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The coolability limits of a reactor pressure vessel lower head
E L S EV] E R
Nuclear Engineering and Design
Nuclear Engineering and Design 169 (1997) 59-76
The coolability limits of a reactor pressure vessel lower head T.G.
Theofanous
*, S. S y r i
Center Jbr Risk Studies and Safety, University of California, Santa Barbara, CA 93106, USA
Received 5 December 1996
Abstract
Configurations II and III of the ULPU experimental facility are described, and results from a comprehensive set of experiments are provided. The facility affords full-scale simulations of the boiling crisis phenomenon on the hemispherical lower head of a reactor pressure vessel submerged in water, and heated internally. Whereas Configuration I experiments (published previously) established the lower limits of coolability under low submergence, pool-boiling conditions, with Configuration II we investigate coolability under conditions more appropriate to practical interest in severe accident management; that is, heat flux shapes (as functions of angular position) representative of a core melt contained by the lower head, full submergence of the reactor pressure vessel, and natural circulation. Additionally, with Configuration III, we examine the effect of a channel-like geometry created by the reactor vessel thermal insulation. Critical heat fluxes as a function of the angular position on the lower head are reported and related to the observed two-phase flow regimes. © 1997 Elsevier Science S.A.
1. Introduction
on the inside, be below what could cause a boiling crisis (B) on the outside. The key features o f the
The purpose o f this paper is to make available the first experimental data directly relevant to establishing the coolability limits (critical heat flux) o f reactor-scale hemispheres submerged in water and heated internally. The situation arises in the m a n a g e m e n t o f severe accidents, and a relatively recent idea that the relocation of molten corium could be arrested, at the lower head o f a reactor pressure vessel, by external flooding, as in Fig. 1 ( T h e o f a n o u s et al., 1994a). F o r this idea to work, it is necessary that the thermal load created by natural convection o f the heat-generating pool
Steel Layer
8o~ (O) / q(O)
* Corresponding author.
~
Corium Pool
~ ]
Fig. 1. Schematic of the in-vessel retention phenomenology.
0029-5493/97/$17.00 © 1997 Elsevier Science S.A. All rights reserved. PII S0029-5493(97)00024-1
60
T.G. Theo/anous, S. Syri Nuch, ar Enghwering and Design 169 (1997) 59- 76
[ Clearance 12
Clearance 2
Fig. 2. Key geometric features of cavity flooding and venting paths. Illustration of the two-phase boundary layer on the lower head. Thermal insulation not shown.
problem can be seen with the help of Fig. 2, depicting the reactor-vessel-cavity configuration in the AP600 design, which is our main current interest (Theofanous et al., 1997). Considering first the 'local' aspects, we will note that the downward-facing geometry lends itself to the formation of a two-phase boundary layer that is 'squeezed' upon the heating surface by gravity forces. This 'squeeze' is moderated by turbulent mixing (including interracial instabilities and entertainment) as buoyancy drives the steam past the heating surface and the surrounding water. The balance between these two mechanisms is quite different in the various regions around the lower head. This gives rise to widely different two-phase flow regimes, and as a consequence, we can expect significant variations of the critical heat flux and perhaps even of the underlying mechanism(s) for it. For example, in the vicinity of the stagnation point (0 ~ 0 °) the vapor velocities are very low and the surface orientation, relative to the gravity vector, drives the maximum phase separation with the vapor 'squeezed' up against the wall. We expect periodic formation of relatively large bubbles, growth, and escape. Under such flow conditions we can expect that boiling crisis will occur if the underlying thin liquid film dries out, within one such period. As a
consequence, surface wetability (controlling the behavior of the thin film), and wall thickness (controlling the rate of heatup following the dryout of this film) should be important. Aspects that could affect the behavior of these bubbles, including any convection in the free-stream water (see global behavior discussed below) and its subcooling, should also be important. On the other hand, near the equator of the lower head hemisphere (0 ~ 90°), we expect that cumulative vapor generation from all the upstream positions will give rise to high vapor, and entrained liquid, flow velocities, and in addition, the vapor "squeeze' effect described above would be minimal in the nearparallel orientation of the surface to the gravity vector. As a consequence, we expect a highly turbulent two-phase flow within a relatively ~diffuse' boundary layer. Correspondingly, the BC mechanism would be convection-dominated.
ULPU-2000
Configuration I
Fig. 3. Schematic of configuration I in ULPU-2000. The heater blocks extend over the region 30 ° < 0 < 30 °.
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76
Vent
II
II Spray
Exit Restriction
\\\
~
\
:>
Condenser
ULPU-2000 Configuration n
DP 2
_ + _ I Riser
I"T~
Downcomer
DPI
Baffle to obtain Configuration Ill
I ~~
Electromagnetic flow meter
Heater blocks Observation windows
I
void drift in the riser, and the void fraction distribution depends on the power input (including shape), the subcooling, and on the flow losses along the natural circulation path--especially on any inlet or outlet restriction. The following clarifications need to be made: (a) power is also supplied long the riser section, it being the radiative heat flow conducted through the side wall of the reactor vessel; (b) our present interest is in wide open paths into the reactor cavity, as in Fig. 2; and (c) key aspects of the flow geometry would be dictated by the thermal insulation design (not shown in Fig. 2), which is a reactor specific feat u r e - h e r e we concentrate on a basically unimpeded geometry, which in principle is a feasible design option (it is in the AP600). In addition, the dynamics of the natural circulation loop, such as flow and pressure oscillations may be important (i.e. loop elasticity), especially on the bubble regime discussed above.
Fig. 4. Schematicof Configurations II and II1 in ULPU-2000. The heater blocks extend over the region 0° < 0 < 90°. However, phase separation and perhaps even capillarity may continue to have some bearing on the mechanism in this regime as well. Evidently intermediate (or mixed) mechanisms can be envisioned as we go from the stagnation region to the equator with the details of this transition region being strongly dependent upon the power (input) shape. Let us turn next to the global aspects. In a fully-flooded cavity, as in Fig. 2, the gravity head would be ~ 7 m, which corresponds to ~ 14°C subcooling with saturated water at the top. The supply of this subcooled water to the cavity depends on any flow restriction on the flow path, and on the gravity head developed due to the two-phase flow in the riser. Moreover, the availability of any subcooled water in the immediate vicinity of the heating surface would depend on the local mixing and flow patterns as described above. This coupling of the local to the global behavior is quite important in that it indicates a truly conjugate behavior. The key mechanism is
61
Fig. 5. The ULPU-2000 configuration II in operation.
62
T.G. TheoJcmous, S. Syri / Nuclear Engineering and Design 169 (1997) 59 76 aspects of the phenomena involved. Based on the above discussion, these key aspects include: 1. Heater length scale and shape. To properly represent the two-phase boundary layer behavior, especially in the intermediate region, we need a full-length geometry, including the correct curvature. 2. Heating surface thermal inertia. In the reactor, the lower head thickness could vary from its initial value of 15 cm (for most of the region), to a low value of a few cent\metres due to thinning by melt attack near the equator (Theofanous et al., 1997). Therefore, a minimum of a few cent\metres in heated thickness is needed to properly represent the thermal inertia. A
//lllll\\
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III
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II II II II II II il / I /I I
Fig. 6. The U L P U - 2 0 0 0 c o n f i g u r a t i o n II test section.
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Fig. 7. One of the h e a t e r b l o c k s i m m e d i a t e l y after m a c h i n i n g .
Mechanistically, these are behaviors unlike any CHF situation considered previously. Considering also the status of fundamental understanding in the extensively investigated problem of BC on a horizontal, upward facing plate (Lienhard, 1994), we chose to pursue an experimental approach that faithfully simulates the reactor in all the key
~
Side view
3.0' mm Fig. 8. The heating zones a n d t h e r m o c o u p l e p o s i t i o n s on a h e a t i n g block. Over the three h e a t e r blocks there are a total of 24 zones. Z o n e no. 1 is at 0 ~ 90 ° a n d zone no. 24 is at 0 ~ 0 ° .
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76 Table 1 Listing of runs carried out in Configuration II
1.o
~"
~
0.8
Group
Subgroup
Remarks
0.6
SF
SF-1 SF-7 SF-13 SF-17 SF-21
3 3 1 3 2
Tests Tests Test Tests Tests
UF
UF-3-180% UF-3-220% UF-8-145% UF-8-180%
1 1 2 1
Test Test Tests Test
0.4
,.e
/ J j J
/ 0.2
•
ti
/
20
J
40
60
Angle
80
100
0
Fig. 9. The reference reactor heat flux shape ( ) (from Theofanous et al., 1994a), and the shape used in the simulations ( - - - ) . 3.5
--- -
3.0
UF-8-180%
, -'
Power Shaping Principle
.
'
i "
2.5
2.0
1.5
~,
63
1.0
0.5
15
30
45
60
75
90
Angle (degrees)
Fig. 10. Illustration of the power shape 'compromise' necessary to force boiling crisis at O = 0 °.
3. Power shape. In conjunction with the above item 1, a correct power history in the upstream region is necessary to represent the two-phase boundary layer conditions in the reactor. In addition, the total power shape, including that in the riser, must be represented to simulate void distribution and flow behavior. 4. Heating surface wetability. The reactor lower head is designed to be made of forged carbon steel (SA 508), machined to 200 r.m.s, mean roughness. Even though it is to be painted in situ, the paint is expected to flake off in boiling water, and the exposed steel surface to be well-wetted. This surface condition must be confirmed and matched experimentally. 5. Loop length scale and hydraulic diameter. To properly represent subcooling due to gravity
head we need a full-length loop (--~ 7 m). The annular gap in the reactor is ,-~ 20 cm ( ~ 15 cm excluding the insulation), and a similarly large cross sectional length scale (diameter) is needed in a one-dimensional geometry to allow proper simulation of the vapor drift. The U L P U facility was built to embody all of the above key characteristics, allowing, therefore, a full-scale simulation capability. The experimental approach is evolving gradually from overall parametric studies and simulations (of the reactor conditions of interest), to detailed investigation of local phenomena (Angelini and Theofanous, 1996), and thus eventually to identification of the crisis mechanism and to an analytical model. The emphasis on simulations in the early part of this program derives from two reasons. One is to afford an early identification and focus on the
1.2
~:~
UF-8-145% -- i 1 ~ UF 3 180% UF-8-L80% UF 3 220%
S
1.0
0,8
0.6
0.4
0.2
15
30
45
60
75
90
Angle (degrees)
Fig. 11. Heat flux profiles imposed on UF-type runs. The shape required for 'simulation' for these runs is shown in Fig. 10.
