ML20082Q635
| ML20082Q635 | |
| Person / Time | |
|---|---|
| Site: | 05200003 |
| Issue date: | 03/31/1995 |
| From: | WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML20046D692 | List: |
| References | |
| WCAP-14327, NUDOCS 9504280329 | |
| Download: ML20082Q635 (88) | |
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t i l Cane <moost NowPaorma:nay Cum 3 I WCAP-14327 - l I 1 ; i l l ! EXPERIMENTAL BASIS FOR THE , I AP600 CONTAINMENT VESSEL HEAT ; AND MASS TRANSFER CORRELATIONS , t t i i March 31,1995 i l i l C 1995 Westinghouse Electric Corporation l All Rights Reserved l l u:\np60th!865-nan:1W1095
TABLE OF CONTENTS Section Title Page !
SUMMARY
I 1.0 Introduction 1-1 2.0 Heat and Mass Transfer Correlations 2-1 2.1 Annulus Region 2-2 , 2.2 Entrance Effects 2-4 2.3 Inside Containment 2-5 l 3.0 Experimental Basis for the Heat and Mass Transfer Correlations 3-1 3.1 ne Hugot Tests 3-1 3.2 De Eckert and Diaguila Tests 3-9 3.3 The Siegel and Norris Tests 3-22 3.4 De Westinghouse S'IU Dry Flat Plate Tests 3-33 3.5 ne Westinghouse Large-Scale Tests - Dry External Heat Transfer 3-36 i 3.6 The Gilliland and Sherwood Evaporation Tests 3-40 3.7 De Westinghouse STC Flat Plate Evaporation Tests 3-46 3.8 The University of Wisconsin Condensation Tests 3-51 3.9 ne Westinghouse Large-Scale Tests - Internal Condensation 3-61 4.0 Assessment of Results and Statistics 4-1 4.1 Convection 4-1 4.2 Evaporation 4-4 I 4.3 Condensation 4-8 4.4 Measurement Uncertainty 4-12 5.0 Conclusions 5-1 6.0 References 6-1 i J l l uwis65-non:ib.o4tws iii
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LIST OF TABLES Table Title _Page 3.1-1 . Hugot Test Data 3-2 3.2-1 Eckert & Diaguila Test Data 3-9 3.3-1 Siegel & Norris Test Data 3-22 3.4-1 Westinghouse STC Dry Flat Plate Test Parameters 3-33 3.5-1 Westinghouse Large-Scale PCS Tests - Dry External Heat Transfer 3-38 3.6-1 Gilliland & Sherwood Test Parameters 3-41 3.7-1 Westinghouse STC Flat Plate Test Parameters 3-46 3.8-1 Wisconsin Condensation Test Parameters 3-52 .- 3.9-1 Westinghouse Large-Scale PCS Tests - Internal Condensation 3-62 { ovoas65..on:imims iv
LIST OF FIGURES Figure Title h 3.1-1 Local Nusselt Number Comparison for Hugot Test 1 3-3 3.1-2 Local Nusselt Number Comparison for Hugot Test 2 3-4 3.1-3 Local Nusselt Number Comparison for Hugot Test 3 3-5 3.1-4 Local Nusselt Number Comparison for Hugot Test 4 3-6 , 3.1-5 Local Nusselt Number Comparison for Hugot Test 5 3-7 l 3.1-6 Comparison of Predicted-to-Measured Local Nusselt Numbers for Hugot Tests 3-8 3.2-1 Local Nusselt Number Comparison for Eckert and Diaguila Test 1 3-11 3.2-2 Local Nusselt Number Comparison for Eckert and Diaguila Test 2 3-12 , 3.2-3 Local Nusselt Number Comparison for Eckert and Diaguila Test 3 3-13 l 3.2-4 Local Nusselt Number Comparison for Eckert and Diaguila Test 4 3-14 3.2-5 Local Nusselt Number Comparison for Eckert and Diaguila Test 5 3-15 3.2-6 Local Nusselt Number Comparison for Eckert and Diaguila Test 6 3-16 3.2-7 Local Nusselt Number Comparison for Eckert and Diaguila Test 7 3-17 3.2-8 Local Nusselt Number Comparison for Eckert and Diaguila Test 8 3-18 3.2-9 Local Nusselt Number Comparison for Eckert and Diaguila Test 9 3-19 3.2-10 Local Nusselt Number Comparison for Eckert and Diaguila Test 10 3-20 3.2-11 Comparison of Predicted-to-Measured Local Nusselt Numbers for the Eckert and Diaguila Tests 3-21 3.3-1 Local Nusselt Number Comparison for Siegel and Norris Test 1 3-24 3.3-2 Local Nusselt Number Comparison for Siegel and Norris Test 2 3-25 3.3-3 Local Nusselt Number Comparison for Siegel and Norris Test 3 3-26 3.3-4 Local Nusselt Number Comparison for Siegel and Norris Test 4 3-27 3.3-5 Local Nusselt Number Comparison for Siegel and Norris Test 5 3-28 3.3-6 Local Nusselt Number Comparison for Siegel and Norris Test 6 3-29 3.3-7 Local Nusselt Number Comparison for Siegel and Norris Test 7 3-30 3.3-8 Local Nusselt Number Comparison for Siegel and Norris Test 8 3-31 3.3-9 Comparison of Predicted-to-Measured Local Nusselt Numbers for the Siegel and Norris Tests 3-32 l 3.4-1 Comparison of Predicted-to-Measured Local Nusselt Numbers for the STC Dry Flat Plate Tests 3-34 3.4-2 Correlated Heat Transfer Data for the STC Dry Flat Plate Tests 3-35 3.5-1 Comparison of Predicted-to-Measured Local Nusselt Numbers for the Westinghouse Large-Scale PCS Tests 3-39 3.6-1 Calculated Steam Partial Pressure vs. Channel Length from a Selected Gilliland and Sherwood Evaporation Test 3-44 3.6-2 Comparison of Predicted-to-Measured Evaporation Rates for the Gilliland and Sherwood Evaporation Tests 3-45 n:\np6(Xh1865-non:lt>o41(95 y
LIST OF FIGURES (Continued) Figure Title Pace 3.7-1 Bulk-to-Film Steam Partial Pressure Differences vs. Channel Length from Selected Westinghouse STC FLt Plate Evaporation Tests 3-48 3.7-2 Comparison of Predicted-to-Measured Sherwood Numbers for the SK Flat Plate Evaporation Tests 3-49 3.7-3 Correlated Mass Transfer Data for the STC Wet Flat Plate Tests 3-50 3.8-1 Bulk-to-Film Steam Partial Pressure Differences vs. Channel Length from Selected Wisconsin Condensation Tests 3-55 3.8-2 Comparison of Predicted-to-Measured Sherwood Numbers for the Wisconsin Condensation Tests 3-56 3.8-3 Comparison of Predicted-to-Measured Sherwood Numbers for the Wisconsin Condensation Tests 3-57 3.8 4 Comparison of Predicted-to-Measured Sherwood Numbers for the Wisconsin Condensation Tests 3-58 3.8-5 Comparison of Predicted-to-Measured Sherwood Numbers for the Wisconsin Condensation Tests 3-59 3.8-6 Correlated Mass Transfer Data for the Wisconsin Condensation Tests 3-60 3.9-1 Comparison of Predicted-to Measured Local Shenvood Numbers for the Westinghouse Large-Scale PCS Tests 3-63 3.9-2 Comparison of Predicted-to-Measured Local Sherwood Numbers for the Westinghouse Large-Scale PCS Tests 3-64 3.9-3 Comparison of Predicted-to-Measured Local Sherwood Numbers for the Westinghouse Large-Scale PCS Tests 3-65 3.9-4 Correlated Condensation Mass Transfer Data for the Westinghouse Large-Scale PCS Tests 3-66 4.1-1 Predicted-to-Measured Nusselt Numbers for Convection as a Function of the Reynolds Number 4-2 4.1-2 Predicted-to-Measured Nusselt Numbers for Convection as a Function of the Grashof Number 4-3 4.2-1 Predicted-to-Measured Sherwood Nun bers for Evaporation as a Function of the Reynolds Number 4-5 4.2-2 Predicted-to-Measured Sherwood Numbers for Evaporation as a Function of the Grashof Number 4-6 4.2-3 Predicted-to-Measured Sherwood Numbers for Evaporation as a Function of the Dimensionless Pressure 4-7 4.3-1 Predicted-to-Measured Sherwood Numbers for Condensation as a Function of the Reynolds Number 4-9 4.3-2 Predicted-to-Measured Sherwood Numbers for Evaporation as a Function of the Grashof Number 4-10 4.3-3 Predicted-to-Measured Sherwood Numbers for Condensation as a Function of the Dimensionless Number 4-11 u:kp60LA1865 mon:lt>-G41095 vi
l
SUMMARY
The AP600 plant design utilizes passive safety systems that rely on natural forces, such as circulation, convection, evaporation, and gravity. His unique design includes a passive containment cooling system (PCS) to remove heat released to the containment following any postulated event and to transfer this heat from the containment to the environment. His system employs natural draft air cooling and the evaporation of a water film from the outside of the steel containment shell to transfer heat from the containment to the environment. Section 1.0 of this report describes the heat and mass transfer phenomena that are required to be modeled for the AP600 containment analyses, and Section 2.0 describes the heat and mass transfer correlations selected for modeling heat removal from the AP600 containment. The Sherwood mass l transfer correlation is used for modeling condensation on the inside and evaporation on the outside of the AP600 containment vessel. The Sherwood mass transfer correlation is derived by dimensional analysis using the Reynolds analogy and Colburn j factors and is, therefore, dependent on the Nusselt number, which is derived from the heat transfer correlation. After examining the literature, conventional free convection (McAdams) and forced convection (Colburn for channels and Schlichting for flat plates) heat transfer correlations were selected to be l used in the analytical model. An entrance effect multiplier, which is important for predicting heat j transfer in short channels, is applied to the Colburn forced convection heat transfer correlation. An approximate method recommended by Churchill was implemented to combine the free and forced convection correlations for the mixed convection regime. A lower limit on the mixed convection correlation for assisting free and forced flows was selected based on work by Eckert and Diaguila. We result is a single heat transfer correlation that gives free convection values at low Reynolds l numbers, forced convection values at low Grashof numbers, and a combination of the two in mixed I convection. Section 3.0 presents comparisons of the heat and mass transfer predicted by the analytical model with j test data from various sources. The method used to calculate heat and mass transfer on the AP600 containment vessel produces good agreement with the available data. Predicted-to-measured test data comparisons are provided that cover a range of natural convection induced Reynolds numbers up to 3.8x10' and Grashof numbers (based on channel diameter) up to 7.2x10' . The AP600 riser Reynolds number ranges up to lx10' and the Grashof numbers (based on channel diameter) range up to 4x10'. Therefore,'the test data that was used to validate the analytical model covers the range of anticipated conditions in the plant. owums65.noowmo95 1
9 1 1 I I 1
1.0 INTRODUCTION
A design basis accident (DBA) in the AP600, such as a loss-of-coolant accident (LOCA), will pressurize the containment shell and potentially challenge the shell design limit. The AP600 passive containment cooling system (PCS) is designed to remove sufficient heat from containment during the limiting DBA to maintain containment pressure below the design limit. Heat is removed from the containment atmosphere by condensation and convective heat transfer to the shell, where it is conducted through the shell and rejected to the atmosphere on the outside of containment. Rejection to the atmosphere is by convection to the buoyant cooling air, radiation to the baffle, and evaporation of the external cooling film to the cooling air. A key aspect of modeling heat removal from the containment is the convective heat transfer correlation. The convective heat transfer correlation and the heat and mass transfer analogy provide the basis for calculating the magnitude of condensation and evaporation, which are the dominant heat transfer mechanisms. The air flow through the annulus region between the baffle and the outside of the containment shell is buoyancy-induced and has a relatively high velocity during a DBA event. The flow along the inside of the containment shell is buoyancy-induced after blowdown. The heat and mass transfer on both the inside and outside of the containment shell are strongly influenced by significant natural-circulation-driven air flows. Consequently, the heat and mass transfer are generally qualified as mixed, turbulent free and forced convection. Because the scale of the AP600 containment heat transfer surface is approximately 100-ft high, over 95 percent of the surface operates in turbulent convection. The heat transfer from the small portion of the inner surface that operates with laminar flow is underpredicted by the turbulent convection heat transfer correlation. 'Iherefore, special laminar flow heat transfer correlations are not included in the analytical model. The containment response to the DBA event is calculated using the Westinghouse version of the GOTHIC code, WGOTHIC. This report describes the heat and mass transfer correlations selected to model condensation and evaporation on the AP600 containment and provides comparisons with experimental data to justify the use of these correlations in the .W_ GOTHIC code. ; l 1 l 1 unap600\1ses-non.ib-041095 1-1
,9 g.
