatively high natural frequencies, the maximum
wall accelerations were taken to be the same as
the maximum floor accelerations as indicated
below. 1.25 (OBE) E-W 0.131 g x 1.25 = 0.164 g N-S 0.122 g x 1.25 = 0.153 g (SSE) E-W 0.157 g N-S 0.157 g
For cantilever walls C and D (figure 3K.G.2-4)
which have relatively low natural frequencies, maximum spectral accelerations were used as
indicated below.
1.25 (OBE) N-S 1.25 x 1.80 g = 2.25 g (SSE) N-S 1.10 g
3K.G.2.4 Load Combinations As required by "Structural Design Criteria for Evaluating the Effects of High-Energy Pipe Breaks on Category I Structures Outside the Containment"-Document (B) of the NRC, the following load
combinations were examined for each postulated break:
- 1) U = D + L + T a + R a + 1.5 P a
- 2) U = D + L + T a + R a + 1.25 P a + 1.0(Y r + Y j + Y m) + 1.25 Feqo
- 3) U = D + L + T a + R a + 1.0 P a + 1.0(Y r + Y j + Y m) + 1.0 Feqs
The values of the input loads D, L, T a , R a , P a , Y r , Y j , Y m , Feqo, and Feqs are described in paragraph 3K.G.2.3.
3K.G.2.5 Description Of Program The program used to analyze the main steam room is a general structural analysis program
originally developed by Edward L. Wilson of the University of California and subsequently improved by Bechtel. This program is called SAP (1.8).
(2)
The purpose of the computer program is to perform linear, elastic analyses of three-dimensional structural systems. The structural systems to be analyzed may be composed of combinations FNP-FSAR-3K
3K.G-5 REV 21 5/08 of a number of structural element types. The present version contains the following element
types:
A. Boundary.
B. Truss.
C. Beam.
D. Curved beam.
E. Plane strain.
F. Membrane (plane stress).
G. Simple plate.
H. Shell.
I. Thick shell.
J. Brick.
K. Axisymmetric ring.
Two elements were utilized in the analysis of the main steam room. These were the shell and
beam elements.
Systems composed of large numbers of joints and elements may be analyzed. There is no limitation in the program on the number of joints, number of elements, number of load cases, or
equation bandwidth. In addition to being able to solve very large structural systems, the
program can also analyze smaller problems with an efficiency comparable to smaller special-purpose programs. The reason for this is the fact that storage requirements of the program are
adjusted dynamically during execution to conform to the actual requirements of the particular problem being considered.
The thin shell element used in this analysis is either a triangular or quadrilateral element of
arbitrary geometry formed from four compatib le triangles. The bending properties of this quadrilateral element are completely described in a paper titled, "A Refined Quadrilateral
Element For Analysis of Plate Bending." (3)
The element employs a partially restrained linear strain triangle to represent the membrane
behavior. As shown in figure 3K.G.2-1, the central node is located at the average of the
coordinates of the four corner nodes. The element has 17 interior degrees of freedom which
are eliminated at the element level prior to assembling; therefore, the resulting quadrilateral
element has 20 degrees of freedom, 5 per node, in the local element coordinate system.
FNP-FSAR-3K
3K.G-6 REV 21 5/08 For flat plates, the stiffness associated with the rotation normal to the shell surface is not
defined; therefore, the appropriate boundary condition must be enforced.
The beam element is a straight, prismatic beam member. Any force and/or moment at either or
both ends of the beam may be released if necessary. The following loads can be directly
applied to the element:
A. Inertia loads.
B. Thermal loads due to uniform temperature difference and temperature gradient.
C. Fixed end forces and moments.
D. Uniformly and linearly distributed loads along the span.
E. Concentrated forces and moments on the span.
Displacements of each node, axial forces, shear forces, and torsional and bending moments at
both ends of the beam are computed.
Each joint in the system may have from 0 to 6 degrees of freedom as required. The user must
ensure that the degrees of freedom specified for a given joint are compatible with the element
types which are adjacent to it. Optimum so lution efficiency is obtained by minimizing the
number of degrees of freedom of the system.
A right-handed orthogonal coordinate system, shown in figure 3K.G.2-2, is used to describe the
geometry of the structure. All joint loads and displacements are defined with reference to this
system. A local coordinate system is used for each element type.
Loads may be applied by means of both point loads acting at the joints and by element loading (e.g. gravity, temperature). Each element may have an unlimited number of loads. Any number of load cases may be analyzed with each load case consisting of an unlimited combination of
element loads and nodal point loads.
There is no size limitation built into the program, so the size of the problem that can be solved
depends only on the machine core capacity. All storage is allocated at the time of execution
and may be adjusted either upward or downward during execution. Therefore, the actual
storage used will conform not only to the size of the structure, but will also conform to the
specific requirements of each phase of the analysis process.
For static analysis, the program is divided into five phases. A machine-dependent overlay
system is used for each phase. These five are executed in the following sequences:
A. Data Input - Joint coordinates and loads are read or generated. As element properties are read or generated, the element stiffness matrices are formed and
placed on tape.
FNP-FSAR-3K
3K.G-7 REV 21 5/08 B. Formation of the global stiffness matrix is accomplished by reading the element stiffness tape and forming the joint equilibrium equations in blocks.
C. Formation of load vectors is accomplished by processing the element loads and nodal loads for each loading case.
D. Equilibrium equations are solved for joint displacements; all load conditions are treated at the same time.
E. From the joint displacements, element stresses are calculated for all load conditions.
The capacity of the program is controlled by the number of joints (nodal points) of the structural
system. All joint data are retained in high-s peed storage during the formation of the element stiffness matrices. For each joint, three coordinates and six boundary condition codes are
required; therefore, the minimum required storage for a given problem is nine times the number
of joints in the system.
Immediately after the joint data are supplied to the program, a relationship between each joint
degree of freedom and the corresponding equation number is established. Each of the six
boundary condition codes for a given joint is replaced by the equation number for that degree of
freedom. Restrained boundary conditions are identified by a zero equation number. Slave degrees of freedom (for beam elements) are identified by a negative joint number of the master
node.