64
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76
particular flow and heat transfer regimes relevant to the problem of practical interest. The other, and perhaps more important one, is that the practical need for reasonably robust estimates of C H F is imminent as reactor-specific accident management schemes are now up against key decision points and regulatory scrutiny (AP600 and Loviisa--see Theofanous et al., 1994a). The results presented in this paper are intended to fulfil this immediate need. In addition, these resuits provide an initial perspective on mechanisms as a starting point for the more detailed investigation of the phenomena at the local level. It should also be noted that both of these reactors have lower heads with no penetrations, so the effects of such complications in geometry are left for future studies. The experiment concept is based on what we call the 'power shaping principle' (Theofanous et al., 1994b). Briefly, it allows us to determine the power shape on a two-dimensional test section (representing a 'slice' of the lower head) needed to create the correct hydrodynamic conditions (matching those in the reactor) at any angular position for which the critical heat flux is sought. So far, the experiment involves three distinct configurations as in Fig. 3 and Fig. 4. Configuration I is for studying saturated, pool boiling in - 30 ° < 0 < 30 °, and especially in the region around O ~ 0 °, which is not as well represented in the other configurations. Configuration I! is for simulating the complete geometry (a one-quarter circle) under loop flow (including the effects of subcooling) conditions. As seen in Fig. 4, this configuration is to represent an open-tothe-cavity geometry. A channel geometry, as it might arise from particular thermal insulation designs with an inlet at the very bottom (0 ~ 0°), can be created by introducing a baffle, as illustrated in Fig. 4, to obtain Configuration III. The results for Configuration I have been presented previously (Theofanous et al., 1994b). Up to an including the present work, the heater blocks were made out of copper and the data were obtained with the surface of it aged. Item 4 from the above-specified similarity requirements was addressed by using a heat block made of prototypic (AP600) steel and painted likewise, as described in Appendix E.4 of DOE/ID-10460.
2. Description of configurations I! and III The overall geometry of the experiment and related terminology are shown in Fig. 4. There are three heater blocks (the primary heater, to be described shortly below) fit on top of a two-dimensional chamber (15 cm wide) with a shape (in the other two dimensions) as shown. This chamber simulates an open lower cavity geometry (no reactor vessel insulation); it is 2 m wide (at the base) and 0.5 m in height at the short end. The chamber is made of stainless steel sheet, 2.5 mm thick, and it is reinforced externally over the flat areas to sustain the substantial hydrostatic forces without any significant distortion. This test section stands bolted on shock absorbers capable of carrying 3000 kg, which in turn are anchored to the floor. The riser and downcomer are assembled from glass piping 6 m long, and 15.2 and 7.6 cm 6.~
, ....
5•C
L o)
I n p u t - - SF-21 P o w e r S h a p i n g Principle
,_2
4.£
3.£ 1,- - N
2,C
,_2
1 £
0
(a)
15
30
45
6O
75
~0
A n g l e (degrees) I.(
~ - - - r
0A
0.t
o.,
0." i O0~ "
(b)
15
30
I
k
45
60
Dehvered Flux 75
J 90
A.glo (degrees)
Fig. 12. (a) The actual power shape for run SF-21 compared to that derived from the power shaping principle. (b) The effect of axial conduction in the heater block in modifying the input flux shape for run SF-21.
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76 5.l - .....
Input -- SF-17 Power Shaping
Principle
4.C
3.0
~=
2,o
7" ,-2
1.0
o~
,
,
i
,
,
15
(a)
F
,
,
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,
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(degrees)
1.0
0.8
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0.2 - ..... '
(b)
115
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3r0
' Angle
415
'
610
Input Delivered '
715
Flux '
90
(degrees)
Fig. 13. (a) The actual power shape for run SF-17 compared to that derived from the power shaping principle. (b) The effect of axial conduction in the heater block in modifyingthe input flux shape for run SF-17. in diameter, respectively. This riser is equipped with a cable heater (referred to as the secondary heater) extending essentially over the whole length, and operating at a fixed power level of 24 kW. The purpose of this heater is to simulate the radiative power from the melt delivered through the wall of the reactor pressure vessel to the water in the annular space between the cavity wall and the reactor vessel. (The riser in the experiment corresponds to this space). Finally, the condenser unit at the top is to minimize coolant losses and allow operation with saturated water at the inlet to the downcomer. In Configuration II (and III) runs, this was accomplished by direct contact condensation; that is, a fine spray of demineralized, room temperature water injected at rates sufficient to make up the steam loss to the environment (let out through a permanently open vent line). The condenser was also equipped with a safety relief valve which, however, did not ever have to energize.
65
The Configuration III specifications as shown in Fig. 4 were obtained from Fig. K.3a in Appendix K of DOE/ID-10460. The baffle simulating the lower part of the reflective insulation was introduced such as to allow the minimum clearance of 9" (22.86 cm) at an angular position of 30 °. The riser is 15.24 cm in diameter as compared to the 15 cm annular gap allowed around the reactor vessel by the insulation. Finally, an exit restriction (10 cm diameter) was introduced as shown to reduce the flow area down to ,~ 40% of that available in the riser (as in the case for the reactor). These riser and restriction dimensions correspond well to the prorated values based on the minimum simulated circumferential fraction in U L P U . This minimum is obtained for a matching angle of 90 ° and corresponds to 1.2% of the reactor circumference. The riser and restriction areas then become 220 cm 2 and 84 c m 2, respectively, with equivalent diameters of 16.7 and 10 cm respectively. The whole facility, in operation, is shown in Fig. 5 and Fig. 6. The three heater blocks were built, each covering nominally a 30 ° arc of a circle, with a radius of curvature of 1.76 m, which is well within the range of typical reactor dimensions. The actual arc is 27 °, with a length of 0.83 m. The block height and width were chosen as 7.6 and 15 cm, respectively, such as to ensure sufficient thermal inertia and to minimize side-wall effects. One such block, just prior to assembly, is shown in Fig. 7. The holes are 9.5 m m in diameter and allow two cartridge heaters (per hole) to be inserted, one from each side. The 'fit' must be very good, while precision machining is also required in the accurate positioning of the holes, the 1.5 m m in diameter holes needed for the thermocouples, and in forming the smooth, curved faces. Copper stock was selected as the raw material. Voltage (217 V) is supplied to the cartridge heaters through 32 relays which are individually computer controlled to cycle in the ' o n - o f f positions so as to obtain any desired power distribution on the heater block 1. In the runs reported here we In all experiments the cyclingtime was 3.52 s, which is small when compared to the conduction time constant of the copper block (51 s). The transient conduction simulation of the heater block shows that the heat flux variation on the surface is less than 0.1%
66
T.G. TheoJkmous, S. Syri / Nuclear Engineering and Design 169 (1997) 59 76
- .....
Input -- SF 7 Pov, er S h a p i n g Principle
- .....
I n p u t - - SF- I P o w e r S h a p i n g Principle
1.(
o~
r~
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o. 15
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3'0
(a)
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....
610
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_ ,.._,_~._,~__ F_J-f-
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90
(a)
A n g l e (degrees)
1l
....
0 (b)
15
30
45
60
~o
;5
'
. . . . . .
~,
Input
F
75
7'5 '----9 o
Angle (degrees)
90
A n g l e (dcgrees)
(b)
0
15
30
45
Angle(deg,-~e,)
69
75
99
Fig. 14. (a) The actual power shape for run SF-7 compared to that derived from the power shaping principle. (b) The input flux shape used for SF-7, and the actual shape obtained (output flux) accounting for axial conduction.
Fig. 15. (a) The actual power shape for run SF-1 compared to that derived from the power shaping principle. (b) The effect of axial conduction in the heater block in modifying the input flux shape for run SF-1.
g r o u p e d the r e l a y - c a r t r i d g e c o n n e c t i o n such as to create eight i n d i v i d u a l l y c o n t r o l l e d h e a t i n g zones per h e a t e r b l o c k (3.375 ° arc each, for a t o t a l o f 24 zones). T h e h e a t i n g b l o c k s can deliver up to 2000 k W m 2 (hence the n a m e U L P U - 2 0 0 0 ) , a n d a t o t a l p o w e r o f 500 k W . T h e t o t a l p o w e r o b t a i n e d by s u m m i n g the c a l c u l a t e d p o w e r o f each cartridge ( f r o m the v o l t a g e applied, the k n o w n resistivity, a n d the f r a c t i o n a l ' o n ' - t i m e ) , a g r e e d to within 1% o f the t o t a l p o w e r s u p p l i e d by the p o w e r g e n e r a t o r . T e m p e r a t u r e s are m e a s u r e d at eight c o r r e s p o n d i n g p o s i t i o n s a l o n g the length o f each block, as in Fig. 8. O n the sides, the blocks o v e r l a p with the 2.5 m m stainless ' c h a m b e r ' walls (Fig. 4), b y a p p r o x i m a t e l y 13 ram, with a r u b b e r g a s k e t in between. A b s e n c e o f steep t e m p e r a t u r e g r a d i e n t s in this region ( u n d e r nucleate boiling c o n d i t i o n s the surface s u p e r h e a t is b e l o w ~ 40°C) assures t h a t a n y ' b y p a s s ' losses to w a t e r are also negligible.
T h e flow regimes c o u l d be o b s e r v e d t h r o u g h the w i n d o w s illustrated in Fig. 4, a n d n o r m a l as well as high speed video r e c o r d i n g s were m a d e for m o r e d e t a i l e d study. W a t e r t e m p e r a t u r e s were m e a s u r e d at the inlet a n d outlet o f the test section, a n d at several a n g u l a r p o s i t i o n s as a function o f d i s t a n c e f r o m the the h e a t e r surface (using
Table 2 Heat flux correction (to the input flux) due to the axial conduction Run no.
CHF Position
Correction Factor
All UF SF-21 SF-17 SF- 13 SF-7 SF-1
0° 15° 30° 45 ° 67.5° 90°
1 0.965 0.992 0.94 0.962 0.998
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76
67
tion. Heat losses were evaluated, both analytically and experimentally, and found to be negligible.
0.65
0.60 b r~
3. Overview of the test program
0.55
•
o
o • • © O
0.50
! 0.45
• o ,
0.4t
i 5
.
.
.
.
.
Boiling Crisis Nucleate Boilin ll5 . . . .
li0
20
Operating Time (hrs)
Fig. 16. Illustration of the surface aging effect in Configuration II, for times greater than 4 h. Data for earlier times are not shown. All data refer to boiling crisis at 0 = 0°. 2.0
1.5
~.~ 1.0 ,'r ~
0.5
Configuration I!
1
• o
A
Boiling Crisis Nucleate Boiling
Configuration I ~ ] • [] 15
30
45
Angle (degrees)
Boiling Crisis Nucleate Boiling
i
i
60
75
,
90
Fig. 17. Collective presentation of all critical heat flux in both Configuration I and Configuration II experiments in ULPU2000. The line shown for Configuration II is Eq. (3). The boiling crisis points at 0 ~ 0° have been displaced to the right for clarity of presentation.
thermocouple trees). There are thermocouples also at the 'back side' of the heater b l o c k s - - t h e y are monitored during operation, on-line, together with the surface thermocouples as the critical safety parameters. Void fractions in the riser were measured, in the lower and upper halves of it, using differential pressure measurement. Boiling/ condensation-induced vibrations wer~ obtained from a pressure transducer mounted on the short end of the test section 'chamber'. Finally, the loop flow rate was obtained from an electromagnetic flow meter installed around the middle of the downcomer section. Data from all these instruments were continuously recorded during opera-
The main series of runs with Configuration II were carried out with sufficient water in the loop to close the natural circulation path. Thus, there were a relatively strong flow, 2 depending on the total power level (typically ~ 120 gpm), and subcooling at the base, corresponding to the gravity head ( ~ 10 K). Although the flow regimes near the heater surface in these runs were qualitatively similar to those observed in Configuration I, here the BC would occur typically within a few minutes or not at all. It appears that the present conditions (flow, subcooling) do not favor a BCdelay-time behavior, as found in Configuration I. Thus, zeroing in the C H F was now much easier; starting from the high end, the power was reduced after successive occurrences of boiling crisis (each of those runs being only a few minutes long), until the last reduction (typically by a few percent) where the heater remained in nucleate boiling. Once it was established that the delay time was not a significant factor here, this nucleate boiling condition was allowed to continue, for most runs, for 10-30 min. A grouping of several successive BCs and a final period in nucleate boiling is referred to as one experiment run. Regarding the power shape, two types of runs were carried out: 'uniform flux' (UF) runs, involving a number of zones around 0 = 0 ° at a uniform power (that is, simulating heat flux at 0 ~ 0 °) while the remaining two blocks were powered according to the power shaping principle; and 'shaped flux' (SF) runs, powered according to the power shaping principle, to simulate BC at specific locations (0-¢ 0 °) along the test section. The reference power shape used here is the same as that employed in Configuration I, and it is shown in Fig. 9. However, certain modifiers from these general descriptions are necessary: 2 However, an auxiliary set was also performed, to examine sensitivity to flow rates and power shapes, as described in Appendix A.