N l l l
2.0 HEAT AND MASS TRANSFER CORRELATIONS Heat transfer is driven by a temperature gradient, and mass transfer is driven by a concentration gradient. Given a mixture of gases A and B, where B is noncondensable and A is transferred to (or from) the bulk gas from (or to) the liquid film, the mass transfer equation for gas A is shown below: G = koM,(py - p,o) (1) where: G = condensing or evaporating mass flux (Ibm /hr-ft.2) ko = mass transfer coefficient (Ibm-mole /hr-ft.2-psi) M4 = molecular weight of gas A (lbm/lbm-mole) pu = partial pressure of gas A at the interface (psi) pAa = partial pressure of gas A in the bulk gas mixture (psi) j He mass transfer coefficient, ko, can be predicted using empirical correlations similar to the heat transfer coefficient, hc. The Sherwood number for mass transfer is analogous to the Nusselt number for heat transfer. De Sherwood number for gas-phase mass transfer is shown below. l 8 Sh = (2) D,P l where: R = universal gas constant (ft.8-psi /R-Ibm-mole) T = boundary layer temperature (Tm + Ta)/2 (R) pm = log mean partial pressure (paa - pij/In[(P - py/(P - p30)] (psi) L = characteristic length (ft.) D, = diffusion coefficient (ft.2/hr) P = total pressure (psi) An empirical correlation for the Sherwood number, which is derived by dimensional analysis using the Reynolds analogy and Colburn j factors for heat and mass transfer, is shown below. Nu Sh = (3) (Pr/Sc)* where: Nu = Nusselt number based on the heat transfer correlation evaluated at the boundary layer temperature Pr = Prandtl number evaluated at the boundary layer temperature Sc = Schmidt number evaluated at the boundary layer temperature uAap600\l865-mon:1t>ot1095 21
This correlation is used in the analytical model to calculate both condensation and evapomtion mass transfer. The heat transfer correlations used in the analytical model to calculate the Nusselt number will be described in the following subsections. 2.1 Annulus Region Heat is rejected from the containment shcIl to the air within the annulus region gap by radiation, convection, and evaporation from an applied liquid film. As described in Section 2.0, the evaporation mass transfer in the annulus is directly proportional to the convective heat transfer through the Nusselt number. This section presents the correlations for modeling convective heat transfer within the annulus region of the AP600 containment. As described in Section 1.0, the convective heat transfer in the AP600 will primarily be turbulent rather than laminar. Heat transfer from the small fraction of area in the laminar heat transfer regime will be underpredicted using the turbulent convection heat transfer correlations and therefore, a laminar I heat transfer correlation has not been modeled. 'Ihe flow regime for turbulent convective heat transfer is typically qualified as either free, forced, or mixed. The combination of free and forced convection in the mixed regime is either assisting (i.e., they work in the same direction, as in upward flow in a hot pipe) or opposed (i.e., they work against each other, as in downward flow in a hot pipe). Based on a review of the literature, the free convection heat transfer correlation for gas mixtures has the form Nu = C (GrPr)", with the value of C varying between 0.09 and 0.15 and the value of N varying between 0.3 and 0.4. The AP600 is expected to operate in either the transition (10' < Gr < 10"') or turbulent (Gr > 10*) free convection range. Data in the transition and turbulent free convection range are best fit using a value of 1/3 for N. A value of 0.13 for C seems to best fit the large-scale PCS data. The McAdams* correlation, shown below, has been selected for calculating turbulent free convection heat transfer in the annulus. Nu g = 0.13(Gr,Pr)'8 (4) This correlation assumes that in turbulent free convection channel flow the local heat transfer coefficient is independent of distance from the leading edge. This correlation is widely used to calculate turbulent free convection heat transfer from both vertical and inclined heated flat plates with both constant temperature and constant heat flux boundary conditions. Since this correlation is used for heat transfer within the annulus, the annulus hydraulic diameter is used as the characteristic length in the Grashof and Nusselt numbers. The work of Vlict* shows that Equation 4 underpredicts the heat transfer from a horizontal flat plate. Even though it may slightly underpredict heat transfer from the small fraction of the dome that can be I u Aap60tM 865-nowlt>-481095 2-2 l l
considered a horizontal surface, Equation 4 will be used to conservatively calculate turbulent free convection heat transfer for the entire shell surface. The Colburn* correlation, shown below, has been selected for calculating turbulent forced convection heat transfer in the annulus. i (5) Nu,, = 0.023Rel# rPn his correlation is applicable to both constant temperature and constant heat flux boundary conditions for fully developed flow in long ducts. His correlation is widely used to calculate turbulent forced convection heat transfer in long tubes and ducts. He annulus hydraulic diameter is used as the characteristic length in the Reynolds and Nusselt numbers. A length or distance dependent multiplier is required to account for the increase in forced convection heat transfer as the boundary layer develops at the entrance of a heated channel. This entrance effect multiplier is described in more detail in Section 2.2. For calculational purposes, a single correlation (or combination of free and forced convection correlations)is needed to cover the entire range of mixed convection. A method for calculating mixed free and forced convection heat transfer was recommended by Churchill
- and is given below. For opposed free and forced convection:
Nu, = (Nul,,+Nu,[,) o (6) and for assisting free and forced convection, Nu, is the larger of the following three expressions: abs [(Nul,,- Nu,[) ; Nu,,, ; 0.75Nu,, (7) he lower limit in the latter equation, which prevents the value of Nu, from going to zero when Nuy,, and Nu,r are equal, comes from Eckert and Diaguila.W Re method for calculating mixed convection heat transfer is asymptotic to both the individual free and forced convection correlations. Consequently,it is unnecessary to a priori choose whether the heat transfer regime is free, forced, or mixed in the analytical model. unap600\1865 con:Ib 041095 2-3
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2.2 Entrance Effects he measured heat transfer coefficient at the entrance to a heated channel or plate is significantly higher than predicted by the Colburn forced convection heat transfer correlation. 'The increase in heat transfer at the entrance is attributed to the development of the boundary layer. The entrance effect is important in modeling heat transfer in short channels or plates (e.g., some test assemblies), but is relatively unimportant for modeling heat transfer from the AP600 containment vessel due to its much larger scale. The correlation and coefficients recommended by Boelter, Young, and Iverson* are used to account for the entrance effect: h" = 1 + F _d (8) h. 8 L where:
- h. = the heat transfer coefficient calculated from the Colburn correlation based on diameter d h, = the mean or length average heat transfer coefficient over length L Fi = a geometry-dependent constant multiplier from Reference 11 An equation is needed that will give a length average heat transfer coefficient between x, and x2-Given an equation for h(x), the average value of h on the interval (x i,x2) is:
1
= (9) h,,> 'h(x)dx x2 -x ,
Analytically, l'1,g could be derived from the above definition over the interval (0, L), but the equation produces a singularity when this is attempted. A modest change to the exponent, however, results in: h, ,,,, _ d (x 2 -X ) (10)
- h. . p* L 3(x2 -Xi
)
a form that has the same average over length L, but with slightly lower values for small values of x, and with slightly higher values for higher values of x. Herefore, the calculated forced convection heat transfer coefficient multiplier input value is dependent on the WGOTHIC model noding structure. u Aar60.lSh5-non:ltwo41095 24
2.3 Inside Containment Heat is transferred from the containment atmosphere to the containment inner shell surface by condensation and convection. Condensation and convective heat transfer take place at die outer surface of a thin liquid film that develops on the inside surface of the containment vessel. The liquid film provides a relatively small resistance to heat transfer from the containment atmosphere to the wall. The containment shell will be heated as air and steam flow along it. Although the bulk of the flow is expected to be downward along the inner shell, some of the flow may be upward and consequently, the local mixed convection heat transfer regime can be either assisting or opposing. The modeling of convective heat transfer to the inner containment shell requires a correlation for vertical and inclined plates in an open geometry. 1 The McAdams correlation, described in Section 2.1, is used for calculating turbulent free convection heat transfer inside containment with the Grashof number length parameter based on the heated length instead of the channel hydraulic diameter. The flat plate correlation,shown below, has been selected for calculating turbulent forced convection heat transfer inside containment. Nu, = 0.0296Re,% 28 (II)
'Ihis correlation is applicable to an open geometry, therefore, the Re, and Nu, numbers are dependent on the heated length and not the channel hydraulic diameter.
