ML20011B075
| ML20011B075 | |
| Person / Time | |
|---|---|
| Site: | Rancho Seco |
| Issue date: | 10/07/1981 |
| From: | Cheverton R, Kam F, Kryter R OAK RIDGE NATIONAL LABORATORY |
| To: | |
| Shared Package | |
| ML19270A275 | List: |
| References | |
| CON-FIN-B-0468, CON-FIN-B-468 NUREG-CR-2083, NUREG-CR-2083-D, NUREG-CR-2083-DRFT, ORNL-TM-8072, NUDOCS 8111040152 | |
| Download: ML20011B075 (75) | |
Text
UCERIM REPORT
.a
!"TREG/CR-2083 DRAFI INL/TM-8072 Dist. Category RG Contract No.
W-7405-eng-26 Instrumentation and Controls Division EVAIRATION OF THE THREAT TO PWR VESSEL DiTEGRITT POSED BT PRESSURIZED THEIL% L SHOCK EVETTS R. C. Kryterl
- 7. J. Burns 2 R. D. Cheverton3 R. A. Hedrick4 F. B. K. Kam5
- c. W. Mayo 4
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Oak Ridge National Laboratory Oak Ridge, El 37830 Manuscripted Completed - October 7, 1981 Data Published -
CAUTION m see.
- e. has a.: noen m n.as peiene cuer,.co end the menessmed n or its informeo is easy for omcies use. Ne seinese to the public shall ne mede wkhout the appro.el of the Lew Departament of Union Car %$e Corporation. Nuclear Di-visies.
' Instrumentation and Centrols Division, Oak Ridge National Laboratory.
2Engineering Physics Division, Oak Ridge National Laboratory.
3Engineering Technology Division, Oak Ridge National Laboratory.
4Science Applications, Inc., Oak Ridge 3 ranch Office.
50perations Research and Development Division, Oak Ridge National Laboratory.
Prepared for the U.S. Nuclear Regulatory Cocunission' Office of Nuclear Regulatory Research Under Interagency Agreements DOE 40-551-75 and 40-552-75 NRC FIN No. 30468 l
Prepared by the OAK RIDGE NATIONAL LABORATORT Oak Ridge, Tennessee 37830 operated by UNION CAR 3 ICE COR? ORATION for the DRAFT DEPULTMDTI 0F CTERGY DRAFT E REPORT 8111040152 811019
- DRADOCK 05000312 PDR l
INTERIM REPORT Evaluation of the Threat to PWR Vessel Integrity Posed by Pressurized Thermal Shock Events l
l Task Coordinator:
R. C. Kryter 2
3 Contributing Aathors:
T. J. Burns, R. D. Cheverton,
5 4
R. A. Hedrick", F. B. K. Kam, and C. W. Mayo
1.0 INTRODUCTION
Pressurized water reactors (. Als) are susceptible to certain types of hypothetical accidents that, under some circumstances, including operation of the reactor beyond a critical time in its life, could result in failure of th pressure vessel as a result of extensive propagation of crack-lika defects in the vessel wall. Accidescs of particular concern are those that result in rapid cooling (thermal shock) of the inner surfaen of the reactor vessel (RV) vall.
particularly if they also involve substantial primary-system pressure.
(Such accidents have been referred to as " overcooling accidents" (exessive cooling for a particular pressure) and/or " pressurized thermal shock."]
For a particular accident and operator and system response, the tendency for preexistant vessel flaws to propagate during thermal-shock loading conditions is a function of the relative magnitudes of the stress field or stress intensity factor (K ) and the material fracture-and arrest-toughness values (KIC and K,).g ".hese toughnesses decrease r
with decreasing temperature and increasing fast-neutron fluence, and K increases with increasing stress and is greater for a surface flaw g
than for a buried flaw. Thus, flaws on the inner surface of the RV wall are of greatest concern for thermal-shock loading.
The positive gradient in temperature that exists within the wall during a thermal transient and the negative gradient in fluence togetbor result in a positive gradient in K, that provides a mechanism for g
arrest of 'a fast-running crack. Ecwever, if the primary-system pressure is high enough, the gradient in Kg may be such that I Instrumentation and Controls Division, Oak Ridge National
~
Laboratory 2Engineering Physics Division, Oak Ridge National Laboratory 3 Engineering Technology Division, Oak Ridge National Laboratory
" Science Applications, Inc., Oak Ridge Branch Office 50perations Research and Development Division, Cak Ridge National Laboratory
~
i 1-2 arrest cake place, and the flaw will then extend through the vessel Depending on the 6 emperature and pressure of the primary system aud the length and orientation of the flaw at the time of its wall penetration, the opening produced could eicher be negligible in i
size or sufficient to preclude adequate cooling of the reactor core.
{'
For instance, previous overcooling-accident calculation 21 indicate that in the event of a double-ended pipe-break loss-of-coolant accident (LOCA), which produces perhaps the most severe of all thermal shocks 1
but also a very low vessel pressure, flave presumably will not be driven through the wall. In another case 2 the internal pressure remains rather high, the coolant temperature remains well above saturation for atmospheric pressure, and RV failure with a sizable 1
opening is predicted.
i As already mentioned, the tendency for crack propagation increases with increasing reduction in meterial toughness and thus with increasing fluenu. An additional factor that influences the extent of the I
radiation-induced reduction ir. toughness of present-day reactor 1
pressure vessel materials is the presencs of impurities such as copper and, to a lesser extent, phosphorous. Within limits, the higher the concentration of these two elements the greater the radiation-induced reduction in toughness for a given fluence.
In '.he context of calculated flaw behavior unde e pressurized. thermal-shock loading conditions, a broad range of copper concentrations exists among PWR pressure vessels currently.in service. Some of the vessels in the high-copper category appear, on the bases of selected hypothetical accidents, assumed in',tial flaws, and presumably conservative analyses, to be suscepti' ole te fail.tre at early dates, while vessels with low copper category are not susceptible to failure for an extensive period.
i Because of the apparent severity of overcooling accidents and the i
obvious complexities associated with defining accidents and their likely frequency of occurrence, performing realistic systems analyses to determine appropriate inpt.t temperature and oressur transients for the vessel integrity studies, and accurately cialuati'.g the mechanical integri :y of the pressure vessel, ::horough plant-specific studies cre in ' o rde r.
In Ms-1981, the U.S. Nuclear Ragalatory Commission ('NRC) requested 3
assistacce from the Oak Ridge National Laboratory (ORNL) in attaining such an unde standing of the severity of the thr:at posed by pressurized thermal shock occurrences, subjset to the constraint that an interim report which would consolidate, evaluate, and summarize all the partinent data s.nd analyses identified and enllected mat be produced in four months time. This short time f rame precluded l
undertaking new studies and calculatious of sigcf1@nt magnitude, se the evalttated results cited '.n this report are necessarily drawn f rom l
kitown previous work and literature search.
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The major goals of this' ORNL iniegrative effort were to (1) identify what is presently known about the pressurized thermal shock problem,.
including the major areas of uncertainty acd the sensitivity of the(2) identify what estimated severity of threat to these uncertainties; is not known about the problem, including suggested means for correcting any such deficiencies; and (3) propose and evaluate possible The work required to meet these goals was divided sitigative measures.
into six principal tasks:-
Define the problem elements that dominate in establishing the overall likelihood of RV failure and develop a scheme for assessing 1.
the relative safety significance and likelihood of occurrence for the spectrum of possible initiating and subsequent events.
Review presently existing thermal-hydraulic analyses of various
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2.
postulated overcooling scenarios and crit.cally assess their
- realism and usefulness in defining a generic spectrum of Identify critical assumptions and input overcooling events.
uncertainties and estimate their probable effects on the predicted temperatures and pressures.
l l Review the function of plant-specific control and safety systems, 3.
Consider system along with procedure-directed operator actions.
modifications which would help to lessen the severity and frequency i
s f overcooling transients.
Estimate the overall severity of threat to RV integrity imposed by 4.
pressurized thermal shock occurrences.
i Propose potential corrective actions which cight be effective in.
5.
Discuss m bable effectiveness reducing the severity of threst.
and relative ease of implementatiou.
Provide recommendations for extending the study in an effective manner in FY 1982 to obtain a broader, more balanced understanding 6.
relates to the spectrum of current plant of the problem as it designs.
representative plant to be studied was The selection of the first somewhat arbitrary but in consideration of an extensive hist.ory of thermal-hydraulic upsets in Babcock and Wilcox (3&W) plants and the low thermal inertia provided by thu B&W once-through steam senerator design, a reactor built by this manufacturer seemed a reasonable Since Oconee-1 has a RV with longitudinal velds having a is the lead S&W plant (commerical choice.
relatively high copper content, operation began in July 1973), and has a larger cumulative power history /.~4.9 EFPY to date) than its sister units, this plant was selected (with NRC concurrence) to provide a basis, so far asOn the othe practical, for our initial study.
hydraulic behavior needs to be further evaluated as recommended later in this raport and because their are special control systems provisions In Oconee-1 limiting transients, more analysis needs to be done before their results are applied to Oconee-1 or generali:ed to other plants.
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1-4 REFERENCES - CHAPTER 1 1.
- 1. D. Cheverton, S. K. Iskander. and S. E. Bolt, Applicability of LEFM to the Analysis of PWR Vessels Under LOCA-ECC Thermal Shock Conditions 01DPJNUEZG-40 (October 1978).
2.
R. D. Cheverton and S. K. Iskander, " Thermal-Shock Investigations i
Heavv-Section Steel Technology Program Quarterly Progress deport",
for Januarv - March 1981, ORNL/TM-7822, pp. 76-83.
3.
Letter, 1. M. Bernero (NRC) to A. L. Lotts (ORNL), " Report on Pressurized Thermal Shock," dated May 11, 1981.
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2.0 OVERC00 LING TRANSIENTS IDENTIFIED AS SAFETY CONCERNS There are three basic. mechanisms for rapidly cooling the primary coolant astem depressurisation of the primary or secondary system, injection of cold fluid, and rapid removal of energy through the steam generator. Four general classes of transients can be identified as encompassing one or more of these cooling mechanisms o Large-Break Loss-of<oolant Accident (DLOCA) i o haf t-Break I4ss-of<oolant Accident (SBLOCA)
I o Main Steam Line Break (MSU )
o Runaway Feedwater Transient (RTT) f The severity and probability of occurrence of each of these transients is dependent on plant-specific characteristics.
The D LOCA produces primary fluid temperature temporal derivatives cn I
the order of 36,000*F/hr, arresting at a base temperature of ~350*F.
To this system is injected 40-85'T high pressure injection system (HPIS) fluid and 90*F core flood tank (CFT) fluid, which results in rapid chilling of the fluid next to the RV and causes an effectively conduction-limited temperature transient in the vessel wall.
The SBLOCA, in contrast, produces order of magnitude lower primary fluid temperature temporal derivatives than the ut0CA, generally less than 2200*F/hr, depending on the size of the break. Also, depending on break size, the CFT system may actuate in addition to the HPIS. A critical difference between SBLOCA and LBLOCA is that the HPIS can j
repressurize the system for many break sizes.
The MSG usually produces primary fluid temperature temporal derivatives that lit, between those of the G LOCA and SBLOCA. These decreasing temperatures result from the rapid primary system energy removal produced by flashing of the fluid on the secondary side of the steam generator. The lowest primary fluid temperature achievable in this trausient is determined by the perfor: nance of the steam generator feed train, HPIS, and CFT.
i The RFT is essentially a variant of the MSU, but without the initial rapid steam generator secondary-side blowdown and the resultant rapid removal of energy from the primary system. The primary system temperature temporal derivatives for the RFT are usually the lowest L
j among the four classes of transients. The progress of the RFT is totally controlled by the performance of the steam generator feed train.
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4 3.0 SEVERITY OF THE THREAT 3.1 Probability of Occurrence of Initiating Events Ultimately, the pro'oability of a thermal-stress-induced challenge to the reactor pressure vessel is dependent on the f requency of requisite initiating events. However, the concern to this study is not the probability of individual initiating events themselves, but rather the total probability that the thermal-hydraulic transients resulting from the initiating events produce pressure and temperature conditions which approach the structural limitations of the RV.
This total probability can be viewed as the :sticiplicative combination of three probabilities:
(1) the probability of an initiating event, (2) the probability that the control and safety systans fail to respond to the transient in an appropriate manner to protec3 the vessel, and (3) the probability that the reactor operator fails to diagnose the exact nature of the transient and therefore fails to take cppropriate action or possibly takes action which actually exacerbates the transia..
t As noted previously, four transients were identified as possible j
thermai shock initiators: a small-break IDCA, a large-break LOCA, a main set am line break, and a runaway feedwater transient. Due to limited time and resources, a detailed characterization of the various factors (i.e., initiating events and system / operator responses) for each transient has not yet been performed. An estimate of the probability of each initiating event which could be a precursor.co conditions having the potential for thermal shock to the RV was made.
Probability Per Reactor Year Estimate Range SBLOCA*
3 x 10
3 x 10-5 go 3 x to-3 13LOCAa t x tow 1 x to-5 to 1 x 10-3 MSI3a 5 x 10-8 1 x 10-6 to 1 x 1o%
RET 1.0 Url Mr To 1.0 The RFT event is the most complex with a very high probability for the initial event but requiring failure in order to produce thermal shock.
This is very plant specific, and for Oconee-1 it appears that multiple independent failures are required (see Section 4.5.4).
- Estimated f rom Reactor Safety Study, WASH-1400.
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o 3-2 Table 3-1.
