ML18086B273
| ML18086B273 | |
| Person / Time | |
|---|---|
| Site: | Salem  |
| Issue date: | 01/31/1981 |
| From: | Nystrom J ALDEN RESEARCH LABORATORY |
| To: | |
| Shared Package | |
| ML18086B272 | List: |
| References | |
| 24-18-M302LM, 24-81-M302LM, NUDOCS 8201290088 | |
| Download: ML18086B273 (64) | |
Text
EXPERIMENTAL EVALUATION OF FLOW PATTERNS IN AN RHR SUMP WITH SIMULATION OF SCREEN BLOCKAGE SALEM NUCLEAR GENERATING STATION by James B. Nystrom Research Sponsored by Public Service Electric & Gas Company
/A\\c-n 1 ALDEN RESEARCH LABORATORY b~ WORCESTER POLYTECHNIC INSTITUTE January 1~
98~
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PDR 24-8 1 /M302LM
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EXPERIMENTAL EVALUATION OF FLOW PATTERNS IN AN RHR SUMP WITH SIMULATION OF SCREEN BLOCKAGE SALEM NUCLEAR GENERATING STATION by James B. Nystrom Research Sponsored by Public Service Electric & Gas Company George E. Hecker, Director ALDEN RESEARCH LABORATORY WORCESTER POLYTECHNIC INSTITUTE HOLDEN, MASSACHUSETTS January 1981 J
';:/
Printed at ARL -
March 1981
ABSTRACT A hydraulic model of the containment building sump for the Salem Nuclear Generating Station, Unit 2, was constructed at a scale.of 1:2.8.
The re-sidual heat removal pumps withdraw water from the sump after a postulated loss of coolant accident for re-injection into the core. To assure accept-able operation of the RHR pumps, the model was tested for a wide range of possible approach flow distributions, bar rack and screen blockage schemes, water depths, and pump operation combinations.
The tests were designed to assure that no air entraining vortices were formed, head losses across the screens and in the inlet were acceptable, and the swirl in the inlet pipe was acceptable.
Test results indicated that the maximum swirl angle was 2.25 degrees, which is less than that caused by a short radius 90 degree bend.
For a flowrate of 4800 gpm per line, loss measurements indicated an average pipe inlet loss of 0.38 ft and bar rack and screen losses ranging from 0.04 ft for clean screens to 0.36 ft for the worst case of 50 percent blockage
- Vortex persistence was determined for six non-uniform approach flow distri-butions and twelve bar rack and screen blockage configurations using both Froude scale velocity and prototype velocity.
The vortex suppression grid at elevation 77 ft 0 inches was effective in breaking up vortex activity and no coherent vortices were seen below the grid.
At no time during the test program were air core vortices evident.
Froude scale velocity result-ed in vortex activity limited to dye core vortices.
Prototype velocity tests showed an increase in vortex activity for some blockage schemes. For example, four screen blockage schemes caused air bubble entraining vorti-cies existing for about 20 percent of the time.
Except for one of these cases, the amount of air entrained in the downward flow below elevation 77 ft O inches was small.
Other conditions tested, such as increased and decreased water levels and one pump operation, decreased vortex activity.
I...
ii Since trash and air bubble entraining vortices were noted with the original design, a retrofitted anti-vortex baffle was tested at elevation 80.ft 7 inches. Tests with prototype velocity indicated all vortex activity greater than type 3, coherent dye core, was eliminated, and no air bubble entrain-ment occurred.
Combination of non-uniform approach flow and screen block-age was found to have essentially the same vortex activity as the screen blockage with a uniform approach flow.
iii TABLE OF CONTENTS ABSTRACT TABLE OF CONTENTS INTRODUCTION PROTOTYPE DESCRIPTION SIMILITUDE Froude Scaling Similarity of Vortex Motion ARL Vortex Activity Projection Technique Dynamic Similarity of Flow Through Screens MODEL DESCRIPTION INSTRUMENTATION AND OBSERVATION TECHNIQUES Flow Measurement Pressure Gradelines Pipe Swirl Vortex Activity Observation of Flow Patterns TEST PROCEDURE TEST RESULTS Vortex Activity with Normal Water Level Froude Velocity Scale Prototype Velocity Tests Vortex Activity with Varying Water Level and Number of Pumps Inlet and Screen Head Losses Swirl Angle Measurements Final Design Test Results Swirl Angle Measurements Combine Blockage and Non-Uniform Approach
SUMMARY
REFERENCES FIGURES PHOTOGRAPHS Page No.
i iii 1
2 4
6 8
10 11 14 16 16 16 16 17 17 18 19 19 20 21 22 24 26 27 29 30 32 34
. INTRODUCTION The reactor containment building of the Salem Generating Station, Unit 2, is provided with a residual heat removal (RHR) system designed to cool the shutdown reactor core and the containment in the event of a Loss of Coolant Accident (LOCA). Water is injected to maintain core cooling and, initially, this water is drawn from the refueling water storage tank.
When the water level in this tank reaches a predetermined level, the residual heat removal system is switched from the injection mode to the recirculation mode.
At this point, wat;er is drawn from the containment sump, which then contains*
water drained from the break and from the containment spray system.