The Churchill method for calculating mixed convection heat transfer in the annulus, which is described in Section 2.1, is also used to calculate mixed convection heat transfer inside containment. I l unap600\l865-nan:Ib-o41095 2-5
l I l 3.0 EXPERIMENTAL BASIS FOR THE HEAT AND MASS TRANSFER CORRELATIONS The heat transfer correlations that have been coded into WGOTHIC have been validated by comparison to test results from various sources. He experimental basis for the method of combining the free and forced convection correlations to model turbulent mixed convective heat transfer and applying this mixed convection heat transfer correlation to modeling heat transfer in the AP600 containment vessel is covered in the first five subsections that follow. Since the calculated local mixed convection Nusselt number has a nonlinear dependence on the Reynolds and Grashof numbers, it is necessary to present a comparison of the predicted-to-measured Nusselt numbers as a function of I the dimensionless height for each of these tests. l The mass transfer correlation that has been coded into WGOTHIC has been validated by comparison to condensation and evaporation test results from various sources. The experimental basis for applying this correlation to modeling mass transfer to and from the AP600 containment vessel is covered in the last four subsections that follow. Since the total evaporation or condensation mass transfer rates were measured in the separate effects tests, comparisons are " local" to the extent of the size of the test apparatus as it represents a portion of the AP600 containment shell. 3.1 The Hugot Testsm Hugot conducted heat transfer tests on a set of heated, parallel, vertical, isothermal plates with closed sides. The channel height was 3.3 m, and the plate separation distance was variable between 5 and 60 cm. The plate temperatures were varied between 40 and 160*C. The tests provide data for validating the assisting mixed convection heat transfer mode for moderate Reynolds and Grashof numbers. Hugot did not report the air flow rate or velocity induced in the heated channel; therefore, it was necessary to use a computer model that could calculate air flow rates as well as heat transfer. De tests were modeled using the WGOTHIC code. The test section was divided into 11 axial volumes. Because most of the rapid changes occur at the entrance, the first 10 volumes were cach 1/15th of the total volume; the last volume was 1/3 of the total volume. He code calculated the velocity, air temperature, and heat transfer coefficient in each of the 11 volumes. De WGOTHIC calculations assumed a combined entrance and exit form loss of 1.5. Consequently, the WGOTHIC heat transfer calculation includes uncenainties on the air flow rate. De local heat transfer coefficient was calculated using the method for combining the free and forced convection correlations described in Section 2.1. Hugot reported heat transfer coefficients based on the (T,- Tu,,) temperatme difference. Consequently, the .WGOTHIC predictions, which are based on the (T, - Tm) temperature difference, were transformed to a basis comparable to the Hugot data for presentation. uAmp60mits65-non:Ib-041o+5 3-1
=
The calculated local Nusselt number results for each of the five tests are compared with the test data and are shown as a function of dimensionless height in Figures 3.1-1 through 3.15. Some relevant test parameters are shown in the Table 3.1-1. TABLE 3.1 1 HUGOT TEST DATA Test Number IJD. Plate Temp. ('C) Gr, Range Re, 1 4.4 68.0 2.40E09 - 2.61E09 35400 2 4.4 160.9 3.31E09 - 3.65E09 42400 3 18.15 172.5 3.30E07 - 4.70E07 12900 4 18.15 101.5 3.25E07 - 4.45E07 12200 5 18.15 72.9 2.71E07 - 3.65E07 11000 A compilation of the predicted-to-measured local Nusselt numbers for all five tests is shown in Figure 3.1-6. The mean predicted-to-measured value of 1.179 is also shown. De standard deviation of the predicted-to-measured values for all five tests is 0.429. Both the mean and standard deviation are strongly affected by the relatively large predicted-to-measured local Nusselt number ratios at the channel entrance. If these entrance values are removed, the mean falls to 1.095 and the standard deviation falls to 0.213. Except for the channel entrance, the predicted local Nusselt numbers are very close to the measured values for tests 1 and 2. These two tests had the highest Re, numbers of the set and were performed with the gap width set to 60 cm. De entrance effect multiplier for the calculated forced convection heat transfer coefficient is height dependent and can have a large value when volumes with small elevation differences are modeled at the entrance. De difference between the calculated and measured local Nusselt numbers near the entrance is due to the relatively large entrance effect multiplier on the calculated forced convection heat transfer coefficient. Ec predicted local Nusselt numbers are slightly higher than measured for tests 3 and 4. The trend in the rate of change for the measured local Nusselt numbers was not predicted by the code. The rate of change in the measured local Nusselt number data for tests 3 and 4 begins to level off around an x/d value of 7, then increases rapidly between the x/d values of 8 and 12, and finally returns to the original rate of change. These phenomena were not observed in any of the other tests. Tests 3 and 4 were performed at relatively high temperatures with the gap width set to 10 cm. De predicted local Nusselt numbers are lower than the measured values for test 5. Although the gap width is the same as tests 3 and 4, the trend in the rate of change of the local Nusselt numbers was not the same. This test was performed at a relatively low temperature. unap60ml865-nan:lt411095 3-2 l 1
700 600 - 500 - , b 400 -
- z S
- 300 - N 8
n 200 - g w 100 - g R 0 ' t t , 0 1 2 3 4 $ Dimensionless Height CX/d)
- Test Data + W CorreIatIon i
Figure 3.1-1 Local Nusselt Number Comparison for Hugot Test 1 a:W1865-nan:It>o41095 3-3 l 1
=
900 800 - 700 - g 500 - Z 500 - S 400 - 0 o J 300 - 200 - g 100 - m 0 O 1 2 3 4 5 Dimensionless Height (X/d) x Test Data + W Correlation l I Figure 3.1-2 Local Nusselt Number Comparison for Hugot Test 2 a:W1865-con ItMH1095 34
600 500 - g A00 - y 0 B a 3 300 - z w c o *
> 200 -
w ww m x 100 - n W D O 5 10 15 20 Dimensionless Height (X/d) x Test Data + W Correlation 1 l i Figure 3.1-3 Local Nusselt Number Comparison for Hugot Test 3 uwaxn1865-ooo:idaims 3-5 _ _.
- - . . . . . . ~
I l 600 500 - W r M , g 400 - o
- 2 3 300 -
2 o o
- J 200 - y l
M R E - r i 100 - W w 0 i O 5 10 15 20 DimensIoniess Height (X/d)
- Test Data + W Correlation i
Figure 3.14 Local Nusselt Number Comparison for Hugot Test 4 j n:W1865-ace:1b-041095 3-6
1 t i I I 600 ; r
?
m I 500 - l M ? W W , t. 400 - m j o Z I w 3 300 - I z w ; c
- O i o m
- J 200 -
, [r E !
- I 100 -
O O 5 10 15 20 f Dimensionless Height (X/d) m Test Data + W Correlation l l h I i Figure 3.15 Local Nusselt Number Comparison for Hugot Test 5 i nAap600u865-noscit441095 3-7
J l t
)
i 4 I h 3.5 - ' w z
- j 3 -
c o O 2.5 - u GI ! b 8 - m t e 8 s 5 S 1.s - y y u [
*
- x x E * "
S v wY ^^ "
- m y * = -
= = :
aar,eEihygg [ 1 - **** a m a ** *
- N m
w w n y y y
' i t 0.5 s 0 I 5 10 15 20 Dimensionless Length (X/d)
- Data Points Mean Value C1 179) i t
l t f s i i Figure 3.1-6 Comparison of Predicted to Measured Local Nusselt Numbers for Hugot Tests a \ uAmp60(A1865-nan:Ibol1095 3-8
3.2 The Eckert and Diaguila Tests"' i Eckert and Diaguila conducted heat transfer tests on a vertical tube that was 13.5-ft. high with a f 23.25-in. inside diameter. Inlet and outlet air pipes and dense screens were located at each end. A 10-ft. steam jacket supplied steam with a few degrees of superheat as the heat source. Sixteen condensation chambers collected and piped condensate to a station where the flow . ate was measured and the local heat flux determined. An air flow at approximately 80'F at pressures from 1 atmosphere to 99 psia was forced through the test section. Tests were conducted with forced flow in both the ; upward (assisting mixed convection) and downward (opposed mixed convection) direction. Thermocouples at the tube center and in the tube wall provided a temperature difference from which f the local heat transfer coefficient could be determined. The tests provided heat transfer measurements to validate the mixed convection heat transfer correlation at prototypic Reynolds and Grashof numbers. 'Ihe calculated local Nusselt number results for each of the 10 assisting convection tests are compared with the measured data and are shown as a function of the dimensionless height in Figures 3.2-1 through 3.2-10. Some relevant test parameters are shown in Table 3.2-1. l TABLE 3.21 ECKERT & DIAGUILA TEST DATA f Test Number Gr,Pr Range Re, . I 1 6.9E09 - 1.lE10 377000 l 2 6.9E09 - 1.lE10 180000 3 6.9E09 - 1.4E10 100000 4 7.5E09 - 1.6E10 36000 5 1.4E10 - 1.8E10 231000 6 1.3E10 - 2.5E10 134000 7 1.4E10 - 3.7E10 55000 8 3.5E10 - 5.lE10 314000 9 3.5E10 - 5.5E10 246000 10 3.4E10 - 7.2E10 77000 u:%p60LAl665-non:1 bot 1095 39
A compilation of the predicted-to-measured local Nusselt numbers for all 10 tests is shown in Figure 3.2-11. De average predicted-to-measured value at each location and mean value over all locations are also shown. The mean value is 1.028 with a standard deviation of 0.272. The Eckert and Diaguila data showed large, unexplained variations in the original report; thus, the standard deviation reported here is not excessive. However, the good agreement with the mean indicates that the significant trends are well represented by the correlation. De calculated local Nusselt numbers are about equal to or slightly higher than the measured values for cases with lower Reynolds numbers (tests 4,7, and 10). The calculated local Nusselt numbers decrease in comparison with the measured values as the Reynolds number is increased. The apparent trend of the Eckert and Diaguila data with the Reynolds number may be due to the fact that the measured centerline temperature is not the same as the bulk temperature, i.e., the difference between the bulk and centerline temperatures change as the flow develops away from the entrance. He data were scaled from figures in the referenced paper and this process may have also introduced some of the scatter, unap6(n1E65-nan:Ib-041oos 3-10
l i l 3500.0 R y M 3000.0 - b 2500.0 - Z M
+a M
y 2000.0 - M M w 2 , h 1500.0 - M O J M 1000.0 - M i 500.0 O.0 1.0 2.0 3.0 4.0 5.0 6.0 Dirnensionless Height (X/D) ,
-1 M Test Data _ w correlation .i
)
1 i i i l l Figure 3.2-1 Local Nusselt Number Comparison for Eckert and Diaguila Test I uAnp6(XA1865-noo:lho41(95 3-11
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2000.0 n w m 1500.0 - n b n 2 w w n
.a g 1000.0 -
w
$ n M z
E 8 500.0 - w E 0.0 O.0 1.0 2.0 3.0 4.0 5.0 8.0 Dimensionless Height (X/D) x Test Data _e_W Correlation i Figure 3.2-2 Local Nusselt Number Comparison for Eckert and Diaguila Test 2 u Anp6(XA1865-oon:Ib.041095 3-12 i
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)
1 g4 i i 1400.o 1200.D - 1000.0 - g a f 2 800.0 - g , , g e m m j 600.0 - m 400.0 - n 200.0 - 1 0.0 ' ' t , , C.0 1.0 2.0 3.0 4D 5.0 8.0 Dimensionless Height (X/0) m Test Data + w correlation l Figure 3.2-3 Local Nusselt Number Comparison for Eckert and Diaguila Test 3 ( i u%e1865-nan:1bosim5 3-13
=
1400.0 1200.0 - K g 1000.0 - N Y Z p s00.0 - to R j 600.0 - C W 8 400.0 - M J m 200.0 - M M N 0.0 O.0 1.0 2.0 3.0 4.0 5.0 6.0 l DirnensIoniess Height CX/ D) l M Test Data + W CorreIatIon l l Figure 3.2-4 Local Nusselt Number Comparison for Eckert and Diaguila Test 4 i u:waniss5-oon:imims 3-14 {
2500.0 K K , M M 2000.0 - M N h M Z 1500,0 - R 6
+J M E #
m N Y 1000.0 - E
- 8 J W 500.0 -
l 0.0 ' ' ' ' ' i 0.0 1.0 2.0 3.0 4.0 5.0 6.0 ' Dimensionless Height CX/D) w Test Date _.e_.W CorreIation t i Figure 3.2-5 Local Nusselt Number Coniparison for Eckert and Diaguila Test 5 owls 65-oon:it>aions 3-15 l i i
. = . . - ,
t E i 2000.0 1500.0 - b , z , , _a w
$ 1000.0 -
, y 8 =
Z w a , w w , O 500.0 - w w w r T 0.0 ' ' ' i e 0.0 1.0 2.0 3.0 4.0 5.0 6.0 Dimensionless Helght (X/D] , w Test Data + W Correlation I I t Figure 3.2-6 Local Nusselt Number Comparison for Eckert and Diaguila Test 6 n:\np6(XA1865-non:lt>.041CR5 3-16
l f i I i 2000.0 i 1500.0 - b
=
2 N W
$ 1000.0 - y 8 = ;
2 W 'n y C U 500.0 - N M N w
- I l
0.0 ' ' ' f a l 0.0 1.0 2.0 3.0 4.0 5.0 6.0 Dimensionless Height (X/D) I m Test Data + W Correlation l 1 i Figure 3.2-7 Local Nusselt Number Comparison for Eckert and Diaguila Test 7 u:Waa1865-nan:15041095 3-17 i
v i I 1 3000.0 2500.0 - N b
- a g
Z 2000.0 - N W
+J 1500,0 -
C O W l a . 1000.0 - K N 500.0 O.0 1.0 2.0 3.0 4.0 5.0 6.0 Dimensionless Height (X/D)- x Test Data _.e_W Correlation Figure 3.2-8 Local Nusselt Number Comparison for Eckert and Diaguila Test 8 u:\ap6(IA1865-non:lt>041095 3-18 i
,w.