68
T.G. TheoJanous, S. Syri /Nuclear Engineering and Design 169 (1997) 59-76
Table 3 Listing of the experimental runs performed in ULPU-2000 Configuration II, heat fluxes, and respective waiting times Test
Nucleate Boiling Flux (kW m -2)
CHF (kW m -2)
B.C. Time (min)
Time (rain)
UF-3-180%
461
15
UF-3-220% UF-8-145%(1)
479
15
UF-8-145% (2)
527
120
UF-8-180% SF-I (1) SF-I (2)
498 1485 1504
15 10 30
SF-1 (3) SF-7 (1) SF-7 (2) SF-7 (3) SF-7 (4)
1579 1231 1246 1282 1299
10 30 30 10 60
SF-13 SF-17 (1) SF-17 (2) SF-17 (3) SF-21 (1)
932 932 914 914 789
30 10 30 30 30
SF-21 (2)
762
30
1. Because of the strong p e a k i n g in the reference shape, the use of power s h a p i n g principle for 0 ~ 0 ° always produces BC at 0 = 90°; that is, u n d e r reactor conditions, the position 0 ~ 90 ° is by far m o r e limiting t h a n the very b o t t o m o f the lower head. Thus, to o b t a i n the BC at 0 ~ 0 ° we h a d to reduce the d o w n s t r e a m power very significantly below that required by the power shaping principle. This was done, up to whatever degree was necessary, while m a i n t a i n ing the shape, however. T h a t is, the highest power in the d o w n s t r e a m (to the u n i f o r m flux section) region is still at 90 ° . This power level was d e n o t e d in the r u n identification n u m b e r as a percentage of the u n i f o r m flux level, ira-
517 498 479
2.5 3.5 14
545 526 508 573 564 555 536
1.5 3.5 47 1.5 1 2 3
1617 1579 1617
1 2 2.5
1335 1317 1335
2 1.5 1
932
2.5
862 843 816 789
1 2 2 13
posed at 0 = 0 °. A n example of what has been described above is given in Fig. 10. 2. F o r S F runs, the same ' c o m p r o m i s e ' , for the same reason described above, was necessary for forcing BC a n y w h e r e o n the lower block (0 < 0 < 30°). 3. F o r BC in the u p p e r block (60 ° < 0 < 90°), a n essentially exact s i m u l a t i o n could be o b t a i n e d for the lower end o f it (0 -~ 67.5°); b u t in order to force BC at 90 °, the u p s t r e a m power profile within the third block h a d to be reduced somewhat. The r e a s o n for this is that in this u p p e r region the reactor flux is rather flat, which translates to a rather u n i f o r m shape also acc o r d i n g to the p o w e r shaping principle, thus yielding BC at 0 < 90 °.
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76
A b o u t naming, we follow the same procedure employed in Configuration I. That is, we use a prefix, U F or SF, to denote uniform and shaped flux, respectively. Following this there is a numerical index that for U F runs indicates the number of zones subject to the uniform flux, while for SF runs, this index indicates the zone at which BC is simulated. Uniform flux runs are only employed to simulate BC at the b o t t o m end of the the test section ( 0 - 0°). Then, there is a third numerical index (as noted above already) for U F runs only, indicating the percentage of increase in power at the upper end of the test section (0 = 90 °) in relation to that on the uniform flux portion. Finally, the number in parenthesis at the end indicates the order in which repeat runs (if any) were performed; repeat runs were not run consecutively, but rather were intermingled a m o n g the various runs.
--
1.5
=
10
.~)
0.5
,
~ L
i
15
~_.,
30
,
45
i
i
60
75
90
A n g l e (degrees)
Fig. 18. Configuration lII results in ULPU compared to proposed CHF correlation, Eq. (3). too0
~
800
,
......
,
k
. . . .
I
. . . .
69
To illustrate the above, UF-8-145% (2) means: (a) a uniform flux run, simulating BC at 0 ~ 0 °, with all eight zones of the lower block (i.e. zones 17-24) at the same power; (b) middle and upper blocks powered according to the power shaping principle (as discussed above) and such that the power level at the upper end (90 ° ) is 145% of that at 0 = 0°; and (c) this is the second run carried out under these conditions. On the other hand, SF-7 (4) means: (a) run with a heat flux distribution according to the power shaping principle; (b) a run simulating BC in zone # 7 (0 ~ 67°); and (c) this is the fourth run carried out under these conditions. A listing of all runs carried out in Configuration II is given in Table 1. The flux shapes employed in all U F runs are shown in Fig. 11. The flux shapes employed in all SF runs are shown in relation to the shapes dictated by the power shaping principle in Figs. 12(a)-15(b). The effect of axial conduction in smoothing the input (stepwise) fluxes is shown in Fig. 12(b)-15(b), and the implied flux corrections' at the points of boiling crisis are summarized in Table 2. These corrections have been applied already to the data reported. This type of correction for all U F runs is negligible. F o r all Configuration II runs, the copper block surface was fully aged. All Configuration III runs were carried out under a reference flux shape corresponding to the base reactor case, as in Appendix A, and with the lower heater block surface in a 'fresh' condition, cleaned to a fine polish (about aging effects, see below).
I
Old Shape
4. Experimental results and discussion ~"
6oo
~"~
400
200
0
. . . .
o
I
20
. . . .
I
. . . .
40 Angle
I
60
. . . .
i
,
8o
(degrees)
Fig. 19. The presently used base case flux shape in relation to that used in the tests reported in the body of the paper.
A rough idea of the surface aging effect on C H F is in Fig. 16. In it, we see that the effect is measurable even after ~ 4 h of operation, but it is clearly leveled out, compared to the very initial value of 450 kW m -2. Also, for 0 = 9 0 ° , the initial value was ~ 1.2 M W m 2, as compared to the fully-aged value of ~ 1.6 M W m 2. As noted already, this work is focused on the fully-aged
T.G. Theofanous, S. Syri /Nuclear Engineering and Design 169 (1997) 59 76
70
Table 4 Summary of the U L P U (Configuration II) experimental program on flux shape and recirculation flow effects ID Number a
Reference b Shape
Z00A
Fig. Q. 1
Maximum Flow C H F (kW/m 2) 508 564
Medium Flow C H F (kW/m 2)
No Flow C H F (kW/m 2)
Angular Position (Actual) c (°)
545
0 0 0 0(6) 0 0 15 15 26 26 37 37 37 37 45 45 45 45 45 45 56
658 (Z00B)
498 564
(Z00C) Z08B
Fig. Q.1
ZI4A
Fig. Q. 1
Z19B
Fig. Q.I
(Z 19C) Z24A
Fig. Q. 1
602 718 770 902 902 1125 1139 1072 1072 1205 1191 1151
649 866 1005
991 1096 1083 1042
(Z24B) (Z24C) Z30A
Fig. Q.1
1179 1228 1116
1064
1104
Z35A
Fig. Q.I
Z35C (Z40A)
Fig. 7.16
Z40B (Z40C) Z42A Z48A
Fig. Q.I Fig. 7.15 Eq. (5.30)
Z48B
Eq. (5.30) with cut
1263 1263 1255 1340 1186 1318 1340 1489 1310 1523 1480 1622
56
1190
66 66 66 66 75(90) 75(80) 75 75 75(85) 78 90 90 90 90
1370
1263 1118
1353
a The number indicates the zone number at which C H F is being simulated. Total 48 equal size zones starting with zone 0 at 0 ° and ending with zone 48 at 90 ° . Numbers in parenthesis show sensitivity variations around the base case. Number in bold print. Parametric cases examined for the reactor, in DOE/ID-10460. b Flux shapes based on calculated reactor shapes. Figures and equations refer to DOE/ID-10460. Fig. Q.I is the base case simulation. The actual boiling crisis position is shown in parenthesis when it is different from the expected one.
surface condition, and these are the results reported and discussed below. The results from all Configuration II runs are in Fig. 17, and in tabular form in Table 3. In Fig. 17, we also show the results from Configuration I.
The following observations
and comments
can be
made: 1. C o n f i g u r a t i o n tolerance
to
Configuration
II exhibits a significantly higher boiling I. T h e s e
crisis
as
compared
to
are clearly the effects
T.G. Theofanous, S. S y r i / N u c l e a r Engineering and Design 169 (1997) 5 9 - 7 6
of flow and subcooling, and they amount to about 50% increase• 2. For Configuration II, the variation of CHF with angular position seems to be composed of two, remarkably linear, regions. We believe these to be reflective to the significantly different flow regimes observed in the respective regions (see discussion in the introduction section). 3. The CHF level at the upper end ( ~ 1.6 M W m - 2 ) is remarkably higher than the 'conventional wisdom value' for a horizontal, upward facing, flat plate in pool boiling, which is ~ 1 M W m 2. Clearly, flow and orientation are important. Further, it should be noted that in Table 3 we find three runs in which BC occurred rather late compared to a few minutes; however, note that all
. . . . ...........
o:
71
I ' ' ',~' I '
I . . . . . . . . Principle Input
1 j
./"
6
," - "'
" / 'i ~i
..-"" ""
;_..
.,./" F~ 4
/" /,
2
....( " , , , I
.... 20
I ....
L,,
I ....
40 60 Angle (Degrees)
80
F i g . 2 2 . The input flux shape, in comparison to that obtained from the power shaping principle for run Z00C.
I .... ............
. . . .
I
]
'
Principle Input Actually Delivered
S '
'
'
I
. . . .
'
'
4 •
!
'
I
'
171171711 Pn.cipl~ Input Actually Delivered
/
2
// /-
20
40 Angle (Degrees)
60
80
2 .
o~, 0
, , I . . . . . . . . EO 40 Angle
I .... 60
I , 80
(Degrees)
F i g . 2 0 . The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z 0 0 A .
.
.
.
.
.
nc,*,opu,
F i g . 2 3 . The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z 0 8 B .
]
.........i
S
three cases were for the lowermost region ( 0 ° < 0 < 15°), and that the differences were rather small (2-4%). This again is another indication, or symptom, of a different mechanism in the lower region identified in the data trends (Fig. 17). The lower envelope for the Configuration II data in Fig. 17 can be expressed by: qcr(0)
=
500-1- 13.30 k W m - 2
for
0 < 15 °
(1) qcr(0) = 20
40 Ang]e
60
80
(Degrees)
Fig: 21. The input flux shape, in comparison to that obtained from the power shaping principle for run Z 0 0 B .