After the coordinates of the joints are supplied and the equation numbers of the degrees of
freedom established, the stiffness and stress-displacement transformation matrices are
calculated for each structural element in the system. Very little additional high-speed storage is
required for this phase since these matrices can be formed and placed on tape storage as the
element properties are read. In addition to the element matrices, the corresponding equation
numbers are written on tape.
The total stiffness matrix is formed by making a pass through the element stiffness matrices and
adding in the appropriate element stiffness coefficients. To minimize the effort in searching
through all the element stiffnesses, the element stiffness matrices for several blocks are
transferred to another storage unit; therefore, in the formation of the next several blocks, the
time required to search for the contributions to these blocks is reduced significantly.
The equilibrium equations (the global stiffness matrix and load vectors) are stored and
transferred in and out of storage in large blocks. The block size is determined automatically at
the time of solution, thus utilizing storage in the most efficient manner for each particular
problem.
The computer program is built around two optional large-capacity linear equation solvers, USOL and SESOL. The procedure used to solve the equations is not significantly different from the method developed by Gauss in 1827. The banded characteristics of the equations are
recognized.
FNP-FSAR-3K
3K.G-8 REV 21 5/08 Operations with zero coefficients are skipped. Data are transferred in and out of high-speed
storage in large blocks; therefore, a small amount of time is lost in the transfer of data. In the
SESOL routine, random access files are used to reduce further the equation solution data
transfer time.
After the joint displacements are calculated, a pass is made through the element
stress-displacement matrix tape, and the element forces and movements are calculated and
printed.
The output for the main steam room analysis includes nodal point displacements and rotations, element membrane force components, and element bending moment components.
3K.G.3
SUMMARY
OF RESULTS AND CONCLUSIONS Sample deflection curves, moment diagrams, and tables of results are shown in figures
3K.G.3-1 through 3K.G.3-5 and tables 3K.G.3-1 through 3K.G.3-7.
The results of this linear, elastic finite element analysis have shown that the walls and slab in
the main steam room are sufficiently strong to resist various combination loads following a
postulated pipe break in the main steam room, with at least an 18-percent margin of safety over
and above the margin provided by the load increases and load factors used in the analysis.
This conclusion is based on an examination of the most critical section of the walls and slab
governing the entire structural strength. Structural capacity is established when this section first
reaches its elastic limit. Due to the great uncertainty involved in a pipe break incident, the
additional strength gained from the structure which, after reaching this elastic limit, would then
undergo a nonlinear process prior to its final collapse, is not taken into consideration in this
evaluation of the structural capability.
A conventional linear elastic analysis of the portion of the containment wall adjacent to the main
steam room indicated that this wall is strong enough to resist the most severe combination of
pressure loads resulting from a postulated pipe break in the main steam room with a 64-percent
margin of safety.
FNP-FSAR-3K
3K.G-9 REV 21 5/08 REFERENCES
- 1. Bijlaard, P. P., "Stresses From Radial Loads in Cylindrical Pressure Vessels," Welding Journal Research Supplement , 1954.
- 2. Structural Analysis Program, Pacific International Computing Corporation.
- 3. "A Refined Quadrilateral Element for Analysis of Plate Bending," Proceedings, (Second) Conference on Matrix Methods in Structural Mechanics , Wright-Patterson AFB, Ohio, 1968.
FNP-FSAR-3K
3K.G-10 REV 21 5/08 NOTES ON VALUES IN TABLES
- 1. The selected points for tabulation are indicated in figures 3K.G.3-1 through 3K.G.3-5.
These locations are considered the possible critical areas when subjected to various
combined loads during a pipe break incident, as described previously.
- 2. For concrete, tensile stress was not considered.
- 3. Except from the operating thermal condition, the membrane stresses were found negligible. Axial compressive forces (Px & Py) indicated in the tables result from pipe
restraint forces in some local areas and were combined with bending moments in stress
calculation.
- 4. The resulting operating thermal stress was found to be compressive across the entire thickness of the walls and slab. To account for the uncertainty involved in the actual
temperature distribution, only maximum thermal compressive stress was added to
concrete, and no reduction was made for tensile reinforcing stress.
- 5. The allowable stress was taken as 85 percent of the specified compressive strength for concrete and 90 percent of the yield strength for reinforcing bars.
REV 21 5/08 THIN SHELL ELEMENT JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.G.2-1
REV 21 5/08 GLOBAL COORDINATE SYSTEM JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.G.2-2
REV 21 5/08 MAIN STEAM ROOM SLAB FINITE ELEMENT MODEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.G.2-3
REV 21 5/08 MAIN STEAM ROOM WALLS FINITE ELEMENT MODEL JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.G.2-4
REV 21 5/08 SAMPLE DEFLECTION AND MOMENT DIAGRAMS WALL A JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.G.3-1
REV 21 5/08 SAMPLE DEFLECTION AND MOMENT DIAGRAMS WALL B JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.G.3-2
REV 21 5/08 SAMPLE DEFLECTION AND MOMENT DIAGRAMS WALL C JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.G.3-3
REV 21 5/08 SAMPLE DEFLECTION AND MOMENT DIAGRAMS WALL D JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.G.3-4
REV 21 5/08 SAMPLE DEFLECTION AND MOMENT DIAGRAMS SLAB JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3K.G.3-5
FNP-FSAR-3L
3L-i REV 21 5/08 3L ASME SECTION III NUCLEAR CLASS AUXILIARY PIPING STRUCTURE ANALYSIS TABLE OF CONTENTS
Page 3L.