Summary of Pressurized Thermal Shock Evaluation Mechanistic Results i
INITIATING EVENTS La rge-B reak Small-B reak -
Main Runaway SUBSEQUENT LOCA LOCA Steaaline Feedwater EVENTS (I2LOCA)
(SBLOCA) 3 reak Transient (MSta ) _.- - -
-(RFT)
The rmal-TRAC simulation (a) Rancho Seco (a) TRAC simula-(a) Rancho Seco Hydraulic (Westinghouse (actual plant tion (b)IRT simula-Information Plant) transient)
(b)IRT simulation tion Sources (b) TRAC simula-tion Operator None (a) Rancho Seco (a) TRAC:
(a) Sea SBLOCA Actions operator over-initiates auxil- (b) None Taken or rode automatic iary feedwater i
Assumed trip of mair.
at 30 s.
Main feedwater pumps feedwater (b) TRAC: ini-ramped at 40 s.
tiates aux. feed (b) IRT: none water at 30 s.
Main feedwater ramped at 40 s.
Thermal-No repres-(a) Rancho Seco:
(a) TRAC:
(n)See SBLOCA i
Hydraulic surization of
- Repressuriza-
- No repres-(b)
Indications primary coolant tion surization Rep ressuriza-or system
- Tain 280*F Tain = 350'T tion Predictions (b) TRAC:
(b) IRT:
Tmin < 150*T analysis
- Repressuriza-terminated tion prematurely
- Tmin < 150*T Vessel Crack initir.aa (a) Vessel fails (a) TRAC:
(a) see SBLOCA F racture and arrest (r at 20 IFPY Analysis not yet (b) IRT: Vessel Mechanics vessel failure)
(b) Insufficient available fails at 3 EFPY Predictions at 20 ETPY information for (b) IRT: Vessel analysis fails at 4 ETPY Limitations No (a) Ranch Seco:
(a) TRAC:
(a) See SBLCCA and rep ressuriza-
- Pressure &
- Mild case (b)
- Concerns tion, so of temperature
- Feedtrain Assumes secondary data not as tables feedwater concern entirely ade-(b) IRT:
control quate '
- Assumes feed-failure water control Feedtrain (b) TRAC:
failure treated with
- Press / temp.
- Feedt:ain tables data incomplace as tables e
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s 3-3 3.2 Pressure Vessel Integrity For the purpose of these studies the integrity of the pressure vessel was considered to be jeopardized if the f racture mechanics analysis indicated that an inner surface flaw would propagate through the vessel wall as a result of an overcooling accident. Four specific overcooling accidents for which f racture-mechanics analyses were performed are the large-break LOCA, Rancho Seco (1978), turbine trip with stuck-open bypass valves, and main steam line break. Results of the prelisivary analyses indicate that the vesael will not fail as a result of the ut.0CA, but failure was predicted for the other three accidents. The calculated threshold times for failure were 20, 3 and 4 EFPYs, 18 respectively, based on a fluence rate of 0.046 x 10 neutrons /ca'/EFPY, which is similar to that for B&W plants.
The sir of the break that results f rom propagation of the flaw through the wall is of utmost importance since it is a factor in determining whether the vessel will be able to retain sufficient water to cool the co re. Because cooling temperatures at some, if net all, locations in the primary system are expected to be well above 212*F at the time of predicted failure (excludes DLOCA), there is a large amount of stored energy that will be released, and thus a potential exists for a rather large opening in the vessel wall. A more. quantitative assessment of the problem awair.s completion of detailed studies.
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4.0 PLANT CONTROL AND OPERATIONS i
4.1 Introduction i
Plant control systees and operator actions were reviewed to define system setpoints and capacities relevant to pressurized thermal shock transients. Potential feedwater control failures and operator actions 4
were investigated and the control system and operator actions which took place in the Rancho Seco overcooling transient were reviewed.
These data were used to evaluate control system response asseptions employed in the thermal-hydraulic analyses available to us and to develop conclusions and recommendations concerning control system modeling for pressurized thermal shock thermal-hydraulic analysis, operator actions, and potential problem areas.
4.2 Reactor Protection Systes The Reactor Protection System (RPS) is a safety-grade system designed to trip the reactor according to the values of a variety of input parameters, and thereby to protect both the core from fuel rod cladding damage and the reactor coolant system from overpressurization.1 Reactor trip directly influences main feedwater control through the Integrated Control Systes2 (ICS) unit load control. The neutron power i
signal obtained from the RPS can also modify main feedwater demand if its mismatch with the ICS reactor demand level exceeds'a set tolerance.
Following a reactor ; rip, the heat generated by the reactor is l
detersined by the shutdown rate.
In order for the remainder of the unit to " follow" the reactor, the unit load demand (and hence the total feedwater demand) will track the actual megawatts generated at a maximum rate of 20% per air.ute. For transients involving initial depressurization, the RPS will trip the reactor at a low pressure set point of 1925 psi. For transients initiated by a turbine trip, the reactor will be tripped at the start of the ttansient.
4.3 High Pressure Injection 4.3.1
System Description
The High PressuJe Injection System (HPIS) injects water into the four reactor vessel inlet pipes upon actuation of appropriate trips in the engineered safety features system. The HPIS comprises three high-pressure pumps; the flow can be controlled manually and the pumps can be aligned manually in several different ways.3 Normal HPIS actuation will inject full flow from two of the three pumps, with suction taken from the borated water storage tank (BWST).
y 4-2 4.3.2 Capacitiesl The HPIS pump characteristic performance curve is shown in Fig. 4-1.
The BWST has a volume of 388x103 gallons. The RPI valves will be fully open within 14 s from an actuation signal, and the pumps will be up to speed within 6 s.
4.3.3.
Set Pointsl The KPIS will actuate when the reactor coolant system pressure drops to 1500 psi.
i 4.4 Low Pressure Injection 4.4.1
System Description
The Low Pressure Injection System (LPIS) injects water into the reactor i
vessel downconer through two pipes located on opposite sides of the.
core and at ninety degrees from the reactor vessel outlet neazles. Low pressure injection is provided by three low pressure pumps (operated in parallel) and two accumulators. Pumps are nornally aligned to draw l'
reactor building sump. The pump flowrate can be controlled manually. -
from the BWST, b te can be manually transferred to take suction from the 4.4.2 Capacitiesl The LPIS pump characteristic performance curve is shown in Fig. 4-2..
The LPI valves will be fully open within 15 s after actuation, and':he pumps will be up to speed within 8 s.
3 As stated previously, the BWST has a capacity of 388x10 gallons. When considering the total inventory of borated water availaole to the LPIS, it must be noted that the containment spray system, if actuated, will also draw from the BWST. The accumulators have a combined capacity of' 21x103 gallons.
4.4.3 Set Pointsl l
The accumulators will discharge water into the reactor vessel when the l
pressure falls below 600 psi, and the LPIS pumps will be actuated when the primary pressur.e falls below 200 psi.
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4.5 Main Feedwater Control 4.5.1
System Description
Main feedwater control is one function of the ICS. A feedwater demand.
signal is developed, based on unit load demand but also ratioed and
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compared to the measured feedwater flowrate so as to regulate the main feedwater control valve. The total corrected feedwater demand signal is modified to control the feedwater pump speed, and so to meet the feedwater demand and to maintain constant pressure drop across the main control valves. Steam generator low-level and high-level limits intercept the feedwater demand to prevent underfill and overfill conditions.
The main feedwater pumps are supplied water from the condenser hotvell, the surge tank, and the condensate storage tank through three condensate booster pumps and three hotwell pumps.
A variety of feedback and trip functions affect the main feedwater ccatrol. Manual control of alA feedwater pumps and valves is also possible.
4.5.2 Capacitiesl The full main feedwater flow ia 6539x103 lbm/hr from the hotwell per-steam generator. The =vN= vater inventory in the feedvater system is 295x103 gallons.
4.5.3 Set Points o The total feedwater demand viil run back at a maximum race of 20% per minute to track scaerated negawatts following a reactor trip.2 Operators are presently requirsd to trip the reactor coolant o
pumps following actuation of the engiseered safety features system. Tripping the reactor coolant pumps will, in turn, cause the feedwater demand to run back to a maximum value of 20% at a rate of 50% per minute.1 o The Oconee-1 unit has a steam generator high-level limit that will trip the main feedwater pumps if the steam generator is filled to this level.a (This trip may not be present on other B&W units).
o The Oconee-1 unit will also trip both sain feedvater pumps if there is a loss ;f ICS power.8 (This trip may not be present on other 3&W units.)
All hotwell pumps will be trippeda when the hotwell level o
reaches " emergency low."
All condensate pumps will be tripped when che condensate o
bocster pump suction header pressure decreases )elow 43 psig.8 i
4.5.4 ICS Failure Modes and Effects The ICS is a complex, nonsafety-grade control system. A failure modes and eff ects analysis (FMEA) review was therefore performed to w---
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I 4-7 investigate potential failures in the 'ICS that might lead to excessive feedwater. This review divided the ICS into three general ~ areas, as shown in Fig. 4-3.
The first area constitutes higher levels in the ICS, where significant feedback and feedwater flow limiting actions win be effected from a variety of other process signals. The second area is characterized by limited ICS feedback, but manual control of the feedwater system is stin nssible. The third area encompasses failures below au control points.
The results of this analysis, which are susearized in Table 4-1, show that single control failures below the manual control points always 1
leave one or more. alternate means by which the feedvater may be manually controlled. Manual control is required, in general, to assure proper feedwater control. ICS failures in region 2 win limit feedwater flow to the steam generator high-level limit. In the case of Oconee-1, if the ICS high-level limit does not act, a separate high-level signal win trip the main feedwater pumps.
Feedvater flow win be limited for ICS failures in region 1 by both the high-level limits and by feedback from other parameters. It should be noted that without the steam generator high-level trip for the feedwater pumps, failure of the startup level signal to a " low" condition can result in unlimited overfeed to the steam generator.
ICS failure or power failure win generally lead to a loss of feedwater, due to a zero dem'and speed signal being presented to tha main feedwater pumps. Oconee-1 also has a trip to stop the main feedwater pumps on loss of ICS power. The presence of feedwater control for a loss of ICS power condition is known to be highly plant specific.
4.6 Rancho Seco Transient A significanti overcooling transient occurred on March 20,1978 at the Rancho Seco reactor fol'.owing a failure of power supplies that fed both control room indicators and the ICS. The initial plant response was a loss of feedwater transient combined with incorrect indications presented to the operators. Auxiliary feedwater was also supplied to one steam generator through an ICS control path (no longer present in B&W units)<
As a result, the operator was presented with an indication of 0% feedwater demand for one loop and 100% feedwater demand for the other. His response was to manually renove the main feedwater pump trips and so restore main feedwater. Upon restoration of the instrumentation power, it was discovered that the reactor had been overcooled and corrective actions were taken.3 The instrumentation and control system power supplies have since been upgraded.
The Rancho Seco incident clearly demonstrates that significant overcooling transients can be induced by operator action.
In this case, the operator's actions were the result of unsuspected and
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FHEA Review for Excessive Feedwater Failures and Control i
i Failure Hode Long-Te rm Cont rol Below ICS Hanual Station Feedwater Pump (s)
Fa,ils to full s' peed ICS will close FW valve to maintain proper flow. Valves can also be controlled manually.
Hain Feedsater Valve Fails Open Block valve and pump speed can be controlled control Valve manually.
Main Feedwater Block Valve Falla Open No effect except for leakage around control Valve valve (when closed).
Pump speed and control valve can be controlled manually.
4 Startup Feedwater Valve Fails Open Startup feedwater block valves can be closed p
4
'Contr;o1Velve manually.
Sta rtup Feedwater Valve Fails Open Two block valves available. Sta rtup. feedwate r Block Valve cont rol valve can also be closed manually.
Below Other Pa rameter Feedback Feudwater Flow Fall Low Cont rol valves open; cont rol valve and block valve can be controlled manually.
(Fills steam generator to bigh-level limit.)
Startup Level Fails Low Feedwater valves fail open independent of feedwater demand. Manual control of valvea and pump speed possible. Oconee-1 has sep-a rnte high-level feedwater pump trip.
Loop Feedsater Falls in direction to Valve can be controlled manually.
(Fills Demand Error open valve steam generator to high-level limit.)
1 99 L.
Table 4-1.
(Continued)
Failure Mode long-Term Control Above_0ther Parameter Feedback
/
Loop or Total Falls in direction to Valves and pump can be con rolled manually.
Feedwater Demand increase demand (Cross limits, BTU limits, and level limits limit increase and control at high level.)
Other ICS AC Power Loss of AC Power to ICS Flow control valves freeze. Feedwater pump goes to low-speed limit. Oconee-1 has feed-water pump trip on loss of ICS power.
4 e
,.-.=-m.a,
--w wm wm--e--=-.-#---
4-10 4
I widespread information failure. It shoul:1 be noted that as a result of this and other power-supply-failure-induced control system transients, all BW units have since been required to review their system power supplies and make modifications and develop procedures as necessary to reduce the probability of such events and to provide operator guidance for controlling the unit.
4.7 Operator Actions Our review of the ICS failure modes and effects analysis (FMEA) indicated that a few single or double control failures cas lead to uncontrolled feedwater supply, but that the majority of potential r
failures have feedback paths that will act to reduce feedwater.
automatically. It is concluded that there are a larger number of ways.
that excessive feedwater can result from operator errors of commission than from errors of omission.
Oconee-1 event sequences for small steam line break and excessive 1
feedwater were reviewed for indicated operator actions. These event-sequences assumed multiple system failures and f ailure to control different systems properly along the event paths. Potential severe consequences.were identified where key problems remained uncorrected.
These event sequences show the necessity for correct operator actions along a number of paths.5 4.8 Review of Thermal-Hydraulic Calculations Available thermal-hydraulic analyses were reviewed for correctness of' assumed control system response. These analyses were found to include operation of the safety systems as designed but to employ rather arbitrary assumptions about feedwater system operation.