540 + 10.70 kW m
15 ° < 0 < 90 °
2
for (2)
However, the data can be fit equally well with one equation:
T.G. Theofanous, S. Syri /Nuclear Engineering and Design 169 (1997) 59 76
72
qcr(0)
. . . .
= 4 9 0 + 3 0 . 2 0 - 8.88- 1 0 - 1 0 2 + 1.35-10 -6.65"10-504kWm
203
2
- ~ ~ ~i~i
I
'
'
....
'
I ....
t.z
(3)
This line is shown in Fig. 17. The Configuration III results are in Fig. 18. As expected, in the upper regions, the effect of a channel-like geometry, i.e., enhanced convection, is seen to be favorable, but quantitatively not very significant. At the pole (0 ~ 0°), we can see more of an effect, as convection is just about enough to compensate for the deleterious consequence of a fresh, poorly wetted, surface. The uncertainty in these results is very small compared to available margins in the case of
I '
Principle Input
ii/',,i J
;::i","i"t",
0 0
,
,
I
20
I
. . . .
40
. . . .
I
60
,
80
Angle (Degrees) Fig. 26. T h e input flux shape in c o m p a r i s o n to that obtained from the p o w e r shaping principle for run Z19B. I ....
. . . . . . . .
I '
Actually Delivered 1.5
,,•,,"• m,
,
,
.:[ ~[[[i~[[
,
I
'i
- ¸•
. . . .
lnputPrinciple
Actually Delivered
05
5
"/,
,
0.0
2
20
/'
Angle
o,,I,, 0
20
b
40
Angle
~
I
q
80
60
. . . .
o~
.............
I
' ' ' ' I
. . . .
' ' I ' ,""i!!
Principle [npul Actually Delivered
/
'i - ' "
2 •._ .-"
A /,'
I 20
60
40 Angle
,
,
t , , , , I , 60 80
(Degrees)
Fig. 27. T h e input and actually delivered flux shapes, in c o m p a r i s o n to that obtained from the p o w e r shaping principle for run Z 2 4 A .
(Degrees)
Fig. 24. T h e input and actually delivered flux shapes, in c o m p a r i s o n to that obtained from the p o w e r shaping principle for run Z 1 4 A .
3
,
40
(Degrees)
Fig. 25. T h e input and actually delivered flux shapes, in c o m p a r i s o n to that obtained from the p o w e r shaping principle for run Z19C.
AP600 (Theofanous et al., 1997) as well as in comparison to other uncertainties that normally enter such assessments. Basically, there are three reasons for this: (1) small measurement error in cartridge power ( ~ 2%); (2) negligible heat losses from the heating blocks; and (3) convenience for back-to-back runs and easy recovery from BC. As noted already, the small measurement error is confirmed by findinbg cumulative power of cartridge heaters to be within 1% of the total power supplied (independently measured). The easy recovery from BC is because of the high thermal inertia of the heater block• Back-to-back experiments allow the approach to the true BC limit by successive fine-step approximations. As in Fig. 17 and Table 3, the data show that this could be done within a few percent.
73
T.G. Theofanous, S. Syri /Nuclear Engineering and Design 169 (1997) 59-76 5. Conclusions
1.5
. . . .
I
. . . .
I . . . .
i~i~.~i~i~:~i~. InputPrinciple
With the completion of the Configuration II and Configuration III experiments reported here, we have established a firm estimate for the coolability limits of the lower head of a reactor vessel submerged in water and heated internally. The scale of the experiment, and the power shaping principle, ensure that these data, and resulting correlation, are directly applicable to the reactor. The correlation reveals a two region behavior (Eq. (1) and Eq. (2)) which has been related to the different flow regimes observed in the two-phase boundary layer in the respective regions.
. . . .
I
. . . .
I
. . . .
I
'
'
, "'~
0.5
0.0
~
.... 20
i ....
I ....
40 60 Ang]e (Degrees)
i, 80
Fig. 30. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z30A. . . . .
o~ I
I
1.0
.......
. . . .
. . . .
Actually Delive~d
1.~
20
I
]
I
. . . .
I
. . . .
I
. . . .
I
,7
Input Principle ] Aclually Delivered
1.o
'
15 ~
0.5
20
0.5
40 60 Angle (Degrees)
B0
Fig. 31. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z35A.
0.0 20
40 60 Angle (Degrees)
80
Fig. 28. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z24B. Acknowledgements 2.0
2
~
1.5
.
.
.
.
I
.
.
.
I
.
Actually Delivered
.
.
.
.
L
.
.
.
.-.,/"'"/'"/,/"
.
I
!
'
']
',........... "
0.5
0.0 20
40 Angle (Degrees)
60
80
Fig. 29. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z24C.
This work was funded by IVO International Ltd. (supporting S. Syri) and the U.S. Department of Energy's Advanced Reactor Severe Accident Program (ARSAP) through ANL subcontract No. 23572401. We are also grateful to Sergio Angelini, T o n y Salmassi, and Mitch Frey, for help with design and construction of Configuration II, and to Jeff Smith and Charlie Liu, for helping with the construction, running of the latest portion of the experiments, and helping with the data reductions. Messrs. Steven Sorrell (DOE) and Stephen Addition ( T E N E R A ) played a key role in programmatically supporting this project.
T.G. TheoJanous, S. Syri /Nuelear Engineering and Design 169 (1997) 59 76
74
Appendix A. Critical heat flux sensitivity to heat flux shapes and recirculation flow rates
t.o
og
--
A series of experiments was carried out in U L P U , to examine sensitivity of the critical heat fluxes to the heat flux shape and the natural circulation flow rates. The experimental facility was in Configuration II, as described above, except for increasing the power control resolution (by adding relays to the control circuit) to 48 zones. In the runs described here each zone corresponds to an arc length of 1.875 ° . We have also changed the nomenclature; now, the zone numbers begin at 0 ° and increase with angular position, so that zone 0 is over 0 ° < 0 < 1.875 ° and zone 48 over 88.125 ° < 0 < 9 0 ° .
]"'-
. . . .
- - - -
'
'
'
I
....
I
'
'
Z '
I
Principle Input ActuallyDelivered
0.6 I
04
02
20
40
60
80
Angle (Degrees)
Fig. 34. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z40B. I0
[.'po, P' 08
....
[0
~
0.8
2
.... ............
''
Principle
,~ . . . .
Inpu~ Actuay De vered
~ , , - ~ : ~
[
0.6
"
J
0.4
0.6 02 04
0.0 3
0.;~
0
20
40
60
Fig. 32. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z35C. ,
,
.....
•0 .
, -~-
_J _. . . ._
I ....
I'
Principle Pr Input
' . 2 ~ - , ~ . / , . ~ -I ,
,
;Z
08
0.4
0,2
20
60
40
Angle
go
(Degrees)
Fig. 35. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z40C.
80
(Degrees)
Angle
60
40 Angle
0. 0
1.0
20
80
(Degrees)
Fig. 33. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z40A.
Flux shape sensitivity is examined using, besides the reference flux shape shown in Fig. 19 (note that this is somewhat different from the shape used in the test series described a b o v e 'old shape'), several variations around it, as well as shapes associated with the parametric cases considered for the reactor (in DOE/ID-10460). Flow rate sensitivity studies were carried out by throttling the d o w n c o m e r so as to obtain about 60 gpm, or about half the flow normally obtained ( ~ 130 gpm) with the flow path completely open. In addition, the case o f zero flow was examined by letting the level drop sufficiently so as to interrupt the natural circulation flow path. The experimental matrix is summarized in Table 4, together with the resulting C H F values.
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76
The second column indicates the origin of the reference flux shape used in each case. The actual flux shapes used are shown, together with the shapes that would be required according to the power shaping principle, in Figs. 20-37. The results are shown in graphical form in Fig. 38 and Fig. 39. Fig. 38 contains the comparison between all present full-flow data with the CHF correlation, Eq. (3). This figure shows that the correlation represents a robust lower envelope even under some not insignificant variations of the heat flux shapes. But this figure also shows that there is a significant additional tolerance to CHF under a strongly peaked heat
t.0
. . . .
I
. . . .
I
. . . .
I ' '
'
~; ',[~
i,'j
P~putpL
20
40
60
80
(Degrees)
Angle
Fig. 36. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z42A. 1.0
. . . .
I
. . . .
I
. . . .
I
I 'j
. . . .
2000
~
1600
~
1200
o~
.>/: / Y
800
400
I
,
,
,
I
20
,
,
,
40
I
,
,
,
60
E
,
,
80
Angle (degrees)
100
Fig. 38. Comparison of all present full flow data with the correlation, Eq. (3). The open circles are for a flux representative of the extreme parametric (discontinuous, high peak at the top).
flux at the top. The sensitivity to the flow rate can be seen in Fig. 39. It is very interesting that the zero flow case produces a negative effect only near the top and even then, it is only a 20% reduction. Note that the zero flow case in effect eliminates the cavity subcooling as it is operated under a significantly lower water level ( ~ 2.0 vs. 5.6 m from the bottom of the lower head). These results indicate that simply immersing the lower head of the AP600 would be sufficient to remove the thermal loads from a base-case-like scenario with very significant margins.
2.0
J
IPlput pie
75
,
0.8 1.6
• [:3 ×
130 G P M 60 GPM 0 GPM
0.6
: ~
1.2
Q
04
0.8
| o
02
0.4
0.0 20
40 Angle
6O
80
(Degrees)
Fig. 37. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z48A.
,
I
20
,
,
,
~
40
,
,
,
i
60
Angle (degrees)
,
,
,
I
80
,
,
,
100
Fig. 39. Sensitivity of measured C H F on natural circulation flow rate.
76
T.G. TheofiTnous, S. Syri /Nuclear Engineering and Design 169 (1997) 59-76
References Lienhard, J.H., Snares of Pool Boiling Research: Putting Our History to Use, in G.F. Hewitt, (ed.), 1994, Proceedings 10th International Heat Transfer Conference, Brighton, UK. Vol. 1, Keynote Papers, Gk-ll, Chameleon Press, London, pp. 333 348. Theofanous, T.G., Liu, C., Angelini, S., Kym/ilS.inen, O., Tuomisto, H. and Addition, S., 1994a, Experience from the First Two Integrated Approaches to In-Vessel Retention Through External Cooling, OECD/CSNI/NEA Workshop
on Large Molten Pool Heat Transfer, Grenoble, France, March 9-11. Theofanous, T.G., Syri, S., Salmassi, T., Kym/ilfiinen, O. and Harri Tuomisto, 1994b, Critical heat flux through curved, downward facing, thick walls, Nucl. Eng. Des., 15, 247258. Theofanous, T.G., Liu, C., Additon, S., Angelini, S., Kymfil~iinen, O. and Salmassi, T., In-vessel coolability and retention of a core melt, Nucl. Eng. Des., (this issue). Angelini, S. and Theofanous, T.G., 1996, A New Boiling Transition Regime, Proceedings, The Japan-US Seminar on Two-Phase Flow Dynamics, July 15-20, Fukuoka, Japan.
Nuclear Engineering and Design
Nuclear Engineering and Design 169 (1997) 59-76
The coolability limits of a reactor pressure vessel lower head T.G.