2500.0 n 2000,0 - g w K N 2 1500.0 - g 2
- e
$ y "
E
- 1000.0 - y E
- 8 J
M 500.0 - N O.0 O.0 1.0 2.0 3.0 4.0 5.0 6.0 Dimensionless Height (X/D) m Test Data _ ,._W Correlation 1 Figure 3.2-9 Local Nusselt Number Comparison for Eckert and Diaguila Test 9 a:\np60A1865-non:1MM1095 3-19
, . . _ ~ . - , . f I f I 2500.0 i r 2000.0 - b f 1500.0 - e a , e w 01
$ W X 1000.0
- 10 R #
O ! W 500.0 - W W K L 0.0 ' ' ' ' O.0 1.0 2.0 3.0 4.0 5.0 , Dimensionless Height (X/D) w Test Data + W Correietion s i Figure 3.2-10 Local Nusselt Number Comparison for Eckert and Diaguila Test 10 u%waisss-non:it,mioos 3-20
r 2.0 N R " i *
- w 3
15 - x m ( x x z
- X m
y g g ,
- w I $;
8 " a " # wa y 3 10 1.0 -
# W 5 hw y N x y N gg 3
a * *
- I**E *
- E*# - x *
- s fj 0.5 -
E a g*: - -
- 2 3
t , , 0.0
- .0 2.0 3,o 4.0 5.0 6.0 Dimensionless Height (X/d) x Data Points Average Value Mean Value C1.028) i Figure 3.211 Comparison of Predicted-to-Measured Local Nusselt Numbers for the Eckert and Diaguila Tests u:W1865-ooanono95 3-21
?
3.3 The Siegel and Norris Tests d' Siegel and Norris conducted heat transfer tests on a set of heated parallel vertical flat plates. De channel height was 5.833 ft., and the plate separation distance was variable between 0.125 to 1.25 ft. A constant, uniform heat flux of approximately 1100 Btu /hr-ft.2 was applied. Only those tests that had the test section open at the bottom were examined for comparison. The tests generated data that > were used to validate the assisting mixed convection heat transfer mode for low Reynolds numbers and moderate Grashof numbers. De tests were modeled using the WGOTHIC code. De test section was divided into 11 axial volumes. Because most of the rapid changes occur at the entrance, the first 10 volumes were each 1/15th of the total volume; the last volume was 1/3 of the total volume. He code calculated the velocity, air temperature, and heat transfer coefficient in each of the 11 volumes. The local heat transfer coefficient was calculated using the method for combining the free and forced convection correlations described in Section 2.1. Siegel and Norris reported heat transfer coefficients based on the (Tw - Tw o) temperature difference. : Consequently, the .W_ GOTHIC predictions, which are based on the (T - Tm) temperature difference, were transformed to a basis comparable to the Siegel and Norris data for presentation. The calculated local Nusselt number results for each of the eight tests are compared with the test data and are shown as a function of dimensionless height in Figures 3.3-1 through 3.3-8. Some relevant test parameters are shown in Table 3.3-1. TABLE 3.3-1 SIEGEL & NORRIS TEST DATA Test Number Ill\ Air Temp. ('F) Gr/r Range Re, Range 1 3.00 80.6 - 86.4 4.23E08 - 6.10E08 1.07EM - 1.13EM 2 4.16 80.9 - 88.5 1.58E08 - 2.42E08 8.73E03 - 9.18E03 3 7.66 81.2 - 93.5 2.40E07 - 4.19E07 5.77E03 - 6.03E03 4 1233 81.5 - 99.4 4.40E06 - 1.05E07 4.01E03 - 4.18E03 5 24.00 82.6 - 114.6 6.43E05 - 1.48E06 2.20E03 - 2.28E03 6 1233 81.5 - 1003 5.42E06 - 1.17E07 3.82E03 - 3.98E03 7 1233 82.1 - 107.6 6.43E% - 1.17E07 2.76E03 - 2.89E03 8 1233 83.4 - 123.4 6.70E06 - 1.29E07 1.65E03 - 1.73E03 una;6XA1865-oon:1be41095 3-22
l l l i i A compilation of the predicted-to-measured local Nusselt numbers for all eight tests is shown m , 1 Figure 3.3-9. The mean predicted-to-measured value of 0.857 is also shown. The standard deviation ; of the predicted-to-measured values for all eight tests is 0.0903. l l 1 As demonstrated in tests 1 through 5, the calculated local Nusselt numbers match the measured data 1 fairly well at lower values of UD,, but increasingly underpredict as the UD, value increases. Tests 4, 6,7, and 8 demonstrate the effect of reduced air flow (at constant UD,) by increasing the channel loss 1 coefficient from 1.5 to 35.6. The calculated local Nusselt numbers increasingly underpredict the measured values as the air flow is reduced. l l 1 l I l l l I r 1 us mnis65.non: e b-041ms 3-23 I l
t s 1 i 300 250 - N g 200 - O a t' 3 150 - z , ca o o J 100 - x ! 50 - X w 0 ' ' ' ' 0 0.5 1 1.5 2 2.5 3 3.5 , Dimensionless Height (X/d) 1 m Test Date + W CorreIation i 1
)
l l 1 i I i i Figure 3.3-1 Local Nusselt Number Comparison for Siegel and Norris Test 1 1 l a:Ws65-oon:tomio95 3-24 i
- _ - - - . . . = _ .
i t i l i I i i i 300 r f 250 - r t 200 - .o ' Oa i z i 3- 150 - z c : U O J 100 - y i i i m 50 - ; i i m O ' ' ' 't 0 1 2 3 4 5 Dimensionless Height (X/d) x Test Data _,_.W Correlation l t 1 Figure 33-2 Local Nusselt Number Comparison for Siegel and Norris Test 2 a:W1865-aco:ll>o41095 3-25 i i l
m 250 M 200 - b 150 - W Q E E 100 - M 3 W 50 - W f f f f f f f 0 1 2 3 4 5 6 7 8 9 DimensIonless Height (X/d) x Test Data + w correlation Figure 3.3 3 Local Nusselt Number Comparison for Siegel and Norris Test 3 uwis65.non:1t>o4to95 3-26 i
. , )
l l I< j
.I 1
l l i l i i i 250 [ t
- i 200 - ;
r L i o , 150 - ; Z t S , C 100 - o J W 50 - W M 0 ' ' O 5 10 1s Dimensionless Height (X/d) W Test Data + w correlation Figure 3.3-4 Local Nusselt Number Comparison for Siegel and Norris Test 4 a:ps65-no.:1b441095 3-27
t t h I r P l t h r 250 l I 200 - W ; L 150 - i 2 m i 3 z 10 100 - u O N J
" ?
50 - w r 0 ' ' ' ' ' f 0 5 10 15 20 25 30 , Dimensionless Height (X/d) l w Test Data _,_w correlation
}
t I I i Figure 3.3 5 Local Nusselt Number Comparison for Siegel and Norris Test 5 n:Wd865-mos:lt>-041095 3-28 -
i . r i i i 2 L t i t (
-l 250 ;
t r 200 - m I L 0 150 - i Z
- 3 !
2 < i I C 100 - u y : 0 ? J !
" I so -
N , o ' ' o 5 10 15 Dimensionless Height (X/d) x Test Data _,_W Correlation l I i i l Figure 3.3-6 Local Nusseh Number Comparison for Siegel and Norris Test 6 u w w a1865-oona b a io95 3 29
l 250 f 200 - r L { W L 150 - 2 X 3 Z 100 - O M J 50 - N i M 0 ' ' O 5 10 15 i Di mens i on l ess He i ght CX/ d) y Test Date +W Correlet ion ! i t Figure 3.3-7 Local Nusselt Number Comparison for Siegel and Norris Test 7 ! i uAmp60(A1865-non:1bo41095 3-30 i i'
i,
?
j' , i
)
C 200 i l M 150 - b z ; 3 100 - l z . e u
- O
.a 50 -
w M 0 ' ' O 5 10 15 Dimensionless Height (X/d) x Test Data + W Correlation i Figure 3.3-8 Local Nusselt Number Comparison for Siegel and Norris Test 8 a:W1865-non:1t>441095 3-31
I Y 1.6 , i b g 1.4 .- 3 z 3 z 1.7 g a o O
.,,J W V 1 ..
O wy b WX
- 8 2 = g*
"" N ~
n f
, 0.8 kW w w w y D K y X e 5 N w e
.o 3 0.s -
k
' t i D.4 i 0 5 10 15 20 25 Dimensionless Length CX/d) x Data Points Mean Va lue (0. 857) l l
l P Figure 3.3-9 Comparison of Predicted-to-Measured Local Nusselt Numbers for the Siegel and Norris Tests , t a:\ap60CA1865-non:ll>-041095 3-32
3.4 The Westinghouse STC Dry Flat Plate Tests A series of forced convection heat transfer tests were performed at the Westinghouse Science and Technology Center (STC)?* The purpose for these tests was to provide heat transfer data for channels with a heat flux and cooling air flow rate representative of the AP600 annulus during a DBA. ! The test section was a vertical, 6-ft. long heated flat steel plate that had been coated with a highly ; wettable inorganic-zine coating. A clear acrylic cover provided a channel for the forced air flow. 'Ihe ; plate temperature and air flow rates were varied for each test. The measured parameters for each test are shown in Table 3.4-1. , The data from the seven STC dry flat plate tests were compared with section average results calculated using the selected heat transfer correlations. The calculated channel averaged predicted-to-measured Nusselt numbers for each of the seven tests are shown as a function of the Reynolds number in Figure 3.4-1. The measured data are compared with the heat transfer correlation in Figure 3.4-2. The mean value is 0.983 with a standard deviation of 0.072. a,b TABLE 3.4-1 WESTINGHOUSE STC DRY FLAT PLATE TEST PARAMETERS l l l l l uAap60ml865-non:lb-041095 3-33
t t 4 L i z s 1.s - e E E ' g a 8 w 1 - ; b , E w 8 h
*M 0.5 -
t 5 e N < 0 30 40 50 60 70 to 90 100 110 Thmasands , i ReynoIds Nutter w Test Data Mean (0.983) r b A-Figure 3.4-1 Comparison of Predicted-to Measured Nusselt Number for the STC Dry Flat Mate Tests a:W s65upf.165 3 34 t
l I
)
)
i I 1 9cco , i w m
- 5 m 100 -
g 3 2 i 10 1.0E,06
'
- 1.0E+05 ReynoIds Nurrber w Test Data correlation 1
1 1 Figure 3.4 2 Correlated Heat Transfer Data for the STC Dry Hat Plate Tests e W m wyt:ibo40s95 3-35
l 1 1 3.5 The Westinghouse Large-Scale Tests - Dry External IIeat Transfer 02) A series of heat transfer tests were performed at the large-scale PCS test facility at the Westinghouse Science and Tecimology Center (STC). The purpose of these tests wn to compile data for developing and validating the analytical models of heat transfer. The heat transfer tests were performed over a range of internal test vessel pressures, bounding the calculated worst-case containment pressure, to obtain heat transfer data at relevant conditions and characterize air cooling velocities developed by natural convection. 'Ihese tests were run without the application of a liquid film on the containment shell. The large-scale PCS test facility uses a 20-ft. tall,15-ft. diameter pressure vessel to simulate the steel containment shell. The geometry is approximately a 1/8-scale of the AP600 containment vesse!. The vessel contains air at I atmosphere when cold and is supphed with steam at pressures up to 100 psig. Steam is injected through a diffuser (to reduce kinetic energy) at the lower elevation and rises upward as a plume. Air is entrained in the rising plume, resulting in a natural circulation flow pattern and various degrees of mixing within the vessel. A plexiglass cylinder is installed around the vessel to form the air cooling annulus. Air flows upward through the annulus via natural convection to cool the vessel resulting in condensation of the steam inside the vessel. A fan is located at the top of the annular shell to provide the capability ofinducing higher air velocities than can be achieved during purely natural convection. Thermocouples are located on both the inner and outer surfaces of the vessel at various angles at each of 10 different elevations to detennine the temperature and flux distribution over the height and circumference of the vessel. Thermocouples are also placed inside G . vessel on a movable rake to measure the bulk temperature at various locations. The cooling air teinr fature and velocity are measured at seveml locations in the annulus. The steam inlet pressure, temperature, flow, and condensate flow and temperature are measured to provide an accurate measurement of the total heat supplied to the vessel. A tremendous amount of data were generated for each test. This data varied over time, angular position, and elevation. Only the time-averaged, steady-state data were used for validating the heat transfer correlation. Interpolated values were used for some locations since all of the required local data measurements were not available, (i.e., either because the local measurements were not all taken at the same angular location or were not measured at all). For this reason, the data were also averaged circumferentially to reduce the uncertainty. The steady-state, circumferentially averaged data from 14 of the 16 dry large-scale PCS heat transfer tests were compared with results calculated using the selected heat transfer correlations. Note, tests RCol5 and RC016 were omitted from this comparison because the forced asymmetric annular air flow u \ap60ml865-ene:Ib-041095 3-36
rate imposed for these tests would have affected the circu:nferential averaging. Some relevant test parameters are shown in Table 3.5-1. A compilation of the predicted-to-measured local Nusselt numbers for all 14 tests is shown in Figure 3.5-1. The average predicted-to-measured value at each location and mean value over all locations are also shown. The mean value is 0.895 with a standard deviation of 0.122. In all the tests, the internal vapor temperature was nearly constant between elevations C and B, the internal wall temperature was nearly constant between elevations B and A and the outer wall temperature increases linearly from elevation D to A. For test RC013, the vertical inner wall temperature increased and peaked at elevation C; this also occurred in tests 14A,15,16,29A, and 29B. The steam flow for these tests was higher than the other tests. The local inner wall heat transfer coefficient was apparently higher at elevation C for these tests, causing the heat flux to be } higher at this elevation. As a result, the correlation underpredicted heat transfer at this elevation in all of the tests. Except for the entrance, the correlation slightly underpredicts heat transfer. The entrance-effect multiplier may be too large for the fairly small fraction of the total heat transfer area at small distances from the entrance. This effect was observed on other tests as well. owwnis65-nonab-m:o95 3-37 - i .