In particular, the TRAC calculations for MS13 and SBLOCA assumed that the main feedwater continues to supply 100% flowrate and is then reduced to zero by operator action over a two-minute period starting 40 s into the transient. Emergency feedwater is assumed to be initiated at 30 s to the remaining intact steam generator for the MSI3 and to both steam generators for SBLOCA. In the absence of such operator actions, automatic control system actions would perform the same fu.setions, but the effect of the possible difference in timing on the thermal-hydraulic transient is not known at this tise.
S
't
--+-y,,.-
w-e,y.,,n r w..y.g.
.y.w,_,,.,,e
,m,9,,yg,--ww,,,,,.y.-,rw,,,.,y_
,,,,,-,..,,._y
.,,,.-,.,,_.~.,,m_-_
r--,--
-w.
o N
4-11 The IRT calculations for MSl3 and RFT assume that the main feedwater continues to supply 100% flowrate to one steam ge.rator for the duration of the feedvater supply. Owing to the wide variety of feedvater control feedback and trip functions present, as described in Sect. 4.5., achievement of 100% flow rate during this tranisent is not possible; the actual achievable flow rate and its effect on vessel integrity are not known, but the lower flow race would result in a less severe transient.
Therefore, the IRT and TRAC feedwater flow assumptions in both the IRT and TRAC can be considered to be approximately bounding cases for excessive feedwater supply. Another perspective would be to view the cases as representing two events of different probability. The TRAC case assumes correct operation of the feedwater control system, whether by ICS or the reactor operator. The IRT case corresponds to an overt operator error (manually supplying full feedwater flow) or a multiple control system failure with lack of corrective operator action. On this basis we consider the IRT transient to be less probable by a factor of 10-3 to 10-8 4.9 Summary The plant control and operations review identifies a variety of potential failures that could possibly result in excessive feedwater supply. Most of these failures have feedback or trips that can be expected to reduce the feedwater in a fairly short period of time. The operator can also take action to terminate main feedwater for all single and double ICS failures identified.
~
The thermal-hydraulic ca1culations reviewed employed somewhat arbitrary but approximately bounding feedwater response assumptions. The assumptions used in the IKI analyses, in particular, are considered to be conservative, with regard to the severity of the transient.
More feedwater supply calculations we /.d require nodeling the feedwater demand runbacks, limits, and feedwater system trip, points that were described in Sect. 4.5.
Using a model of this type, it would be necessary to investigate several of the potential control system failures to identify the worst credible case. The probability of excessive feedwater is likely to be dominated by the probability of overt operator error, since the control system failure re' quires two independent failures plus lack of operator corrective action.
5 e
e.w+w
,r
.-m..,,.-
v.e,,,,.--.-.--.-.c
.-y,
,---.--y-,4.,--,,.-%,,,y---...,--7---p+m---.v-c.,.
REFERENCES - CHAPTER 4 1.
Oconee-1 Final Safety Analysis 'eport.
R 2.
Oconee-1 Instruction Book for Integrated Control and Non-Nuclear Instrumentation Systems.
3.
Sacramente Municipal Utility District Followup Report to Reportable Occurrence 78-1, Re Docket No. 50-312, operating License DPR-54, dated March 31, 1978.
4.
Oconee Unit 3 Piping Drawings:
PO-101-A, B Righ-Pressure Injection; PO-102-A Low-Pressure Injection; PO-121-A-3, B-3A, 3-38, D-3 Condensate, Feedwater, Emergency Feedwater Systems (from Duke Power Company).
5.
Oconee Small Steamline Break and Excessive Main Feedwater Event Trees and Safety Sequence Diagrams (from Duke Power Company).
6.
Transient Response of Babcock & Wilcox - Designed Reactors, NUREG-0667 (May 1980).
7.
Integrated Control System Reliability Analysis, BAW-1564.
8.
Duke Power Company, personal communication.
9.
"oss of Non-Class-1-E Instrumentation and Control Power Systems Bus During Operation," USNRC IE Bulletin 79-27.
s N
5.0 THERMAL-HYDRAULICS 5.1 Literature Search for Accumitlated Experience Data Bases. Thirteen data bases x
o BHRA o
CIM o COMPENDEI o
CONF o EDR o
EIA o FEDEI o
ISMEC o NSA o
RSI o WELDASEARCH were searched for thermal-hydraulic system data relevant to thermal shock tranisents. Capsule descriptions of the data bases are given in Appendix A.
Approximately 600 thermal-shock-related entrius were found; however, the majority of the entries were LBLOCA emergency coria cooling injection studies and generally contained no system thermal-hydraulic infarmation. There were also a significant number of licensee event reports (LERs); however, they focus on the incident precursors, not the transient data, and so are of limited usefullness to this study.
Oconee Licensee Event Reports. Five LIRs of interest, covering four events, were found for the Oconee Nuclear Power Station (see Appendix 3).
The four events were distributed as one each for Units 1 and 2, and two for Unit 3.
Three events were in the class of steam generator overfeeds.(RFT), and the other was an open power-operated relief valva (SBLOCA).
None had major consequences, since the operators took timely action.
Soecific Documents. Twenty specific documents, as listed in Appendix C, were also reviewed.
The first seven of these references were particularly interusting since they contain system thermal-hydraulic data.
5.2 Thermal Shock Plant Transient Data The major source of actual plant data for transient overcooling is the March 20, 1978 Rancho Seco event (Appendix C., Rafs.
I and 2).
The pressure and temperature data employed by ORNL as input for f ract;ure mechanics calculations (Raf. 2) are shown in Figs. 5-1 and 5-2, respectively. The pressure surges contained in the original data (Ref.
- 1) were removed for simplicity (they may or may not be "real").
Owing to the locatir.,ns of the inlet temperature measuring instruments,
which are placed in wells at the suction side of the reactor coolant pumps and are therefore upstream of the HPIS injection, the use of these actual plant data as the RV forcing functions for a 3&W plant introduces some uncertainty.
The temperature of the fluid at the RV inlet nozzle might therefore be expected to be lower than measured by the instrumentation, unless two-phase thermal equilibrium flow exists.
p-e 5-2 O
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-8 9
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5-4 On the other hand, for transients in which the vent valves open (not likely for the Rencho Seco event, cince the reactor coolant pumps are running, but quite likely if the pumps are shut off), mixing in the downcomer could be significant and the fluid temperature at the RV wall could be higher than the measurement (again, barring two-phase thermal equilibrium flow). Due to the unknown nature of these competing effects, the RV wall temperature forcing function is not easily derived from standard instrumentation in B&W plants during overcooling events.
5.3 Overcooling Simulations 5.3.1 Results of Analyses To our knowledge, two computer codes, IRT and TRAC, have been used to i
predict the thermal-hydraulic characteristics of overcooling transients for Oconee-type plants. The pressure and temperature predictions for five scenarios analyzed to date are shown in Figs. 5-3 and 5-4.
Note that primary system repressurization is predicted for all cases except case 4, which corresponds to MSLB as ' simulated with TRAC, and 4
case 5, for which the transient predictions are incomplete.
Repressurization does not occur in case 4 because of an assumption of thermal equilibrium made in modeling the press.urizer. The initial s
depressurization is similar in all cases, except that IRT does not predict so sudden a primary sy, stem contraction as TRAC. The lower worth assumed for the control rods in case 2 is re'sponsible for the quick return of system pressure, as compared to case 1.
Fig 2re 5-4 shows the degree of primary system overcooling to be similar for cases 1 and 3.
The lower rod worth of case 2 is again evident in the rate of cooldown. Case 4, TRAC MSLB, is the only one which does not show great overcooling; this difference is attributable to ass.ned operator termination of main feedwater flow at 40 s into the transient and the use of,csergency feedwater to the intact steam generator.
L 5.3.2 Limitations All the current simulations possess limitations which give concern for the realism of the thermal-hydraulic predictions. These limitations are, in part, inherent in the codes sud also result from modeling deficienciu and questionable input assumptions, as discussed below.
' Feed Train. Owing to the multiplicity of heaters and punps and the various. input conditions and f eedbacks present, it is no t, easy to c.alculate the steam generator secondary-side inlet conditions.
Accordingly, they are supplied by look-up tables in both IRT and TRAC, and the values entered are the result of simplifying assumptions.
Moreover, since this tabular input is fixed for the duration of the transient, the course of the RIT and the latter stages of the. MSL3 cases are almost completely determined by the values entered initially in the look-up tables.
- - - ~.. -
r n
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59 0 4 0 3 1
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I.EGEND
{
B a'.f 14RT, Twt.ine Talp,Cc am=.0001 i
2-lRT,Twbine Tsip,8crem=.0345 2
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5 TRAC,SBLOCA o
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I I
I I
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100 20e 300 400 56J 000 700 000 000 1000 1100 1200 1
TIME (SEC) 1
', Fig. 5-4.
Composite Temperature Profiles from Simulations.
~
i i
l
5-7 Fluid Mixing.
It is obvious that the-degree of mixing between the HPIS and primary-system fluids will have a marked effect on the RV van.
temperature in the downconer region. Further, as shown in Appendix C, Ref. 3, for cases where the vent valves open, the degree of fluid mixing assumed to occur in the upper downcomer completely determines the severity of these transients.
i Neither T!!AC nor IRT contains models which are suitable for analyzing these special cases.
IRT has only equilibrium capability for.the HPIS and vent valve mixing, whereas TRAC has non-equilibrium capability.
It is our understanding that the cases modeled thus far do not predict vent valve opening, so some basic code differences may not be manifested in the simulation results.
Control Systems. The performance of the control system was assumed and specified before the cases were run; the assumptions were rather -
arbitrary and, in most cases, quite conservative. Direct feedback l
control system modeling for these transients is not currently possible in either TRAC or IRT.
Reoressurization. Repressurization is a key phenomenon, both as to its
~
eventual occurrance and the specific time at which it occurs within the transient, since this relates to likely operator actions. Only bounding cases have been run thus far, i.e., fun equilibrium or noninterchasge pressurizar models have been used. ActuaHy, more than the pressurizer is involved here; the upper head and the entire hot leg act as "pressurizers" during these transients and their performance during primary system refill win also affect the rate of rep ressurization.
Flow Distribution. An ability to calculate flow in the primary system over conditions ranging f rom full forced flow to natural circulation is required to treat these transients properly. How wen the codes perform such calculations has not been determined, although to date the IRT calculations have ap,,arently been " driven" by input and have not utilized a. momentum equation solution.
Existence of Two Phases. Owing to void formation at the top of the
" candy cane," loss of natural circulation will occur in B&W plants during many overcooling transients IRT cannot treat this phenomenon correctly, whici' obviously limits the transients to which this code is applicable.
Prima ry Metal. The effect of heat transfer to the primary fluid f rom the primary metal has not been fuHy evaluated. Such heet transfer has been included in some cases (IRT-RFT, TRAC W.SI3 and S3LCCA) and not in others (IRT-MSLB ).
The effect could be significar:t in some transients, and so should be included in realistic evaluations.
IRT Cases 1, 2, and 3.
These are,two cases of turbine trip with open bypass valves (Cases 1 and 2) and a main steam line break (Case 3).
In au cases fun main feedwater flow is assumed. This is a very conservative assumption. With one st.eam generator flooded and the
--s*cw'-ri,+@-amW-cyyw 3pr+w tw e '-
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erq wr
'"'N'*1 q-------m-+-p-me rw---5.-mes.re-,qs se--gut---%a--w m-
-wttr*=e---p--+-wew++w-
5-8 intact steam generator isolated, the operator would have to go to extremes to keep the main steaa-driven feed psumps running, having lost his primary staatt source (the steam generators). h hot well and condensate booster pumps do not have enough head to maintain full flow (at least in Cases I and 2).
In adda. tion, a multiplicity of ICS failures would have to be assumed to prevent automatic runbad and trip of the main feedwater train.
h feedwater temperature was assumed to be a simple ramp down to the hot well temperature, 91*F, over one minute. This is again conservative. h several pumps and heaters and the length of piping will all have considerable capacity to hold the temperature up.
In addition, the high-pressure heaters take their steam from the main steam lines and not the turbine, so they will not be isolated by the trip.
h rate at which the cooled primary system fluid is transferred to the pressure vessel is not proper 1.v calculated. With the primary pumps tripped and the system depressurized, voiding in the " candy cane" will inhibit natural circulation until the system is refilled by the HPIS.
h repressurization by the HPIS is also overpredicted. h non-interchange models used for the upper head and pressurizar result in a steam compression calculation producing the pressure.
Since the steam and water are assumed not to interact, the steam " bubble" is compressed at a volume reduction rate equal to the HPIS charging capacity.
TRAC Case 4 This is a main steam line break and is more realistic in 1 :s assumptions of feed train performance. h operator is asstz:ed to start auxiliary feedwater at 30 s and terminate main feedwater at 40 s.
The ICS would have performed the same functions in the same time frame had the operator not acted. h auxiliary feedwater comes straight f rom the hot well, so its temperature is more easily determined.
h adequacy of TRAC to treat " cat.ay cane" voiding a'nd consequent loss of natural circulation and, therefore, the transport of cooled fluid to the pressure vessel before repressurization by the HPIS is not known.
However, the repressurization rate is known to be too slow. Full e quilibrium is assumed in the node above the water level in 7.he p ressuriza r.
This results in condensation of the steam. The refo re, repressurization will not occur until all steam has been condensed and the pressurizar is full; this is unrealistic.
By comparison, case 4 is more realistic and also much milder than the IAT-MSI.3, Case 3, as can be seen in Figures 5-3 and 5-4.
Under the limitations noted, these two. cases could be considered bounding i
analyses.
5.4 "3est Judgment
- Pressure and Temperature Forcing Functions 1
Owing to the above deficiencies, the available simulated forcing functions must be regarded as approximations to the true functions; the magnitude and sign of the error is not presently known. Whateve r their deficiencies, the Rancho Seco data represent a recorded event, j
not a simulation, and so provide the closest realistic vessel forcing functions currently available.
4 m.--
.m_s__r,_.