Theofanous
*, S. S y r i
Center Jbr Risk Studies and Safety, University of California, Santa Barbara, CA 93106, USA
Received 5 December 1996
Abstract
Configurations II and III of the ULPU experimental facility are described, and results from a comprehensive set of experiments are provided. The facility affords full-scale simulations of the boiling crisis phenomenon on the hemispherical lower head of a reactor pressure vessel submerged in water, and heated internally. Whereas Configuration I experiments (published previously) established the lower limits of coolability under low submergence, pool-boiling conditions, with Configuration II we investigate coolability under conditions more appropriate to practical interest in severe accident management; that is, heat flux shapes (as functions of angular position) representative of a core melt contained by the lower head, full submergence of the reactor pressure vessel, and natural circulation. Additionally, with Configuration III, we examine the effect of a channel-like geometry created by the reactor vessel thermal insulation. Critical heat fluxes as a function of the angular position on the lower head are reported and related to the observed two-phase flow regimes. © 1997 Elsevier Science S.A.
1. Introduction
on the inside, be below what could cause a boiling crisis (B) on the outside. The key features o f the
The purpose o f this paper is to make available the first experimental data directly relevant to establishing the coolability limits (critical heat flux) o f reactor-scale hemispheres submerged in water and heated internally. The situation arises in the m a n a g e m e n t o f severe accidents, and a relatively recent idea that the relocation of molten corium could be arrested, at the lower head o f a reactor pressure vessel, by external flooding, as in Fig. 1 ( T h e o f a n o u s et al., 1994a). F o r this idea to work, it is necessary that the thermal load created by natural convection o f the heat-generating pool
Steel Layer
8o~ (O) / q(O)
* Corresponding author.
~
Corium Pool
~ ]
Fig. 1. Schematic of the in-vessel retention phenomenology.
0029-5493/97/$17.00 © 1997 Elsevier Science S.A. All rights reserved. PII S0029-5493(97)00024-1
60
T.G. Theo/anous, S. Syri Nuch, ar Enghwering and Design 169 (1997) 59- 76
[ Clearance 12
Clearance 2
Fig. 2. Key geometric features of cavity flooding and venting paths. Illustration of the two-phase boundary layer on the lower head. Thermal insulation not shown.
problem can be seen with the help of Fig. 2, depicting the reactor-vessel-cavity configuration in the AP600 design, which is our main current interest (Theofanous et al., 1997). Considering first the 'local' aspects, we will note that the downward-facing geometry lends itself to the formation of a two-phase boundary layer that is 'squeezed' upon the heating surface by gravity forces. This 'squeeze' is moderated by turbulent mixing (including interracial instabilities and entertainment) as buoyancy drives the steam past the heating surface and the surrounding water. The balance between these two mechanisms is quite different in the various regions around the lower head. This gives rise to widely different two-phase flow regimes, and as a consequence, we can expect significant variations of the critical heat flux and perhaps even of the underlying mechanism(s) for it. For example, in the vicinity of the stagnation point (0 ~ 0 °) the vapor velocities are very low and the surface orientation, relative to the gravity vector, drives the maximum phase separation with the vapor 'squeezed' up against the wall. We expect periodic formation of relatively large bubbles, growth, and escape. Under such flow conditions we can expect that boiling crisis will occur if the underlying thin liquid film dries out, within one such period. As a
consequence, surface wetability (controlling the behavior of the thin film), and wall thickness (controlling the rate of heatup following the dryout of this film) should be important. Aspects that could affect the behavior of these bubbles, including any convection in the free-stream water (see global behavior discussed below) and its subcooling, should also be important. On the other hand, near the equator of the lower head hemisphere (0 ~ 90°), we expect that cumulative vapor generation from all the upstream positions will give rise to high vapor, and entrained liquid, flow velocities, and in addition, the vapor "squeeze' effect described above would be minimal in the nearparallel orientation of the surface to the gravity vector. As a consequence, we expect a highly turbulent two-phase flow within a relatively ~diffuse' boundary layer. Correspondingly, the BC mechanism would be convection-dominated.
ULPU-2000
Configuration I
Fig. 3. Schematic of configuration I in ULPU-2000. The heater blocks extend over the region 30 ° < 0 < 30 °.
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76
Vent
II
II Spray
Exit Restriction
\\\
~
\
:>
Condenser
ULPU-2000 Configuration n
DP 2
_ + _ I Riser
I"T~
Downcomer
DPI
Baffle to obtain Configuration Ill
I ~~
Electromagnetic flow meter
Heater blocks Observation windows
I
void drift in the riser, and the void fraction distribution depends on the power input (including shape), the subcooling, and on the flow losses along the natural circulation path--especially on any inlet or outlet restriction. The following clarifications need to be made: (a) power is also supplied long the riser section, it being the radiative heat flow conducted through the side wall of the reactor vessel; (b) our present interest is in wide open paths into the reactor cavity, as in Fig. 2; and (c) key aspects of the flow geometry would be dictated by the thermal insulation design (not shown in Fig. 2), which is a reactor specific feat u r e - h e r e we concentrate on a basically unimpeded geometry, which in principle is a feasible design option (it is in the AP600). In addition, the dynamics of the natural circulation loop, such as flow and pressure oscillations may be important (i.e. loop elasticity), especially on the bubble regime discussed above.
Fig. 4. Schematicof Configurations II and II1 in ULPU-2000. The heater blocks extend over the region 0° < 0 < 90°. However, phase separation and perhaps even capillarity may continue to have some bearing on the mechanism in this regime as well. Evidently intermediate (or mixed) mechanisms can be envisioned as we go from the stagnation region to the equator with the details of this transition region being strongly dependent upon the power (input) shape. Let us turn next to the global aspects. In a fully-flooded cavity, as in Fig. 2, the gravity head would be ~ 7 m, which corresponds to ~ 14°C subcooling with saturated water at the top. The supply of this subcooled water to the cavity depends on any flow restriction on the flow path, and on the gravity head developed due to the two-phase flow in the riser. Moreover, the availability of any subcooled water in the immediate vicinity of the heating surface would depend on the local mixing and flow patterns as described above. This coupling of the local to the global behavior is quite important in that it indicates a truly conjugate behavior. The key mechanism is
61
Fig. 5. The ULPU-2000 configuration II in operation.
62
T.G. TheoJcmous, S. Syri / Nuclear Engineering and Design 169 (1997) 59 76 aspects of the phenomena involved. Based on the above discussion, these key aspects include: 1. Heater length scale and shape. To properly represent the two-phase boundary layer behavior, especially in the intermediate region, we need a full-length geometry, including the correct curvature. 2. Heating surface thermal inertia. In the reactor, the lower head thickness could vary from its initial value of 15 cm (for most of the region), to a low value of a few cent\metres due to thinning by melt attack near the equator (Theofanous et al., 1997). Therefore, a minimum of a few cent\metres in heated thickness is needed to properly represent the thermal inertia. A
//lllll\\
Illflll\\ /Ill\liNk II II1/\ III I I l \ \
III
I\\ I\ I\ /\
II II II II II II il / I /I I
Fig. 6. The U L P U - 2 0 0 0 c o n f i g u r a t i o n II test section.
I\ I/ 1/ I\ I\ I
/ II / iI iI / III /1 / /
8 ] 7 I 6
Top view
1 3 ¶-t
4~_
4
13.75 cm
/
/ / / /
///
3 I 2 \ l l'~ Z o n ; # / /
•
¢ 8 / 7 / 6 ] 5 ~ 4
10.00 cm
/
It \ \\ q \\ \\ 1\\ \\\ll 5
I
- 5.00 cm
Tc
-3 I 2 X 1
\
Fig. 7. One of the h e a t e r b l o c k s i m m e d i a t e l y after m a c h i n i n g .
Mechanistically, these are behaviors unlike any CHF situation considered previously. Considering also the status of fundamental understanding in the extensively investigated problem of BC on a horizontal, upward facing plate (Lienhard, 1994), we chose to pursue an experimental approach that faithfully simulates the reactor in all the key
~
Side view
3.0' mm Fig. 8. The heating zones a n d t h e r m o c o u p l e p o s i t i o n s on a h e a t i n g block. Over the three h e a t e r blocks there are a total of 24 zones. Z o n e no. 1 is at 0 ~ 90 ° a n d zone no. 24 is at 0 ~ 0 ° .
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76 Table 1 Listing of runs carried out in Configuration II
1.o
~"
~
0.8
Group
Subgroup
Remarks
0.6
SF
SF-1 SF-7 SF-13 SF-17 SF-21
3 3 1 3 2
Tests Tests Test Tests Tests
UF
UF-3-180% UF-3-220% UF-8-145% UF-8-180%
1 1 2 1
Test Test Tests Test
0.4
,.e
/ J j J
/ 0.2
•
ti
/
20
J
40
60
Angle
80
100
0
Fig. 9. The reference reactor heat flux shape ( ) (from Theofanous et al., 1994a), and the shape used in the simulations ( - - - ) . 3.5
--- -
3.0
UF-8-180%
, -'
Power Shaping Principle
.
'
i "
2.5
2.0
1.5
~,
63
1.0
0.5
15
30
45
60
75
90
Angle (degrees)
Fig. 10. Illustration of the power shape 'compromise' necessary to force boiling crisis at O = 0 °.
3. Power shape. In conjunction with the above item 1, a correct power history in the upstream region is necessary to represent the two-phase boundary layer conditions in the reactor. In addition, the total power shape, including that in the riser, must be represented to simulate void distribution and flow behavior. 4. Heating surface wetability. The reactor lower head is designed to be made of forged carbon steel (SA 508), machined to 200 r.m.s, mean roughness. Even though it is to be painted in situ, the paint is expected to flake off in boiling water, and the exposed steel surface to be well-wetted. This surface condition must be confirmed and matched experimentally. 5. Loop length scale and hydraulic diameter. To properly represent subcooling due to gravity
head we need a full-length loop (--~ 7 m). The annular gap in the reactor is ,-~ 20 cm ( ~ 15 cm excluding the insulation), and a similarly large cross sectional length scale (diameter) is needed in a one-dimensional geometry to allow proper simulation of the vapor drift. The U L P U facility was built to embody all of the above key characteristics, allowing, therefore, a full-scale simulation capability. The experimental approach is evolving gradually from overall parametric studies and simulations (of the reactor conditions of interest), to detailed investigation of local phenomena (Angelini and Theofanous, 1996), and thus eventually to identification of the crisis mechanism and to an analytical model. The emphasis on simulations in the early part of this program derives from two reasons. One is to afford an early identification and focus on the
1.2
~:~
UF-8-145% -- i 1 ~ UF 3 180% UF-8-L80% UF 3 220%
S
1.0
0,8
0.6
0.4
0.2
15
30
45
60
75
90
Angle (degrees)
Fig. 11. Heat flux profiles imposed on UF-type runs. The shape required for 'simulation' for these runs is shown in Fig. 10.