1.0 INTRODUCTION
....................................................................................................3L-1
3L.2.0 NUCLEAR CLASS 1..............................................................................................3L-1
3L.2.1 PIPING CONSIDERED..........................................................................................3L-1
3L.2.2 METHODS OF ANALYSIS.....................................................................................3L-2
3L.2.2.1 Class 1 Lines (Including Accumulator Lines).....................................3L-2 3L.2.2.2 Safety Injection Lines (Except Accumulator Lines)............................3L-2
3L.2.3 BREAK POINTS AND WHIP RESTRAINT LOCATIONS.......................................3L-6
3L.2.3.1 Class 1 Lines (Including Accumulator Lines).....................................3L-6 3L.2.3.2 Safety Injection Lines (Except Accumulator Lines)............................3L-6
3L.3.0 NUCLEAR CLASS 2..............................................................................................3L-6
3L.3.1 PIPING CONSIDERED..........................................................................................3L-6
3L.3.2 METHODS OF ANALYSIS.....................................................................................3L-7
3L.3.2.1 Piping System Analysis (ME 632)......................................................3L-7 3L.3.2.2 Linear Elastic Analysis of Piping Systems (ME 101).......................3L-11 3L.3.2.3 Local Stresses in Cylindrical Shells due to External Loadings (ME 210)..........................................................................3L-12
3L.3.3 BREAK POINTS AND WHIP RESTRAINT LOCATIONS.....................................3L-12
FNP-FSAR-3L
3L-1 REV 21 5/08 APPENDIX 3L ASME SECTION III NUCLEAR CLASS AUXILIARY PIPING STRUCTURAL ANALYSIS
3L.
1.0 INTRODUCTION
This appendix was prepared in response to question MEB-2.3A, transmitted by a letter from K.
Kniel (NRC) to A. Barton (APC) on June 14, 1974. The appendix presents a summary of the
analysis for branch lines in the containment which are ASME Section III, Nuclear Class 1 and 2.
The Design Specification for ASME Nuclear Class 1 auxiliary piping requires that a stress
analysis be performed according to the ASME Boiler and Pressure Vessel Code,Section III, Nuclear Power Plant Components, 1971 Edition (including applicable addenda).
3L.2.0 NUCLEAR CLASS 1 This section contains the structural evaluation of ASME III Nuclear Class 1 piping connected to
the reactor coolant piping and inside the containment building and all fittings connecting the
above piping under postulated loading conditions. These loads result from thermal expansion, pressure, weight, earthquake, design basis accident, and plant operational thermal and
pressure transients. Criteria for postulated break locations are specified in subsection 3.6.2.3.
3L.2.1 PIPING CONSIDERED The Class 1 piping considered in this appendix consists of the following lines:
Size (in.) Line 14 Pressurizer surge line 12 Residual heat removal line, loop 1 12 Residual heat removal line, loop 3 12 SIS accumulator line, loop 1 12 SIS accumulator line, loop 2 12 SIS accumulator line, loop 3 3 CVCS normal charging line 3 CVCS alternate charging line 3 CVCS normal letdown line FNP-FSAR-3L
3L-2 REV 21 5/08 In addition to the lines listed above, the safety injection lines (except accumulator lines) were
also considered. The stress analysis results for these lines are given in the applicable Class 1
stress reports.
3L.2.2 METHODS OF ANALYSIS 3L.2.2.1 Class 1 Lines (Including Accumulator Lines)
The analytical methods used in this analysis are described in subsection 5.2.1.10. They consist
of the transfer matrix method and stiffness matrix formation for the static structural analysis, the
response spectrum method for seismic dynamic analysis, and a structural analysis for the effect
of a reactor coolant loop pipe break. The complexity of the piping systems requires the use of a
computer to obtain the displacements, forces, and stresses in the piping and support members.
The computer codes used for the Class 1 piping systems are capable of performing an elastic
analysis of redundant piping systems subjected to thermal, static, and dynamic loads. A
detailed description, the extent of application, and the verification and qualification of the
WESTDYN 7 computer code can be found in topical report WCAP-8252 , Documentation of Selected Westinghouse Structural Analysis Computer Codes (April 1974).
Emergency core cooling system (ECCS) branch lines are analyzed for the effects of postulated
reactor coolant pipe breaks as a faulted condition. Emergency core cooling system lines
attached to both the unbroken loops and to the unbroken legs of the broken loop were
considered. By comparison of the magnitude of the reactor coolant loop (RCL) and reactor
pressure vessel (RPV) LOCA displacements, it was determined that the effects of a crossover
leg break would be the most severe condition for the safety injection system (SIS) accumulator
lines. (The RPV inlet and outlet nozzle breaks impose less severe loading conditions on the
accumulator lines; therefore, the crossover leg break is presented as the limiting case.) A
dynamic analysis was performed on a linear, elastic basis by applying the time-history displacement output of the RCL analysis to the ECCS lines using program FIXFM. The
resultant stresses were combined with other faulted condition stresses to satisfy the ASME
code equation 9 faulted stress intensity limit of 3.0 S
- m.
3L.2.2.2 Safety Injection Lines (Except Accumulator Lines)
The following are descriptions of the computer programs used in stress analysis of the safety injection lines (except accumulator lines). In addition to the following, ME632, described in
subsection 3L.3.2.1, was also used.
3L.2.2.2.1 Thermal Stress Program (ME 662)
Purpose To determine the temperature and stress distributions within a body as a function of time when subjected to thermal and/or mechanical loads. The program is valid for axisymmetric or plane
structures.
FNP-FSAR-3L
3L-3 REV 21 5/08 Method of Analysis The program consists of two parts, each of which can be used separately. The first part calculates steady-state or transient temperature distributions due to temperature or heat
flux inputs. The method used is the finite element technique coupled with a step-by-step time
integration procedure. The program adopts a stepwise description of environmental
temperatures and heat transfer coefficients if they are time dependent. Transient temperature
distributions are calculated from the specified initial temperatures and the step function heat
inputs.
The second part of the program is built on the displacement method of the matrix theory of
structures, which calculates the displacements and stresses within the solids with orthotropic, temperature-dependent nonlinear material properties.
The user has the option of saving the results from part 1 on an external tape. After reviewing
the printout, he can specify the transient states for the stress evaluations. Part 2 then picks up
the necessary information from the tape and performs the calculations.
References
- 1. Wilson, E. and Nickell, S. R. "Application of the Finite Element Method to Heat Conduction Analysis," Nuclear Engineering and Design , Vol. 4, 1966.
- 2. Wilson, E., "Structural Analysis of Axisymmetric Solids," AIAA Journal , Vol. 3, No. 12, December 1965.
Program Verification The program has been verified by comparing its output with the "ASME Program Verification
and Qualification Problem Library," standard thermal problem. The results were acceptable.
