ML20032C076
| ML20032C076 | |
| Person / Time | |
|---|---|
| Site: | Midland |
| Issue date: | 09/18/1981 |
| From: | Taylor J BABCOCK & WILCOX CO. |
| To: | Berggren J Office of Nuclear Reactor Regulation |
| References | |
| NUDOCS 8111060674 | |
| Download: ML20032C076 (75) | |
Text
Babcock & Wilcox Nuclear Power Generation Division a McDermott company 3315 Old Forest Road P.O. Box 1260 Lynchburg, Virginia 24505 (804) 384-5111 September 18, 1981 o
b Ve. John R. Berggren O
1 Standardization and Special ANk.
Projects Branch b'y N 0
~
8 U. S. Nuclear Regulatory Commission g) 7887 7920 Norfolk Avenue A
Bethesda, MD 20814
-/
S
Dear Mr. Berggren:
t Enclosed is a copy (34 additional copies have been sent to you under separate cover) of B&W's responses to NRC questions on BAW 10128 Topical Report describing our TRAP-2 code. This is a partial submittal.
Two more submittals will be made to complete our responses to these ques tions. The next set of responses will be issued in mid-November with completion of all responses by the end of this year.
These questions were transmitted January 26, 1979 in a letter from S. A. Varga to J. H. Taylor.
This topical report is under review as part of Dr.141 ton Jensens' review of Consumer Power Co.'s Midland plant FSAR Lhapter 15. Please forward the appropriate number of copies of these responses to Dr. Jensen.
Very truly yces, THE BC0CK & WILC0X COMPANY l
J. H. Taylo, Manager l
Licensing
Enclosure:
As s ta ted.
JHT:efc l
cc: w/o attachment Mr. Darl Hood (NRC) l Dr. Walton L. Jensen (NRC)
R. B. Borsum l
l 01 e
sd ju'?e=*egm PDR
4 QUESTIO.iS AND RESPONSES PART I BAW-10128, TRAP-2
" FORTRAN PROGRN4 FOR DIGITAL SI!!ULATION OF THE TRANSIENT BEHAVIOR OF THE ONCE THROUGH STEAM GEtiERATOR AND ASSOCIATED REACTOR COOLANT SYSTEM" O
SEPTEMBER 15, 1981 i
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l Part I provides response to NRC Questions l
l 1, 2, 3, 4, 5, 7, 8, 13, 14, 16, 17, 18, l
19, 20, 21, 22, and 25 submitted 1/26/79.
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- 1) BAW-10128 states on Page 111 that TRAP 2 is used for analysis of main steam line breaks and main feedwater line breaks. Provide a discussion of any additional proposed usage of the TRAP 2 code for transient or accident analysis.
RESPONSE
TRAP 2 primary usage is to provide mass-energy release data during a steam line or feedwater line break in a conservative manner. In addition, TRAP 2 will be used to analyze the following transients:
- Steam Generator Tube Rupture
- Loss of Feedwater
- Main Feedwater Overfeed '(in conjunction with turbine trip / reactor trip).
- Steam Regulator Malfunction
- Non-symmetric Secondary and/or Primary System Transients
- Accidents Leading to Two-Phase Conditions OV
3(Q
- 2) Provide the results of noding and time step studies which indicate that the solutions obtained by the TRAP 2 code are convergent.
RESPONSE
In order to provide these studies, it was necessary to choose a base case to use as a reference for evaluating the impact of the varied parameter. The case chosen for these studies is a double ended rupture of a main steam line for a typical 205 FA B&W NSS. The specific initial conditions and other significant assumptions are listed in Table 2-1.
Figure 2-1 is a noding diagran of the base case TRAP 2 model used and Table 2-2 is a description of the nodes and paths.
It should be noted that for the purposes of these studies the particular selection of parameters is not critical. The point of the analyses is to show the impact of varying particular parameters of interest. The results of the base case run are presented in Table 2-3 (sequence of events) and in figures 2.2 (which shows the dynamic response of variousimportantsystemparameters).
The time step used for the base case analysis was.0008 sec. This is also the p'
time step used for almost all analyses conducted with the 10 node SG model V
described above.
In order to show that this time step selection leads to convergent solutions, two 35 sec. runs were made with time steps of.001 sec.
and.0004 sec. The resulting sequence of events was found to be identical to the base case; and, the resulting system response for these two runs was found to be the same as for the base case run (i.e., variation in important system parameters shown in Figure 2-2 was less than 1%). Since the double ended rupture of a main steam line is the accident with the greatest rate of change of important system variables and the results are convergent using a
.0008 sec. time step, it is concluded that this selection of time step will yield convergent solutions for other less severe transients mentioned in response to Question 1.
In order to determine the sensitivity of the results to noding variations, two different noding arrangements were run. One of these cases overlaps with another question, and will only be discussed briefly here.
A case was run which has substantial noding differences. For this case, the nodirg diagram used is shown in Figure 2-3, and Table 2-4 gives a description of t'e nodes and paths. This model is a considerably simplified representation n
of the system. The base case assumptions were used and a 65 second run made.
The resulting sequence of events and system dynamic response are given in Table 2-5 and Figure 2-4. Comparing these to those of the base case reveals that the response is very close to that predicted by the more detailed base case model. Due to the reduced number of steam generator nodes, one would expect some difference in results. The point of the comparison is to show that even substantial noding changes yield quite consistent results. Again, 3
it should be noted that the case involved here is a very rapid transient. For less severe transients, the impact of noding variation should be less.
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Question 7 requests a sensitivity study in which the core is modeled as a separate node. That study was performed and the noding diagram and results presented under Question 7.
Referring to those results, it can be seen that this minor noding modification yielded results almost identical to the base case results (less than 1% variation in important system variables).
The base case already employs nearly the full noding capacity of the TRAP 2 code. Accordingly, the base case represents the most detailed noding possible.
Based on the time step and noding sensitivity studies described above, it is concluded that the solutions obtained by TRAP 2 analyses are convergent.
O O
TABLE 2-1 BASE CASE ANALAYSIS ASSUMPTIONS Parameter Power Level 102%
T F
601
- ave, RCS Operating Pressure (at Pressurizer tep),
psig 2195 Pressurizer Level (indicated), in.
185 RPS Trip Signals High Flux, % FP 110 Low Pressure (core outlet), psig 1950 ESFAS Trip Setpoints Low RC Press., psig 1585 Low SG Press. psig 585 ESFAS Trip Delay, sec.
2.5 MSIV Closure Time, sec.
5
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MFWIV Closure Time (linear ramped area, sec.)
15 Auxiliary Feedwater Design Capacity Turbine, gpm 1620 Motor, gpm 810 (one per generator)
Temperature, F 40 Initiation Time After ESFAS, sec.
With Offsite Power 15.0 Main Feedwater Temperature, F 465 HPI System Design Capacity per Pump, gpm 700 I
Temperature, UF 40 1
Boron Concentration, ppm 2270 Initiation Time After ESFAS, sec.
With Offsite Power 20 OTSG Outlet Pressure, psig 1050 0
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TABLE 2-2 TRAP 2 PATH DESCRIPTION Path Number Description 1
Core path 2
Core bypass 3, 17 Paths from the reactor to the hotlegs 4, 18 Paths from the hotlegs to the steam generators 5-13, 19-27
' Steam generator primary paths 14, 28 Paths from the steam generators to the cold legs, including the RCS pumps 15, 29 Paths from the cold legs to the RC downcomer 16 Path from the downcomer to the RC lower plemum 30 Pressurizer surge lines 51-54 Pressurizer flow paths 31-39, 41-49 Steam gener,ator secondary flow paths 40, 50 Paths from the SG heat transfer regions to the rises 55.58 Paths. from the SG risers to main steam lines (MSL)
- 1-#4 60, 61 MSL #2 & #3
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59, 62-64 MSIV #1-#4 65, 66 Paths from the junction of MSL #1, 3 and 2, 4 to the turbine stop valves (TSV's) 67-70 TSV #1-#4'
'I 71 Cross connection 72, 73 MFW pumps 74 Path from the MFW pumps to MFW HTR banks #6 & #7 75 Path from MFW banks #6 & #7 to the MFW branch 76, 77 Path from the MFW branch to the MSLV's 78, 79 MFIV "A" & "B" 80, 81 Paths from the MFIV's to the SG downcomer 82, 83 Paths from the SG downcomer to the SG heat transfer region 84, 85 AFW flow paths 86 HPI 87 LPI 88, 89 SLB paths
TABLE 2-2 (CONTINUED)
TRAP 2 NODE DESCRIPTION Node Number Description 1
RV lower plenum up thru the bottom of the core support sheets.