64
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76
particular flow and heat transfer regimes relevant to the problem of practical interest. The other, and perhaps more important one, is that the practical need for reasonably robust estimates of C H F is imminent as reactor-specific accident management schemes are now up against key decision points and regulatory scrutiny (AP600 and Loviisa--see Theofanous et al., 1994a). The results presented in this paper are intended to fulfil this immediate need. In addition, these resuits provide an initial perspective on mechanisms as a starting point for the more detailed investigation of the phenomena at the local level. It should also be noted that both of these reactors have lower heads with no penetrations, so the effects of such complications in geometry are left for future studies. The experiment concept is based on what we call the 'power shaping principle' (Theofanous et al., 1994b). Briefly, it allows us to determine the power shape on a two-dimensional test section (representing a 'slice' of the lower head) needed to create the correct hydrodynamic conditions (matching those in the reactor) at any angular position for which the critical heat flux is sought. So far, the experiment involves three distinct configurations as in Fig. 3 and Fig. 4. Configuration I is for studying saturated, pool boiling in - 30 ° < 0 < 30 °, and especially in the region around O ~ 0 °, which is not as well represented in the other configurations. Configuration I! is for simulating the complete geometry (a one-quarter circle) under loop flow (including the effects of subcooling) conditions. As seen in Fig. 4, this configuration is to represent an open-tothe-cavity geometry. A channel geometry, as it might arise from particular thermal insulation designs with an inlet at the very bottom (0 ~ 0°), can be created by introducing a baffle, as illustrated in Fig. 4, to obtain Configuration III. The results for Configuration I have been presented previously (Theofanous et al., 1994b). Up to an including the present work, the heater blocks were made out of copper and the data were obtained with the surface of it aged. Item 4 from the above-specified similarity requirements was addressed by using a heat block made of prototypic (AP600) steel and painted likewise, as described in Appendix E.4 of DOE/ID-10460.
2. Description of configurations I! and III The overall geometry of the experiment and related terminology are shown in Fig. 4. There are three heater blocks (the primary heater, to be described shortly below) fit on top of a two-dimensional chamber (15 cm wide) with a shape (in the other two dimensions) as shown. This chamber simulates an open lower cavity geometry (no reactor vessel insulation); it is 2 m wide (at the base) and 0.5 m in height at the short end. The chamber is made of stainless steel sheet, 2.5 mm thick, and it is reinforced externally over the flat areas to sustain the substantial hydrostatic forces without any significant distortion. This test section stands bolted on shock absorbers capable of carrying 3000 kg, which in turn are anchored to the floor. The riser and downcomer are assembled from glass piping 6 m long, and 15.2 and 7.6 cm 6.~
, ....
5•C
L o)
I n p u t - - SF-21 P o w e r S h a p i n g Principle
,_2
4.£
3.£ 1,- - N
2,C
,_2
1 £
0
(a)
15
30
45
6O
75
~0
A n g l e (degrees) I.(
~ - - - r
0A
0.t
o.,
0." i O0~ "
(b)
15
30
I
k
45
60
Dehvered Flux 75
J 90
A.glo (degrees)
Fig. 12. (a) The actual power shape for run SF-21 compared to that derived from the power shaping principle. (b) The effect of axial conduction in the heater block in modifying the input flux shape for run SF-21.
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76 5.l - .....
Input -- SF-17 Power Shaping
Principle
4.C
3.0
~=
2,o
7" ,-2
1.0
o~
,
,
i
,
,
15
(a)
F
,
,
30
i
,
45 Angle
,
i
,
60
7c5
'
90
(degrees)
1.0
0.8
0.6
=
0.4
0.2 - ..... '
(b)
115
'
3r0
' Angle
415
'
610
Input Delivered '
715
Flux '
90
(degrees)
Fig. 13. (a) The actual power shape for run SF-17 compared to that derived from the power shaping principle. (b) The effect of axial conduction in the heater block in modifyingthe input flux shape for run SF-17. in diameter, respectively. This riser is equipped with a cable heater (referred to as the secondary heater) extending essentially over the whole length, and operating at a fixed power level of 24 kW. The purpose of this heater is to simulate the radiative power from the melt delivered through the wall of the reactor pressure vessel to the water in the annular space between the cavity wall and the reactor vessel. (The riser in the experiment corresponds to this space). Finally, the condenser unit at the top is to minimize coolant losses and allow operation with saturated water at the inlet to the downcomer. In Configuration II (and III) runs, this was accomplished by direct contact condensation; that is, a fine spray of demineralized, room temperature water injected at rates sufficient to make up the steam loss to the environment (let out through a permanently open vent line). The condenser was also equipped with a safety relief valve which, however, did not ever have to energize.
65
The Configuration III specifications as shown in Fig. 4 were obtained from Fig. K.3a in Appendix K of DOE/ID-10460. The baffle simulating the lower part of the reflective insulation was introduced such as to allow the minimum clearance of 9" (22.86 cm) at an angular position of 30 °. The riser is 15.24 cm in diameter as compared to the 15 cm annular gap allowed around the reactor vessel by the insulation. Finally, an exit restriction (10 cm diameter) was introduced as shown to reduce the flow area down to ,~ 40% of that available in the riser (as in the case for the reactor). These riser and restriction dimensions correspond well to the prorated values based on the minimum simulated circumferential fraction in U L P U . This minimum is obtained for a matching angle of 90 ° and corresponds to 1.2% of the reactor circumference. The riser and restriction areas then become 220 cm 2 and 84 c m 2, respectively, with equivalent diameters of 16.7 and 10 cm respectively. The whole facility, in operation, is shown in Fig. 5 and Fig. 6. The three heater blocks were built, each covering nominally a 30 ° arc of a circle, with a radius of curvature of 1.76 m, which is well within the range of typical reactor dimensions. The actual arc is 27 °, with a length of 0.83 m. The block height and width were chosen as 7.6 and 15 cm, respectively, such as to ensure sufficient thermal inertia and to minimize side-wall effects. One such block, just prior to assembly, is shown in Fig. 7. The holes are 9.5 m m in diameter and allow two cartridge heaters (per hole) to be inserted, one from each side. The 'fit' must be very good, while precision machining is also required in the accurate positioning of the holes, the 1.5 m m in diameter holes needed for the thermocouples, and in forming the smooth, curved faces. Copper stock was selected as the raw material. Voltage (217 V) is supplied to the cartridge heaters through 32 relays which are individually computer controlled to cycle in the ' o n - o f f positions so as to obtain any desired power distribution on the heater block 1. In the runs reported here we In all experiments the cyclingtime was 3.52 s, which is small when compared to the conduction time constant of the copper block (51 s). The transient conduction simulation of the heater block shows that the heat flux variation on the surface is less than 0.1%
66
T.G. TheoJkmous, S. Syri / Nuclear Engineering and Design 169 (1997) 59 76
- .....
Input -- SF 7 Pov, er S h a p i n g Principle
- .....
I n p u t - - SF- I P o w e r S h a p i n g Principle
1.(
o~
r~
755 r
o. 15
"
3'0
(a)
' -'
415
....
610
7iS
_ ,.._,_~._,~__ F_J-f-
3o
90
(a)
A n g l e (degrees)
1l
....
0 (b)
15
30
45
60
~o
;5
'
. . . . . .
~,
Input
F
75
7'5 '----9 o
Angle (degrees)
90
A n g l e (dcgrees)
(b)
0
15
30
45
Angle(deg,-~e,)
69
75
99
Fig. 14. (a) The actual power shape for run SF-7 compared to that derived from the power shaping principle. (b) The input flux shape used for SF-7, and the actual shape obtained (output flux) accounting for axial conduction.
Fig. 15. (a) The actual power shape for run SF-1 compared to that derived from the power shaping principle. (b) The effect of axial conduction in the heater block in modifying the input flux shape for run SF-1.
g r o u p e d the r e l a y - c a r t r i d g e c o n n e c t i o n such as to create eight i n d i v i d u a l l y c o n t r o l l e d h e a t i n g zones per h e a t e r b l o c k (3.375 ° arc each, for a t o t a l o f 24 zones). T h e h e a t i n g b l o c k s can deliver up to 2000 k W m 2 (hence the n a m e U L P U - 2 0 0 0 ) , a n d a t o t a l p o w e r o f 500 k W . T h e t o t a l p o w e r o b t a i n e d by s u m m i n g the c a l c u l a t e d p o w e r o f each cartridge ( f r o m the v o l t a g e applied, the k n o w n resistivity, a n d the f r a c t i o n a l ' o n ' - t i m e ) , a g r e e d to within 1% o f the t o t a l p o w e r s u p p l i e d by the p o w e r g e n e r a t o r . T e m p e r a t u r e s are m e a s u r e d at eight c o r r e s p o n d i n g p o s i t i o n s a l o n g the length o f each block, as in Fig. 8. O n the sides, the blocks o v e r l a p with the 2.5 m m stainless ' c h a m b e r ' walls (Fig. 4), b y a p p r o x i m a t e l y 13 ram, with a r u b b e r g a s k e t in between. A b s e n c e o f steep t e m p e r a t u r e g r a d i e n t s in this region ( u n d e r nucleate boiling c o n d i t i o n s the surface s u p e r h e a t is b e l o w ~ 40°C) assures t h a t a n y ' b y p a s s ' losses to w a t e r are also negligible.
T h e flow regimes c o u l d be o b s e r v e d t h r o u g h the w i n d o w s illustrated in Fig. 4, a n d n o r m a l as well as high speed video r e c o r d i n g s were m a d e for m o r e d e t a i l e d study. W a t e r t e m p e r a t u r e s were m e a s u r e d at the inlet a n d outlet o f the test section, a n d at several a n g u l a r p o s i t i o n s as a function o f d i s t a n c e f r o m the the h e a t e r surface (using
Table 2 Heat flux correction (to the input flux) due to the axial conduction Run no.
CHF Position
Correction Factor
All UF SF-21 SF-17 SF- 13 SF-7 SF-1
0° 15° 30° 45 ° 67.5° 90°
1 0.965 0.992 0.94 0.962 0.998
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76
67
tion. Heat losses were evaluated, both analytically and experimentally, and found to be negligible.
0.65
0.60 b r~
3. Overview of the test program
0.55
•
o
o • • © O
0.50
! 0.45
• o ,
0.4t
i 5
.
.
.
.
.
Boiling Crisis Nucleate Boilin ll5 . . . .
li0
20
Operating Time (hrs)
Fig. 16. Illustration of the surface aging effect in Configuration II, for times greater than 4 h. Data for earlier times are not shown. All data refer to boiling crisis at 0 = 0°. 2.0
1.5
~.~ 1.0 ,'r ~
0.5
Configuration I!
1
• o
A
Boiling Crisis Nucleate Boiling
Configuration I ~ ] • [] 15
30
45
Angle (degrees)
Boiling Crisis Nucleate Boiling
i
i
60
75
,
90
Fig. 17. Collective presentation of all critical heat flux in both Configuration I and Configuration II experiments in ULPU2000. The line shown for Configuration II is Eq. (3). The boiling crisis points at 0 ~ 0° have been displaced to the right for clarity of presentation.
thermocouple trees). There are thermocouples also at the 'back side' of the heater b l o c k s - - t h e y are monitored during operation, on-line, together with the surface thermocouples as the critical safety parameters. Void fractions in the riser were measured, in the lower and upper halves of it, using differential pressure measurement. Boiling/ condensation-induced vibrations wer~ obtained from a pressure transducer mounted on the short end of the test section 'chamber'. Finally, the loop flow rate was obtained from an electromagnetic flow meter installed around the middle of the downcomer section. Data from all these instruments were continuously recorded during opera-
The main series of runs with Configuration II were carried out with sufficient water in the loop to close the natural circulation path. Thus, there were a relatively strong flow, 2 depending on the total power level (typically ~ 120 gpm), and subcooling at the base, corresponding to the gravity head ( ~ 10 K). Although the flow regimes near the heater surface in these runs were qualitatively similar to those observed in Configuration I, here the BC would occur typically within a few minutes or not at all. It appears that the present conditions (flow, subcooling) do not favor a BCdelay-time behavior, as found in Configuration I. Thus, zeroing in the C H F was now much easier; starting from the high end, the power was reduced after successive occurrences of boiling crisis (each of those runs being only a few minutes long), until the last reduction (typically by a few percent) where the heater remained in nucleate boiling. Once it was established that the delay time was not a significant factor here, this nucleate boiling condition was allowed to continue, for most runs, for 10-30 min. A grouping of several successive BCs and a final period in nucleate boiling is referred to as one experiment run. Regarding the power shape, two types of runs were carried out: 'uniform flux' (UF) runs, involving a number of zones around 0 = 0 ° at a uniform power (that is, simulating heat flux at 0 ~ 0 °) while the remaining two blocks were powered according to the power shaping principle; and 'shaped flux' (SF) runs, powered according to the power shaping principle, to simulate BC at specific locations (0-¢ 0 °) along the test section. The reference power shape used here is the same as that employed in Configuration I, and it is shown in Fig. 9. However, certain modifiers from these general descriptions are necessary: 2 However, an auxiliary set was also performed, to examine sensitivity to flow rates and power shapes, as described in Appendix A.