3L.2.2.2.2 LOTEMP Program (ME-913)
Purpose This program is used to calculate piping stresses in accordance with the simplified method of
NB-3650 of the ASME Section III Code.
Method of Analysis In order to calculate the stresses and usage factors according to the rules of NB-3650, the
program requires the following input data:
A. Moments due to thermal expansion, deadweight seismic, and seismic movement.
B. Thermal gradient data t 1 , T 2 , T a and T b.
FNP-FSAR-3L
3L-4 REV 21 5/08 C. Material properties, cross-section, pressures, weld information, and component type at each data point of the pipe.
D. Allowable stresses and number of cycles.
The stresses are calculated in accordance with equations 9 through 14 defined in
Section NB-3650 of the ASME Code. The stresses and the usage factor are printed out for
each data point in the analysis.
References
- 1. ASME Sec. III Boiler and Pressure Vessel Code, 1971 Edition.
- 2. ASME Sample Program for Analysis of Class 1 Piping.
Program Verification LOTEMP has been verified by comparing its output with the "ASME Sample Problem," using the identical input information. The LOTEMP results were identical to the results derived in the
sample program.
3L.2.2.2.3 Pipe Thermal Transient Program "DELTA T" (ME 912)
Purpose To calculate temperature gradient across the pipe wall and along the axis of the pipe, per ASME
Section III code.
Method of Analysis ME 912 is developed to calculate the reduced thermal transients along the axial direction of
piping. It also calculates the thermal transient radially across the pipe wall at various locations.
It allows nonuniform initial temperature distribution and time-dependent temperature inputs.
ME 912 prepares all the thermal input, T 1 , T 2 , (T a-T b), etc. for program ME 913, Nuclear Class 1 piping stress analysis, per ASME Section III code.
Program Verification The temperature gradient for various pipe sizes and for various temperatures has been
calculated manually and verified with the results from ME 912. Also, the results from ME 912
have been compared with many commerc ially available programs. The results were very close.
FNP-FSAR-3L
3L-5 REV 21 5/08 3L.2.2.2.4 Local Stress Analysis at Lug Supports on Piping Systems (ME 916)
Purpose To calculate stress intensities and fatigue analysis at the junction of the integral attachment of
lugs and stanchions to the pipe, per ASME Section III, NB-3600 criteria.
Method of Analysis In ME 916 input data needed are lug and pipe size, stress indices, material properties, loading conditions, and input loadings for pipe and lugs. For fatigue analysis, properties of cyclic loads
or load pair set also are defined. The output from this program provides allowable stress
equations, per MB-3600 and a cumulative usage factor.
Verification ME 916 has been verified by a set of hand calculations of stresses as shown in NB-3600, Class I analyses. All stress and usage factors agree with hand calculations and with the results
from many standard commercial programs.
3L.2.2.2.5 ANSYS Program (Rev. 2)
Purpose This program is used to calculate stress displacement and load history as a function of time, caused by transient displacement in the reactor coolant loops during a loss-of-coolant accident
and during major breaks in the loops.
Method of Analysis A time-history displacement profile at the various nozzle connections for each of the postulated
reactor coolant loop break cases was obtained from Westinghouse on computer tapes. Pipings
were modeled as finite element elastic stick and elastic plates. Transient displacements were
applied at the nozzle connections, and reduced linear transient dynamic analysis was
performed.
The results were extracted in three different steps, as follows:
A. Displacement Pass: Displacement time-history at each node of the geometry was obtained.
B. Stress Pass: Stresses, forces, and moments were obtained at each nodal point.
C. Max Pass: Maximum and minimum val ues of displacement, stress, forces, and moments, independent of time, were obtained in this pass.
FNP-FSAR-3L
3L-6 REV 21 5/08 Verification
The ANSYS program has been developed and verified by Swanson Analysis Systems, Inc.
References
- 1. ANSYS Theoretical Manual
- 2. ANSYS User's Information Manual
- 3. ANSYS Verification Manual
3L.2.3 BREAK POINTS AND WHIP RESTRAINT LOCATIONS Stress analysis results utilized in the criteria for determining pipe break locations are
documented in the applicable piping stress calculation for each piping system. Whip restraint
locations based on postulated pipe break locations are shown on applicable civil design
drawings.
3L.2.3.1 Class 1 Lines (Including Accumulator Lines)
The criteria for postulated break locations are specified in subsection 3.6.2.3. It has been
determined that in all cases for these lines, the governing criterion for postulated break locations
is primary and secondary stress intensity range.
3L.2.3.2 Safety Injection Lines (Except Accumulator Lines)
The criteria for postulated break locations are specified in subsection 3.6.2.3.
3L.3.0 NUCLEAR CLASS 2 This section provides the information related to Class 2 piping.
3L.3.1 PIPING CONSIDERED The Class 2 piping considered in this appendix consists of the main steam and main feedwater
piping in the containment.
The stress analysis results utilized in the criteria for determining pipe break locations are
documented in the applicable piping stress calculation for each piping system.
FNP-FSAR-3L
3L-7 REV 21 5/08 3L.3.2 METHODS OF ANALYSIS The following is a description of the computer program used in stress analysis of the systems
listed in 3L.2.2.2 and 3L.3.1, and also a brief description of the programs' assumptions and
theory. All programs conform to the design and control measures required by Appendix B of 10
CFR Part 50.
3L.3.2.1 Piping System Analysis (ME 632)
Purpose The stresses and loads in piping systems due to restrained expansion, deadweight, seismic
movement, and earthquake are calculated using the static analysis computer program.
Method of Analysis The stiffness method of finite element analysis has been used in this program. In this method, the displacements of the joints of a given structure are considered to be the basic unknowns.
The dynamic analysis of the program utilizes the general theory of response analysis by the
modal synthesis methods. The modal synthesis , in principle, exploits known maximum accelerations produced in a single degree of freedom model of certain frequency. The method
is described in detail in the references. The program's principal assumptions are:
A. Linearly elastic structure.
B. Simultaneous displacement of all supports described by a single time dependent function.
C. Lumped mass model satisfactorily replaces the structure.
D. Modal synthesis is applicable.
E. Rotational inertias of the masses have negligible effect.
Static Analysis For gravity, thermal, and seismic movement analyses, the static load and displacement matrices
were formed in addition to the stiffness matrix of the mathematical model. These matrices
included the applied joint forces and displacements, the distributed loading on the mathematical
model, and the thermal forces developed in the me mbers of the model, whichever is applicable.