...~
-,,-m...,_,,-,
-_...mm,,,,,,
.~,....____,-..-.,_.-.,,,,m,.-.,_.c_.
l 5-9 References, By Legend Number, for Figs. 5-3 and 5-4 1.
" Runaway Feedwater Af ter Turbine Trip Report," letter, M. Levine to N. Zuber dated July 2,1980.
2.
Ibid.
3.
" Analysis of a Steam Line Break with Primary System Overcooling for a Typical B&W Reactor," letter, R. Carbone to R. Kryter dated August 14, 1981.-
4.
"Comolation of Scheduled Analysis on Pressurized Thermal Shock Scenarios," letter, S. Fabic to C. Serpan dated June 22, 1981.
5.
Ibid.
6.
" Parametric Analysis of Rancho Seco Overcooling Accident," letter, R. Cheverton to M. Vagins dated March 3,1981.
s
.a O
e e
O i
1 i
l l
I I
l
6.0 ESTIMATION OF NEUTRON FLUENCE AT THE REAcr0R VESSEL WALL 6.1 Intr'oduction A realistic evaluation of a postulated " pressurized thermal shock" accident requires a knowledge of the neutron fluence (E 11.0 MeV) and its uncertainty throughout the reactor vessel wall. The fluence values must be known at the location of those materials (welds or places) which are most likely to be affected by the conditions attained during the transient.
Currently the Code of Federal Regulations (10CFR50, Appendices G and H) and Regulatory Guide 1.99 (Ref.1) require licensees to provida estimates of the neutron fluence in the reactor vessel beltlina region as a part of their surveillance programs. The methodologies adopted by different vendors and service laboratories to obtain the fluences can be separated into three parts:
a.
Neutron transport calculations b.
Dosimetry measurements c.
Analysis of uncertainties The technicues for accomplishing parts "a" and "b" do not vary significantly from vendor to vendor. However, the uncertainty analysis involves ecsbining differential and integral dosimetry data, both measured and calculated, in a consistent 'ashion so as to obtain absolute fluence values (and their uncertainties) in the RV wall as a function of energy. These fluences must be extracted from an analysis of measurements performed at the location of a surveillance capsula, which may be distant from the RV vall.
Although considerable work has been expended in developing methodologies 1-5 to achieve part "c", the application to power reactors has just begun.
A preliminary list of uncertainties affecting the calculation and measurement of neutron flux and fluences in LWRs was compiled by the AS'Di E10.05.01 Task Group on Uncertainty Analysis. This list, with a f ew additions, is given in Table 6.1.
6.2 Babcock and Wilcox M.achodology 6.2.1 Neutron Transport Calculacions The calculated energy group fluxes are determined using a discrete i
ordinates solution of the Bolt:mann transport equation. The codes used are ANISN6 and DOT.7 Table 6.2 below gives the steps *in the B&W
~
transport calculational procedure.
O w.,
mm..
..m--
m
t
.a l
T 6-2 Table 6.1.
Uncertainties 'for Calculation and Dosimetry Measurement Procedures in LWRs Source of Uncertainty Estimated Uncer-tainty (%)
I.
Calculational Procedure,b a
A.
Source Tars
- 1. Fuel management - uncertainty in the fuel cycle 10
- 2. Fuel position
<5
- 3. Burnable poison
'O
- 4. Core power distribution (cycle and cycle-to-cycle variation) 30
- 5. Power / time history (cycle and cycle-to-cycle variation) 10
- 6. Local power at cure periphery vs. total power a.
axial 20 b.
radial and azimuchal 20
- 7. Control rod position 5
B. Transport c
1.
Flux synthesis (reduction of 3-D to 1-D-and 2-D calculations) e 2.
Energy group structure c
3.
Quadrature (S and anisotropic scattering P )
e n
g 4.
Cross sections e
5.
Spatial mesh c
6.
Geometry c
7.
Time-averaging vs. changing core condition e
8.
Extrapolation (lead factor) c D
. II. Dosimetry Measurement Procedures c
A.
Nuclear data (reaction rate cross sections, branching rations, etc.)
e B.
Competing reactions e
C.
Photofission corrections e
D.
Flux / time history c
E.
Counting calibration c
- Values listed compiled' by ASIM E/0.05.01 Task Group od Uncertainty Analysis.
D3AW-1485 discusses several of the sources of uncertainty, but does not i
specify the effect on fluence estimates at the RV welds.
Currently unavailable.
~
C
. ~. -.
6-3 Table 6.2.
B&W Method.for Obtaining Neutron Yluences Obtain a pin-to pin time-averaged power distribution a.
12-group CASK cross section sets for 1-D b.
Obtain P3 and Pg and 2-D calculations, respectively discrete ordinates transport calculation c.
Perform a 1-D, P, Sg 3
discrete ordinates transport calculation d.
Perform a 1-D, P, SS g
e.
Obtain a P Pg correction factor from the 1-D calculations to 3f apply to a 2-D calculation f.
Perform a 2-D,.x-y calculation with the surveillance capsule Perform a 2-D, it-y calculation without the surveillance capsula g.
h.
Obtain a capsult-perturbation factor from the 2-D calculations to correct the measured activity or calculated fluxes 1.
Perform a 2-D, P, M calculation g
J.
Perform an axial 2-D, P, r-s calculation 3
k.
Obtain synthesized 3-D group fluxes in the reactor vessel 1.
Correct estimates of the group fluxes, based on the P /Pg and, 3
capsula perturbation factors 6.2.2 B&W Surveillance Dosimetry Measurements The surveillance program for Oconee-1 comprises eight surveillance capsules designed to monitor the effects of neutron and thermal environment on the saterials of the reactor pressure vessel core region. The capsules, which were inserted before initial plant startup, are positioned between the thermal shield and the RV wall.
Six of the capsules, placed two in each holder tube, are positioned near the expected peak axial and aximuthal neutron flux. The remaining two capsules (designed to monitor thermal aging) are placed in an area of essentially zero neutron flux.
Capsule OCl-F was removed during the first refueling shutdown of 1
l Unit 1, and capsula OCl-E was removed during the second refueling i
shutdown.
I Four activation detectors with reaction thresholds in the energy range l
of interest were placed in each surveillance capsule. The properties of interest for the detectors are given in Tabla 6.3, and the results of the averaged measured reaction rates are given in column 3 of Table 6.4.
Table 6.3.
Surveillance Capsule Detector Data l
[
Detector Threshold Energy Product Half-lif e (MeV) l 59Co(n,Y)60Co Thermay 5.26 y 237Np(n,f)l37Cs 0.3 30 y 238 (n,f)l37Cs 1.5 30 y U
54Fe(n p)S4Mn 2.0 313 d 58Ni(n,p)58co 2.5 71.3 d l
l v=--
t am.e-
+-g 7,-
=,
6,-.i----
9 ty 9we1 e+v g--p g
93-
,,y-w.,,--wy-n,, -
,,y.%--g' me-4-.
--gm.---+---
.r--P,a y,,
- - -,. -. + - - + -
---+-+--s
---si
.e 6-4 Table 6.4.
Conparison of Calculated to Experimental Reaction Rates for Ocoras-1 s
Capsule Reaction Experiment, E Calculated, C Ratio,d E/C (uci/g)c (uci/g)
OCl-F" 3"Fe(n.p)54Mn 17.6 2 0.95 21.0 0.84 58Mi(n.p)58Co 495.0 2 2 422.0 1,.17 258U(n,f)l37Cs 1.3620.21 0.58
.2.34 2373p(g,f)137Cs 7.86 2 0.10 2.90 2.70 b
OCl-E 54Fe(n,p) 5'!!n 536 2 62 729.3 0.74 58Ni(n,p)38Co 975 i 163 1,266.0
~0.77
~
~
238 (n,f)l37Cs 1.94 2 0.18 1.719 1.13 U
237Np(n,f)l37Cs 9.32 2 1.18 8.799 1.06-a3Ag_t421 b3AW-1436 c54Mn and 58Co values are given in units of per gram of target for OCl-E and per gram of dosimeter for OCl-F.
dNormalization constant t
O f
6-5 6.2.3 Determination of the Neutron 71uence (E)1 MeV) at the Reactor
~
Vessel The energy-dependent neutron flux is not directly available from activation detectors because they provide only the integrated flux on the target material as a function of both irradiation time and neutron energy. To obtain an accurate estimate of the average neutron flux incident upon the detector, the following parameters must be known:
(1) the operating history of the reactor, (2) the energy response of the given detector, and (3) the neutron energy-group fluxes at the detector location. Two means are available to obtain the desired spectrum: iterative unfolding of experimental foil data and neutron transport methods. Due to sr lack of sufficient threshold foil detectors satisfying both the threshold energy and half-life requirements necessary for a sutveillance program, the neutron energy spectrum was obtained by the transport method (Sec 6.2.1), instead of by spectrum unfolding.
The calculated spectrum is used in the following equations to obtain the calculated activities used for comparison with the experimental values. The basic equation for the activity, D (in uC1/g) ist
-l tjg \\ -l @ rj)
M g
l
@i t{e(E)+(E)[F D
f 1-e
'e
=
g 3
3,7, go jt g
j (6.1) i where C
normalizing constant
=
Avogadro's number N
=
atomic veight of target material i A
=
g t
F
.=
either weight fraction of target isotope in the i 1
material or fission yield of desired isotope e (E) y,
group-averaged cross sections for material i
=
$(E) group-averaged fluxes calculated by DOT analysis
=
F) fraction of full power during j" time interval, t)
=
ch
(
decay constant of the i material
=
t) interval of ' power history
=
sum of total irradiation time, i.e., residual ci=e in reactor T
=
and wait time between reactor shutdowa and counting Eh t) cumulative time from reactor startup to end of j time
=
period, i.e, tj=
k e
k=1 l
l l
~ _,.~ _._
6-6 The normalizing constant, C, can be 'obtained by equating the right hand side of Eq. (6.1) to the measured activity. With C specified, the neutron fluence 1 1 MeV can be calculated from
$(E > 1.0 MeV) = C 15[MeV $(E) j 1 F ed,
(6.2) j E=1 where M is the number of irradiation time intervals.
The last col:ma, of Table 6.4 (Ratio E/C) shows the spread of the normalizing constant as a function of the threshold reaction used in the measurements.
BAW-1436 states that the 238U and 237Np reactions from the OCl-F capsule have been deleted in all current evaluations on a basis of inconsistency.
6.3 Results.:.nd Uncertainty Analysis The estimated fluences (E > 1 MeV) at the axial welds (Tame 6.5) were
~
determined from Tables 6.6 and 8.1 of BAW-1436.
The procedures used in obtal'11ng these estimates are given in BAW-1485 (proprietary). The artimated uncertainties in the surveillance capsule ' analysis are also provided in Sect.
4 (Tables 4-I and 4-2) and Appendix F of BAW-1485 (proprietary). Fluences at the vessel wall locations'may be as high as 250%.8 The proceduce outlined in BAW-1485 identified many of the sources of uncertainty stated in Table 6.1 but did not specify all their values..
6.4 Conclusions and Recomnendations The methodology used by B&W is similar to that used by other vendors and service lab' oratories. One weakness in the methodology is the analysis of uncertainties. This analysis should provide not only estinates for the sources of uncertainty identified in Table 6.1, but a statistically sound technique for combining all the individual estimates to arrive at an overall uncertainty for the fluences at the RV wall. This task requires considerable effort and funding on the part of the vendors, and only recently has work been done in this area.1-5 Another weakness relates to surveillance Josimetry measurements and the subsequent analysis to obtain fluxes 1 1 MeV or other suitable neutron exposure parameters. To address this problem, o
O e
..p
,-e----
=. -
8 Table 6.5.
Pred19.ted Fast Fluence in Oconee-1 RV at the Axial Welds for 8 EFPY Beltline Location Mate rial Surveillance luside of RV T/4 3T/4 Outside of locations Wall RV Wall tipper long. weld SA-1073 4.95E+18 2.93E+18 1.63E+18 3.70E+17 1.39E+17 tilddle long. weld SA-1493 4.83EF18 2.86E+18 1.59Etl8
- 3. 6 5E+17.
1.36E+17
- 1. owe r long. ' weld SA-1430 6.42Etl8 3.80E+18 2.!!E+18 4.79E+17 1.81E+17 Peak flux location 7.30EF18 4.32E+18 2.40E+18 5.45E+17 2.07E+17 i
~
an W-1436, Tables 6-6 and 8-1 O
e
6-8 i
a pressure vessel benchmark facility for power reactor surveillance dosimetry validation and certification is.needed to help identify and reduce uncertainties.
It is thought that,,with care,_an overall uncertainty of 210-30% for the fluences at the vessel wall should be achievable.
x
{
REFERENCES - CHAPTER 6 1.,
W. N. McElroy, Ed., LWR Pressure Vessel Surveillance Desimetry Improvement Program: PCA Experiments and Blind Test, NUREG/CR-1861 (1981).
2.
W. N. McElroy, F. 3. K. Ean, E. D. McGarry, A. Fabry, LWR Pressure Vessel Surveillance Dosimetry Improvement Program, NUREG/CR-1747 (1980).
3.
R. E. Maarker. J. J. Wagschal, 3. L. Broadhead, Development and i
Demonstration of an Advanced Methodology for LWR Dosimetry Applications, (EPRI eport to be published in 1981).
4.
J. J. Wagschal, R. E. Maerker, Y. Teivin, " Extrapolation of Surveillance Desimetry Information to Predict Pressure Vessel Fluences," pp. 631-632 in Trans. Amer. Nucl. Soc., 1980 Annual Meeting, Vol. 34.
i i
W. N. McElroy, et al., Surveillance Dosimetry of Operating Power Plants, Hanford Engineering Development Laboratory, HEDL-SA-2546 (1981).
6.