68
T.G. TheoJanous, S. Syri /Nuclear Engineering and Design 169 (1997) 59-76
Table 3 Listing of the experimental runs performed in ULPU-2000 Configuration II, heat fluxes, and respective waiting times Test
Nucleate Boiling Flux (kW m -2)
CHF (kW m -2)
B.C. Time (min)
Time (rain)
UF-3-180%
461
15
UF-3-220% UF-8-145%(1)
479
15
UF-8-145% (2)
527
120
UF-8-180% SF-I (1) SF-I (2)
498 1485 1504
15 10 30
SF-1 (3) SF-7 (1) SF-7 (2) SF-7 (3) SF-7 (4)
1579 1231 1246 1282 1299
10 30 30 10 60
SF-13 SF-17 (1) SF-17 (2) SF-17 (3) SF-21 (1)
932 932 914 914 789
30 10 30 30 30
SF-21 (2)
762
30
1. Because of the strong p e a k i n g in the reference shape, the use of power s h a p i n g principle for 0 ~ 0 ° always produces BC at 0 = 90°; that is, u n d e r reactor conditions, the position 0 ~ 90 ° is by far m o r e limiting t h a n the very b o t t o m o f the lower head. Thus, to o b t a i n the BC at 0 ~ 0 ° we h a d to reduce the d o w n s t r e a m power very significantly below that required by the power shaping principle. This was done, up to whatever degree was necessary, while m a i n t a i n ing the shape, however. T h a t is, the highest power in the d o w n s t r e a m (to the u n i f o r m flux section) region is still at 90 ° . This power level was d e n o t e d in the r u n identification n u m b e r as a percentage of the u n i f o r m flux level, ira-
517 498 479
2.5 3.5 14
545 526 508 573 564 555 536
1.5 3.5 47 1.5 1 2 3
1617 1579 1617
1 2 2.5
1335 1317 1335
2 1.5 1
932
2.5
862 843 816 789
1 2 2 13
posed at 0 = 0 °. A n example of what has been described above is given in Fig. 10. 2. F o r S F runs, the same ' c o m p r o m i s e ' , for the same reason described above, was necessary for forcing BC a n y w h e r e o n the lower block (0 < 0 < 30°). 3. F o r BC in the u p p e r block (60 ° < 0 < 90°), a n essentially exact s i m u l a t i o n could be o b t a i n e d for the lower end o f it (0 -~ 67.5°); b u t in order to force BC at 90 °, the u p s t r e a m power profile within the third block h a d to be reduced somewhat. The r e a s o n for this is that in this u p p e r region the reactor flux is rather flat, which translates to a rather u n i f o r m shape also acc o r d i n g to the p o w e r shaping principle, thus yielding BC at 0 < 90 °.
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76
A b o u t naming, we follow the same procedure employed in Configuration I. That is, we use a prefix, U F or SF, to denote uniform and shaped flux, respectively. Following this there is a numerical index that for U F runs indicates the number of zones subject to the uniform flux, while for SF runs, this index indicates the zone at which BC is simulated. Uniform flux runs are only employed to simulate BC at the b o t t o m end of the the test section ( 0 - 0°). Then, there is a third numerical index (as noted above already) for U F runs only, indicating the percentage of increase in power at the upper end of the test section (0 = 90 °) in relation to that on the uniform flux portion. Finally, the number in parenthesis at the end indicates the order in which repeat runs (if any) were performed; repeat runs were not run consecutively, but rather were intermingled a m o n g the various runs.
--
1.5
=
10
.~)
0.5
,
~ L
i
15
~_.,
30
,
45
i
i
60
75
90
A n g l e (degrees)
Fig. 18. Configuration lII results in ULPU compared to proposed CHF correlation, Eq. (3). too0
~
800
,
......
,
k
. . . .
I
. . . .
69
To illustrate the above, UF-8-145% (2) means: (a) a uniform flux run, simulating BC at 0 ~ 0 °, with all eight zones of the lower block (i.e. zones 17-24) at the same power; (b) middle and upper blocks powered according to the power shaping principle (as discussed above) and such that the power level at the upper end (90 ° ) is 145% of that at 0 = 0°; and (c) this is the second run carried out under these conditions. On the other hand, SF-7 (4) means: (a) run with a heat flux distribution according to the power shaping principle; (b) a run simulating BC in zone # 7 (0 ~ 67°); and (c) this is the fourth run carried out under these conditions. A listing of all runs carried out in Configuration II is given in Table 1. The flux shapes employed in all U F runs are shown in Fig. 11. The flux shapes employed in all SF runs are shown in relation to the shapes dictated by the power shaping principle in Figs. 12(a)-15(b). The effect of axial conduction in smoothing the input (stepwise) fluxes is shown in Fig. 12(b)-15(b), and the implied flux corrections' at the points of boiling crisis are summarized in Table 2. These corrections have been applied already to the data reported. This type of correction for all U F runs is negligible. F o r all Configuration II runs, the copper block surface was fully aged. All Configuration III runs were carried out under a reference flux shape corresponding to the base reactor case, as in Appendix A, and with the lower heater block surface in a 'fresh' condition, cleaned to a fine polish (about aging effects, see below).
I
Old Shape
4. Experimental results and discussion ~"
6oo
~"~
400
200
0
. . . .
o
I
20
. . . .
I
. . . .
40 Angle
I
60
. . . .
i
,
8o
(degrees)
Fig. 19. The presently used base case flux shape in relation to that used in the tests reported in the body of the paper.
A rough idea of the surface aging effect on C H F is in Fig. 16. In it, we see that the effect is measurable even after ~ 4 h of operation, but it is clearly leveled out, compared to the very initial value of 450 kW m -2. Also, for 0 = 9 0 ° , the initial value was ~ 1.2 M W m 2, as compared to the fully-aged value of ~ 1.6 M W m 2. As noted already, this work is focused on the fully-aged
T.G. Theofanous, S. Syri /Nuclear Engineering and Design 169 (1997) 59 76
70
Table 4 Summary of the U L P U (Configuration II) experimental program on flux shape and recirculation flow effects ID Number a
Reference b Shape
Z00A
Fig. Q. 1
Maximum Flow C H F (kW/m 2) 508 564
Medium Flow C H F (kW/m 2)
No Flow C H F (kW/m 2)
Angular Position (Actual) c (°)
545
0 0 0 0(6) 0 0 15 15 26 26 37 37 37 37 45 45 45 45 45 45 56
658 (Z00B)
498 564
(Z00C) Z08B
Fig. Q.1
ZI4A
Fig. Q. 1
Z19B
Fig. Q.I
(Z 19C) Z24A
Fig. Q. 1
602 718 770 902 902 1125 1139 1072 1072 1205 1191 1151
649 866 1005
991 1096 1083 1042
(Z24B) (Z24C) Z30A
Fig. Q.1
1179 1228 1116
1064
1104
Z35A
Fig. Q.I
Z35C (Z40A)
Fig. 7.16
Z40B (Z40C) Z42A Z48A
Fig. Q.I Fig. 7.15 Eq. (5.30)
Z48B
Eq. (5.30) with cut
1263 1263 1255 1340 1186 1318 1340 1489 1310 1523 1480 1622
56
1190
66 66 66 66 75(90) 75(80) 75 75 75(85) 78 90 90 90 90
1370
1263 1118
1353
a The number indicates the zone number at which C H F is being simulated. Total 48 equal size zones starting with zone 0 at 0 ° and ending with zone 48 at 90 ° . Numbers in parenthesis show sensitivity variations around the base case. Number in bold print. Parametric cases examined for the reactor, in DOE/ID-10460. b Flux shapes based on calculated reactor shapes. Figures and equations refer to DOE/ID-10460. Fig. Q.I is the base case simulation. The actual boiling crisis position is shown in parenthesis when it is different from the expected one.
surface condition, and these are the results reported and discussed below. The results from all Configuration II runs are in Fig. 17, and in tabular form in Table 3. In Fig. 17, we also show the results from Configuration I.
The following observations
and comments
can be
made: 1. C o n f i g u r a t i o n tolerance
to
Configuration
II exhibits a significantly higher boiling I. T h e s e
crisis
as
compared
to
are clearly the effects
T.G. Theofanous, S. S y r i / N u c l e a r Engineering and Design 169 (1997) 5 9 - 7 6
of flow and subcooling, and they amount to about 50% increase• 2. For Configuration II, the variation of CHF with angular position seems to be composed of two, remarkably linear, regions. We believe these to be reflective to the significantly different flow regimes observed in the respective regions (see discussion in the introduction section). 3. The CHF level at the upper end ( ~ 1.6 M W m - 2 ) is remarkably higher than the 'conventional wisdom value' for a horizontal, upward facing, flat plate in pool boiling, which is ~ 1 M W m 2. Clearly, flow and orientation are important. Further, it should be noted that in Table 3 we find three runs in which BC occurred rather late compared to a few minutes; however, note that all
. . . . ...........
o:
71
I ' ' ',~' I '
I . . . . . . . . Principle Input
1 j
./"
6
," - "'
" / 'i ~i
..-"" ""
;_..
.,./" F~ 4
/" /,
2
....( " , , , I
.... 20
I ....
L,,
I ....
40 60 Angle (Degrees)
80
F i g . 2 2 . The input flux shape, in comparison to that obtained from the power shaping principle for run Z00C.
I .... ............
. . . .
I
]
'
Principle Input Actually Delivered
S '
'
'
I
. . . .
'
'
4 •
!
'
I
'
171171711 Pn.cipl~ Input Actually Delivered
/
2
// /-
20
40 Angle (Degrees)
60
80
2 .
o~, 0
, , I . . . . . . . . EO 40 Angle
I .... 60
I , 80
(Degrees)
F i g . 2 0 . The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z 0 0 A .
.
.
.
.
.
nc,*,opu,
F i g . 2 3 . The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z 0 8 B .
]
.........i
S
three cases were for the lowermost region ( 0 ° < 0 < 15°), and that the differences were rather small (2-4%). This again is another indication, or symptom, of a different mechanism in the lower region identified in the data trends (Fig. 17). The lower envelope for the Configuration II data in Fig. 17 can be expressed by: qcr(0)
=
500-1- 13.30 k W m - 2
for
0 < 15 °
(1) qcr(0) = 20
40 Ang]e
60
80
(Degrees)
Fig: 21. The input flux shape, in comparison to that obtained from the power shaping principle for run Z 0 0 B .