Once these matrices were formed, the joint displacements of the mathematical model were
found by solving the following equation:
R - Kr = 0 Eq. (1) in which:
R = Joint load matrix FNP-FSAR-3L
3L-8 REV 21 5/08 K = Stiffness matrix of pipe loop
r = Joint displacement matrix
After the joint displacements were determined, the individual member forces were obtained by
using the member stiffness properties, and, finally, the support reactions were calculated.
Dynamic Analysis The dynamic analysis of flexible piping system s is performed using the response spectrum method. A flexible piping system is idealized as a mathematical model consisting of lumped masses connected by massless elastic members. The lumped masses are carefully located so
as to adequately represent the dynamic and elastic properties of the piping system. The three
dimensional stiffness matrix of the mathematic al model is determined by the direct stiffness method. Axial, shear, flexural and torsional deformations of each member are included. For
curved members, a decreased stiffness is used in accordance with ASME Section III. The mass
matrix is also calculated.
After the stiffness and the mass matrix of the mathematical model are calculated, the natural
frequencies of piping system and corresponding mode shapes are determined using the
following equation:
0)MWK(2 n= Eq. (2) where:
K = stiffness matrix W n = natural circular frequency for the nth mode M = mass matrix
N = mode shape matrix for the nth mode 0 = zero matrix
The Givens or the Jacobi method is used in the solution of the above equation. The mode
shapes are normalized as follows:
1M n t n= Eq. (3) A generalized mass matrix is calculated, and should correspond to:
IM t= Eq. (4) where FNP-FSAR-3L
3L-9 REV 21 5/08
= matrix of mode shapes t = transposition of I = identifies matrix
If any one of the off-diagonal terms in the generation of the left-hand side of Equation (4) is
greater than 1 x 10
-4 , the problem is aborted. This occurs when poor or improper modeling of the piping system exists.
The response spectrum method is them used to find the maximum response of each mode:
n 2 n n max (t)nMWSaDM Y n= Eq. (5) where
Sa n = spectral acceleration value for the nth mode
D = earthquake vector matrix, used to introduce earthquake direction to the response analysis
O = transposition of the nth mode shape
M = generalized mass of the nth mode; equals one by Equation (2-2)
Y n = generalized coordinate for the nth mode
Using the maximum generalized coordinate for each mode, the maximum displacements
associated with each mode are calculated:
max n n(t)YV= Eq. (6) Once the appropriate maximum modal displacements have been determined for each mass
point, the effective inertia forces for each mode are computed:
n nVKQ= Eq. (7) where:
Q n = effective inertia force matrix due to nth mode
V = displacement matrix due to nth mode
The effective acceleration for each mode is calculated:
FNP-FSAR-3L
3L-10 REV 21 5/08 n 1 nQMa= Eq. (8) where:
a = effective acceleration matrix due to nth mode
M-1 = the inverse of mass matrix
After the effective inertia forces have been determined, the internal forces and moments for
each mode are also calculated:
nQb n S= Eq. (9) where:
S = internal force and moment matrix due to the nth mode
b = force transformation matrix
The modal stresses are then calculated from the modal internal forces and moments in
accordance with ASME Section III. The analysis is made three times: once for the vertical
direction and once for each of the two principal horizontal directions of the building. The
method of combining the modal responses (i.e., displacements, effective inertia forces, effective
accelerations, internal forces and moments, support reactions, and stresses) is the square root
of the sum of the squares.
References
- 1. Gere, J. M. and Weaver, W. Jr., Analysis of Framed Structures , D. Van Nostrand Co., Inc., 1965.
- 2. Weaver, W. Jr., Computer Program for Structural Analysis, D. Van Nostrand Co., Inc., 1967.
- 3. Roark, R. J. Formulas for Stress and Strain , McGraw-Hill, 1965.
- 4. Morris, D. L. "Curved Beam Stiffness Coefficients," Struct. Div. Journal, ASCE , May 1968. Verification
The program has been verified by comparing its output with the "ASME Program Verification and Qualifications Program Library," standard problems. The results were acceptable.
FNP-FSAR-3L
3L-11 REV 21 5/08 3L.3.2.2 Linear Elastic Analysis of Piping Systems (ME 101)
Purpose This program serves the same purpose as ME 632. In addition it forms the stress equations, as
defined in ANSI B31.1 and ASME Section III, from individual loading conditions and satisfies
them.
Method of Analysis
This program replaces the program ME 632 and has almost the same features. The basic
method of analysis is the same as ME 632 discussed in subsection 3L.2.2.2.3.
The development of ME 101 is intended to produce a more efficient and systematic piping
program. ME 101 is structured so as to allow easy incorporation of changes and any further
enhancements.
ME 101 has the capability of performing stress combinations, per ASME Section III and
ANSI B31.1 codes and of satisfying appropriate equations. In analysis it incorporates NRC
Regulatory Guide 1.92. It prepares load summary sheets and stress summary sheets for stress
reports.
Verification
The program has been verified by a series of hand calculations and by comparing the results of the program with the results from commercially available standard computer programs.
3L.3.2.3 Local Stresses in Cylindrical Shells due to External Loadings (ME 210)
Purpose To calculate local stresses caused in pipe walls due to external loading on lugs or stanchions
attached integrally to the pipe.
Method of Analysis
ME 210 is based on WRC Bulletin 107 for local stresses in cylindrical shells due to external
loading. In this program induced stresses in the pipe walls, due to loads applied on lugs and
stanchions, are calculated and combined with stresses obtained in ME 101 analysis, to satisfy
stress equations, per ASME Section III, NC and MD-3600, and ANSI B31.1.
Program Verification
The program has been verified through a set of hand calculations using procedures outlined in
WRC Bulletin No. 107.