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2 Reactor core, core bypass, upper plenum, and outlet nozzles.
3 Hot leg A, includes SG A upper plenum and upper ~.ube support sheets.
4-13 SG A tube region between tube support sheets.
14 A cold legs, including lower tubesheet, SG A low plenum, and RCS pumps Al & A2.
15 RV annulus, includes inlet nozzles.
16 Hot leg B, includes SG B upper plenum and tube support sheets (upper).
17-26 SG B tube regions (primary) between the rube support sheets.
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28-30 Pressurizer 31-40 SG A secondary side heat transfer region.
41-50 SG B secondary side heat transfer region.
51 SG A steam riser.
52 SG B steam riser.
53 Main Steam Line #1 to the MSIV (SG A).
54, 57 Main Steam Line #2 to the MSIV (SG A) - Split to model a DESLB.
55, 58 Main Steam Line #3 to the MSIV (SG B) - Split to model
- DESLB, 56 Main Steam Line #4 to the MSIV (SG B).
59 MSL #1 & 3 from the MSIV's to the cross pairing (denoted here as "X").
60 MSL #2 & 4 from the MSIV's to the cross pairing (denoted as "Y").
61 From cross pairing X to the turbine stop valves on X includes 1/2 cross connection.
62 From cross pairing Y to the turbine stop valves on Y includes 1/2 cross connection.
Q 63 Turbine
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s TABLE 2-2 (CONTINUED)
TRAP 2 NODE DESCRIPTION Node Number Description 64 From the MFW pumps to HTR bank #6.
65 FW HTR banks.# 6 & 7, and associated piping.
66 From FW HTR bank #7 to the FW fork (does not include any fork piping).
67 From the FWL fork to MFIV "B".
68 From the FWL fork to MFIV "A".
69 From MFIV B to SG B inlet.
70 From MFIV A to SG A inlet.
71 SG B downcomer.
72 SG A downcomer.
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TABLE 2-3 SEQUENCE OF EVENTS Event Time, s Double Ended Rupture of 28" Steam Line Between SG and MSIV 5.0 Closure of Turbine Stop Valves 5.15 High Flux Trip Setpoint Reached 8.8 Rods Begin to Drop 9.2 ESFAS Signal on Low Primary System Pressure 15.2 Main Feedwater Isolation Valves Begin to Close 17.7 MSIV's Close 22.7 Auxiliary Feedwater Begins to Flow to Intact SG 30.2 Main Feedwater Isolation Valves Close 32.7 HPI Flow Begins 35.2 O
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Core 2
Core bypass 3, 11, 12 Hot leg 5, 6, 7, 13, 14, 15 Steam generator primary tubes 8, 16 RC pumps 9, 17 Cold leg piping 10 Downcomer, reactor vessel 18 Pressurizer surge line 19 Steam piping crossover 20, 27 Main feedwater pumps 21, 28 Feedwater piping 22, 23, 24, 29, 30, 31 Secondary heat transfer region steam generator p
36, 37 Steam rise, SG V
25, 26, 32, 33 Main steam piping 34, 35 HPI pumps 40, 41 FW Piping (MSIV's) 39 LPI 42, 43, 44, 45 Pressurizer O
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Node Number Description 1
RV lower plenum to bottom of support plates.
2 Core, core bypass, upper plenum, outlet nozzles.
3 Hot leg 4,5,6,7 Steam Generator primary tube region 8
Hot leg 9, 10, 11, 12 Steam Generator primary tube region.
13 Cold leg 14 Cold leg 15 Reactor vessel downcomer 16, 34, 35 Pressurizer 17, 18, 19, 32 Steam Generator secondary tube region.
20, 21, 22, 33 Steam Generator secondary tube region.
23 Feedwater piping
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24 Feedwater piping 25 Steam riser 26 Steam riser 27 Steam piping 28 Steam piping 29 Turbine 30 Atmosphere 31 Feedwater piping i
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TABLE 2-5 SEQUENCE OF EVENTS Event
' Time, s Double Ended Rupture of 28" Steam.
Line Between SG and MSIV 5.0 Closure of Turbine Stop Valves 5.15 High Flux Trip Setpoint Reached 8.8 Rods Begin to Drop 9.2 ESFAS Signal on Low Primary System Pressure 15.1
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Main Feedwater Isolation Valves Begin to Close 17.6 MSIV Close 22.6 b
Auxiliary Feedwater Begins to Flow to Intact SG 30.1 Main Feedwater Isolation Valves Close 32.6 HPI Flow Begins 35.1 l
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Question 3: We understand that the RADAR code will be utilized to predict the minimum DNBR for the core hot channel. Discuss how the 4
TRAP-2 output will be utilized by the RADAR code and provide any conservative assumptions that will be applied to the TRAP output in generating input for RADAR. Discuss how hot channel flow reduction factors will be applied.
Response
The RADAR computer code (BAW-10069A) is one pirticular code which may be used in certain applications in conjunction with TRAP-2 output to describe core conditions during a system transient. An evaluation of the capability of TRAP-2 to simulate the transient system responses of core flow, power, pressure, coolant conditions, etc., for a given set of input parameters and boundary conditions should not be tied to one particular method of core condition evaluation. Flexibility must be maintained for core evaluations based upon the use of j
several existing codes (RADAR, COBRA 3C, COBRA 4, THETAIB, LYNXT, etc) or future company and industry developed models.
1 The TRAP-2 output of the transient system parameters of core flow, power, pressure, and coolant temperature are normally used directly in conjunction with conservative design power distributions for core conditions evaluations. This type of analysis assumes that the core condition itself has negligible effect on the input parameters to TRAP-2 and that no unusual q
local conditions exist in the core which adversely affect the Q
system response. These assumptions result in conservatively accurate evaluations of most types of transients analyzed; however, it is recognized that for certain types of severe transient conditions (ie. stuck rod, ejected rod, rapid primary depressurizations, etc.) it may be desirable to account for local core nuclear and themal hydraulic con-ditions to provide feedback into the system response and for use in the core modell These severe transients, with unusual local perturbations 5 the core, may require an 7
l iterative analysis technique between a spatial kinetics core model and TRAP-2 to obtain more realistic core power distributions and system feedback parameters. Detailed descriptions of analysis techniques are provided to the NRC staff when any-thing other than the direct use of the systems code output in combination with design power distributions have been employed for core evaluations.
Thermal hydraulic licensing analyses and core condition evaluations conservatively account for core power distribution uncertainties, system pressure uncertainties, coolant temperature uncertainties, coolant flow maldistributions, bundle spacing uncertainties, subchannel flow reduction factor, fuel densification, hot channel factor on average pin power (accounts for rod and fuel lot enrichment variations, fuel stack length variations, etc.),
and hot channel factor on local heat flux (accounts for variations i
in pellet density, cross-sectional area, weight per length, local enrichment, clad diameter, etc). All themal-hydraulic uncertainties and conservatisms have historically been accounted g
for by direct inclusion in analyses. For certain new plants, a statistical core analysis technique will be used and some of the i
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\\J minimum design DNBR. A topical report describing the statistical core analysis technique (BAW-10145P) was filed with the NRC staff in October 1980. Regardless of the method employed to account for the thennal-hydraulic uncertainties, the conservatisms are always applied to the system transient response for core evaluations, unless the analysis description indicates otherwise.