W.-W. Engle, Jr., A User's Manual for ANISN, A One-Dimensional Discrete Ordinates Transport Code with Anisotropic Scattering, Oak Ridge Gaseous Diffusion Plant, USAEC Report K-1693 (1967).
7.
W. A. Rhoades, D. 3. Simpson, R. L. Childs, and W. W. Engle, The DOT-IV Two-Dimensional Discrets Ordinates Transport Code with Space Dependent Mesh and Quadrature, Oak Ridge National Laboratory, ORNL/'Di-6529 (1979)_.
8.
W. N. McElroy et al., Surveillance Dosimetry of Operating Power Plants, HEDL-SA-2546 (October 1981).
e 9
e s,-
eea w-
-. m
-,,.-.e
--,y
--r-
z
. s
' },
x t
x-s e-7.0 PRESSURE VESSEL MATERIAL PROPERTIES
$s a
\\
7.1 Material Properties Required for Vessel Integrity Studies The material properties required for studying vessel integrity can be grouped in accordance with the three types of analyses that must be performed (Table 7.1).
Table 7.1 Reactor Vessel Material Properties Needed for Vessel Integrity Studies 4
Thermal Analysis Thermal conductivity (k)
Specific heat (c,)
Density (p) c.a Stress Analysis
~
s.-
Linear thermal coefficient of expansion (a)
Modulus of elasticity (E)
Poisson's ratio
'(v)
Yield and ultimate strengths (oy, a )
u Fracture-Sechanics Analysis Static crack-initiation coughness (Kre)
(I,)
Static crack-arrest toughness t
Reference nil ductility temperature (AniDT) l
- 7. 2 Dependence of Material Properties on Chemical i
Composition, Temperature, and Fast-Neutron 21uence Generally. speaking, all of the material properties in Table 7.1 are functions of material chemical composition, temperature, and fast-neutron fluence and must be treated accordingly in carrying out the
~
vessel integrity studies.
~
7.2.1 Material Chemical Comoosition and Heat Treatment With the exception of a few of the earliest reactor pressure vessels, all belt-line regio-s of U.S. commercial reactor vessels presently in service were fabricated from the three materials described in Table 7.2.
Two additional materials that must be considered are the weld filler material (used to join sections of rolled plate and forging rings) and the vessel cladding (applied by depositing weld metal). The weld filler material is similar to the base sacerial, whereas the cladding is an austenitic stainless steel (18 Cr-8 Ni).
o 7-2 s
A chemical element of special'ine.erestita both the vessel base material and-associated ralds it. copper, since it enhances radiation damage, uptch results 1n reduced, fracture toughness. Righ concentrations of
~
copper exist in rose welds becavse the weld wire was plated with
' N : copper; iLis 'elen:ent is also present as an impurity in the base
~
~
,s material.
~',i 4
i s
~
t, The -chemical ecaposition $1 the 'vess11 materials influences all the
\\
J I
parameters in Table 7.1, while the vtrious vessel heat treatments
- (tempering and stress relieving) affect primarily e, o, and i,
y u
.\\
17.2.2 Temperature Dependence of Properties
~
The temperature dependences of.the parameters in Table 7.1 are reatocably well known, sad e4ch parameter (with the exceptica of RTNDT and v which har a ns:;11gible* temperature dependence) or an appropriate combidatica thereof,' is included in.the ASt2 Pressure Vessel Code as a s
functiori of temperature. Values for k, k/pe, e, E, a, and a y
s p
y u
s.
t.
w Tabis 7.2 Materials Used'in the Fabrication of U.S.
Commerciaq.RasetorVessels N
1 g
Wgt. Eere sos'icion for Designated Materials Place SA 302 Plate SA 533 Forging SA 508 Element GR B GR B, C1 1 C1 1 s.
O.25 max 0.27 max C
0.25 max
?
-Cr 0.25-0.45
[
s
'Ni
-0.40-0.70 0.50-1.00 Mo 0.45-0.60 0.45-0.60 0.55-0.70 1.15-1.50 1.15-1.50 0.50-0.90 Mn Si 0.15-0.30 0.15-0.30
,0.15-0.35 P
0.035 max 0.035 max 0.025 max S
0.040 mx 0.040 max 0.025 m.x Fe balance-balance balance s
mM 4
i j
'o 7-3 as functions of temperature are inc1'uded in ASME Section III, while i
KIc and Kg, as functions of T - RTNDT are included in ASME Sectica KI for temperatures (T) less than that associated with the 1
upper shelf,2 The uccertainty in the parameters included in ASME Section III is not large and is expected to have a negligible effect on the evaluation of vessel integrity. However, the uncertainties in Kge and Kg, vs T - RTNDT is substantial, and there is also a significant uncertainty (approximately 220*F) in the determination of RTNDT.
The curves for Kge and Kg, vs T - RTNDT in ASME Section II represent the lower bound of a limited amount of data that were obtained sou years ago for A533 and A508 asterial. The use of this lower bound would appear to be conservative.; however, recent 3
j experiments with large test cylinders *" indicate that long cracks in latge structures will usually behave is accordance with the lower bound of data obtained from a large number of small (IT to 3T ~;T) specimens.
The uncertainty in the ASME Code lower bound is under investigation at this time.
The uncertainty associated with the use of RTNDT as a normalization factor is also under investigation. An alternative to relying on RTNDT j
is to determine', through laboratory testing, Kre and Kgg.vs T for each vessel. Howevet, this epproach appears to be impractical i
because of the large number of specimens required, since RTNDT is a
]
function of fluence.
During a reactor ves.el thermal transient of medium duration, the out ar portion of the vessel wall remains at temperaturas corresponding to ductile behaviour (i.e., the upper-shelf portion of the Kre vs T i
curve).
It is not likely that cleavage (brittle) /racture can proceed through this zone; however, the crack may tear at relatively low crack-tip velocit,ies to a depth at which plast.I: instability is achieved.
E Tearing-resistance material property data are required for an accurate analysis, and such data are, at present, very limited. An alteractive
' approach currently used is to assume what appears to be a conservative upper-shelf toughness that is essent'. ally independent of camperature and flaence; this upper shelf value is then compared to the calculated strees intensity factor. For very severe accidents such an approach is probably adequate, but for less severe cases the tearing resir tance may terminate crack propagation. The degree of -:onservatism in the present model for the less severe cases is not. known.
l 7.2.3 Dependence of Material Prooerties on Fast-Neutron Fluence l
Of the material properties listed in Table 7.1, those that have a l
significant dependence on fast-neutron fluence are o, o ' K u
Ic' Kg,, and RTNDT. Radiation damage increases a and, ko a lesser y
o, while Kre and Kg, are dtcreased and RTNDT is
- extent, u
increased. The average of o and a is used at the conclusion of y
u the f racture-mechanics anslysis-to see if the uncracked ligament has become plastic under pressure loading. Usually, strength values for
-+~amy+e-m,-,g
--y-,
,-.mu~.-y,~,,.,.-_--,..,-e.--,-e- -.,.. _ - _, -.
,,,_.,.-+,,.,..--,,,3..,--.,_,,.._v._,.,y., -. -, -.
+-c
7-4 i
i the unirradiated material are used, and this approac~u introduces some conse rvatism. Some data on elevated strength are available, and their use in the analysis would result in a higher permissible pressure during asthermal transient.
In accordance with ASM. code procedure, the decreases in K and Kg, due to radiation @ange are estimated by shifting the
,e and KIa vs T curves along the temperature axis by an amount ARTNDT = f(F, Cu, P), where F = fast-neutron fluence (E >_1 MeV) and Cu and P are the copper and phosphorous concentrations, respectively.
4 Values for ARTNDT = f(F, Cu, P) are included in Reg. Guide 1.99, Rev.
1,5 which was thought to be conservative at the time 1 its issuance.
A more recent evaluations of the available data indicates that the ARTNDT for materials containing a high concentration of nickel, which, appears to enhance the effect.of copper on radiation damage, agrees rather well with Reg. Guide 1.99, while lower concentrations of nickel.
result in conservative values for the reference temperature shift.
There are indications that many of the welds in the older PWR reactor vessels have high concentrations of nickel, and thus estimates of ARTNDT from Reg. Guide 1.99 presumably are not excessively conservative. However, the data base is much smaller than would be desired and will take some time to increase substantially, even though surveillance specimens from power reactors are becoming available and irradiation programs at materials testing reactors are under way.
I To date, radiation damage to the cladding is of little concers to the analysis of overcooling accidents because it has been assumed that the initial flaw will extend through the vessel cladding into the base sacerial, and also that the flaw will be very long on the surface so that cladding resistance to crack extension is not important. However, an assumption of long initial flaws may be unnecessarily conservative, since the presence of cladding may prevent short flaws f roa extending, i
particularly if the cladding retains its high taaring resistance at high fluences. There is a limited amount of data 7 for veld cladding i
i that indientes a substantial reduction in Charpy upper-shelf energy
(~100 to 30 fr-lb) at a fluence of ~8 x 1018 n/cm2 and an irradiation
(
temperature of 550*Y.8 Thus, it is not clear that the cladding will prevent short flaws from growing long.
REFERENCES - CHAPTER 7 l
1.
ASME Boiler and Pressure Vessel Code,Section XI.
{
2.
T. U. Marsten, Ed., " Flaw Evaluation Procedures, ASME Section II, Electric Power Research Institute, EPRI NP-719-SR (Aug 2st 1978).
3.
R. D. Cheverton and S. K. Iskahder, HSST Program Quarterly Progress l
Report for October-December 1980, NUREG/CR-1941, pp. 37-50 (March 1981).
l O
.-w-e,w_,.v.,-
or.-
-._,e.--,.-.-m_m_
..,, ~,, - - - _.,,
_-.-..r,,._-
_,----,_,_m-
- -.. - - -, - + -
7-5 References (Continued) 4.
R. D. Cheverton et al., HSST Program Quarterly Progress Report for January - March 1980, NUREG/CR-1477, pp. 16-28 (July 1980).
5.
U.S. Nuclear Regulatory Commission, " Effects of Residual Elements on Predicted Damage to Reactor Vessel Materials," Regulatory Guide 1.99, Rev. 1 (Sept. 1976).
6.
P. N. Randall, U.S. Nuclear Regulatory Commission, personal communication (Sept. 1981).
7.
F. J. Loss, Naval Research Laboratory, personal cemmunication (July 1981).
8.
It is of interest to note that be::2use of a recent change in core loading, the anticipated fluence by the end of 32 EFPY for Oconee-1 is ~1.1 x 1019 neutrons /cm2 (personal communication, Duke Power Company, October 1981).
O e
4 e
t dP e
O n--
-n
-s---
re m.
e-v,-
e-
- ~ - - - -, - - -
~<
-~~-------vn-
i INTZURITY Q REACTOR VESSELS DURING OVERC00 LING ACCIDENTS 8.1 Description of Basic Problems During an overcooling transient in a PWR the reactor pressure vessel is subjected to a thermal shock in the sense that thermal stresses are created in the vess61 van as a result of rapid removal of heat fros 4
its inner surface. The thermal stresses are superimposed on the pressure stresses, with a result that the not stresses are positive (tensile) at and near the inner surface of the van and are substantiany lower and perhaps negative elsewhern, depending on the magnitude of the pressure stress. The concern over the high tensile stresses near the inner surface is that they result in high stress intensity factors (K ) for any inner-surface flaws which any be g
p resent. To compound the matter, the reduced temperature and the relatively high fast-neutron fluence near the inner surface result in relatively low fracture toughness values (K and Kh) foe the vessel gg material in the same area. Thus, there is a possibility of crack p ropagation. The positive gradient in temperature, combined with the negative gradients in stress and fluence through the wall, tends to provide a mechanism for crack arrest deeper in the van. However, if the crack is very 1)ng on the surface and propagates deep enough, the remaining vessel ligament win become plastic and the vessel internal pressure win ultimately result in rupture of the vessel,. Thus, fo r i
each thermal transient there will be a aaw4=n= permissible pressure that is a functics of time..
)
Crack propagation may also be limited by a phenomenoa referred to as warm prestressing (WPS), which has been demonstrated in the laboratory i
with saan specimensl and also in a rather large, thick-waned cylinder l
during a thermal shock experiment.2 In such cases, WPS simply refers to the inability of P. crack to initiate while Kg is decreasing with
- time, i.e., while the crack is closing.
While this special situation is encountered during some specific overcooling accidents, caution must be exerciseii in taking credit for WPS because changes in the pressure that affect little else can delay or eliminate the requisite conditions for WPS.
The integrity of a reactor vessel during a postulated overcooling event is evaluated in terms of the centinued ability of the vessel to contain the coolant in such a way that melting of the reactor fuel win not occur. Generany speaking, this means that the water level must be maintained above ghe core, and to do this there must not be a significant breach in the vessel wall. below the level established by the top of the core. Therefore. it is necessary to determine if a preexistent flaw will propagate through the wall, and'if it will, to estimate the probable size of the breach and its resistance to leakage.
e
,y--
,.c,--,-,-.-,,,
--.-,-y,,v.---,,c-vy-w,
+. - - - - - -.
-.,,,m._,,
,,...-,,.~------m.._.~-.
-,,-,m.-.-_.-..
8-2 The investigative effort thus far has been directed at understanding the behavior of flaws during thermal, pressure, and thermal plus-pressure loading conditions.
It has been assumed on the basis of limited available data that if the temperature of a major portion of the coolant in the primary system is well above 212*F at the time a long flaw penetrates the wall, the final opening may be excessive in the sense that flooding of the core could not be maintained. Methods for estimating the size of the opening more accurately will be evaluated in the near future.
In the following paragraphs a calculational model for predicting, crack behavior during. overcooling accidents is described, and a summary cf results for specific accidents is presented and discussed. The reactor plant analyzed for thase studies is Oconee-1, and the postulated accidents include a main steam line break, a turbine trip followed by stuck-open bypass valves, a small-break LOCA, a Rancho Seco-type t ransient, and a large-break LOCA.