FNP-FSAR-3L
3L-12 REV 21 5/08 References
- 1. Forsythe, G. E. and Wasow, W. R., Finite-Difference Methods for Partial Differential Equations , John Wiley, 1960, pp 101-107, 119-125.
- 2. Holman, J. P., Heat Transfer , Third Edition, McGraw-Hill, 1972, Eqs. (6.1) and (6.29).
- 3. McNeill, D. R. and Brock, J. B., "Charts for Transient Temperatures in Pipes," in Heating/Piping/Air Conditioning, November 1971, pp 107-119.
- 4. Tung, T. K., "Thermal Gradients in Pipe Walls Due to Ramps in Fluid Temperature," Report BR-5853-T-012 , Bechtel Power Corporation, San Francisco, 1975.
- 5. Tung, T. K., "Analysis on Axial Discontinuity Temperature Difference in Pipe Walls," to be reported.
3L.3.3 BREAK POINTS AND WHIP RESTRAINT LOCATIONS
Break points are postulated in accordance with the requirements set forth in attachment A of
appendix 3K. Whip restraint locations based on postulated pipe break locations are shown on
applicable civil design drawings.
FNP-FSAR-3M
3M-i REV 21 5/08 3M REACTOR PRESSURE VESSEL SUPPORT LOADS TABLE OF CONTENTS
3M.1 INTRODUCTION.........................................................................................................3M-1
3M.2 INTERFACE INFORMATION......................................................................................3M-1
3M.3 LOADING CONDITIONS.............................................................................................3M-2
3M.4 REACTOR VESSEL AND INTERNALS MODELING..................................................3M-2
3M.5 ANALYTICAL METHODS...........................................................................................3M-3
3M.6 RESULTS OF THE ANALYSIS...................................................................................3M-4
FNP-FSAR-3M
3M-ii REV 21 5/08 LIST OF TABLES 3M-1 Maximum Reactor Vessel Displacements at Reactor Vessel Centerline
3M-2 Maximum Reactor Vessel Support Loads for Postulated Pipe Rupture Conditions
3M-3 Maximum Reactor Vessel Support Loads for Combined Pipe Rupture Condition, Safe Shutdown Earthquake, and Deadweight
FNP-FSAR-3M
3M-iii REV 21 5/08 LIST OF FIGURES
3M-1 Reactor Vessel Support Shoe
3M-2 Reactor Vessel Support Box
3M-3 Mathematical Model for Horizontal Response
3M-4 Mathematical Model for Vertical Response
FNP-FSAR-3M
3M-1 REV 21 5/08 APPENDIX 3M REACTOR PRESSURE VESSEL SUPPORT LOADS
3M.1 INTRODUCTION This appendix presents the method of computing the reactor pressure vessel loss-of-coolant accident (LOCA) support loads and displacements. The structural analysis considers
simultaneous application of the time history loads on the reactor vessel resulting from the
reactor coolant loop vessel nozzle mechanical loads, internal hydraulic pressure transients, and
reactor cavity pressurization (for postulated breaks in the reactor coolant pipe at the vessel
nozzles). The vessel is restrained by reactor vessel support pads and shoes beneath each
nozzle, and the reactor coolant loops with the primary supports of the steam generators and the
reactor coolant pumps. The objective of this analysis is to obtain reactor vessel displacements
and reactor vessel support loads.
Pipe displacement restraints installed in the primary shield wall limit the break opening area of
the vessel nozzle pipe breaks to less than 100 in.
2 for the inlet nozzle and 30 in.
2 for the outlet nozzle. These areas were determined to be an upper bound by using worst case vessel and
pipe relative motions based on similar plant analyses. Detailed studies have shown that pipe
breaks at the hot or cold leg reactor vessel nozzles, even with a limited break area, would give
the highest reactor vessel support loads and the highest vessel displacements, primarily due to
the influence of reactor cavity pressurization. By considering these breaks, the most severe
reactor vessel support loads are determined. For completeness, a break outside the shield
wall, for which there is no cavity pressurization, is also analyzed; specifically, the pump outlet
nozzle pipe break is considered. In summary, three loss of coolant accident conditions are
analyzed:
A. Reactor vessel inlet nozzle pipe break.
B. Reactor vessel outlet nozzle pipe break.
C. Reactor coolant pump outlet nozzle pipe break.
3M.2 INTERFACE INFORMATION Bechtel Power Corporation performed the reactor containment design and analysis. Stiffness of
the primary shield wall beneath the reactor vessel supports and asymmetric cavity
pressurization loading was provided by Bechtel to Westinghouse. Cavity pressure loads were
provided as force time histories acting on the reactor vessel.
All other input information was developed within Westinghouse. These items are reactor
internals properties, loop mechanical loads and loop stiffness, internal hydraulic pressure
transients, and reactor support stiffnesses. These inputs allowed formulation of the
mathematical models and performance of the analyses, as will be described.
FNP-FSAR-3M
3M-2 REV 21 5/08 3M.3 LOADING CONDITIONS
Following a postulated pipe rupture at the reactor vessel nozzle, the reactor vessel is excited by
time history forces. As described, these forces are the combined effect of three phenomena:
reactor coolant loop mechanical loads, reactor cavity pressurization forces, and reactor internal
hydraulic forces.
The reactor coolant loop mechanical forces are derived from the elastic dynamic analyses of the
loop piping for the postulated break. This analysis is described in subsection 5.2.1.10.1.1. The
dynamic reactions on the nozzles of all the unbroken piping legs are applied to the vessel in the
RPV blowdown analysis.
Reactor cavity pressurization forces arise for the pipe breaks at the vessel nozzles from the
steam and water which is released into the reactor cavity through the annulus around the
broken pipe. The reactor cavity is pressurized asymmetrically with higher pressure on the side
adjacent to the break. These differences in pressure horizontally across the reactor cavity
result in horizontal forces applied to the reactor vessel. Smaller vertical forces arising from
pressure on the bottom of the vessel and the vessel flanges are also applied to the reactor
vessel. The cavity pressure analysis is described in section 6.2.