l l i i
- - ali TABLE 3.51 i
WESTINGHOUSE LARGE FCALE PCS TESTS - DRY EXTERNAL HEAT TRANSFER i i 1 5 l h
.r 7
I l t i
?
)
b t l f 1 3 a:\npNEA1865-oon:Ib-041095 3-38
, a. c -._. - _ . _ . -- --
t h i i i s a L i G g E > z w j 1.5 - E s z , E W w ! n n w
- y 1 -
l ,
= = =
4 y x , N - i
! g ,
$ o.s - E e w n >
u 5 5D E o ! O 5 10 15 20 Length (ft)
, Test Data + Elevation Average
+ Mean (0. 895) i t
Figure 3.5-1 Comparison of Predicted-to-Measured Local Nusselt Number for the Westinghouse , Large-Scale PCS Tests u:\np600\l865-nan:lbo41095 3-39
=
l l 1 3.6 The Gilliland and Sherwood Evaporation Tests l Gilliland and Sherwood'"' examined isothermal evaporation mass transfer in a vertical pipe. Films of various liquids were applied to the inside wall of the pipe. Air could flow either upward or downward through the test section so evaporation for both countercurrent and cocurrent flow could be studied. The test section was a 117-cm long vertical pipe with a 2.67-cm mside diameter. Calming sections j were added at both ends of the test section. A falling liquid film covered the inside surface of the test l section. The film flow rate was held constant in all tests at approximately 790 cc/ min while the air flow rate was varied. The inlet air and liquid film teenperatures were maintained within 3'C. The reported parameters for each test are shown in Table 3.6-1. Relatively low air flow rates, compared to the mass transfer rates, were used in these mass transfer tests. The lower air flow rates caused a large difference in the bulk steam partial pressure from inlet-to-outlet, as shown in Figure 3.6-1. Because of the large change in the bulk air / steam mixture properties over the length of the test section, average properties could not be used to evaluate the test data. A simple 10-cell model wr.s developed to evaluate the test data. A compilation of the predicted-to-measured evaporation rates for all 71 tests is shown as a function of the Reynolds number in Figure 3.6-2. The mean value for ell tests is also shown. The mean value is 0.925 with a standard deviation of 0.072. I Since the liquid and air temperatures were nearly the same in these tests, the evaporative mass transfer was almost entirely driven by the difference in partial pressure between the liquid film surface and bulk mixture. l l i I i uwmisss-amtwuious 3-40
TABLE 3.6-1 GILLILAND & SHERWOOD TEST PARAMETERS Air T,,-Top T,, Bottom T Top T.. Bottom Pressure Flow Evap Test 'C 'C 'C 'C mm Hg g/ min ec/ min 1 30.8 27 31.1 26.1 770 250 3.8 3 29.6 27.9 29.8 27.8 772 243 3.6 5 32.1 25.8 32.5 24.9 777 125 1.6 9 32.6 28.3 33.1 27.4 777 143 1.8 11 32.1 28.1 33 27.4 770 54 0.73 13 32.2 26.8 32.9 26.2 770 214 2.5 15 28.7 27.3 29.4 26.7 772 324 4.7 17 30 27.8 30.6 27.1 770 218 3.2 19 32.6 28.9 33.1 28 3 770 158 2.7 21 323 28.6 32.9 28.2 770 103 1.8 23 32.4 29.2 32.9 28.5 770 74 1.4
- 25 41.1 38.4 41.7 37.8 785 220 6.2 27 40.4 41.3 41.4 40.8 800 80 2.4 29 38.9 35.8 39.5 35.2 782 48 1.5 31 40.5 28 30.1 26.5 775 502 7.4 33 42.4 37.2 42.6 36.6 777 111 3.7 35 42.4 38.8 43.5 38.1 775 141 4.4 37 42 36.9 43.1 36.7 777 174 5.4 39 41.7 37.4 42.2 36.9 760 215 7.4 41 41.1 38.5 41.9 38 760 197 5.7 43 40.7 38.9 413 383 770 143 4.1 45 40.7 39.9 413 39.5 770 96 3.2 47 45.2 39.7 44 3 39.5 765 51 2 49 44.1 39.7 44.7 38.6 765 457 12.9 51 43.8 40 44 38.4 765 625 17 unap6MIR65.non:lb-641095 3-41
=
I TABLE 3.6-1 (Cont.) GILLILAND & SHERWOOD TEST PARAMETERS Air T,,,-Top T ,,-Bottom T. -Top T,.. Bottom Pressure Flow Evap Test 'C 'C 'C 'C mm-Hg g/ min ec/ min 53 56.1 50.9 55.2 50 760 46 33 55 52.8 46.1 53.5 45.4 802 96 5.9 57 51.6 49.2 50.1 48.4 767 88 5.3 - 59 52.6 43.2 53.2 42.7 767 119 6.5 61 53.9 48.2 53.5 47.8 785 248 13.2 63 54.6 49.1 54 3 48.5 785 168 9.6 65 53.1 483 533 47 787 475 22.8 67 52.1 49.6 54.7 49.1 782 126 8.1 2 34.9 33.4 35.6 33.2 772 63 1.6 4 33.8 32.1 34.5 32 114 60 12 6 36.2 34.9 36.4 34.6 2006 66 0.65 8 31.8 26.7 32.5 26.3 407 125 4.4 10 32.6 27.6 32.8 27.1 1480 127 1.1 12 31.9 28.1 32 27.9 269 122 6 14 33.1 28.6 31 28 941 123 1.6 16 31.4 28.8 32 28.5 1966 121 0.75 18 32.8 28 3 32.5 28 556 217 4.8 20 31 28.1 32 27.5 1419 216 1.75 22 30.8 26 31.9 25.2 424 214 6.7 24 25.8 25.1 36 34.6 2325 65 0.44 27 31.9 27.6 33.6 27.1 1248 213 23 29 32.1 26.7 32.7 26.2 1958 218 1.5 l 31 32.9 29.1 34.1 28.7 919 374 4.8 33 43.8 38.9 44 3 39.5 765 51 1.9 uwins.noo:st,-o.iims 3-42 1
'^ !
1 l TABLE 3.61 (Cont.) GILLILAND & SHERWOOD TEST PARAMETERS Air T.,-Top T,.-Bottom T,... Top T, ..-Bottom Pressure Flow Evap j Test 'C 'C *C 'C mm-Hg g/ min cc/ min ! 35 40.8 37 41.9 36.5 112 47 15 37 42.8 40.4 43.6 40.2 1695 51 0.8 39 40.6 37.1 41 35.9 249 48 63 41 42.1 37.7 42.8 37.2 992 49 *5. 43 42.7 39.4 433 38.9 1922 50 0.9 45 42.6 38.4 43.2 37.9 505 138 7.0 47 42.8 39.7 43.2 39.2 1385 137 2.4 49 41.8 35.8 423 35.3 396 134 9.1 51 41.8 37.8 42.5 37.2 1183 147 3.0 53 41.8 38.1 42.6 37.7 607 187 7.2 55 41.2 37.1 41.7 36.6 411 187 11 3 57 403 363 40.9 36 1235 195 3.6 59 42.7 38.6 43.1 37.9 IMS 339 7.1 61 51.8 49.6 52.3 49 762 88 5.2 63 51.9 47.9 53.1 48.8 1321 88 3.2 65 53.4 49.4 56 49 320 121 20.8 67 55.1 49.9 55.6 49.6 1418 123 4.0 69 53.9 47.6 55.2 48.5 518 120 13.1 71 53.6 46.8 54.2 46 757 201 11.8 73 54.6 47.7 55.2 47.1 951 201 8.7 75 55.9 47.1 56.2 46.9 574 354 24.8 u:W1865-oon:1b-Gilcos 3-43
, , . _ - . _ . , ~. _ - .. .-
i 6 I i l 40 35 - 0 30 - e b 25 - b 20 - b _ 15 -
+J !