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4 Question 4: Our previous evaluation of the RADAR' code dated September 19, 1974 was limited to the use of the code for loss of -
flow transients. BAW-10064 describes RADAR and states that the code is designed for the analysis of slow reactor tran--
sients. Justify the applicability of RADAR for the analysis of core response during steam line breaks and feedwater line breaks.
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Response
This question is not directly applicable to the TRAP-2 code.
As stated in the response to question 3, several codes, such as RADAR, LYNXT, 'and COBRA 3C could be used in conjunction with TRAP-2 to predict DNBRs during a transient. The major lJmiting assumption in RADAR is that the mass velocity along L
a channel remains constant. Rapid changes in the void fraction during a transient can cause inaccuracy in RADAR predicted 4
DNBRs. Before applying RADAR along with TRAP-2 in analyzing a steam line or feedwater line break, the margin of con-servatism will be verified with a transient code which con-siders variable mass velocity along a channel.
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- 5) Provide comparisons between the predictions of the RADAR code with those of the THETA-1B code for the hot channel transient for a double ended steam line break and a double ended feedwater line break.
RESPONSE
The interface between TRAP 2 and RADAR is discussed in response to questions 3 and 4.
THETA-1B code is not used with any of the TRAP 2 results for hot channel analysis. Comparisons between RADAR and THETA-1B are outside the scope of the TRAP 2 Topical Report.
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- 7) The TRAP noding diagram shown in Figure 1-1 of BAW-10128 indicates that the core is not modeled as a separate fluid volume. Justify by neans of a sensitivity study that conservative results are obtained for the hot channel analysis using this approach.
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RESPONSE
a A TRAP 2 model with a separate core node was constructed to demonstrate the sensitivity to core modeling. The noding diagram is shown on Figure 7-l.
The only difference between this model and the base case model (see question
- 2) is the addition of node-74 (the core) and a new path (path 90). The volume of the core, which was previously included in node 1, was subtracted from the node 1 volume so the new configuration results in the same total system volume. A case was analyzed out to 65 seconds using this model and the results are presented on Table 7-1 and in Figure 7-1.
Comparing these to the base case results, it can be seen that the system response is almost -
identical.
In particular, the parameters of importance for hot channel analysis (system pressure, flux and core inlet temperature) are essentially j
unchanged (i.e., variation is less than 1%). From this, it is cons uded
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that results of hot channel analyses using the base case core model are the same as for the more detailed model.
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Time (s)
Base Case Sensitivity Study 5.
567.9 567.9 10.
559.7 559.7 15.
539.3 539.4 20.
523.8 524.3 25.
505.4 505.8 30.
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485.0 484.9 45.
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- 8) Figure 1-1 indicatas that the upper plenum and upper head region are described with a single node, thus, the region of relatively stagnant flow above the upper plenum cover plate may not be adequately described.
During a rapid plant cooldown, the reactor coolant in this upper volume may flash and act to pressurize the primary system. Provide sensitivity studies which provide the effect of modeling this upper volume on the core DNBR following a main steam line break.
RESPONSE
The reason the upper head has been left out in past analyses is because it was judged to be a conservative ar
- h te steam line break analyses.
It was anticipated thct this region k a be relatively stagnant and therefore not available for energy transfer to the secondary side. Also, if the mass had been included and allowed to circulate with the rest of the primary system fluid the result would have been a less severe cooldown transient.
In addition, if the volume were modeled as a separate node, it was anticipated that it would flash and thereby reduce the depressurization.
In order to verify this, and respond to the specific request of this question, a TRAP 2 model O-was set up in which the upper head was included as a separate node (see noding diagram figure 8-1).
This model is identical to the base case mo41 (see question 2) except for the addition of the upper head volume node 74 ar.; i.he path connecting it to the upper plenum.
This case was ~run out to 65 seconds and the results presented in figure 8-2.
The most significant difference in the results is in the surge line flow rates and the system pressure. It is noted that these track the base case response very well until 12.5 seconds at which time the depressurization is arrested by upper head flashing. As expected, the depressurization results are less severe due to this flashing. Since the flux and RC temperature responses are almost the same and the pressure response is favorable from a DNB point of view, the DNBR r..sults for this case are less severe than for the base case. From this, it is concluded that analyses in which the upper head is not modeled give conservative DNBR results.
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Figure 8-2.
SYSTEM RESPONSE e
~
~
5 i
"o 7
1,4 620 1
M n
1.2 600 d'
4 3
1.0 580 I
w 0.8 E
560 E
E.
?
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0.6 g 540 1
e 3
0.4 520 I
0.2 500 i
1 0,0 480 0
10 20 30 40 50 60 O
0 10~
20 30 40 50 60 i
L d
Time, sec Time, sec j
i 1
1 4,
4000 6000 j
3500 UPPER HEAD MODEL 5000 UPPER HEAD MODE;.
,j
--- BASE CASE E
BASE CASE 9,
a
=
I3000 3
4000 2
e
=>
I
~=
0 2500 2
3000 3
I j
w af
=
as y 2000 5 2000 -.h l
3 m
m I
j 5 1500
\\
1000 1000
\\
0 e
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0 10 20 30.
40 50 60 0
10 20 30 40 50 60
~
Time, see Time, sec W
I
)
s/
-13) Page 1-40 indicates that a slip flow model is utilized in the steam i
generator ^ for steady-state operation and a bubble rise model is utilized for transient analysis, i
a)' Provide the slip flow equations utilized for steady-state operation, b) Describe how switching is accomplished between the two models at the start of
+ransient.
c) Provide the results of an analysis demonstrating that the transient modeling is ca'pable of maintaining the steady-state level in the absence of a break or utilizing a break of negligible size.
RESPONSE
a) A slip-flow mode'l is not utilized in the steam generator for steady-state i
operation. Instead, the steady-state equations in the steam generator utilize a zero bubble rise velocity. This is a valid assumption which is verified by comparing steam generator inventory with a steady state stean l
generator code. It allows the steam generator inventory to be conservatively
.large for steam line breaks or conservatively small for feedwater line breaks.
(See question 14) b) The switch between the steady-state and trarsient mode is done abruptly within one time step. This change is made automatically by the code based on input time for the rupture to take place.
c) An analysis was performed using the base case but reducing the break size.
to an insignificant size. The resulting system response plotted on the same scale as the base case results, are presented in Figure 13-1. The 4
i break occurs at 5. seconds. As can be seen, this transient model maintains the steady-state level for all of these variables.
i l
i, l
1 e
n,.m-,~
-,,.n.,
...-,i,,.,-.-.-.-,,,,nnv.,w,,...--n.,nr.-,-,.-,,
..nm..,,,
.,..~.,n n -, - ~ ~,
-,-m.-,,-
F i gu r e 13-1.
SYSTEM RESPONSE n.
1.4 o
w 600 n
- 1. 2 O
Q.
IO 580 w
=
J 0.8 a 560
~
=
D n
w f
0.6 E.540 g
2
~
0.4 g 520 am 0.2 500 i.
0.0 i
480 i
0 2-4 6
8 10 12 14 0
2 4
6 8
10-12 Time, see Time, sec j
i t
i t
a 4000 3500 i
=
D.
e i
o-3000 1
u l
5 l
2500 3
l 1
m 2000
=
O l
1500 d
l 1000 I
500 t
i i
0 2
4 6
8 10 12 14 Time, sec r
t t
nQ
- 14) Equation 1-88 is used to calculate the flow weighted enthalpy for each secondary steam generator node. The actual enthalpy for the seccndary nodes will be lower than those predicted by equation 1-88 since the steam will travel faster than the liquid. Discuss how secondary control volume mass and energies are calculated in the TRAP 2 code to account for slip between the steam and water phases. Discuss how these values of mass and energy are made to correspond to those of the design for the steam generator.