The internals reaction forces develop from asymmetric pressure distributions inside the reactor
vessel. For a vessel inlet nozzle break and pump outlet nozzle break, the depressurization
wave path is through the broken loop inlet nozzle and into the region between the core barrel
and reactor vessel. (See figure 3.9-1.) This region is called the downcomer annulus. The initial
waves propagate up, down and around the downcomer annulus and up through the fuel. In the
case of an RPV outlet nozzle break, the wave passes through the outlet nozzle and directly into
the upper internals region, depressurizes the core, and enters the downcomer annulus from the
bottom of the vessel. Thus, for an outlet nozzle break, the downcomer annulus is
depressurized with much smaller differences in pressure horizontally across the core barrel than
for the inlet break. For both the inlet and outlet nozzle breaks, the depressurization waves
continue their propagation by reflection and transmission through the reactor vessel fluid but the
initial depressurization wave has the greatest effect on the loads.
The reactor internals hydraulic pressure transients were calculated including the assumption
that the structural motion is coupled with the pressure transients. This phenomena has been
referred to as hydroelastic coupling or fluid-structure interaction. The hydraulic analysis
considers the fluid structure interaction of the core barrel by accounting for the deflections of
constraining boundaries which are represented by masses and springs. The dynamic response
of the core barrel in its beam bending mode responding to blowdown forces compensates for
internal pressure variation by increasing the volume of the more highly pressurized regions.
The analytical methods used to develop the reactor internals hydraulics are described in
WCAP-8708.(1) 3M.4 REACTOR VESSEL AND INTERNALS MODELING The reactor vessel and internals general assembly is shown in figure 3.9-1. The reactor vessel
is restrained by two mechanisms: the three attached reactor coolant loops with the steam
generator and reactor coolant pump primary supports, and six reactor vessel supports, one FNP-FSAR-3M
3M-3 REV 21 5/08 beneath each reactor vessel nozzle. The reactor vessel supports are described in
subsection 5.5.14 and are shown in figures 5.5-7, 3M-1, and 3M-2. The support shoe provides
restraint in the horizontal directions and for downward reactor vessel motion.
The reactor vessel model consists of two separate nonlinear elastic models connected at a
common node. One model represents the dynamic ve rtical characteristics of the vessel and its internals, and the other model represents the translational and rotational characteristics of the
structure. These two models are combined in the DARI-WOSTAS code (2) to represent motion of the reactor vessel and its internals in the plane of the vessel centerline and the broken pipe
centerline.
The model for horizontal motion is shown in figure 3M-3. Each node has one translational and
one rotational degree of freedom in the vertical plane containing the centerline of the nozzle
attached to the broken pipe and the centerline of the vessel. A combination of beam elements
and concentrated masses are used to represent the components including the vessel, core
barrel, neutron panels, fuel assemblies, and upper support columns. Connections between the
various components are either pin-pin rigid links, translational impact springs with damping, or
rotational springs.
The model for vertical motion is shown in figure 3M-4. Each mass node has one translational
degree of freedom. The structure is represented by concentrated masses, springs, dampers, gaps, and frictional elements. The model includes the core barrel, lower support columns, bottom nozzles, fuel rods, top nozzles, upper support columns, upper support structure, and
reactor vessel.
The horizontal and vertical models are coupled at the elevation of the primary nozzle
centerlines. Node 1 of the horizontal model is coupled with node 2 of the vertical model at the
reactor vessel nozzle elevation. This coupled node has external restraints characterized by a
3 x 3 matrix which represents the reactor coolant loop stiffness characteristics, by linear
horizontal springs which describe the tangential resistance of the supports, and by individual
nonlinear vertical vessel support dynamic elem ents (spring dashpot system) which provide restraint only in the vertically downward direction. The supports as represented in the horizontal
and vertical models (figures 3M-3 and 3M-4) are not indicative of the complexity of the support
system used in the analysis. The individual supports are located at the actual support pad
locations and accurately represent the independent nonlinear behavior of each support.
3M.5 ANALYTICAL METHODS The time-history effects of the cavity pressurization loads, internals loads and loop mechanical
loads are combined and applied simultaneously to the applicable nodes of the mathematical
model of the reactor vessel and internals. The anal ysis is performed by numerically integrating the differential equations of motion to obtain the transient response. The output of the analysis
includes, among other things, the displacements of the reactor vessel and the loads in the
reactor vessel supports. The loads from the postulated pipe break on the vessel supports are
combined with other applicable faulted condition loads and subsequently used to calculate the
stresses in the supports. Also, the reactor coolant loop is analyzed by applying the reactor
vessel displacements to the reactor coolant loop model. The resulting loads and stresses in
the piping, components, and supports are then combined with those from the loop dynamic FNP-FSAR-3M
3M-4 REV 21 5/08 blowdown analysis, and the adequacy of the system is verified. Thus, the effect of vessel
displacements upon loop response and the effect of loop blowdown upon vessel displacements
are both evaluated.
3M.6 RESULTS OF THE ANALYSIS As described, the reactor vessel and internals were analyzed for three postulated break
locations. Table 3M-1 summarizes the displacements and rotations of and about a point
representing the intersection of the nozzle centerline of the nozzle attached to the leg in which
the break was postulated to occur and the vertical centerline of the reactor vessel. Positive
vertical displacement is up, and positive horiz ontal displacement is away from and along the centerline of the vessel nozzle in the loop in which the break was postulated to occur. These
displacements were calculated using an assumed break opening area for the postulated pipe
ruptures at the vessel nozzles of 144 in.
2(a) and a double-ended rupture at the pump outlet nozzle. These areas are estimated prior to performing the analysis. Following the reactor
coolant system structural analysis, the relative motions of the broken pipe ends are obtained
from the reactor vessel and reactor coolant loop blowdown analyses. These motions resulted in
an average break opening area of less than 85 in.
2 (100 in.2 , peak) for the vessel inlet nozzle break and 15 in.
2 (23 in.2 , peak) for the vessel outlet nozzle break. Since these areas are less than the areas used to generate the applied loads, the system structural analysis is
conservative.