8.2 Calculational Model The calculational codel consists of three basic parts: a thermai analysis of the vessel vall, s. stress analysis, anc ai f racture-mechanics ' analysis. The " thermal analysis is performed for cylindrical geometry, is one-dimensional (radial direction), includes an insulated outer surface, and accepts a transient coolant temperature at the wall's inner surface. A time-independent inner-surface thermal resistance is used that is the sum of the fluid-film resistance and the cladding resistance, in which case the heat capacity of the cladding is igno red.
'Since an importat.t input to the thermal analysis is the temperature of the coolant in the downcomer region, some of the assumptions used in obtaining this temperacure need to be centioned. Depending on the nature of the overcooling accident, the temperatures of the coolant entering the downcomer may be different for the different inlet coolant pipes. Thus, there can be a:imuchal and axial variations in the downcomer coolant temperature. An accurata determination of the temperature distribution as a function of time would be very difficult, and the subsequent use of two-or three-dimensional stress and f racture-mechanics cwputations would be impractical for a parametric-type analysis. An additional complication in this regard for B&W reactors is that relatively warm water f rom the core outlet may enter the upper portion of the downcomer region through the vent valves and -
may thus raise the temperature of the downcomer coolant. For the purpose of the present anal-sis the coolant temperature (temperature t'ransiect) used 4.s input to the vesse] ther=al analysis corresponds, with one exception, to the lowest of ths cold legs (inlet lines). The degree of conservatism associated with this assumption is unknown.
e
.,.m e
8-3 The stress analysis is also one-dimensional and is performed for a cylinder; the cladding is excluded, as is the flaw, consistent with the method used for calculating the stress intensity factor. Loads on the cylinder consist of a radial temperature distribution and internal pressure, both of which are treated as functions of time.
Linear elastic fracture mechanics (LETM) is used for predicting flaw behavio r.
The flaw assumed for this particular study is an inner-surface, long, axially-oriented, sharp crack of unifors depth along its length. Thus, an accurate, two-dimensional (radial and azimuthal) model can be used. Consideration of a long axial surface flaw is realistic and necessary in the senso that short flaws will tend to become long flaws under thermal-shock loading, and the stress intensity factor is greater for long axial surface flaws than for any others.
The two-dimensional model s'oes, however, introduce some conservatism since there exists an axial gradient in fluence (and hence in toughness) that is ignored but which will provide some additional resistance to crack propagation. Furthe rmo re, the cladding, which tends to be a much tougher material than the base material, may suppress the surface extension of short flaws that has been predicted
)
and observed in the absence of cladding.
i As already mentioned, although the cladding is included in the ther:nal analysis it is not included in the fracture-mechanics analysis. the presence of cladding reduces slightly the tensile stress in the base material during a thermal transient. However, if the flaw extends through the cladding, the Kr value is significantly greater than if the cladding were ignored particularly for shallow flaws. Thus, the
- minimum critical crack depth for crack initiation would be less and the threshold flunce for crack initiation and vessel failure would also be less. A detai?ed quantitative assessment of this cladding effect is not yet available.
Fracture-to'ughness curves (K and K vs T - RTND9 for this study re r
were Ctken from Sect. XI of the ASME, ode,3 and an upper-shelf C
toughness of 200 ksiffE was added for both K and K rq Input to the f racture-mechanics analysis incluces (1)r,.the temperature and fast-neutron fluence distributions through the wall, (2) the thermal and pressure stresses without the presence of the flaw, and (3) the copper (Cu) and phosphorous (P) concentrations. The temperature and fluence distributions, coupled with the Cu and P concentrations, are used to calculate K and K, radial distributions se various times rg r
in the transient, and the stresses are used to calculate K values for r
a number of crack depths, ranging from ~3 to 90% of the wall
~
thickness.
O e
N*'
T-TW dui--erw+-m-vmy
"="r
'w a+g
wmiet---t g.-eqrwe-w w-ei-
-w e
sim-
-+
w
- wwm-m ew wmy-t
'---t-
j o
8-4 i
The thermal, stress, and f racture-nechanies analyses were performed witb the OCA-I computer cod a, which uses a superposition technique to
'i accurately calculate Et values for long flaws using stresses for the unflawed cylinder. The purpose in using the superposition technique was to achieve the accuracy of a finite-element analysis at a f raction of the usual cost, thus making parametric studies feasible. A block diagram describing the code is shown in Fig. 8.1.'
Required input for the OCA-I analysis is also indicated in Fig. 8.1, and specific values used for the Oconee-1 studies included herein are given in Table 8.1.
l For these preliminary studies only five trnasients were analyzed, but for each transient several inner-surface-fluence values were included so that the threshold fluence (and thus the number of years of operation) for incipient crack initiation could be astimated.
8.3 Transients Considered for Oconee-1 8.3.1 Main Steam Line Break Time-dependent pressure and temperature curves for this case were submitted to ORNL by Brookhaven National Laboratory (3NL) on August 14, 19815 and are given in Fig. 8.2 and Table 8.2 (corresponds to case 3 in Section 5).
For the thermal analysis of the vessel the fluid-film 2
heat-transfer coefficient was assumei to be 1000 5tu/hr*f t..y, which corresponds to full-flow conditions (primary system) and a total 2
surface conductance of 330 Bru/hr*f t..y, 8.3.2 Rancho Seco Transient The March 1978 Rancho Seco transient was made available to ORNL by NRC 6
i in January 1981 and is shown graphically in Fig. 8.3 and in tabular form in Tab 13 8.3.
As described in Sect. 5, the coolant temperature is nessured upst eam of the injection point for the HPIS and is thus 4
probably somewhat higher than actually exists at the entrance to the downcome r.
The fluid-film heat-transfer coefficient at the vessel wall was assumed to be 1000 stu/hr*f t2..y, 8.3.3 Turbine Trip Followed by Stuck-Ocen 3ynass Valves This transient, which corresponds to case 1 in Section 5, was submitted 7
to NRC by BNL in July 1980 and to ORNL by 3NL in August 1980a and is shown graphically in Fig. 8.4 and in more detail in Table 8.4.
The portion of the transient beyond 1240 s was added by ORNL, assuming that the HPIS would pump against the relief valve setting (2500 psi) and that the temperature of the coolant in the downconer would remain at l
140*F.
The fluid-film heat-transfer coefficient was again assumed to be 1000 3tu/hr*ft2..y,
\\
l
e.
s l
i l
~
4
~
ORNL-0WG SI-1863R ETO k
CYL FITNDT F = fla/w) o K,K*
ID h
p p
g 8
ARTNOT AND o
= f(T) j T,
c Cu, P
= f(F, Co, P) p OD I
i i f I f j f j r j P a
CO3LANT TEMP THEllMAL ANALYSIS K,K, g
g WALL TEMP yg y
93 VS l
TIME TIME AND a/w b
TIME l.
K /K, K,/K,,
a,/w g
a i
1P VS
=
VS TIME AND a/w TIME 1
PflESSullE STRESS ANALYSIS K
g VS TilERMAL AND VS TIME PRESSURE STRESSES l
TIME AND a/w h REQUIRED INPUT J L J L TYPICAL DATA INCLUDED, i
a CYL OPTIONAL INPUT i
E ID K*. f(a. al
'j TYPICAL DATA INCLUDED, NO OPTION OD CALCULATION AND OUTPUT i
{
FJg. 8-1.
Block Diagrain of the OCA-I Computer Code Indicating Basic Input, l
Calculations, and Output.
I i
e f
5 5
?
0 8-6 Table 8.1.
Input to OCA-1 for Oconee-1 Analysis Vessel Dimensions, in.
Outside Dia.
189 Inside Dia.
172 Coolant Temp vs Time Specific Accident Pressure vs Time Specific Accident 2
Heat Transfer Coeff. (h), Stu/hr*f t.ey Large-break I.0CA 200 1
Others 330 Initial Wall Temp, *F 550 RTNDQ,'T 40 Copper Concentration (Cu),'%
0.31 Phosphorous Concentration (P), %
0.012 e
and K, Upper Shelf,
.KIe r
ksi Tin.'
200 Fluence at Inner Surface (F )
Range of Values o
- RTNDT = zero-fluence RTNDT (initial value) l o
i l
e i
er m
m m
8-7 _.
oRNL.QWG 81-18490 ETD 600 3000
' ~~~
soo.
2sco 400 200o C
O b
e 3 :oo TEMPERATURE E
isoo b
z w
h t
200 PRESSUR E 1000 100 soo o
o o
200 400 soo 800 1000 1200 TIME (si e
i Fig.'8-2.
Main Steam line break. temperature and pressure
' transients (HPIS r em*na active).
O e
-e r
---e
8-8 Table 8.2.
Main Steam I.ine Break Temperature and Pressure Transients (HPIS Remains Activated)
P2R Primary Time Level Pressure (s)
(ft)
(psia) 0 18.23 2157.8 1
17.44 2139.1 5
11.01 1964.35 10 2.10 1677.02 15 0.0 1440.25 20 0.0 1323.79 25 0.0 1216.75 30 0.0 1106.18 35 0.0 904.86 40 0.0 760.18 45 0.0 743.5 50 0.0 725.19 60 0.0 689.04 70 0.0 650.19 80 0.0 610.61 90 0.0 578.52 100 0.0 559.57 120 0.0 539.47 140 0.0 513.61 160 0.0 523.63 180 0.0 517.91 200 0.0 507.52 220 0.0 497.72 240 0.0 487.96 260 0.0 478.35 280 0.0 468.92 300 0.0 459.61 350 0.0 438.92-400 0.22 429.69 450 1.40 443.36 500 2.69 459.39 550 4.12 478.53 600 5.67 501.16 650 7.30 527.56 700 9.04 558.72 750 10.84 595.07 800 12.69 637.46 850 14.56 686.83 900 16.43 747.'1 950 18.30 818.56 1000 2L.14 903.93 1050 21.96 1006.48 1103.58 23.84 1139.87 1123.18 24.51 1195.84
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0 20 40 60 80 100 TIME (ment Fig. 8.3.
Temperature and pressure transients for Rancho Seco.
4 9
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-.--.-----v
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8-10 Rancho S' co Temperature and Table 8.3.
e Pressore Transients Time Tempe rature Pressure (min)
(*F)
(psi) 0 590 1500 10 490 1710 20 412 1880 30 356 2020 40 318 2110 50 296 2130 60 282 2100 70 280 2050 80 284 2000 90 299 1950 100 320 1900 l
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TEMPERATURE y
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100 500 PRESSUFIE O
0 (s) 0 200 400 600 800 1000 1200 1400 1800 I
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Fig. 8-4 Turbine trip vi:h stuck-open' bypass values (Scram 'Jorth = 0.061 Ak/k): Te=perature and Pressure Transients.
e er
8-12 Table 8.4.
Turbine Trip" with Stuck-Open 5ypass Valves (Scram Worth = 0.061 Ak/k):
Temperature and Pressure Transients x
Time Core Inlet Primary System (s)
Temperaturi (*F)
Pressure (psia) 0 556.00 2,192.00 5
563.64 2,289.31 55 528.73 1,570.92 125 444.09 523.95
~
215 328.42 320.58 340 271.34 275.21 630 197.45 365.93 800 167.70 463.54 900 157.08 561.81 1000 149.97 723.80 1050 147.34 847.90 1100 145.20 1,020.84 1150 143.45 1,273.30 1200 142.06 1,663.70 1240 141.17 2,149.54 O
e O
4 i
j 8-13
)
1 8.3.4 Small-B reak LOCA -
The SBLOCA case was defined and the thermal-hydraulic analysis was i
performed by Ias Alamos National Laboratory; it was reviewed by NRC in June 19818 (ease 5 in Section 5). The break was assumed to be in the i
cold les downstream of the main circulating pump and ahead of the HPIS injection nomsle and was sized at 10% of the pipe area. The thermal-hydraulic analysis was performed for the first 250 s only, and the resulting pressure and temperature transients are shown in Fig. 8.5 and Table 8.5.
The temperature given is that calculated for the top of the downconer; since the main circulating pumps were assumed to be tripped at 15 s into the transient and the HPIS was assumed to inject coolant for the duration of the transient, this coolant temperature is probably higher than would actually exist locally at the vessel. vall, i.e.,
i channeling of the HPIS coolant would occur.
8.3.5 Large-Break LOCA The 13LOCA has been under detailed investigation for several years 10 and differs f rom the other transients considered in that no repressurizaton of the primary system takes place, and the downcomer temperature transient consists of an essentially step change in temperature f rom normal operating temperature (550*F) to'~70*F.
The fluid-film heat-transfer coefficient used in the thermal analysis of the vessel corresponds to free onvection and was estimated to be 300 2
S tu/hr* f t..F.2which corresponds to a total surface conductance of i
~200 Btu /hr* f t..y, 4
J 8.4 Results of Fracture-Mechanics Analyses t
The fracture-me.chanics analysis will indicate one of three possible l
results for each specific case:
(1) there will be no crack iritiation for any reasonable assumed preexistent crack depth; (2) crack l
initiation will occur, but the crack will arrest permanently; cr (3) crack initiation will occur and the crack will penetrate the wall.
i The results of the analysis are quite sensitive to the radiation-induced reduction in toughness and thus to the fast-neutron fluence, which is a function of the operating time of the reactor. Therefo re, the resu.1.ts are summarized in terms of the threshold fluence for l
incipient crack initiation and for failure of the vessel, and this is done for two cases:
(1) assuming WPS to be effective (if appropriate conditions exist), and (2) assuming WPS not to be effdctive (even if
~
appropriate conditions do exist). A summary of results for the five i
overcooling accidents analyzed is presentud in Tables 8.6 and 8.7.
Table 8.7 indicates the total number of EFPYs that a 3&W-type reactor l
can operate before the overcooling transients considered would likely result in vessel failure.