RESPONSE
The computer code PIE 0TSG2 (NPGD-TM-350 January,1980) is used to predict the mass and energy for each steam generator node. The PIE 0TSG2 code breaks the steam generator down into 100 nodes. When the TRAP 2 deck is set up, each TRAP 2 steam generator nade is made to correspond to ten PIE 0TSG2 nodes. The steady-state secondary side specific enthalpies (h in equation 1-88) are s,n generated by the TRAP 2 code on the basis of input feedwater enthalpy and the primary to secondary heat transfer rate (Qn ia equation 1-88). The Qn, which are input in card series 7111, are chosen to yield specific enthalpies f) equal to those predicted by PIE 0TTG2. Finally, the proper ::. ass inventory for V
each TRAP 2 node is created by adfJsting the TRAP 2 node area (the heights of the no.es are always made to tota' to the actual steam generator height and the " ievation of the particular nod::, correspond to the. PIE 0TSG2 data). As di. cussed in response to question 13, the steady-state calculations assume a bubble rise velocity of zero ~(i.e., a slip model is not used in steady-state).
However, during the transient, the bubble rise velocity model discussed in question 13 is initiated.
l
.()
- 16) Section 1.9 describes two options that can be used in calculating the bubble rise velocity. Discuss which model is used for analysis of feedwater line breaks and steam line breaks. Provide the results of a sensitivity study of the effect of bubble rise velocity in the steam generator'on the hot channel analysis for a. main steam line break.
d r
4 i
i
RESPONSE
The bubble rise velocity is input as a specific value depending on application.
Minimum DNBR results are obtained using a zero bubble rise velocity. As the bubble rise velocity increases, the two phase mixture height in the OTSGs decreases resulting in slower primary system depressurization. Hence, less severe DNBR results.
The base case (see question 2) uses a steam generator bubble rise velocity of zero.
Sensitivity studies using bubble rise velocities of 2.0 ft/sec. and 9.0 ft/sec.
were performed. Table 16-1 and Figure 16-1 present the results for the 2.0 ft/sec.
case and Table 16-1 and Figure 16-2 present the results for the 9.0 ft/sec. case. The O
cooldown and depressurization rates of the primary system are less severe as the bubble rise velocity is increased. The flux response is about the same.
In addition, the return to power experienced in the higher bubble rise velocity cases is-less than for a zero bubble rise velocity.. All of this leads to the conclusion that the hot channel DNBR results are less severe for a case with a bubble i
rise greater than zero and that it is conservative for hot channel analysis purposes to use a zero bubble rise velocity. With respect to mass / energy I
releases to containment, using the higher bubble rise velocity eliminates liquid entrainment and may result in a somewhat higher prediction of total energy releases than the zero bubble rise velocity situation. However, when containment analyses are performed any liquid entrainment predicted by TRAP is removed by raising the enthalpy of the released mass to that of saturated steam. This is a conservative approach to containment analysis.
1 I
i r
r
_,,_,,.,,,,___,_.-._m._-._,,__,--,._,-,,,____._,
.~,__mm,m
,m.,_m.-,,,,.._,+.,m
- m m -,__,
y
,.._m.,_~.--_,__,,
(~N TABLE 16-1 V
SEQUENCE OF EVENTS Bubble Rise Velocity, (ft/sec) 0.
2.
9.
Event Time, s Time, s Time, s Double Ended Rupture of 28" Steam Line 5.0 5.0 5.0 Between SG and M'ilV Closure of Turbine Stop Valves 5.15 5.15 5.15 High Flux Trip Setpoint Reached 8.8 8.8 8.8 Rods Begin to Drop 9.2 9.2 9.2 ESFAS Signal on Low Primary or Low SG Pressure 15.2 12.5 9.3 Main Feedwater Isolation Valves Begin to Close 17.7 15.0 11.8 MSIV Closes 22.7 20.0 16.8 o(-
Auxiliary Feedwater Begins to Flow to Intact SG 30.2 27.5 24.3 Main Feedwater Isolation Valves Close 32.7 30.0 26.8 HPI Flow Begins 35.2 32.5 29.3 O
L
A Figure 16-1.
BUBBLE RISE VEL = 2.0 FT/SEC
- 1. 4 640 "o
~
1.2 620 E
1.0 600 B
m 20.8 k 580 j
e
- 0.6 560 E
B E'
E 0.4 540 B
0.2
[
520 f
500 i
0.0 0
5 10 15 20 25 30 35 0
5 10 15 20 25 30 35 Time, sec Time, sec O
l 4000 35 "o
3500
- 30 o
2_
3000
- - 25 E
5 3
g 2500 g 20 E
5 5 2000 15 B
b1500 5
10 u.
1000 5
i 500 0
e 0
5 10 15 20 25 30 35 0
5 10 15 20 25 30 35 Time, see Time, sec
--.,-,,-,.,.---.--..n.--,.,,,.,e.,..-...---,
n.-,.n_.,
.-.,,,gw,,..,m.__m,,
-.,,,.,-,-~.,-_-,n.m.,m-pe.,
.-.w.-,...,-,e.,---.--v,.
s 4
Figure 16 2.
BUBBLE RISE VELOCITY = 9.0 FT/SEC Os i
- 1. 4 620 mcs 1.2 600 0u l.0
, 580 l
E
=
i y
4 n
0.8
% 560 I
a as a
g 0.6
$540 m.
as E
0.4
- 520
=
0.2 500 I
?
0.0
' 480 f
0 10 20 30 40 50 60 70 0
10 20 30 40 50 60 Time, see Time, see
,I O
I 4
t 4;
4000 35 R
5 m
3500 S 30
)
i i
a c2 I
3000
- 25 g
E N
i
=
m 2500 g 20 l
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! 15 7
a o
E 3
1
=
1500 10 E
C-1000 5
5 500 O
i i
i 0
10 20 30 40 50 60 0
10 20 30 40 50 60 Time, see Time. sec 1
4
~..,__-r_
___,_q,y_.._______,._,.,,_,y,_
g,.
,__,,m,
___,,.7,
_,m._,mr,_._m
-_,m.,,
F 10'
- 17) Provide a comparison to experimental data demonstrating the accuracy of the bubble rise model used in the TRAP 2 code for analysis of steam generator blowdown.
RESPONSE
Analyses presented in response to question 16 show that use of a zero bubble rise velocity results in the most severe transient results. A zero bubble rise velocity is used in TRAP analyses. Since the actual bubble rise velocity must be at least zero and a zero velocity results in the most severe results, use of a zero bubble rise velocity is conservative. Since the results are conservative and, the goal of all accident analyses is to provide conservative system response information, no comparisons with experimental bubble rise velocity data have been made.
(m OO
,O)
V
- 18) Page 1-52 states that tra boron concentration in the core that is used to compute core reactivity is based on the average concentrations in the upstream and downstream control volumes. Since no core node is provided, the boron concentration would be based on the average of the upper and lower plenum volumes. Justify that this approach is adequate for determining the boron concentration in the core for steam and feedwater line breaks.
RESPONSE
As shown in Figure 1-1 boron is injected upstream of the core. As boron flows into node 15 (of Figure 1-1) it is diluted and carried on into node 1 then to node 2.
The boron concentration grows progressively smaller and smaller. The average concentration between nodes 2 and 1 will be less than the concentration of node 1.
Since the core volume is included in node 1, the calculated boron concentration (average between nodes 1 and 2) will be less than the actual concentration in the core. Accordingly, the method used will result in a smaller negative reactivity insertion than would actually obtain. This is conservative for all analyses for which TRAP
((';/
is used.
O
{o)
- 19) Provide a comparison of the TRAP 2 code predictions with applicable reactor transient data.