The maximum loads induced in the vessel supports due to the postulated pipe break are given
in table 3M-2. These loads are per vessel support and are applied at the vessel nozzle pad. It
is conservatively assumed that the maximum horizontal and vertical loads occur simultaneously and on the same support, even though the time-history results show that these loads do not
occur simultaneously on the same support. The peak vertical load occurs for a vessel inlet
nozzle break; the peak horizontal load occurs for the vessel outlet nozzle break. Note that the
peak horizontal load is an extremely conservative value since the break opening area for the
vessel outlet nozzle break is only 15 in.
2 instead of 144 in.
2(a) area used to generate the applied loads. If additional analysis were performed using the lower break opening area, the load would
be considerably reduced. Furthermore, the peak vertical load and peak horizontal load do not
occur on the same vessel support. The largest vertical loads are produced on the supports
beneath and opposite the broken nozzle. The largest horizontal loads are produced on the
supports which are the most perpendicular to the broken nozzle horizontal centerline.
The LOCA loads are combined with other applicable faulted condition loads, and the total
applied loads are obtained. These total loads on a per support basis are summarized in table
3M-3. This total combined load is applied to the reactor vessel supporting structure, which is
analyzed into two independent components: the U-shaped vessel shoe (figure 3M-1), and the
cooling box, which is the structure between the shoe and the concrete (figure 3M-2). Final
analyses have been performed on the support shoe and the cooling box structure, and the
results are presented in subsection 5.2.1.10.1.1(M).
- a. The maximum break opening area of the inlet nozzle was redetermined to be 100 in.
2 and the maximum break area of the outlet nozzle was redetermined to be 30 in.
2 Only the inlet nozzle break was reanalyzed since it bounds the smaller break in the outlet nozzle.
FNP-FSAR-3M
3M-5 REV 21 5/08 The reactor coolant loop piping was evaluated for the primary membrane plus bending stress
intensity against the faulted- condition stress limit, equation 9 of subarticle 3650 of the ASME
Section III, Nuclear Power Piping Code. The loads included in the evaluation result from the
SSE inertia loading, deadweight, pressure, LOCA loop hydraulic forces, and reactor vessel
motion. Individual loadings at critical stress locations were combined, and primary stress
intensities were calculated for the combined load sets. The primary stress intensities at all
locations were under the faulted condition stress limit. It is therefore concluded that the reactor
coolant loop piping of the unbroken loop or the unbroken legs of the broken loop meets the
faulted condition requirements of ASME Section III and is capable of withstanding the
consequences resulting from a break at the reactor vessel inlet or outlet nozzle.
For the evaluation of the design adequacy of equipment, the maximum loads at the primary
equipment nozzles resulting from the analysis of each loading condition were determined. The
external loads imposed upon primary equipment by the reactor coolant loop produce stress intensities which are below the faulted condition allowable values.
The effects of the postulated breaks at the reactor vessel inlet and outlet nozzle locations on the
CRDM's, reactor vessel internals, ECCS branch lines, RCS component supports, and the
reactor core are presented in subsection 5.2.1.10.1.1 (N), subsection 3.9.3.8, appendix 3L, subsection 5.2.1.10.1.1(M), and subsection 4.2.1.3.2, respectively.
The results of these analyses verify that the integrity of the safeguards systems is assured
during a loss of coolant accident and that the reactor can be safely shut down and maintained in
a safe condition.
FNP-FSAR-3M
3M-6 REV 21 5/08 REFERENCES
- 1. Takeuchi, K., et al, "MULTIFLEX - A Fortran-IV Computer Program for Analyzing Thermal - Hydraulic-Structure System Dynamics," WCAP-8708 , February 1976.
- 2. WCAP-8252, "Documentation of Selected Westinghouse Structural Analysis Computer Codes," April 1976.
FNP-FSAR-3M
REV 21 5/08 TABLE 3M-1 MAXIMUM REACTOR VESSEL DISPLACEMENTS AT REACTOR VESSEL CENTERLINE Maximum Horizontal Maximum Vertical Maximum Displacement Displacement Rotation (in.) (in.) (radians) 144 in.2(a) 0.078 0.030 0.00025 RPV inlet 0.0 -0.038 -0.0004 144 in.2(a) 0.086 0.016 0.00007 RPV outlet 0.0 -0.020 -0.00026 Double-ended 0.049 0.004 0.00031
pump outlet
-0.028 -0.036 -0.00029
- a. Physical restraints limit the maximum circumferential break in the inlet nozzle to 100 in.
2 and to 30 in.2 in the outlet nozzle. The maximum displacements and rotations for these breaks were verified to be less than those listed here for a 144 in.
2 break.
FNP-FSAR-3M
REV 21 5/08 TABLE 3M-2 MAXIMUM REACTOR VESSEL SUPPORT LOADS FOR POSTULATED PIPE RUPTURE CONDITIONS (a) LOCA Maximum Vertical Load LOCA Maximum Horizontal Load Per Support Including Per Support Deadweight 2150 Kips 1050 Kips
- a. Physical restraints limit the maximum circumferential break in the inlet nozzle to 100 in.
2 and to 30 in.2 in the outlet nozzle. The maximum loads for these breaks were verified to be less than those listed here for a 144 in.
2 break.
FNP-FSAR-3M
REV 21 5/08 TABLE 3M-3 MAXIMUM REACTOR VESSEL SUPPORT LOADS FOR COMBINED PIPE RUPTURE CONDITION, SAFE SHUTDOWN EARTHQUAKE, AND DEADWEIGHT (a) Maximum Combined Vertical Ma ximum Combined Horizontal Load Per Support Load Per Support 2392 Kips 1326 Kips
- a. Physical restraints limit the maximum circumferential break in the inlet nozzle to 100 in.
2 and to 30 in.2 in the outlet nozzle. The maximum loads for these breaks were verified to be less than those listed here for a 144 in.
2 break.
REV 21 5/08 REACTOR VESSEL SUPPORT SHOE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3M-1
REV 21 5/08 REACTOR VESSEL SUPPORT BOX JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3M-2
REV 21 5/08 MATHEMATICAL MODEL FOR HORIZONTAL RESPONSE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3M-3
REV 21 5/08 MATHEMATICAL MODEL FOR VERTICAL RESPONSE JOSEPH M. FARLEY NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 3M-4