L r g 10 - 5 - r O O 11,7 23.4 35.1 46.8 Se.5 70.2 81.9 93.6 105.3 117 ChanneI Length (cm) _ ,_ steam _ _saturatton P i i Figure 3.61 Calculated Steam Partial Pressure vs. Channel Length from a Selected - Gilliland and Sherwood Evaporation Test u:W865-non.1t*041095 3-44 -
]
2 i 1,5 - _o Y I' t. S I i l
~
_N 4 8 " 2w w 3 r.# ng-N
. m - ;
e 8 D e D.S -
.U_
E O ' ' ' ' ' O 5 10 15 20 25 30 Thousands AeynoIds Number m Test Data Mean (0.925) i l Figure 3.6-2 Comparison of Predicted-to-Measured Evaporation Rates for the Gilliland j and Sherwood Evaporation Tests u:WEA1865 non:lt41095 3-45
3.7 The Westinghouse STC Flat Plate Evaporation Tests A series of liquid film evaporation tests were performed at the Westinghouse Science and Technology Center (STC)."* The purpose for these tests was to observe the behavior of the liquid film and to provide data on both sensible heat transfer and evaporative mass transfer. The test conditions were selected to simulate the outside of the AP600 steel containment vessel with the PCS in operation. The test section was a vertical, 6-ft. long heated flat steel plate that was coated with a highly wettable inorganic-zine coating. A clear acrylic cover provided a channel for the forced air flow and allowed for observation of the applied liquid film. The plate temperature, applied liquid film temperature, and both the liquid and air flow rates were varied for each test. The measured parameters for each test are shown in Table 3.7-1. Relatively high air flow rates, in comparison to the mass transfer, were used in these tests. As a result, as shown in Figure 3.7-1, the change in the bulk-to-film steam partial pressure difference from inlet-to-outlet was fairly small and decreases as the air flow rate increases. Therefore, channel average properties were used to calculte the predicted Sherwood number for comparison with the test data. The data from the 23 STC flat plate evaporation tests were compared with section average results calculated using the selected heat and mass transfer correlations. The calculated average Sherwood numbers for each of the 23 tests are compared with the measured test data and are shown as a function of the Reynolds number in Figure 3.7-2. The measured data are compared with the heat transfer correlation in Figure 3.7-3. The mean value is 0.936 with a standard deviation of 0.139. ,_ _ a,b TAllLE 3.71 WESTINGIIOUSE STC FLAT PLATE TEST PARAMETERS i l u:\ar600\l865-oon:lt,-041095 3-46
_- _. a,b JABLE 3.71 (Cont.) WESTINGHOUSE S1C FLAT PLATE TEST PARAMETERS
=
v:'ap60tA1865-oon:lt@l095 3-47
a,b Figure 3.7-1 Bulk-to-Film Steam Partial Pressure Differences vs. Channel langth from Selected Westinghouse STC Flat Plate Evaportation Tests u:\apMKAl865-oon:1b041095 3-48
l l i
-l s
c
, r 3 1.5 -
2 b ! b w , I a
- b 1 -
, m y m #
- i E
$ E E g w ,
T > j o.s - a E E
' ' ' ' i c i D 20 40 60 30 100 120 MO Thoumaruss ;
Reynolds Number a Test Data _Mean (0.936) 3 t 5 , Figure 3.7 2 Comparison of Predicted-to-Measured Sherwood Numbers for the STC i Flat Plate Evaporation Tests l t u%ww865- onso41o95 3-49
'l 0
b t a i t 1000 i i a W N N > m Y I D i
< t 100 -
g 2 vs 5 F W L I
, i 10
- 1. 00E+ M 1.00E+05 1.00E+06 Reynolds Number m Test Data correlation .
Y k Figure 3.7-3 Correlated Mass Transfer Data for the STC Wet Fiat Plate Tests avanis65- on:Imioos 3-50
3.8 The University of Wisconsin Condensation Tests 1 A series of condensation tests were conducted at the University of Wisconsin."" The purpose of these tests was to provide data on condensing heat and mass transfer in the presence of a noncondensible gas at various inclination angles, velocities and steam / air concentrations. De test conditions are similar to what could be expected following a loss-of-coolant accident (LOCA) or steamline break transient within the AP600 containment vessel. He test section was 6.25-ft. long with a 2.75-ft. entrance length and a 3.5-ft. condensing surface length. He channel cross section was square with an area of 0.25 ft2 . He top of the test section was covered with a thick aluminum plate coated with a highly wettable, protective, inorganic-zinc coating. Seven 0.5-ft. long coolant plates were attached to the back of the aluminum test plate to remove heat. Each coolant plate had both flux meters and cooling coils with thermocouples to provide redundant, diverse energy measurements. The test section could be inclined from any angle (0-90 degrees from horizontal). Plate number I was located at the end nearest the air / steam source and was always at the lowest level when the test section was inclined. Some relevant test parameters are shown in Table 3.8-1. Relatively high air flow rates, in comparison to the mass transfer, were used in these tests. As a l resu't, the change in the bulk-to-film steam partial pressure difference from inlet to outlet was small, as shown in Figure 3.8-1. Herefore, channel average properties were used to calculate the predicted Sherwood number for comparison with the test data. He data from the 59 University of Wisconsin condensation tests was compared with section average results calculated using the selected heat and mass transfer correlations. The calculated channel average Sherwood numbers for each of the 59 tests are compared with the measured test data and are shown as a function of the inclination angle in Figure 3.8-2; as a function of the steam / air Reynolds number in Figure 3.8-3; as a function of the bulk air / steam concentration in Figure 3.8-4; and as a function of heat flux in Figure 3.8-5. The measured data are compared with tin mass transfer correlation in Figure 3.8-6. The mean value is 0.968 with a standard deviation of 0.203. He predicted channel average Sherwood number appears to have a dependence on the inclination angle and the Reynolds number, increasing as they are increased; and a small dependence on the heat flux and steam concentration, decreasing as they are increased. Therefore, condensation heat and mass transfer would be underpredicted at the conditions expected near the top of the AP600 dome during a LOCA or steamline break transient (higher steam concentrations, higher heat fluxes and small angles of inclination). i n.womsenab-es 3-51
l TABLE 3.8-1 WISCONSIN CONDENSATION TEST PARAMETERS Avg. Ileat Flux Temp In Temp Out T-wall Velocity Test W/m2 *C *C *C m/s Angle 83 27342 95.6 95.7 45.9 1 90 82 27493 95.4 95.7 49.5 1 45 80 26395 95.2 94.9 45.2 1 12 81 27117 95.1 94.8 45.6 1 12 , 78 27257 94.9 95.1 44.5 1 0 79 27189 94.6 94.6 47 1 6 86 27260 94.5 94.2 443 1 45 74 16913 90.1 89.9 29.2 1 12 73 16675 90.1 89.8 29.3 1 45 75 17681 90.3 89.7 30.7 1 6 76 16615 90.2 89.7 28.9 1 0 : 72 17178 89.7 89.7 31.3 1 45 71 14651 90 89 29.9 1 90 77 16645 89.3 89.7 29.5 1 0 94 25223 90.5 89.3 39.5 2 0 55 11692 80.6 80.4 29.9 2 0 70 14171 80.5 79.7 29.9 3 90 57 8592 80.1 80 29.8 1 12 48 8166 80.4 79.6 29.9 1 90 43 10140 80.4 79.6 29.5 1 6 50 10168 80.1 79.6 29.6 2 90 69 14069 80.6 79.9 29.7 3 0 68 14537 79.9 79.9 29.7 3 6 u:W1865-noe:1h0410)$ 3-52
i l l TABLE 3.8-1 (Cont.) WISCONSIN CONDENSATION TEST PARAMETERS Avg. IIeat Flux TempIn Temp Out T-wall Velocity Test W/m' 'C 'C 'C m/s Angle 64 14553 79.8 79.8 30.2 3 90 44 10589 80 79.6 30.1 1 12 51 10515 79.9 79.5 29.7 2 45 52 10807 80.1 793 29.6 2 12 56 7983 80 79 293 1 12 47 7973 80.2 78.8 30.1 1 45 54 10939 79.7 79.4 30 2 6 34 9931 80.6 78.8 30.1 1 0 53 10023 79.9 79 29.6 2 6 67 13100 79.5 793 29.8 3 12 66 14523 79.4 79.2 30.2 3 45 85 15974 79.4 793 30.4 3 0 31 9881 78.7 78.7 30 1 0 32 9924 77.4 77.4 30.1 1 0 36 8194 72.2 72.1 30.1 2 0 62 8733 71.7 71.8 29.9 3 45 42 8170 72.5 70.9 29.6 2 6 61 9286 71.5 713 303 3 12 59 8888 71 71.6 293 3 6 39 7960 71.8 71.4 302 2 0 46 6743 71 71.4 30.6 2 45 49 6561 70.8 70.8 293 2 90 41 8677 71.7 70.6 29.6 2 0 87 5572 71 70.2 30.2 1 0 91 4736 70.6 69.9 29.7 1 45 u-\apMal&65-non:lb-Ni@5 3-53
=
TABLE 3.8-1 (Cont.) WISCONSIN CONDENSATION TEST PARAMETERS Avg. Heat Hur Temp In Temp Out T. wall Velocity Test W/m 2 *C 'C 'C m/s Angle 63 8100 70.7 69.7 29.7 3 90 90 5173 69.7 70.5 29.7 1 12 92 4457 69.9 70.1 29.4 1 88 88 53M 70 to.2 29.8 1 6 40 9353 70 70 30.1 2 0 45 7868 69.4 69.4 29.5 2 12 60 9624 69.6 69.6 30 3 0 89 5411 69.2 69 29.9 1 12 58 4869 61.1 61.9 29.9 2 6 93 2769 60.4 60.4 29.7 1 0 65 M49 59.2 60.4 29.9 3 45 l uwaxAt865-non:1t>411095 3-54
12 2
- 11 -
O 10 - g 8
- e -
he e 7 b 8 s'- 2 5 s - D L f 4'- 5 3 g - M <: i , , O a O 1 2 3 4 Channel Length (f t")
-e Test 83 + Test 74 + Test 55 # Test 36 e._ Test 58 Figure 3.8-1 Bulk-to-Film Steam Partial Pressure Differences vs. Channel Length from Selected Wisconsin Condensation Tests a:\np600u865-nan:1M)41005 3 55
.f
}
5 i l [ v t e f s l a r L . S i V 1.5 - O F k a
- i
. - i :
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, I -
o w e e 3 c.s - t E e ' ' ' ' ' ' i i 4 ! 0 10 20 30 40 50 60 70 80 90 100 f Pla*.e inclination from Horizontal (Degrees) f a Test Data Mean CD,958) i I t Figure 3.8-2 Comparison of Predicted-to-Measured Sherwood Numbers for the ' i Wisconsin Condensation Tests j u:Wis65-oon: baio95 3-56
e b 1.s -
]
W b %" N w Y5 u
- m
- gL
- 1 -
), gs - - -
c 5, n'-
- w
?, 5 m
- y c.s -
t E a --- S 10 15 20 25 20 Thousands Reynolds Number
. Test Data Mean C0.968)
Figure 3.8-3 Comparison of Predicted-to-Measured Sherwood Numbers for the Wisconsin Condensation Tests u:wn65-nan:thwims 3-57
i
- I b
z j u 1.5 - 8 fo xx ' 5 $ w ww a ; u w - o fw w : y 1 -
- 3 >
9 w (=a"
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= ;
4
- yt
" 5 i O w m a w w ,
i w t 3 a 0.s - i
.o_
u o k
- 0 O.1 0.2 0.3 0.4 0.5 0.5 0.7 Steam Concen* ration (mole %)
x Test Data Mean C0.968) f f f i 6 Figure 3.N-4 Comparison of Predicted-to-Measured Sherwood Numbers for the i Wisconsin Condensation Tests u v uxts65 oon:1bo4toos 3 58 !
~
T
'i 1
i i i i 2 L Z ' V 1.5 - 8, i. L NE mg b = *
- I ui % =**
m mm D m
-- M I 1 m ' NL h -- -
g m a mm*
- j m
- 5 ,m g K i'
O i $ * *Em g * , s a m t I -i j o.5 - a . 2 I u ; o k
, , , , , , , , , i a
O 1 2 3 4 5 5 */ e 9 10 Thounencia Heat Fl ux (BTU /hr-f t2) m Test Data Mean (0.958) ! 3 i i i i > i l Figure 3.8 5 Comparison of Predicted to-Measured Sherwood Numbers for the Wisconsin l l Condensation Tests l 1 l a:Wann65-mo :Imiov5 3-59 l t
a . . .-. .. - - -. . . .. . _, , - = . . , . f k p k f F 1000 , I F t t i O . N t-
< M
~
h g E , M ' W. t I
.. i
- l l
10 1.00E+03 1.00E+04 1.00E+05 , ReynoIds Nurnber a Test Data correlation t 6 i Figure 3.8-6 Correlated Mass Transfer Data for the Wisconsin Condensation Tests I f
?