RESPONSE
- 1) An asymmetric auxiliary feedwater overcooling event occurred at Crystal River 3 during startup testing. Table 19-1 provides a sequence of events based on data available from the site. Using this information in conjunction with other available plant data, a TRAP 2 analysis was performed to simulate the transient. The TRAP 2 model used is shown on Figure 19-1 and the description of the nodes and paths is on Table 19-2. The results are shown on Figures 19-2 to 19-4.
In view of the complex asymmetric nature of the feedwater transient, the TRAP 2 analysis results are considered to compare well with the plant data. The RCS temperature response compares particularly well considering the length of transient time analyzed and the swing in temperature which was observed.
- 2) An analysis was performed to simulate a reactor trip / turbine trip which occurred at Oconee Unit 1 on June 11, 1979. A loss of both feedwater pumps caused the reactor and turbine to trip. Auxiliary feedwater was provided within six seconds after the feed-pump trip via three auxiliary feedwater p) u pumps. Plant data available from this transient was used as input to a TRAP 2 analysis using a TRAP model similar to the MINI TRAP model described in response to Question 2.
The TRAP response for the system pressure and steam generator pressure were then compared to the plant data (Figure 19-5 and19-6). The reason for the.large difference in primary pressure response shown after 100 seconds is that in the TRAP analysis HPI pumps were not initiated while they were during the actual plant transient. Although not much data is available for comparisen, the pressure response data which is presented compares quite well (or the differences can be accounted for as explained on Figures 19-5 and 19-6).
- 3) A Loss of Feedwater (LOFW) event which occurred at Oconee Unit 1 (Trip No. 21) on July 14, 1973 was analyzed using both the equilibrium and non-equilibrium pressurizer models. The plant was operating at 76% power when a LOFW occurred.
This led to a turbine trip followed by a reactor trip on high RC pressure.
See Table 19-3 for the sequence of events. The resulting system response data is shown in Figures 19-7 through 19-10. The curves labled TRAP 2 Version 6.1 employed the non-equilibrium pressurizer model while the curves labeled TRAP 2 Version 4.1 used the equilibrium pressurizer model. See the response to Question 22 for a discussion of these pressurizer models. As can be seen from the figures, the non-equilibrium pressurizer model gives extremely good results.
The primary system pressure and pressurizer volume (hence the surge rate) compare quite well with the plant data. The more slugglish pressure response predicted by the equilibrium pressurizer model is to be_ expected. The steam generator response is not sensitive to the choice of pressurizer model.
O
Based on the three transients discussed above, TRAP 2 can be used to yield O
realistic and, with suitable assumptions, conservative results for both overcooling and overheating type transients.
' O O
tj TABLE 19-1 TRANSIENT #1 SITE SEQUENCE OF EVENTS Time, Minutes Event 0
Lost power, all feedwater to both generators stopped.
1 Steam-driven emergency feedwater'(EFW) pump up to speed and feeding both steam generators.
2 Started electric-driven emergency feedwater pump.
"A" steam generator is being perferentially fed, but both are getting water.
3 Stopped steam-driven EFW pump.
"A" steam generator is being filled (startup level indication),
"B" steam generator startup and operating range level indicators are both apparently at the bottom of their range.
(A.)
the bottom of its range (as verified by subsequent 4
"B" loop startup feeduater flow indication is at calibration checks).
9 1/2 Restarted steam-driven EFW pump, started feeding "B" steam generator again.
10 Opened a parallel valve (EFV-162) in the feedwater train to "A" steam generator by mistake.
12 1/2 Shut EFV-162 and opened a paralled valve (EFV-161) in the feedwater tain to "B" steam generator.
14 "B" steam generator filling, on its, way to recovery.
TABLE 19-2 gm
's,)
TRAP 2 N0DE DESCRIPTION Node Number Description 1, 33 Reactor Vessel, Lower Plenum 2, 34 Reactor Vessel, Core 3, 35 Reactor Yessel, Upper Plenum
.a' 4,, 10 Hot leg Piping
,507, 11-13 Primary, Steam Generator 8, 14 Cold Leg Piping 9, 32 Reactor Vessel Downcomer 15 Pressurizer 16, 24 Steam Generator Downcomer 17, 25 Steam Generator Lower Plenum t
18-20, 26-28 Secondary, Steam Generator
- 21. Ek Steam Risers 22, 30 Main Steam Piping
()
23 Turbine N_/
31 Containment t
TRAP 2 PATH DESCRIPTION Path Number Description 1, 2 Core 45, 46 Core Bypass 3, 4, 5, 11, 12, 44 Hot Leg Piping 6, 7, 13, 14 Primary, Steam Generator 8, 15 RC Pumps 9, 16 Cold Leg Piping 10, 43 Downcomer, Reactor Vessel 17 Pressurizer Surge Line 18, 19, 26, 27 Steam Generator Downcomer
-/
20, 21, 28, 29 Secondary, Steam Generator
/
22, 30 Aspirator 23, 31 Steam Riser
(~ ).
24, 32 Steam Piping m-4.
J l
1 5
TRAP 2 PATH DESCRIPTION (CONTINUED)
Path Number Description i
25, 33 Turbt.;e Piping
- 34. 35 Break (orleak) Path i
36, 37 HPI 38, 39 AFW 40, 41 Main Feed Pumps 42 LPI I
47, 48 Reactor Vessel Plenums l,
4 i
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4 l
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1 i
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(
I.
.- _.. - - - = -. ~.. -_ - - _ - - -
d 1
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TABLE 19-3 i
TRAN51ENT #3 1
SIGNIFICANT EVLNTS i
Time 3
Description Plant Data TRAP _
0 (1)
"A" main feedwater pump trip on 4 sec 4 sec either low bearing oil pressure or loss of hydraulic oil supply-I-
this initiated ICS run back at l
50%/ min.
t I
r t
(2)
"B" main feedwater pump trip 13 see 13 sec on high discharge pressure-losing & c second feedwater pump caused the turbine to trip i,
22 sec 22 see (3) Reactor trip on high RCS pressure i
i 1
NOTE: " Rupture" occurs in the TRAP 2 run at 1 second. This begins the transient calculation.
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Time, f,linutes i
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.j
~l 2200 i
2l00
%:i i
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(
i O.N.S.
2 RCS PRESSURE INCREASING -
i E
DUE TO HPl BEING
- 2000 INITIATED
,1 C
c a-h E
I 1900 TRAP
/
I I
NOTE: HPl IS NOT INITI ATED IN THE TRAP ANALYSIS I
1800 t
t i
e 0
50 100 150 200 250 300 t
Time, see l
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1200 1
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1150 i
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1050
- i E.
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5 1000 ;-
y OCONEE DATA g
a.
,f d
950 l
NOTE: THE REASON THAT THE OCCNEE l t
g i
DATA INDICATES N.S. PRESSURE
[
900 -
DECREASING RAPIOLY ~25 50 SEC IS DUE TO THE BLOW 00 Vill ON THE
+
VALVES (TRAP 00ES NOT f.00EL THIS).
j 850 -
i i
[
200 i
e i
r,
i 50 100 150 200 250 300 t :
i Time, sec 6t f
~
l s
--g
---,--r,
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Figure 19-1 PZ LIQUID VOL. VS. Tl!AE l
1 a
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i figure 19-9 SG PRESSURE VS TIME i
i e
- l 1
l 1100 l
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l 1050 3
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=
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- 20) Following a steam line break, feedwater flow will increase for the affected steam generator. This will occur'. shen the pressure in the steam generator with the break is reduced so that the total feedwater flow is diverted.
In addition, the reduced pressure will cause flow through the feedwater pumps to approach runout conditions and flashing of the heated feedwater will cause the feedwater inventory in the lines to flow into the affected steam generator. Discuss how these processes are analyzed in the TRAP 2 code. Discuss the affect of increased feed-water flow on the core temperature transient.