The summary of results presented in Table 8.6 shows that for all cases analyzed the minimum critical crack depths for initiation t.rs in the we w - =vvv
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TEMPERATURE 2000 400 C
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PRESSURE w
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Small-Break LOCA temperature and
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e 8
8-15 Table 8.5.
Small-Break LOCA Temperature and Pressure Transients t
Time P
T (s)
(ksi)
(*F)
~ ~ ~~~~~ - ~~- ~ ~- ~
0 1.63 553 12 1.45 558 576
~
17.3 25 1.20 570 50 1.06 553 75 0.930 537 100 0.777 516 125 0.590 486 150 0.457 459 175 0.383 439 200 0.348 430 225 0.322 423 250 0.290 4 15 a
e O
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i 8-16 4
range of 0.17-1.3 in.
This implies 'that at least some of the flaws, because of their size (upper end of the range), might have a high probability of being detected by nondestructive means. However, for fluences 'somewhat greater than those associated with incipient initiation and failure the upper and of the range is such lower.
The calculated critical crack depth would be further reduced relative to the values in Table 8.6 by including the offect of' cladding in the f racture mechanies analysis, assuming, as we are, that the crack extends through the Padding. As mentioned earlier, the inclusion of cladding in the analysis wiu also result in smaller threshold fluences. Thus, in this respect the results in Table 8.6 and 8.7 are somewhat optimistic.
In evaluating the data in Tables 8.6 and 8.7 there is a distinction that must be made between the large-break LOCA, and the other cases.
- Since the LBLOCA does not involve repressurization, a long, axiauy.
oriented flaw presumably would not extend completely through the wall, and even if it did the crack would remain tight and thus leakage of coolant presumably would be negligible. For the other cases the primary system pressure is high enough, consistent with ~ assumptions made, to force the crack all the way through the wall. Furthe rmo re, since the system is pressurized, the temperature of the coolant could l
be high enough to result in srfficient energy rulessa.during. blowdown to open the crack substantially. As mentioned earlier, a detailed i
analysis of crack opening under these cizuwstances has not yet been pe rformed.
4 As indicated in the tables, WPS was predict d for each of the' cases
+
conside red.
It may be reasonable to take advantage of WPS for' the LBLOCA since, by definition, there is no repressurization; however, for the other cases, variations in the represeurization could preclude conditions for WPS without making significant d!!ferences in the results otherwise. Thus, one cannot necessarily depend on WPS to reduce the consequen:es of the transient.
- The fluences listed in Table 8.6 correspond to those at the inner surface of the vessel wall at locations having the specified copper i
concent rations. Thus, to determine the number of U?Ys in Table 8.7 it l
is necessary te know the fluence rate (fluence per UPY) at the same locations. To establish the most limiting Iccation one must consider the combined effect of fluence, copper concentration and initial RTNDT (RTNDT,). The location that would tesd to have the highest RTNDT (RTNDT ARTNDT) would be the likely choice. Such a location was estab12s+hed for Oconee-1 in Raf.11, and the corresponding fluence rate-is 0.046 x 1018 neutrons /cm / U PY.
According to the information in 2
Sect. 6 the uncertainty in this value is 250%. The nean value was used to obtain the threshold times (UPYs) to vessel failure listed in Table i
8.7.
t e
O w
s-m~
,---+,.,-~_m-
.m-nn--nr-r
en-v.-
Table 8.6.
Summary of Flaw Behavior Characteristics for Several Ilypothetical Oconee-1 Overcooling Accidencs (refer to the list of nomenclature for this table on the following page)
F a
t F
a t
(a/w}
11 (a/w)* A if (a/w)lf if 11 2
(n/cm x 10I9)
(min)
(n/cm2 x 10I9)
(min)
Large-Break LOCA WPS 0.4 NO.04
- 5. 5,
0.15 0.9 0.02-0.22
- 1. 5-8 Without WPS 0.15 0.1 30 0.5 0.2 0.04-0.18 14-45 Rancho Seco WPS 1.5 0.1 40 1.0 1.5 0.1 40 Without WPS 0.9 0.1 65 1.0 0.9 0.1 65 Turbine Trip /Open Bypass Valves WPS" 0.2*
0.06 22 1.0 0.2" 0.06 22 A on.
d d
Withuut WPS 0.13 0.15 60 1.0 0.13 0.15 60 Main Steamline Break WPS 0.4 0.04 9
1.0 0.4 O.04 9
d Without WPS 0.2 0.06 18 1.0 0.2 0.06 18 Small-Break LOCA No initiation
- alf a range of crack sizes is not shown, shallower and deeper. flaws will initiate at higher fluences.
bFor this case, failure refers to crack penetration beyond a/w = 0.9.
Presumably the crack will not actually penetrate the outer surface.
CCrack depthe greater than a/w = 0.2 ignored.
dIf the transient time were extended -beyond 60 min., Fi would be less.
Snuration of thermal-hydraulic transient simulated too ahort to permit meaningful fracture mechanics analysis.
L i.
e
o l
8-18
~
Nomenclature for Table 8.6 (a/w)gg f ractional crack depth for incipient initiation (a/w)di f ractional crack depth for final arrest following incipient initiation (a/w)g f ractional crack dapth for first initiation that results in vessel failure (corresponds to threshold fluence for failure)
F threshold fluence at inner surface of vessel wall for incipient crack gg initiation F
threshold fluence at inner surface of vessel wall for incipient vessel gf failure t
time in transient for incipient crack initiation gg t
time in transient for incipient vessel failure gf e
e 0
e e
e I
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1 l
~
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R.-
Table 8.7.
Estimated Threshold Times for Vessel Failure for a B&W-type Reactor for Various Overcooling Events i
Threshold Timea b
Comments and Qualifications La rge-B reak LOCA 20 WPS assumed effective (4 EFPY w/o WPS). Crack not expected to penetrate RV well; vessel integrity l
expected to be maintained, -since repressurization does not occur.
1 Mancho Seco 20 WPS not assumed effective (33 EFPY if it were).
Through-wall crack predicted.
h Turbine Trip with Open 3
WPS not assumed effective (4 EFPY if it were).
Bypass Valves (RFT) c Through-wall crack predicted.
~
l Hain Steam Line B reak c 4
WPS not assumed effective (8 EFPY if it were).
Through-wall crack predicted.
[
j e
Small-B reak LOCA Thermal-hydraulic forcing functions calculated to only 250 s (at which time temperature and pressure are still decreasing), thereby preventing meaningful analysis of crack propagation in RV.
a aEffective full power years at which failure of the vessel is predicted, given the pressure and thermal driving functions presently predicted for the transients, and the assumptions used in thlm study.
b II of 0.046 x 1018 n/cm3/EFPY. Depending on the Based on.a fluence accumulation rate value specifica of surveillance programs and fuel management schemes, this value may have an associated uncertainty of as much as +50%.
cFa11ure predictions based on thermal-hydraulic calculations containing conservative i
assumptions.
8-20 REFERENCES - CHAPTER 8 1.
F. J. Loss, R. A. Gray, Jr., and J. R. Hawthorna, Significance of Wars Prestress to Crack Initiation During Thermal Shock, NRL/NUREG Report 8165 (Sept. 29, 1977).
2.
- 1. D. Cheverton and S. K. Iskander, HSST Program Quarterly Progress Report for October-Decembe,1980, NUREG/CR-1941, pp. 37-50 (March 1981).
3.
ASME Boiler and Pressure-Vessel Code,Section II.
4.
S. K. Iskander, R. D. Cheverton and D. G. Ball, OCA-I. Code for Calculating the Behavior of Flaws on the Inner Surface of a Pressure Vessel Subjected to Temperature and Pressure Transients, ORNL/NUREG-84 (August 1981).
5.
Letter to R. Kryter, ORNL, from R. J. Carbone, BNL, Aug. 14, 1981.
6.
Letter to R. D. Cheverton, ORNL, from J. Strosnider, NRC, January 30, 1981.
7.
Letter to Novak Zuber, NRC, from M. M. Levina, BNL, July 1, 1980.
8.
Letter to R. D. Cheverton, ORNL, from M. M. Levine, BNL, August 11,. 1980.
s i
9.
Letter to C. Z. Serpan, NRC, from S. Fabic, NRC, June, 22, 1981.
10.'
R. D. Cheverton, S. K. Iskander and S. E. Bolt, Anolicability of LIFM to the Analysis of PWR Vessels Under LOCA-ECC Te.armal-Shocic Conditions, ORNL/NUREG-40 (October 1978).
l
- 11. J. Strosnider, NRC, personal communication (September 1981).
l l
e e
e 9
t e
.c 9.0 POSSIBLE MITIGATIVE MEASURES 9.1 Operator Actions In the case of the Oconee-1 control system, reactor operators clearly have a manual capability to terminate excessive feedwater for a wide variety of failures. However, such actions require that the operator correctly recognize overcooling problems at a time early enough in the transient that his resulting response is effective. Diverse sources of information are available to the operators, from which a determination of overcooling can be made. It is therefore possible that operator response can be an effective deterrent to this probles.
It is conceivable that the operator might be able to control the outlet subcooling for certain accidents, as B&W proposes. How ptactical this is, considering instrument locations and fluid transport times (particularly if the reactor coolant pumps are tripped), needs to be evaluated.
The questions of whether and/or when to trip the reactor coolant pumps in overcoolinE upsets need evaluation. Tripping the pumps will raise the temperature and delay the influx of cold water to the vessel for steam generator-driven transients; however, maintaining a properly cooled core and promoting good downconer mixing any necessitate leaving the psaaps on.
It is evident that the operator needs a definitive indicator of adequate core cooling and vessel wall temperatures to achieve a proper balance between conc, erns for core cooling and vessel overcoolings.
4' 9.2 System Changes
(
9.2.1 Oconee Changes Our review of the Oconee-1 control system revealed that several of the upgrades already perfor:ned will act to reduce the likelihood and extent of excessive feedwater transients in this plant. Among these changes are upgrades to the instrument and control system power supplies, identification of diverse information channels for operator use during power supply failures, automatic feedvater pump trip on loss of ICS power, and stesa generator high level limit tri;;s for the main feedwater pumps. This reduces the probability of excessive feedwater being supplied to that of two failures of the control system.and failure of the operator to take corrective action or overt operator error in manually actuating the feedwater.
e 9
e I
l
9-2 Other potential changes can be identifed. An analysis of the condensate booster ptap response following main feedwater pump trip could be used to determine whether or not other feedwater system control actions should be initiated on loss of ICS power. Similarly, analysis could be performed to determine whether a main feedvater trip based on a specified pressure-temperature envelope could be used to prevent inadvertent overcooling. A more detailed systems analysis could possibly identify other alternatives.
9.2.2 General Chanses There are several other changes to the reactor system which deserve evaluation for their effectiveness, cost, and practicality of implementation. These are outlined below.
Borsted Water Storate Tank (BWST) Tesoerature. Increasing the temperature of the 3WST fluid would reduce the degree of overcooling caused by actuation of the safety injection systems. However, this measure obviously has limited effectiveness, since some tanks probably cannot be heated above ~80*F and, in any case, maintenance of temperatures above ~200*F would be impossible without tank p ressu rization.
Feedwater Train Storase. 't.imiting the amount 'of water available to the f eedwater system would obviously reduce the severity of the steam-generator-driven transients. However, there is a trade-off here with practical requirements for normal plant operation.
Containment Floodina. This "fix" has been proposed and discussed in connection with other major accidents.
It would obviously mitigate the consequences of a reactor vessel breach; however, inadvertent actuation of such a system might itself produce a vessel breach through rapid and extreme overcooling of the RV outer wal,1.
9.3 Restoring Pressure vessel Fracture Toughness By Annealing From a vessel design point of view, the most desirable solution to the i.
overcooling-accident vessel-integrity problem is to restore the i
vessel material f racture toughness, which is gradually reduced during l
reactor operation as a result of exposure of a vessel wall to fast neutrons. Studies that have been underway for several, years indicate that the toughness can be restored by annealing at temperatures in the range of 750-850*F for a period of approximately 200 hours0.00231 days <br />0.0556 hours <br />3.306878e-4 weeks <br />7.61e-5 months <br />,2 The l
results of studies conducted by Wastinghouse, under contract to EPRI, indicate that conducting the annealing treatment of the irradiated vessel is practical for some and perhaps most of the vessels in service l
today.
_. -. - - - -.. ~ - - -,,.. - - - - - -
f.
REFERENCES - CHAPTER 9 1 T. R. Mager (Westinghouse), personal communication to R. D. Cheverton (ORNL), September 29,.1981.
2 F. Loss et al. (NRL).
I i
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.,...., _ _. _ _. -., _,,,. _, - _ _. -. _,, _, _, _ _., - _ -....,.. -., _, _ _,, _ _ _..,...., -., _,....., ~. _
a 10.0 CONCLUDING REMARKS A. J RECOMMENDATIONS FOR TURTHER liORK V
10.1 Ceneluding Ramarks
}'
Despite a fair degree of recent effort in the study of pressurized thermal shock phenomena by a maber of knowledgeable groups, the true severity of the threat is, at present, very difficult to ascertain with confidence. The principal problems contributing to uncertainty sre:
h computer codes presantly being used to sisulate the hypothetical overcooling transients were cot designed to treat some of the phenomena that take place and hence produce inaccurate (sometimes nonphysical) thermel-hydraulic forcing functions under certain circumstances. h results of the f racture-mechanics analysis used to predict vessel failure are known to be sensitive to the temporal behavior of these forcing functions.