I u:\ap60thl865-oce:ltW1095 3-60 I
3.9 The Westinghouse Large-Scale Tests - Internal Condensation "') The PCS phase 2 (confirmatory) heat and mass transfer tests were performed at the large-scale PCS test facility at the Westinghouse Science and Technology Center (SK). The phase 2 tests provide data on the transient heat transfer and distribution of noncondensable gas in a geometry similar to the AP600 containment vessel. De purpose of these tests was to provide data to be used in developing and validating the analytical models for heat and mass transfer. The large-scale PCS test facility uses a 20-ft. tall,15-ft. diameter pressure vessel to simulate the steel containment shell. The geometry is approximately a 1/8-scale of the AP600 containment vessel. The vessel contains air at one atmosphere when cold and is supplied with steam at pressures up to 100 psig. Steam is injected through a diffuser (to reduce kinetic energy) at the lower elevation and rises upward as a plume. Air is entrained in the rising plume, resulting in a natural circulation flow pattern and various degrees of mixing within the vessel. A plexiglass cylinder is installed around the vessel to form the air cooling annulus. Air flows upward through the annulus via natural convection to cool the vessel resulting in condensation of the steam inside the vessel. A fan is located at the top of the annular shell to provide the capability of inducing higher air velocities than can be achieved during purely natural convection. A liquid film is applied to the outside of the test vessel to provide additional, evaporative cooling. Dermocouples are located on both the inner and outer surfaces of the vessel at various angles at each of 10 different elevations to determine the temperature and flux distribution over the height and circumference of the vessel. Thermocouples are also placed inside the vessel on a movable rake to measure the bulk temperature at various locations. The cooling air temperature and velocity are measured at several locations in the annulus. The steam inlet pressure, temperature, flow and condensate flow and temperature are measured to provide an accurate measurement of the total heat supplied to the vessel. A tremendous amount of data was generated for each test. His data varied over time, angular position and elevation. Only the time-averaged, steady-state data were used for validating the heat transfer correlation. Interpolated values were used for some locations since all of the required local data measurements were not available, (i.e., either the local measurements were not all taken at the same angular location or not measured at all). For this reason, the data were also averaged circumferentially to reduce the uncertainty. The steady-state, circumferentially averaged data from 7 of the 25 phase 2 tests were compared with results calculated using the selected heat and mass transfer correlations. Only tests with film coverage greater than 90 percent were included in the comparison because lower film coverage would have affected the circumferential averaging. This eliminated 17 of the tests. Test RC062 (blind test for
.W, GOTHIC validation) was also omitted from the comparison because the data has not yet been
- released. Some relevant test parameters are showr, in Table 3.9-1.
u:\np(OA1865-non Ib-041095 3-(> l
, a,b TABLE 3.9-1 WESTINGHOUSE LARGE-SCALE PCS TESTS . INTERNAL CONDENSATION
[ A compilation of the predicted-to-measured local Nusselt numbers for all seven tests is shown in Figure 3.9-1. The average predicted-to-measured value at each location and mean value over all locations are also shown. He mean value is 1.000 with a standard deviation of 0.106. Figure 3.9-4 provides a comparison of the measured data with the mass transfer correlation as a function of the Grashof number. He correlation ma:ches the trend in the data. The comparison in Figure 3.9-1 suggests that condensation is overpredicted at the lower elevations and underpredicted at the higher elevations. The bulk steam concentration and condensing heat flux are higher at the higher elevations and lower at the lower elevations, and as shown in Figures 3.9-2 and 3.9-3, mass transfer is apparently underpredicted at higher heat flux and steam concentrations. . The Wisconsin condensation test data showed only a weak dependence on steam concentration and heat flux, so these are unlikely the cause of the observed behavior. He Wisconsin condensation test data did however show a small dependency on inclination angle; condensation was underpredicted on , horizontal surfaces and overpredicted on vertical surfaces, his is consistent with the observed trends in the large-scale test vessel. uAmooosis65- wiso41095 3-62
. . . _ m . _ _ _ - _ . _ m , f 6 i 4 3
?
3 - 1 1.5 - y 3r b # r ., . - w I w -
~
!' ; ; T: .
$ i h 0.5 -
0 2 i O ' ' ' O 5 10 15 20 i Heated Length (ft)
, Test Data + Elevation Average Values
+ Mean (1.00')
Figure 3.9-1 Ccmparison of Predicted-to-Measured Local Sherwood Number for the Westinghouse Large-Scale PCS Tests u:\np60(N 865-nan:!404les 3-63
i i t I i i 1 L 2 i 3o
- 1.s -
a ! I b 6i : m I :
- = , ,
0 N y 3 P w* * < v e
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u" m
*
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" = m 8 ,
0.5 - E i e S u . e > g , , , , , , , , , o 1 2 3 4 s s 7 e o Thousando Heat Flux (BTU / hr-f t2) , a Test Data l
?
Figure 3.9-2 Comparison of Predicted to-Measured Local Sherwood Number for the Westinghouse Large-Scale PCS Tests
- a uAnpmA18td _m.1tW1095 3-64 s
i f
)
L a b u 8 m 1.s - m
- M w n u* g .
8 mm * * ,
,I !
w
.g .., .
e ,
- 1 b M N N N g g g EY i
- w E ,
j w i f f o.s '- ! E B_ E c.1 c.: 0.3 o.4 c.s o.s o.7 i Steam Concentration y Test Data . i 1 i i l l l Figure 3.9-3 Comparison of Predicted 40-Measured Local Sherwood Numberc for the . Westinghouse Large-Scale PCS Tests unap60LN E65-nos:lt4W1095 3-65 I J
~ - - - . _ .
h i 4 F l 1 l
?
s i 10000
?
Y w ** , W - tn MM I N V W ( 1000 .- ,x l m um , 2 w w w , I X X ., W i W Kw " 100 ' .. 1.00E+09 1.00E+10 1.00E+11 1.00E+12 1.00E+13 r Grasho." Number f x Test Data ..__correfatIon ! t i t 4 i v i Figure 3.9-4 Correlated Condensation Mass Transfer Data for the Westinghouse < Large-Scale PCS Tests u:waxxis65 m.:it>.o4io95 3-66 i
4.0 ASSESSMENT OF RESULTS AND STATISTICS Re ratio of the predicted-to-measured Nusselt number (for convection) or Sherwood number (for condensation and evaporation) was calculated and comparisons for each of the individual tests were presented in Section 3.0. His data was soned and combined by mode of heat transfer (convection, evaporation and condensation) for an assessment of the overall statistics to demonstrate that the heat and mass transfer correlations in WGOTHIC represent the phenomena in the AP600 containment vessel over the expected ranges of the various dimensionless groups during a DBA event. 4.1 Convection ne combined convection data consists of the llugot, Eckert and Diaguila, Siegel and Norris, Westinghouse STC flat plate, and Westinghouse large-scale PCS dry tests. He predicted-to-measured Nusselt number ra lo is shown as a function of the Reynolds number in Figure 4.1-1 and as a function of the Grashof number in Figure 4.1-2. He mean predicted-to-measured Nusselt number ratio is 0.976 with a standard deviation of 0.278. The mean predicted-to-measured Nusselt number value near 1.0 indicates that the heat transfer correlation fits the measured data very well. Condensation and evaporation heat and mass transfer result in much higher heat fluxes than convection alone. Although convection is not the dominant heat removal mechanism, the convective heat transfer correlation serves as the basis for the prediction of condensation and evaporation heat and mass transfer. As shown in Sections 4.2 and 4.3, the standard deviation for the predicted-to-measured evaporation and condensation heat and mass transfer is much lower. Therefore, the relatively large standard deviation in the predicted-to-measured convective heat transfer is most likely due to the larger uncertainty in the thermocouple delta-T measurement uncertainty at lower heat fluxes. The comparison with test data in short channels indicates that the entrance effect multiplier overpredicts heat transfer at small distances from the channel entrance. Although the entrance effect becomes increasingly small for large geometries, an upper limit for the entrance effect multiplier will be established for code input. During a DBA event, the Reynolds number can be as high as 1.0E5 and the Grashof number can be as high as 1.0E9 in the annulus region between the steel containment shell and the concrete shield building of the AP600. The convection test data covered a Reynolds number range or between 1.0E3 and 5.0E5 and a Grashof number range of between 1.0E6 and 1.0 Ell. Therefore, the test data cover the expected range of both dimensionless groups within the annulus. i l l u AagwxAIS65-noo:lt>o41005 4-1
i l l 1 3 w w e fj w 2 15 - *
- w w
w # E h*
* # w 3 * ,l
( g= * ! r- w w i " wII EX
- y I
w w""I ,4 li , *I " 1 - 8 E g 6 it w /,", 5
- I , *l m ,
2 g 0.5 - "I 5 e Nu w N O 1000 10000 100000 1 DOC 100 ReynoIds Number Figure 4.1-1 Predicted-to-Measured Nusselt Numbers for Convection as a Function of the Reynolds Number uap6(XAl865-ewItWl095 4-2
3 M W G f 2 1.5 - a M M f w n % q" y L $ I W e.g z yg , w, 1,% ( zL s-* W l ? ,, - a "* st 4 p j h*i
# #
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/,
- I #, $ ! l e
*l y5
- g* f, \
t t 1 g , * *
*y l 4
U 0.5 - Dg O NE k 0 ' ' ' + , 1000000 10000000 1.0E+08 1.0E+09 1.0E+10 1.0E+11 1,0E+12 Grashof Number l Figure 4.12 Predicted to-Measured Nusselt Numbers for Convection as a Function of the Grashof Number / ! n:\ap600\l865 con:ll>o41095 43
l 4.2 Evnporation For most evaporation heat and mass transfer tests, the mass flow rate is measured only at the entrance and exit of the test assembly. An average Sherwood number can be determined only if the panial pressure difference between the film surface and air does not change significantly. For the evaporation tests evaluated in this report, only the Westinghouse ETC flat plate evaporation tests met this criterion , (i.e., because the variatie1 in partial pressure differences within the test channel was not too large, an l average Sherwood number could be determined using the measured total evaporation rate). ) I The predicted to-measured Sherwood ratio for the Westinghouse STC flat plate evaporation tests is i shown as a function of the Reynolds number, Grashof number and dimensionless steam concentration in Figures 4.2-1 through 4.2-3. The mean predicted-to-measured Sherwood number ratio is 0.936 with a standard deviation of 0.139. The evaporation test data covered a Reynolds number range up to 1.2E5 and a Grashof number range up to 7.0E10. The evaporation test data covers the expected range of both the Reynolds and Grashof numbers in the annulus during a DBA event. The Gilliland and Sherwood evaporation tests provided a comparison of the measured and predicted total evaporation rates at relatively low Reynolds and Grashof numbers. As shown in Section 3.5, the heat and mass transfer correlations predicted the measured total evaporation rates very well. However, local evaporation measurements were not made and internal variations in panial pressure difference vary too much to be able to determine an average Sherwood number for these tests. Evaporation data from the Westinghouse large-scale PCS tests was not measured on a local basis. The large-scale PCS tests will be evaluated in an integral setting using the WGOTIIIC code to provide added confidence that evaporative heat and mass transfer is being well-modeled. The results of this evaluation will be presentec the WGOTHIC Code Verification & Validation Report. I l l I shpMA1865 non ItMMlW5 4-4
l i I 2 h b U 1.s . 8 e i m a V w
~
M M g
. w U N #
.8 E E m 8
T j 0.5 . t e N O ' ' , , , 0 20 40 60 go 100 120 140 - N umende Aeynolds Number w Test Data Mean C0.936) l Figure 4.21 Predicted-to-Measured Sherwood Numbers for Evaporation as a Function of the Reynolds Number i u:W1865-mon:1NM1095 4,5 _
.~ - . . . - - - - _ - . - . = - -
t h 9 I t i 2 L V 1.5 - I L
'r m , a V W N , *
- b #
- W W y E ,
$i W g R
O
+a I
D.5 -
+J v
@ i k
' ' ' ' ' t 0 i 30 35 40 45 50 55 60 BS 70 i
Times 1E*DS Grashof Number m Test Data Mean (0.930 t
?