RESPON5E:
A)
Feedwater line inventory:
Two methods have been used to account for flashing of feedwater line inventory -
- 1) In most recent analyses, the feedwater line inventory has bee, modeled by placing a node of the appropriate size between the MFWIV and the steam generator. The water in this node flashes when the pressure O
drops below saturation pressure and the steam is forced through the steam generator and out the break. This approach will usually be used in the future.
- 2) In some previous analyses, mair, feedwater has been allowed to flow for a period of time after the main feedwater valves w uld have closed.
The amount of time is determined by dividing the total feedwater line inventory by the feedwater flow rate. This approach assures that'all of the feedwater line inventory will be forcad into the steam generator as is a conservitive approach compa.ed to (1) above.
B)
Main Feedpumps:
The mdn feedwater pump model used is described in Section 1.7.3.
This model allows one to use the pump characteristics of the actual pump in the plant and therefore the approach to runout condition is adequately modeled.
C)
The increased feedwater flow will result in a decrease in the temperature of the primary system coolant and its impact is included in the analysis by the above means.
O
i
- 21) Flow through the feedwater isolation valves will be reduced during the time when the isolation valves are closing. This reduction in flow will be a function of the losses in the partially closed valve relative to the total line losses. Discuss how flow through partially closed isolation valves will be calculated in the TRAP 2 code.
RESP 0tlSE:
The flow path containing the feedwater isolation valves is input with an area versus time table such that, upon receiving the appropriate safety signal, the area closure represents valve closure. Also, the K factors for that flow path are input as form loss coefficients (the friction factor is set equal to zero).
Therefore, TRAP 2 calculates a new K factor for each time step based on the input area and the flow rate. See equation 1-27.
O o
a
b(%
- 22) Provide the equations and assumptions used to r:odel the prcssurizer in the TRAP 2 code.
Include the treatment of the spray and heaters. Discuss the accuracy of the equilibrium pressurizer model in TRAP 2 for ana'ysis of feedwater line breaks when the system pressure is calculated to increase and cause a surge into the pressurizer. Provide comparisons to any appropriate experimental data.
RESPONSE
Three methods have been used to model the pressurizer:
1)
A single TRAP 2 control volume has been und in some analyses. This node is treated exactly the same way as any other TRAP 2 node. See pages 1 A through 1-6 of the TRAP To~ il for the equations used. Use of this approach results in an equi 'arium model. This model is used only in overcooling transients when the surge flow is out of the pressurher.
2)
In order to more closely approximate' a non-equilirbium pressurizer for O
overheating transients, a three node pressurizer model has been used.
The pressurizer is divided up into three vertical nodes of equal height.
The bottom node is slightly subcooled liquid, the top node is slightly superheated steam and the middle node is about half full of saturated liquid (equilibrium node). This model results in m. ore realistic pressure increases associated with insurge transients.
Ordinarily the pressurizer heaters and sprays are ignored in TRAP 2 analyses.
Some analyses have been performed where the feedwater heater model has been used to model the pressurizer heaters (this model allows the addition of heat to any node via a heat addition vs. time table). Also, pressurizer sprays have been modeled in certain analyses by providing an area vs time path from the co'd leg node to the pressurizer. The flow is adjusted by selecting the proper path area and the timing of the flow is varied by restarting cases as needed and opening / closing the path.
3)
TRAP 2 Version 6 incorporates an optional non-equilibrium pressurizer model.
The equations and assumptions used in this model are described on tne following pages taken from NPGD-TM-414, Rev. 3 (Revision J), May, 1980, the TRAP 2 users menual.
A comparisor, to experimental data and between the three node and non-equilibrium pressurizer model is given in response in questions 19.
(
v I..
1.15 Non-equilibrium Pressurizer Model An optional nonequilibrium preesurizer model is available in TRAP 2.
The model simulates pressurizer performance using two thermodynamic systems, one for stratified steam and one that contains either subcooled itquid or two-phase mixture. The option includes models for simulating pressurizer sprays, heaters, safety valves, and steam-mixture interface heat and mass transfer. Surge line flow is computed using a linear momentum balance, while relief valve perfor-mance is approximated via input mass flow rate versus pressure tables or the l
k HEM isentropic expansion model.
1.15.1. Geometry The pressurizer is modeled by a right circular cylinder of diam-3 eter, D, with chopped hemispherical ends of height, H,, and radius of curva-ture, g. Given the total pressurizer volume, V, the length of the cylindri-T cal section of the pressurizer, H, is computed as follows:
e
~
H = (V - 2 V ) /A, T
V E volume of each hemispherical end vH* [3 D2 + 4 H ),
I 2
=
24 e
e A E area of cylindrical section 2
= 3D /4, C
Thus, the total height of the pressurizer is H
=H
+ 2
- R,.
T l
I
(
Babcock 8. Wilcox 1-54 REV. J (5/30/80)
3 l
Revision 3 (5/30/80)
O
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1.15.2. Initialization The state of the fluid in each of the pressurizer regions is defiaed by the input pressurizer pressure, P. and the mixture level, H,.
The p
fluid in the mixture region is assumed to be at saturated liquid conditions, while the fluid in the steam region is assumed to be at saturated steam condi-tions. The total mass and energy inventories in the mixture region are ini-1 tialized as follows:
h M,= V,lv,
g U
=M
+u m
a f
v E satut.ted liquid specific volume, g
E saturated liquid specific internal engrgy, uf V, E volume of mixture region
~
~
= { H (3 + R 2
h - H,): 0 5 H, 5 H, O
4
=V
+ (H -H)=A:
H < H
(.
m IL - H 3
T e
e a
e c
e w(H - H,)*
T T
3 b
T m* b ~ e' (H ( =V ~ m Similarly, the initial occam region mass and energy invencories____. are given by M = V /v, a s g U =M .u s a g where v E saturated steam specific volume, u E saturated steam specific internal energy, V, E volume of steam region = v,- v,. 1.15.3. Coverning Equations The equations which gevern transient pressurizer performance ( Q. Babcock & Wilcox 1-55 REV. J (5/30/80) ---,-,_--m ,-...,._,,,_y.--.,,,,,,r., 4 .7 .m. -p.- y,,,- ,,,m .,,,.<r +---.,w_.----,,-, m.,
Revision 3 (5/30/80) k h conservation of Mass. dM* = IV, de s dM* = IV de m where IW, I sum of mass flows into steam region. IV E sum of mass flows into mixture region. m l l Conservation of Energy" 1 dU dV = I(Wh)s + I4 - P de a dt, dU dV dt }m + b - dc where I(Wh), I sum of energy advection rates into stean region, j l I(Wh), 3 sum of energ-advection rates into mixture region, 3' I4, E sum of thermal convection rates into steam region. I( E sum of thermal convection rates into mixture region. Models are provided to compute rates of mass and energy advec-tion resulting from insurge, outsurge, bubble rise, steam-mixture interface fj ccr.dtnsation, spray, condensation on spray droplets, and relief valve discharge. I Ther=al convection codels for heaters, steam-mixture interface, and spray drop-1 ts are also provided. 9 1.15.4. Advancement Procedure The pressurizer governing equations are solved using a forward Euler advance =ent procedure. Thus, the equations of e as and energy conserva-tion become
- 1 M
- H" + (Zw)" At and .n ,E U"+1 - U" + (Imi)" + (" - P" AL,. at l I (. Babcock s.Wilcox l-56 REV. J (5/30/80)
. ~ 3evision 3 (5/30/80) O ( where 'g' ", Y -V n M ,A t, At and subscripts n-1, n, and n+1 are sequential time step indices. Given the mass and energy in each of the pressurizer regions at the new time ste'p, n+1, the volumes of the regions are found by solving the folleving equation iteratively: V - [I P',h - [I vP.h)=0 v T e m, s s i where the specific volume, v, for each region is'found by solving the follow-ing equations: h '3 = u + P ) i i 3 - v(p.h '3) = 0. i v The superscript i refers to the outer interation, which is terminated when the I relative error between two successive iterates, P and.P , is less than 10-6 The superscript j refers to the inner iteration, which is satisfied when the 3 3 relative error between two successive ' iterates, v3 and v "I, is less than 10-6, The value of h '$ computed using the converged specific volume is defined as i i h. 1.15.5. Surge Line Model A specialized flow path model similar to that used in existing TRAP 2 inertial flow paths is provided. Mass flow rates are computed based on the following momentum balance: Wjt dW r L' g y y de HL PB + #SL k PB 20 A{- ~ qB, fA ~ ~ 'SL 3L 1 K v2 '1 1' st gg jg "SL '3Ll +Ajg ~ o o 2p A g p3 I (fL/D)gg l i ~ gL jL SL SL 2p A Ik l Babcock & Mcox 1-57 REY. J (5/30/80) I . ~. -.. - -.