The temperature indications in the lone set of actual plant data available (Rancho Seco) provide only a nominal indication of true RV wall conditions (thay could be too high or too low by an indeterminate amount), and the chart-recorded pressure traces are made suspect by the presence of large " spikes" of unknown and possibly nonphysical origin.
h thermal / stress / fracture-mechanics analyses presently used to predict' crack propagation resulting from the temperature and pressure forcing functions have limitations (e.g.,1-D thermal-and stress analysis; lack of treatment of azimuchal and axial variations in downcomer. coolant temperature; inability to account for the axial gradient in wall fluence; lack of treat-ment of vessel cladding in f racture-mechanics analysis) which introduce uncertainty of an unknown magnitude in the results.
A h fluence at the vessel wall and at critical welds is probably known only to an accuracy of 130% (perhaps 150%), and this implies an uncertatuty of like msguitude in the vessel " life remaining" figures.
h probabilities of occurrence for various overcooling accident initiating events have associated uncertainties of at least plus-or-minua one order of magnitude, and the conditional probabilities for correct subsequent operator. diagnosis of a transient, timeli-ness and correctness of operator response, appropriate. automated control and safety features responses, and the like ars at present undetermined.
O am
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n.,,-w n.
.m.
m-
=_ -..
s, d
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1 i
Nonetheless, for all their shortcoAings, the analyses at hand are the best i
presently available 'on a no1 proprietary basis, and, cwing to the apparent
. severity of the outcoats predti: cod fros, the -limited number of overcooling scenarios seudied, it 1.;, our'ottnion that p sseurized thermal shock suet be regarded as a serious potential ~ threat and geri;s a great deal more study using refined techniques. -
m s
10.2 Recommendations for Turther Work
, sg s
In ' order to reduce the angnitudes of the uncertainties described above, we recommand' that additional work be undertaken in the following areast
~
o;,
~
Refinement of thermal-hydraulic simulation codas and associated models (in particular, treatment of the feed train, fluid mixing, the control systes, primary coolant system repressurization and
~
flow distribution, tim phase phenamena, and the heat capacity of i'
heavy primary metal).
Reficament of vessel thermal / stress /f recture-sechanics analysis techniques (in particular, a corsideration of higher dimension-ality in several of the variables tteated,,and inclusion of the vessel cladding in the frecture mechanics).
Refinement of'the, analytical methods and surveillance capsule 1
data assessment procedures required to estimate fast-neutron fluence in selected 4rass of the RV wall, in order that state-of-the-art accuracias (+10%) may be r.alized.
A thorough study of the probability structure of the various intertwined occurrences (among them, normal plant maneuvers, p.buce events, equipment and operator failures, plant recovery actions, etc.) that are necessary to prodece the severe thermal shock conditions that constitute a serious threat to RV integrity.
Some facets of this recommended work are known to be in progress by the NRC and reactor owner's groups or will be initiz.ted in FY 1982.
9 l-,....--.--....--.-_.---------.-..
-- ~
r
.i s' -
APPENDIX A CAPSULE DESCRIPTIGN OF DATA BASES EIAMLNED
~
1.
BHRA - Fluid Engineering 7,
BRRA Fluid Engineering provides indexing and' abstracting of world-wide information on all aspects of fluid engineering, including statics and dynamics, and laminar and turbulent flow. Theoretical esearch is covered, as well as the latest technology and applicatiew. Data are taken from 3RRA's ten secondary abstract publications, which abstract over 550 a.echnical journals as well as books, proceedings, standards, technical reports, and 3ritish patents. Fajor fields covered include civil engineering hydraulics, industrial aerodynamics, dredging, fluid flow, fluid power, fluid coaling, fluidics feedback, and tribology.
2.
CIM - Inven*ory of Models The Central Inventory of Models data base is maintained by ORNL for DOE, and 12cludes energy-related bibliographic and numeri.c data-bases, graphics packages, integrated hardware / software systems, and models frort DOE laboratories.
3.~
COMPENDEI - Engineering COMPENDEI covers significant world-wide engineering literature (1970 to date) from ~2,000 serials and over 900 monographic publications (including books and conference proceedings). Fields of engineering and related subject areas include: aerospace engineering, agricultural engineering and food technology, auccmotive engineering, bioengineering, chemical engineering, c'vil engineering, computers and data processing, construction materials, control engineering, electrical engineering, electrontes and comnanications engineering, engineering geology, engineering physics, fluid flow, and heat and thermodynamics. Also covered are industrial and management applications, inst:uments and measurerents, light and optical technology, marine engineering, material properties and testing, mechanical engineering, metallurgical and mining engineering, nuclear technology, ocean and underwater technology, petroleum engineering, railroad engineering, transportation, water ard waterworks engineering, and pollution,
~
sanitary engineering, and waste.
4.
CON" - Conference Papers, This includes scientific and technical papers (1973 to date) in the life sciences, physical sciences, and engineering areas that are presan.:ed at regional, national, and international meetings, including small meetings having a cross-disciplinary focus.
ew,,..--,,m.-,
.,.,-,,-.#.y_.-..--e,
...e
.,w,
,,,v-,-
,,v.,--
+,.%,
+.,- -. _.,
e-m.-,
,,_m.--,,--,
,, -, -,, _,,,,. - -__,-c,,,-
e,s-
.e---
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---y-
.c4 2
"~
5.
EDS - Energy
. ' ~
DOE Energy is one of the world's largest sources of literature references on all apsects of energy and related topics. It includes references to journal articles, report literature, conference papers, books, patents, dissertations, and translations. All manner of energy topics are included:
nuclear, wind, fossil, geothermal, tidal, etc., as well as the related topics of environment, policy, and conservation.
6.
EIA - Energy Information EIA Citations are drawn from Energy Information Abstracts, and are compiled by the Environmental Information Center.
7.
FEDEX - Federal Government Activities Federal Index contains information (1976 to date) on federal government activities drawn from the Congressional Record, the Federal Register, The Weekly Compilation of Presidential Documents Commerce Business Daily, and the Washington Post. Additional sources from the F & S Index are also included, beginning in 1979.
The citations provide' access to the Code of Federal Regulations, the U.S. Code, Public Laws, Congressional Bills, Kosolutions and Reports. The information is indexed by acting government agency, affected industry, or institution and type of government action or function.
8.
ISMEC - Mechanical Engineering ISMEC covers mechanical engineering, production engineering, and engineering management.
Subjects covered include ptoduction processes, tools and equipment, energy and power, transport and handling, management and production, measurement and control, and mechanics, materials and devices. References (1973 to present) ar gathered from journal articles, technical reports, conference proceedings, and books.
9.
NSA - Nuclear Science The Nuclear Science Abstracts base presently contains more than 500,000 citations, covering the period 1967 to June 1976.
10.
NSC - Nuclear Safegg The nuclear safety information data base is maintained by the Nuclear Safety Information Center, ORNL, under the joint sponsorship of DOE and NRC.'
I d
O t-
)
4
~
3 11.
KRIS - Covernment Soonsored Research l
NTIS covers U.S. government-sponsored research and development technical reports from over 200 Federal agencies and some reprines, federally-sponsored translations, an/. foreign-language reports in major areas of technical interest. Its multi-disciplinary scope includes aeronautics, agriculture, astronomy and astrophysics, atmospheric sciences, behavioral and social sciences, biological and medical sciences, chemistry, earth sciences and oceanography, electronics and electrical engineering, energy conversion (non-propulsive), materials, and mathematical sciences. Also covered are mechanical, industrial, civil, and marine engineering, methods and equipment, ailitary sciences, missile technology, navigation, communications, detection methods and counter-measures, nuclear science and technology, ordnance, physics, propulsion and fuels, and space technology.
i 12.
RSI - Radiation Shielding Information The Radiation Shielding Information Center data base is maintained by ORNL and contains citations to literature describing computer codes that have been designed L.c perform radiation analysis and shielding calculations, neutron cross-section processing, and I
experimental data analysis.
l 13.
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y APPENDIX B OCONEE LICENSEE EVENT-REPORTS (LERs)*
1.
Cooldown Rate Limit Exceeded Following Loss of ICS Power at Oconee-3
- The RCS cooldown rate limit was exceeded after power to the ICS was lost for about two and one-half minutes. No ES actuation setpoin:s were reached, ard adequate RCS inver.cory was maintained.
No damage was incurred.
T. ass of ICS power resulted from blown fuses in normal invarter (II) and failure of trar.sfer switch to transfer automatically to regulated AC power. When ICS power was restored, excessive feedwater flow caused a rapid RCS cooldown. A redundant transfer switch has since been installed, and personnel have been instructed on how to respond properly to loss of ICS pcwer.
l 2.
Reactor Coolant System Cooldown Rate Excessivs at Oconee-3 During a routine shutdown for maintenance, a minor system transient occurred, which resulted in opening a power-actuated pressurizer relief valve with reactor power at IbL The vau.ve remained open and the RC system depressurizaton c.outinued until the isolation valve was closed. TIa shutdown continued with a i
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cooldown rate of 100*F/hr.
However-when the initial drop in temperature from depressurization. was included, the rate exceeded the 100*F/hr tech spec limit by 1*F/hr. It was determined that boric' acid crystal buildup on the connecting pis of the lever arm of the pilot valve had caused the valve to remain open.
4 3.
Additional Information on Excessive Cooldown Race at Oconee-3 Reactor, power was being reduced from 100% to 15% by the integrated control system for a maintenance shutdown. When 15% was reached, unit load demand was 65 We and power generation was 115 MW4.
This diffe'rence existed because the reactor was operating at its lower liate of 15% and could not f allow load demand. A transient
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occurred that tripped the reactor.
During the transient, a relief valve opened and failed to close. This transient was terminated by closing the isolation valve. Cooldown race was 101*F/hr during the first hour. The relief valve failed because of heat.
expansion, boric acid crystal buildup on the valve lever, and bending of the solenoid spring bracket.
- Text has been modified slightly in some instances to improve clarity and readability.
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4.
Faedvater Transient Following Scram Actuates HPI at Oconee-1 andaninvestigationwasin$Ethe T "Iated.'statalarm began to act arratically On' December 13, 1978,'
Latring invest.igation (12/14/78) recorder shorted, causing an apparent (not real) drop in Ta power cord supplying T,,I 13*F and ICS attempted o
correct T,,.
Unittrippedonhfghpressure/ temperature.
Feedwater transients during cooldown allowed OTSG T* to go dry.
When it was refilled it caused RCS pressure to drop below 1500 psi, which actuated the HPIS. The cause of the T,,, cord short has not been identified. The feedwater transients were probably caused by improper valve operation. The power supply cord was replaced.
5.
Reactor Coolant System Cooldown Rate Exceeds Limits at Ocnnee-2.
When a spurious signe.1 in the 230 kV switchyard circuit breaker failure relay circuitry resulted in the isolation of the switch-yard, the reactor scrammed from 75% power. The scram tripped the feedwater pumps. The emergency feedwater pumps started and filled the steam generators to the 95% level as designed. This high water level, plus norual required steam, resulted in a cooldown rate of 140*F/hr in one loop and 135.5'F/hr in the other, which exceeds the 100*F/hr limit. Reduction
'.n water level set point is being studied.
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,.8-APPENDII C SPECIFIC DOCUMENT REFERENCES ' ^ ~~ "' ~
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1.
" Reactor Coolant Pressure and Temperature Data for the March 20, 1978 Cooldowt. Event at the Rancho Seco Power. Plant,". letter, S. Fabic to C. Serpan dated November 25, 1980.
2.
" Parametric Julysis of Rancho Seco Overcooling Accident," letter, R. Cheverton to M. Vagins dated March 3,1981.
3.
Effect of HPI on Vessel Integrity for Small Break LOCA Event with 2 Extended Loss of Feedvater, BAW-1648 (November 1980).
- 4. " Runaway Feedwater Af ter Turbine Trip Report," letter, M. Levine to N. Zuber dated July 2,1980.
5.
" Transmittal of Preliminary Calculations of a Steam Line Break Accident," letter, S. Fabic to C. Serpan dated May 14, 1981.
6.
" Analysis of a Steam Line Break with Primary System Overcooling for a Typical B&W Reactor", letter, R. Carbone to R. Kryter dated August 14, 1981.
7.
" Completion of Scheduled Analyses on Pressurized Thermal Shock Scenarios," letter, S. Fabic to C. Serpan dated June 22, 1981.
8.
Analysis of Capsules OCl-F from Duke Power Comnany Oconee-1 Reactor Vessel Materials Surveillance Program, BAW-1421 (August 1975).
i 9.
"Oconee Nuclear Station Docket Nos. 50-269, -270, -287," Lette r Report, W. O. Parker, Jr. to H. R. Denton dated July 23, 1980.
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- 10. Final. Safety Analysis Report, Oconee Nuclear Station Units 1, 2, and 3, Rev. 19, Duka Power Company (May 5,1972).
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11.
IRT - A Pressurized Water Reactor System Transient Code, B rookhaven National Laboratory draf t report dated December 1980.
12.
" Thermal Shock to Reactor Pressure Vessels," Letter Report, R. W. Jurgensen to D. G. Eisenhut dated May 14, 1981'.
t 13.
" Reactor Vessel B rittle' Fracture," Letter Report, J. J. Mattimoe l
to H. R. Denton dated. May.12,1981.
14.
" Reactor Vessel Pressurized Thermal Sheck," Letter Report, l
K. P. Baskin to D. G. Eisenhut (undated).
15.
EPRI Research on the Pronerties of Irradiated Materials Pertinent to the Overcooling Transients, T. V. Marston, Ed. (April 1981).
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2 References (Cont'd) 16.
" Pressurized Themal Shock," letter, D. F. Ross to R. Bernero dated May 19, 1981.
e 17.
" Rancho Seco Data," letter, J. Strosnider to R. Cheverton dated January 30, 1981.
18.
"PV Themal Shock," letter, W. B. Cottrell to G. D. Whitman dated' October 21, 1980.
19.
"IRT Output," letter, M. M. Levine to R. D. Chave.rton dated August 11, 1980.
20.
"IRT Results," letter, M. M. Levine to N. Zuber dated July 2',
1980.
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