Figure 4.2-2 Predicted-to-Measured Sherwood Numbers for Evaporation as a Function of the Grashof Number r a:wwmiss5 non:ibosims 46
!~ i 2 Z g 1.s - 2 L
$ a m ,
m 3 ' ~ b
- wm
- w i E N" m w l 3
f
*w ** m 1
3 a.s - l v 5 0 E O 1 1.1 1.2 1.3 1.4 1.5 1.6 P/PBM m Test Data _ Mean (0.936) F Figure 4.2 3 Predicted-to-Measured Sherwood Numbers for Evaporation as a Function of the Dimensionless Pressure nAap600us65-non:1bo41095 4-7
I 4.3 Condensation ne combined condensation data consists of the Wisconsin condensation tests and internal condensation data from the Westinghouse large-scale ICS tests. De predicted-to-measured Sherwood ratio is shown as a function of the Reynolds number, Grashof number and dimensionless sMam concentration in Figures 4.3-1 through 4.3-3. He mean predicted-to-measured Sherwood number ratio is 0.983 with a standard deviation of 0.187. Note, local Reynolds number values could not be determined from the measured internal condensation data from the large-scale PCS tests, therefore, only the Wisconsin condensation test data are shown on Figure 4.3-1. The combined test data covered a Reynolds number range of between 5.0E3 and 2.5E4 and a Grashof number range of between 1.0E10 and 1.0E13. The Reynolds number will vary with time and position inside the AP600 containment vessel daring a DBA event. During the relatively short blowdown phase, the velocity and corresponding Reynolds number will be largest on the wall nearest the break location and decrease as the flow moves away from the break. A natural circulation flow pattern is expected to develop during die depressurization phase when the PCS is in operation. The Reynolds number along the wall will be small during natural circulation. The top end of the Grashof number range (calculated using the AP600 inner vessel wall heated length as the length parameter) is estimated to be 1.0E15 during PCS operation. Even though the large-scale PCS test data do not extend into this range, the large-scale PCS test data show that the condensation mass transfer coefficient is independent oflength and that the correlation matches the trend in the data over the three decades of measured Grashof numbers. Other investigators"7 20 have also concluded that the turbulent-free convection condensation heat transfer coefficient is independent of the length. Therefore, it is recommended that this correlation be extrapolated to full-scale contaimnent modeling. uNp60tM E65-non:lt>otil95 4-8
2 )
- L $
2 y 1.5 - C y " wm m n w f V5
~
n W y K y c g w w 3 # , b Y W* w *
=
^
u 0.5 I _u , E k \ 0 ' ' ' ' S 10 15 20 25 30 Thousands ReynoIds Number i Figure 4.3-1 Predicted-to-Measured Sherwood Numbers for Condensation as a Function of the Reynolds Number u:be60tAl&65-oon:lt>o41095 49
._ m m.... . . . _. - ..m ._.m._. .
l i J t l 2 h h 1,5 - g 8
= , i W W W h
E N N " W y M g
" / "
E y a h
* # hIM" "
," " =
y y" " N y g s a a "a w , I = n ag ,
- w 3
- E y 0.5 -
E 1 J 0 1.0E+09 1.0E*10 1.0E+11 1.0E+12 '1.0E+13
)
Greshof Number i i Figure 4.3 2 Predicted to Measured Sherwood Numbers for Evaporation as a Function of the Grashof Niunber i i ump 60tA1865-non:1tW1095 4-10 !
I i 3 b y 1.s - w 1 . g ww . s ",4 = u % *
; h, "% .. e, 1
w
#g E*
- 3 m
- w
- w m
u 3 0.5 - 2 7 k
)
4 4.s a 25 8 '5 P/ PEN l I l l 1 l l l Figure 4.3 3 Predicted-to-Measured Sherwood Numisers for Condensation as a Function of i Dimensionless Pressure i l l a:\npectA1665-noo:1b-041095 4-II
)
l
. _ _ _ _ _ _ _ = _ _ - . ___-_:
=
j 4.4 Measurement Uncertainty The measured temperatures, partial pressures and flow rates which are used to determine the measured Nusselt or Sherwood numbers all have some uncenainty associated with them. The Nusselt number is determined from measurements of the bulk fluid temperature, the wall surface temperature and wall heat flux. The Sherwood number also requires measurements of the component pressures and film flow rate. Both require physical propenies to be determined based on the measured temperature and pressure, i.e., density, thermal conductivity, specific heat and diffusivity. The instrument uncenainties for the large-scale PCS tests were presented in reference 16 and will be used to determine an estimate for the Nusselt and Sherwood number measurement uncertainties. It is believed that these uncertainties are typical and apply to the other heat and mass transfer tests as well. The thermocouple measurement uncertainty for the large-scale PCS tests is 2*F. For these tests, the fluid and wall temperatures were typically greater than 200 F, so the thermocouple measurement uncertainty would be less than 1 percent. A thermocouple pair is calibrated for heat flux measurements and typically the pair has an uncertainty of less than 0.25'F. For the large-scale PCS tests, the wall temperature differences varied between 1 and 35'F, so the wall delta-T measurement uncenainty would be approximately 1 percent at the higher delta-T (corresponding to higher heat fluxes) and could be as high as 25 percent at the lower delta-T (corresponding to lower heat fluxes). Typically, the lowest wall temperature differences occurred in the dry large-scale PCS tests and at the lower elevations in the wet phase 2 tests. The relatively large wall delta-T measurement uncenainty for the lower wall temperature differences will have an impact on the Nusseli and Sherwood number uncenainties at the lower heat fluxes. The air panial pressure measurement uncenainty for the large-scale PCS tests is 1 psi. For these tests, the steady state air partial pressures varied between 20 and 80 psi, so the measurement uncertainty would be between 1 and 5 percent. The film flow rate measurement uncertainty for the large-scale PCS tests was 0.1 gpm (0.015 lbm/sec.). The condensate flow rate varied between 0.1 and 0.6 lbm/sec, and the external liquid film flow rate varied between 0.2 and 2.5 lbm/sec. Therefore, the measurement uncertainty would be between 1 and 15 percent with the higher uncenainty at the lower flow rate. Physical properties, with the exception of diffusivity, are assumed to have an uncertainty of 1 percent or less. Diffusivity appears to have an uncenainty of about 10 percent based on a comparison of values from three different textbooks. This will have a relatively large impact on the Sherwood number measurement uncertainty, Clearly, the delta-T measurement uncertainty is the main contributor to the measured Nusselt number uncertainty and the liquid film flow rate measurement uncertainty is the main contributor to the una;wou s65-non:sta to95 4-12 j
e [ f measured Sherwood number uncertainty. 'Iherefore, the error bar on the measured data will be longer at locations with low heat fluxes (located near the bottom of the vessel for the wet large-scale PCS I tests) and for tests with low liquid film flow rates. "Ihe data scatter for the large-scale PCS tests fits j this expected trend. The data are scattered over a wider range near the entrance where the heat flux is ; I lower. i t l [ 1 F k i i l l I i l
.I I
l l i a:W1865-ace:ltW1095 4-13
l l l
5.0 CONCLUSION
S l l The heat and mass transfer correlations that are used in the ,WGOTHIC code to model heat removal l from the AP600 containment have been presented in this report. These correlations have been widely used and accepted for the calculation of heat transfer for similar geometries. A number of separate effects tests utilizing geometries representative of the AP600 have been j examined. These tests cover the range of expected conditions for heat and mass transfer within the AP600. The method of calculating the heat and mass transfer coefficients in the WGOTHIC code has been compared with these test data and yields acceptable results. The comparisons to the available test data show that the calculated local heat transfer coefficients demonstrate the proper trends over the entire operating range of the appropriate dimensionless groups for AP600. Die capability to model the AP600 DBA pressure and temperature response and associated uncertainties in an integral setting will be assessed using results from the large-scale tests, and considered when performing DBA analysis. l uAap60lA1865-non:IM41095 5-1
6.0 REFERENCES
i
! 1. McAdams, W. H., Heat Transmission, Third Edition, McGraw-Hill,1954.
- 2. Colburn, A. P., "A Method of Correlating Forced Convection Heat Transfer Data and a Comparison With Fluid Friction," Transactions of the AIChE, Vol. 29 1933, p.174. ;
- 3. Churchill, S. W., " Combined Free and Forced Convection Around Immersed Bodies" (Section 2.5.9) and " Combined Free and Forced Convection in Channels" (Section 2.5.10),
Heat Exchanger Design Handbook, Hemisphere Publishing Corp.,1983.
- 4. Eckert, E. R. G. and A. J. Diaguila., " Convective Heat Transfer for Mixed, Free, and Forced Flow 'Dirough Tbbes," Transactions of the ASME, May,1954, pp. 497-504.
[
- 5. Boelter, L. M. K., G. Young, and H. W. Iverson, NACA TN 1451,1948.
- 6. Siegel, R. and R. H. Norris, " Test of Free Convection in a Partially Enclosed Space Between Two Heated Vertical Plates," Transaction of the ASME, Journal of Heat Transfer, April 1957.
- 7. Hugot, G., " Study of the Natural Convection Between Two Plane, Vertical, Parallel, and Isothermal Plates," derived from doctoral dissertation University of Paris,1972, translated by D. R. de Boisblanc, Ebasco Services Incorporated, June 1991.
- 8. Schlichting, H., Boundary layer Theory, Sixth Edition, McGraw-Hill.
- 9. Vliet, G. C., " Natural Convection Local Heat Transfer on Constant-Heat Flux Inclined Surfaces," Journal of Heat Transfer, November 1969, pp. 511-516.
- 10. Woodcock, J., et. al., Westinghouse-GOTHIC: A Computer Codefor Analysis of Thermal Hydraulic Transientsfor Nuclear Plant Containments and Auxiliary Buildings, WCAP-13246, July 1992.
- 11. Burmeister, L. C., Convective Heat Transfer, John Wiley & Sons,1983.
- 12. Heavy Water Reactor Facility (HWRF) Inrge Scale Passive Containment Cooling System Confinnatory Test Data Report, HWRF-RPT-93-001, July 1993.
- 13. Gilliland, E. R. and T. K. Sherwood, " Diffusion of Vapors into Air Streams," Industrial and l Engineering Chemistry, Vol. 26, No. 5, pp. 516-523.
I l I 1 l l u:\ap600\1865-noo:Ib-041&)5 6-1 l _____ _ - - - - - - - _ - - - ----__---------_----_----------------.-----------------------------------------_-.l
b k-l 14. Westinghouse Electric Corporation, Tests of Heat Transfer and Water Film Evaporation on a Heated Plate Simulating Cooling of the AP600 Reactor Containment, PCS-T2R-001, > April 1992.
- 15. Westinghouse Electric Corporation, Condensation in the Presence of a Noncondensible Gas:
ExperimentalInvestigation, PCS-TZR-009, April 1991.
- 16. Final Data Reportfor PCS Large-Scale Tests, Phase 2 and Phase 3, WCAP-14135 (Proprietary) PCS-12R-032, July 1994.
- 17. Peterson, P. F., V. E. Schrock, T. Kageyama, " Diffusion Layer 'Iheory for Turbulent Vapor Condensation with Noncondensible Gases," Journal of Heat Transfer, November 1993, Vol.115, pp. 998-1003.
- 18. Jakob, M., Heat Transfer, John Wiley & Sons,1%7. -
I u.W1865-son:lt>o41095 6-2 _ _}}