Revision 3 (5/30/80) where g E surge line inercia coefficient, I 7<st 6 W I surge line mass flow rate, g P I pressure at hot leg nozzle, g P E pressure at surge line diffuser, yg p E average surge line density, g Q I elevation of het leg nozzle, i "pg E pressurizer bottom elevation. A E surge line area at hot leg nozzle, g b5 ""#3' p E density of fluid at hot leg nozzle, g p, I density of fluid at surge line diffuser, p A I average flow area, g O ,,,!,orm 1oss e-fact.r. (fL/D)3L E b etion k-factor. i The temporal friction factor is computed based on the following equation: f = 0.0055[1 + (20000 c/D + 10 /Re )l! I 6 D c/D E relative roughness. Re surge line Reynolds number. D 1.15.6. Relief Valve Logic _ A model is provided to simulate the operation of as many as four pressurizer relief valves. All valves are assumed to experience the same inlet conditions; however, valve logic para =eters, setpoints, and characteristics may vary from one valve to another. The mass flux for each valve is computed based on either the HEM isentropic expansion model or input tables of mixture / steam The valve mass flow rates are co=puted as follows: mass flux versus pressure. ( 0 t Babcock s. Wilcox l-58 REV. J (5/30/80)
..-.w Revision 3 (5/30/80) A W G eA = valve valve valve where O I **1I*f "*1"* **** ' "** valve A,yy, = relief valve flow area. y An option is provided to compute an effective flow area for each valve based on HDi and rated valve characteristics. If this option is selected, A is definc.! as followed: gy, valve " R HEM (R'Y A G 5 mass flux computed using HEM, g h I rated enthalpy, R P E rated pressure. g W E rated mass flow rate. R Mass and energy removed from the pressurizer through the relief valves may be added to a TRAP 2' volume or disposed of completely. O(- 3 1.15.7. Sprav Model An option is provided te simulate spray injection at the top of the pressurizer. The sprays can either draw fluid from a TRAP 2 control volume or fluid at user specified enthalpy can be drawn from a source outside the sys-t em". Spray flow rates are either input by the user _or_co=puted using the fol-loving quasi-equilibrium occhanical energy balances /. = /2g 'sp0 ! sp W3p e AP = P,pg - P - p,p (ZPB + b ~ spi
- p I
P,pg pressure at spray line inlet nozzle, = p, density of fluid in spray line, a 2 R, spray line resistance, k/A, = Z elevation of spray line inlet nozzle. = The user may opt to input a constant value for de spray flow rate to be used in lieu of the calculated value. If no upstream centrol volume is specified, conste.nt flow input is used regardless of the flow option chosen. The spray ( flow rate is limited to non-negative values. i abcock M cox l 1-59 l REV. J (5/30/80) y-Vv w v'h-a w-e,ge--e
o-w s
g-wy - v,.9.vvr---,wy 3,-o-n w y, y~tyeaww a&+- w = v rw + w yay
==v wyey-v-,-ry-&-,--e g www-w -= wwe w =y v-vvw-ew wm e
Revision 3 (5/30/80) ~. l l v The spray is turned on and off based on input pressure setpoints. The spray control pressure is taken to be the pressure in a specified TRAP 2 control volume or the pressurizer pressura. I i Spray flow entering the pressurizer drops directly to the mix-ture region.' If the mixture region is absent, then spray flow goes directly I to the steam region. If both regions exist, sensible and condensation heat transfer from the steam to the spray droplets is computed. The total heat transfer rate is given by spr) q =W c @f -h sp cp s where 's = spray efficiency, h E saturated liquid enthalpy, g h E spray fluid enthalpy. sp Negative values for q,p are not permitted. When c, = 1, this" to h while traversing equation implies that the spray fluid is heated from h,p g the steam region. The efficiency of heat transfer to the pressurizer sprays l is computed as follows: 3 sh h
- c
=c spo a so so(h /hspo) h<h
- c
=c a spo a where c,, - eximum spray efficiency, h = mini =um steam region height required to achieve caxi=us sp efficiency. ] The rate of condensation on the spray droplets is given by a f +q esp sp y esp (h, - h ) g f E input fraction of heat transfer resulting in condensation. 1.15.8. Heater Model A pressurizer heater model allowing si=ulation of as many as five heater banks is provided. The heater bank control pressure, P,, (pres-g surizer pressure or the pressure in a user specified control volume), is [ { g b ( Babcock s.Wilcox 1-60 REv. J (5/30/80)
c... _;_.. _. Revision 3 (5/30/80) O compared with input heater bank turn-on and turn-off pressure setpoints, P, y and F , respectively, to determine the current value of electrical power. The electrical power E. for heater banks two through five is a step function, p while for heater bank one the electrical power may be controlled proportionally as follows: 1 P SP1p: p pmax E =E con E = G (P -Pcon) F1p < Pcon Phigh: p p high P E low pressure limit of proportional control band, gp E I maximum electrical power rating for heater bank number Pmax C E proportional gain for' variable output of heater bank P number one. It should be noted that input of proportional gain other than E'*** G = P P high 1p 3 will result in a discontinuity in electrical power output when the control pressure is equal to the low pressure limit of the proportional centrol band. The thermal power in: parted to the fluid, q , at time t is con:puted as follows: = q"er + E" - q"her,(t - t")/t q h ,p her htr Y I heater rod ther=al time constant (MC /UA), htr p MC, E total heat capacity of heater rods, UA I overall conductance from heating ele =ent to the surrounding.
- fluid, n = time step index.