540 + 10.70 kW m
15 ° < 0 < 90 °
2
for (2)
However, the data can be fit equally well with one equation:
T.G. Theofanous, S. Syri /Nuclear Engineering and Design 169 (1997) 59 76
72
qcr(0)
. . . .
= 4 9 0 + 3 0 . 2 0 - 8.88- 1 0 - 1 0 2 + 1.35-10 -6.65"10-504kWm
203
2
- ~ ~ ~i~i
I
'
'
....
'
I ....
t.z
(3)
This line is shown in Fig. 17. The Configuration III results are in Fig. 18. As expected, in the upper regions, the effect of a channel-like geometry, i.e., enhanced convection, is seen to be favorable, but quantitatively not very significant. At the pole (0 ~ 0°), we can see more of an effect, as convection is just about enough to compensate for the deleterious consequence of a fresh, poorly wetted, surface. The uncertainty in these results is very small compared to available margins in the case of
I '
Principle Input
ii/',,i J
;::i","i"t",
0 0
,
,
I
20
I
. . . .
40
. . . .
I
60
,
80
Angle (Degrees) Fig. 26. T h e input flux shape in c o m p a r i s o n to that obtained from the p o w e r shaping principle for run Z19B. I ....
. . . . . . . .
I '
Actually Delivered 1.5
,,•,,"• m,
,
,
.:[ ~[[[i~[[
,
I
'i
- ¸•
. . . .
lnputPrinciple
Actually Delivered
05
5
"/,
,
0.0
2
20
/'
Angle
o,,I,, 0
20
b
40
Angle
~
I
q
80
60
. . . .
o~
.............
I
' ' ' ' I
. . . .
' ' I ' ,""i!!
Principle [npul Actually Delivered
/
'i - ' "
2 •._ .-"
A /,'
I 20
60
40 Angle
,
,
t , , , , I , 60 80
(Degrees)
Fig. 27. T h e input and actually delivered flux shapes, in c o m p a r i s o n to that obtained from the p o w e r shaping principle for run Z 2 4 A .
(Degrees)
Fig. 24. T h e input and actually delivered flux shapes, in c o m p a r i s o n to that obtained from the p o w e r shaping principle for run Z 1 4 A .
3
,
40
(Degrees)
Fig. 25. T h e input and actually delivered flux shapes, in c o m p a r i s o n to that obtained from the p o w e r shaping principle for run Z19C.
AP600 (Theofanous et al., 1997) as well as in comparison to other uncertainties that normally enter such assessments. Basically, there are three reasons for this: (1) small measurement error in cartridge power ( ~ 2%); (2) negligible heat losses from the heating blocks; and (3) convenience for back-to-back runs and easy recovery from BC. As noted already, the small measurement error is confirmed by findinbg cumulative power of cartridge heaters to be within 1% of the total power supplied (independently measured). The easy recovery from BC is because of the high thermal inertia of the heater block• Back-to-back experiments allow the approach to the true BC limit by successive fine-step approximations. As in Fig. 17 and Table 3, the data show that this could be done within a few percent.
73
T.G. Theofanous, S. Syri /Nuclear Engineering and Design 169 (1997) 59-76 5. Conclusions
1.5
. . . .
I
. . . .
I . . . .
i~i~.~i~i~:~i~. InputPrinciple
With the completion of the Configuration II and Configuration III experiments reported here, we have established a firm estimate for the coolability limits of the lower head of a reactor vessel submerged in water and heated internally. The scale of the experiment, and the power shaping principle, ensure that these data, and resulting correlation, are directly applicable to the reactor. The correlation reveals a two region behavior (Eq. (1) and Eq. (2)) which has been related to the different flow regimes observed in the two-phase boundary layer in the respective regions.
. . . .
I
. . . .
I
. . . .
I
'
'
, "'~
0.5
0.0
~
.... 20
i ....
I ....
40 60 Ang]e (Degrees)
i, 80
Fig. 30. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z30A. . . . .
o~ I
I
1.0
.......
. . . .
. . . .
Actually Delive~d
1.~
20
I
]
I
. . . .
I
. . . .
I
. . . .
I
,7
Input Principle ] Aclually Delivered
1.o
'
15 ~
0.5
20
0.5
40 60 Angle (Degrees)
B0
Fig. 31. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z35A.
0.0 20
40 60 Angle (Degrees)
80
Fig. 28. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z24B. Acknowledgements 2.0
2
~
1.5
.
.
.
.
I
.
.
.
I
.
Actually Delivered
.
.
.
.
L
.
.
.
.-.,/"'"/'"/,/"
.
I
!
'
']
',........... "
0.5
0.0 20
40 Angle (Degrees)
60
80
Fig. 29. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z24C.
This work was funded by IVO International Ltd. (supporting S. Syri) and the U.S. Department of Energy's Advanced Reactor Severe Accident Program (ARSAP) through ANL subcontract No. 23572401. We are also grateful to Sergio Angelini, T o n y Salmassi, and Mitch Frey, for help with design and construction of Configuration II, and to Jeff Smith and Charlie Liu, for helping with the construction, running of the latest portion of the experiments, and helping with the data reductions. Messrs. Steven Sorrell (DOE) and Stephen Addition ( T E N E R A ) played a key role in programmatically supporting this project.
T.G. TheoJanous, S. Syri /Nuelear Engineering and Design 169 (1997) 59 76
74
Appendix A. Critical heat flux sensitivity to heat flux shapes and recirculation flow rates
t.o
og
--
A series of experiments was carried out in U L P U , to examine sensitivity of the critical heat fluxes to the heat flux shape and the natural circulation flow rates. The experimental facility was in Configuration II, as described above, except for increasing the power control resolution (by adding relays to the control circuit) to 48 zones. In the runs described here each zone corresponds to an arc length of 1.875 ° . We have also changed the nomenclature; now, the zone numbers begin at 0 ° and increase with angular position, so that zone 0 is over 0 ° < 0 < 1.875 ° and zone 48 over 88.125 ° < 0 < 9 0 ° .
]"'-
. . . .
- - - -
'
'
'
I
....
I
'
'
Z '
I
Principle Input ActuallyDelivered
0.6 I
04
02
20
40
60
80
Angle (Degrees)
Fig. 34. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z40B. I0
[.'po, P' 08
....
[0
~
0.8
2
.... ............
''
Principle
,~ . . . .
Inpu~ Actuay De vered
~ , , - ~ : ~
[
0.6
"
J
0.4
0.6 02 04
0.0 3
0.;~
0
20
40
60
Fig. 32. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z35C. ,
,
.....
•0 .
, -~-
_J _. . . ._
I ....
I'
Principle Pr Input
' . 2 ~ - , ~ . / , . ~ -I ,
,
;Z
08
0.4
0,2
20
60
40
Angle
go
(Degrees)
Fig. 35. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z40C.
80
(Degrees)
Angle
60
40 Angle
0. 0
1.0
20
80
(Degrees)
Fig. 33. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z40A.
Flux shape sensitivity is examined using, besides the reference flux shape shown in Fig. 19 (note that this is somewhat different from the shape used in the test series described a b o v e 'old shape'), several variations around it, as well as shapes associated with the parametric cases considered for the reactor (in DOE/ID-10460). Flow rate sensitivity studies were carried out by throttling the d o w n c o m e r so as to obtain about 60 gpm, or about half the flow normally obtained ( ~ 130 gpm) with the flow path completely open. In addition, the case o f zero flow was examined by letting the level drop sufficiently so as to interrupt the natural circulation flow path. The experimental matrix is summarized in Table 4, together with the resulting C H F values.
T.G. Theofanous, S. Syri / Nuclear Engineering and Design 169 (1997) 59-76
The second column indicates the origin of the reference flux shape used in each case. The actual flux shapes used are shown, together with the shapes that would be required according to the power shaping principle, in Figs. 20-37. The results are shown in graphical form in Fig. 38 and Fig. 39. Fig. 38 contains the comparison between all present full-flow data with the CHF correlation, Eq. (3). This figure shows that the correlation represents a robust lower envelope even under some not insignificant variations of the heat flux shapes. But this figure also shows that there is a significant additional tolerance to CHF under a strongly peaked heat
t.0
. . . .
I
. . . .
I
. . . .
I ' '
'
~; ',[~
i,'j
P~putpL
20
40
60
80
(Degrees)
Angle
Fig. 36. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z42A. 1.0
. . . .
I
. . . .
I
. . . .
I
I 'j
. . . .
2000
~
1600
~
1200
o~
.>/: / Y
800
400
I
,
,
,
I
20
,
,
,
40
I
,
,
,
60
E
,
,
80
Angle (degrees)
100
Fig. 38. Comparison of all present full flow data with the correlation, Eq. (3). The open circles are for a flux representative of the extreme parametric (discontinuous, high peak at the top).
flux at the top. The sensitivity to the flow rate can be seen in Fig. 39. It is very interesting that the zero flow case produces a negative effect only near the top and even then, it is only a 20% reduction. Note that the zero flow case in effect eliminates the cavity subcooling as it is operated under a significantly lower water level ( ~ 2.0 vs. 5.6 m from the bottom of the lower head). These results indicate that simply immersing the lower head of the AP600 would be sufficient to remove the thermal loads from a base-case-like scenario with very significant margins.
2.0
J
IPlput pie
75
,
0.8 1.6
• [:3 ×
130 G P M 60 GPM 0 GPM
0.6
: ~
1.2
Q
04
0.8
| o
02
0.4
0.0 20
40 Angle
6O
80
(Degrees)
Fig. 37. The input and actually delivered flux shapes, in comparison to that obtained from the power shaping principle for run Z48A.
,
I
20
,
,
,
~
40
,
,
,
i
60
Angle (degrees)
,
,
,
I
80
,
,
,
100
Fig. 39. Sensitivity of measured C H F on natural circulation flow rate.
76
T.G. TheofiTnous, S. Syri /Nuclear Engineering and Design 169 (1997) 59-76
References Lienhard, J.H., Snares of Pool Boiling Research: Putting Our History to Use, in G.F. Hewitt, (ed.), 1994, Proceedings 10th International Heat Transfer Conference, Brighton, UK. Vol. 1, Keynote Papers, Gk-ll, Chameleon Press, London, pp. 333 348. Theofanous, T.G., Liu, C., Angelini, S., Kym/ilS.inen, O., Tuomisto, H. and Addition, S., 1994a, Experience from the First Two Integrated Approaches to In-Vessel Retention Through External Cooling, OECD/CSNI/NEA Workshop
on Large Molten Pool Heat Transfer, Grenoble, France, March 9-11. Theofanous, T.G., Syri, S., Salmassi, T., Kym/ilfiinen, O. and Harri Tuomisto, 1994b, Critical heat flux through curved, downward facing, thick walls, Nucl. Eng. Des., 15, 247258. Theofanous, T.G., Liu, C., Additon, S., Angelini, S., Kymfil~iinen, O. and Salmassi, T., In-vessel coolability and retention of a core melt, Nucl. Eng. Des., (this issue). Angelini, S. and Theofanous, T.G., 1996, A New Boiling Transition Regime, Proceedings, The Japan-US Seminar on Two-Phase Flow Dynamics, July 15-20, Fukuoka, Japan.