The ther=al power it: parted to the fluid ia distributed in the pressurizer based . on the fraction of the heater bank exposed to each pressurizer fluid region. 1.15.9. Phase Stratification Models A constant velocity bubble rise codel is provided to si=ulate stratification of steam in the pressurizer. When both pressurizer regions b. x Babcock & Wilcox 1-61 REY. J (5/30/80) l
Revision 3 (5/30/80) l O are present, the model computes that part of the mass and energy traasport rates attributable to bubble rise as follows: Wb c bub's 's ~ i* and l (Wh), = W,h, i l I "I'** v bubble rise velocity, bub o E void fraction of mixture region. a y E saturated steam specific volume,. j' 8 h I saturated steam specific enthalpy. g ) When only one region is present, the distribution of steam in the pressurizer is determined from the follbwing bubble balance: I i uh
- g i
-~E) - (E #)(E ~ t")/v 3 -AVc bub (t m where I mass of bubbles in mixture. M E total mass of steam in pressurizer, g (I XW) I sum of steam flows _into_ stratified steam region. A constant velocity mos!>l for the removal of condensate droplets from the steam region is also provided. The fraction of mass and energy trans-port attributable to condensate rain is computed as follows: W =AVr & (1 ~ "s)/#f r and (Wh)r = W hrf where y = teminal velocity of condensate droplets, rah a, E void fraction of steam region, v I saturated liquid specific volume. g h I saturated liquid specific enthalpy. g C,I. (- '.t Babcock M cox l-62 REV. J (5/30/80) g i i. =amm.a ____.__.m ________m_
~ Revision 3 (5/30/80) O (') The average void fraction for the mixture in each region is given by bub's
- gub*g + (1 - X)Mv g gub E mass bubbles in region
= total steam mass if both regions are present, M E total fluid mass in region, K = mixture quality. 1.15.10. Interface Mass and Energy Transport -Two options for simulating the interaction between the steam and mixture regions are described in this section. Imbedded in the first option is the assumption that the pre-dominant mechanism by which the steam and mixture regions interact is sensible heat transfer, the rate of which is proportional to the difference in bulk fluid temperature between the two regions. The rate of heat transfer is com-puted as follows 3 kg"UA(T,-T,) gg where U = overall heat transfer coefficient, A I area of steaemhture interface. g T, I bulk temperature in steam region. ~ ~ ~ T,E bulk temperature in mixture region. The second option accounts for both sensible heat transfer and condensation at the interface. For this model the total heat transfer rate is given by k " U A (T,,g - T,) g gg where T,,, E saturation temperature. The rate of condensation is defined as .W - f,k M, - h ) e g g 0 t Babcock s Wilcox 1-63 REV. J (5/30/80)
~ s.. t Revision 3 (5/30/80) r, [ } where g I fracti n of heat transfer resulting in condensation, e h, E specific enthalpy in steam region. 1.15.11. Mixture I.evel and Interface Area The calculation of transient mixture level and interface area is co= plicated by the chopped hemispherical geometry at either end of the pres-surizer. One of the following calculations is applicable: 1. For V,< Y
- h S
e H, = 2R cos(o/3 - 37,/2sy), o = cos-I[1 - 37,/2nR ], 3 where 2 - H /2]. and A - 2n[gH g
- 2. For V IV IVh+V:
E, = H + (V - V )/A ' h m c e m h c and A =A' g c 2nQ) 3. For th+V
- Y' 8
=H h * * (#! -3 c m m T m 3 where ,, e,,-g(g _ 3(y ,y )f2, ). g and A = 2r[R ( - H,) - (y - H,) 2/2]. g h 1.15.12. Region Deletion and Recreation The removal of regions from the pressurizer calculation is ae-a co=plished by comparing the mass of each region with an input deletion switch. Thus, when the mass in the mixture region falls belov 6 =aVh l lT g where a E input mixture region deletion switch, v E initial saturated liquid specific volu=e, f the mixture region is removed from the calculation and residual mass and energy is added to the steam. region. If the advancement of the mixture mass equations would result in a negative esss being calculated for the region after a ti=a step of size at, then the time step is split into two substeps, f I I ( (_ Babcock & Wilcox l-" REV. J (5/30/80)
Revision 3 (5/30/80) / \\ At = -M /*f.51)a = nonequilibrium step, () 1 m At - At = equilibrium step. At 3 Similar logic is used to remove the steam region from pressurizer calculations. In this case, however, the deletion criterion is j V/ 42 ~ "2 T 's O '*
- a 3 input steam region deletion switch, 2
y E initial saturated steam specific volume. , A region that has been removed from the pressurizer calculations will be reinstated when user specified conditions are met. The mixture region is recreated when the mass of mixture in the steam region, Mas, exceeds a user specified value, or M, > a 63 3 where a = input mixture region creation switch. Similarly, the steam region 3 3 I is recreated when the mass of stratified steam in the mixture regiott, M,,, exceeds a user specified value, or M "8, 02 sm where a, I input steam region creation switch. The following modes of execution are,possible in the TRAP 2 pressurizer model: 1. 2 = mixture and steam regions present. 2. 4 E only steam region present. 3. 5 I only mixture region present. 1.15.13. TPAP2 Interfaces The interfaces between the pressurizer model and the overall TRAP 2 solution include mass and energy extraction and deposition through the surge line, spray line and relief valves. These interfaces are included in the control volume mass, energy, bubble, and boron inventory moduls. i I O l F. 2bcock s, Wilcox 1-65 REV. J (5/30/80
.a Revision 3 (5/30/80) l I Pressurizer pressure and liquid volume may be used as setpoints in TRAP 2 trip lot,.:. by selecting the maximum allowable control volume number 3 as the detector volu=e (see page vii)*. E l I I l .I, l s 1 f { ) ( s Babcock r. Wilcox 1-66 REV. J (5/30/90)
I
- 25) To permit the NRC to perform audit calculations, provide the area, thickness, and location of the primary and secondary metal :.3at slabs used for analysis of BSAR-205.
Justify the omission of any metal not considered in the analysis.
RESPONSE
Primary and secondary metal heat slabs were not included in the BSAR-205 analysis. For the time span of M/E releases, heat slabs do not affect the results and it is conservative not to include primary metal heat in determination of the minimum subcritical (overcooling) margin. An analysis was perforned to demonstrate this using the heat slab data listed on Table 25-1. The resulting system response,is_shown in Figure 25-1. I' can be seen by comparing tnis figure to the base case figure (see question 2) that inclusion of the metal heat slabs has no' impact in the base case results (i.e., in the time frame of the analysis the resulting dynamic response with metal, heat slabs differs by less than 1% from the results without the heat slabs). O s m m. . - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - ' - " - - - - ~ - - - - ~ ~ ~ ^ - ^ - '~ - - - " - - - ' - - - - - - - - - - ' - - " ^ ' - ' ^ ~ ~ ' ' - - - - - - -
C) HEAT SLAB DATA TABLE 25-1 Heat Capacity Heat Transfer Coefficient Node BTU /Deg. F BTU /Sec-Deg. F 1 31502. 7.894 2 58442. 60.021 3 39385. 70.615 4* 0. O. 5* 0. O. 6* 0. O. 7* 0. O. 8* 0. 0. 9* 0. O. 10* 0. O. 11* 0. O. 12* 0. O. 13* 0. O. 14 28922. 70.797 15 85975. 65.19 16 39385. 70.615 17* 0. O. 18* 0. O. 19* 0 0. 20* 0. O. 21* 0. O. 22* 0. O. 23* 0. O. 24* 0. O. 25* 0. O. 26* 0. O. 27 28922. 70.797 28 13216. 6.92 29 13216. 6.92 30 13216. 6.92
() k- / HtAT SLAB UATA TABLE 25-1 (Continued) Node Heat Capacity Heat Transfer Coefficient BTV/Deg. F BTU /Sec-Deg. F 31 5757.4 .29 32 5757.4 .29 33 5757.4 .29 34 5757.4 .29 35 5757,4 .29 36 5757.4 .29 37 5757.4 .29 38 5757.4 .29 39 5757.4 .29 40 5757.4 .29
- s. /
41 5757.4 .29 42 5757.4 .29 43 5757.4 .29 44 5757.4 .29 45 5757.4 .29 46 5757.4 .29 47 5757.4 .29 48 5757.4 .29 49 5757.4 .29 50 5757.4 .29 51 -- 73** 0. O. The steam generator tube model already models these heat capacities and heat transfer coefficients. The steam piping and FW piping heat capacity a:e not es ( ) included. These are plant specific items and are not considered to be of importance in modeling this transient. ~
x h, .t Figure 25-1. System Response O I r is C ] 3,4 620 l if i2 600 3 a } C 5 1.0 580 i g 560 1{ E 0.8 = {- = u D-1 o 0.6 .S. 540 t. m l S 520 h 0.4 = g,2 500 i 1 l 0.0 -480 4 0 10 20 30 40 50 CD 0 10 20 30 40 50 60 i Time, see Time, sec O } t t.': 35 j 4300 '-- 2 ) 3500 30 3 j = n. 3000
- 25 4,
o a a 2500 J 20 = c = W 2000 $ 15 h a-n .a a E t i .S 1500 10 I S l e 1000 .o 5 J; u. 'h. 0 = = 500 0 10 20 30 40 50 60 0 10 20 30 40 50 60 Time, sec Time, sec 4 e 9 _..._m., -, - _ _,, _. _ - _ - _,,.. _, _. -... - - _,. _,.,,, -, _ _ _. ,_-.m-...r. -}}