Flow approaching the sump is affected by the geometry of the flow path including various appurtenant structures and equipment.
Water level, pump discharge, and water temperature could vary.during the recirculation mode, which lasts for an extended period to. provide sufficient heat removal.
- The Alden Research Laboratory -CARL) of Worcester Polytechnic 1nstitute (WPI) was authorized by Public Service Electric and Gas Company (PSE&G) to
\\.*
construct and test a model of the Salem Generating Station containment* sump with the object of investigating free surface vortex formation, swirl in the inlet piping, inlet losses, or any other flow conditions that could adverse-ly affect the performance of the residual heat removal *(RHR) pumps in *the recirculation mode.
Operating conditions involving a wide range of possible approach flow distributions, screen blockages (due to debris) 1 water depths, and pump operating combinations were tested in the model.
This report presents the. findings of the study and includes a description of the prototype and the model, and summarizes conditions investigated, similitude considerations, test procedures, instrumentation, and inter-pretation of results.
2 PROTOTYPE DESCRIPTION The entrance to the containment sump for the Salem Generating Station, Unit 2, is located at elevation 78 ft O inches in the reactor containment build-ing close to the containment wall, as shown in Figure 1.
The sump, shown in plan in Figure 2 and cross-section in Figure 3, encompasses a plan area
\\
7 ft 6 inches by 6 ft and descends to elevation 70 ft.
Vertical screens, gratings, and a plate partition the sump into three separate areas.
The vertical plate begins l ft 2 inches below the top of the sump and divides the sump into two 7 ft 6 inch by 3 ft areas, one containing two residual heat removal (RHR) pump suction outlets and the other accepting a 16 inch drain line from the building drain trenches.
The RHR sump is further di-vided into two equal areas by a vertical grating, constructed of 1-1/4 inch deep by 3/16 inch thick bars on 1-3/16 inch centers.
Within each RHR sump, two horizontal grids designed to suppress vortices and made from bars 1-1/4 inch deep by 3/16 inch thick on 1-3/16 inch centers are located at elevations 71 ft 6 inches and 77 ft 0 inches.
Vertical fins, 4 inches deep by 1/4 inch thick on 4 inch centers, are located on the periphery of the sump and extend from elevation 70 ft O inches to 71 ft 6 inches.
The two 17 inch inside diameter RHR suction pipes exit the bottom of the sump (elevation 70 ft 0 inches)* and descend vertically approximately 10 dia-meters before reducing to 14 inch diameter and continuing horizontally with various bends to the RHR pumps.
The maximum flow in the suction lines is 4800 gpm/line.
The minimum water level for operation of the RHR pumps is elevation 81 ft 9 inches, resulting in a submergence of approximately 11 ft 9 inches at the suction inlets.
A 9 inch high curb surrounds the sump and supports a trashrack and screen, both designed to assure no debris is entrained in the pump system.
A 3 ft high screen, 5 ft 0 inches by 9 ft 6 inches in plan, is centered over the RHR sump.
The screen has a 1/8 inch mesh and the top surface is screen covered as well as the sides.
The trashrack is 3 ft 2 inches high and
3 8 ft 6 inches by 11 ft 6 inches in *plan,*and is made of 1-1/4 inch deep bars 3/16 inches thick on 1-3/16 inch centers.
The top of the trashrack is a solid plate to protect the sump from falling debris.
A horizontal anti-vortex grate was retrofitted inside the screens at elevation 81 ft 5-1/2 inches, as shown in Figure 15.
The grate had the same bar geometry as the trashrack.
The 16 inch drain line enters the other half of the sump horizontally at elevation 76 ft 10 inches.
Normally, sump pumps remove.any small leakage*
flow to a storage vessel. A 1/8 inch mesh screen above the top of the ver-tical plate dividing the sump allows larger.flows to enter*the*RHR sump.
In the recirculation mode after a postulated loss-of-coolant accident, water primarily approaches the sump by *flowing through four openings in the inner containment (crane) wall into the annular channel formed by the crane and containment walls, in which the containment suinp is located {see Figure 1 for flow paths).
The flow path is relatively unobstructed close to the sump, except for an elevator shaft located near the side of the sump.
A stairway and various vertica*1 support columns provide only minor obstructions to the flow path.
A secondary source of flow to the RHR pumps is the drain line from the inner and outer containment drain trenches.
_J
4 SIMILITUDE The study of dynamically similar fluid motions forms the basis for. the design of models and the interpretation of experimental data.
The basic concept of dynamic similarity may be stated as the requirement that* two systems with geometrically similar boundaries have geometrically similar flow patterns at corresponding instants of time (3).
Thus, all indivi-dual forces acting on corresponding fluid elements of mass must have the same ratios in the two systems.
The condition required for complete similitude may be developed from Newton's second law of motion:
where F.
1 F p F g F v Ft F.
1 inertia force, defined as mass, M, times the acceleration, a pressure force connected with or resulting from the motion gravitational force viscous force force due to surf ace tension (1)
Additional forces may be relevant under special circumstances, such as fluid compression, magnetic or Coriolis forces, but these had no influence on this study and were, therefore, not considered in the following develop-ment.
Equation (1) can be made dimensionless by dividing all the terms by F..
1 Two systems which are geometrically similar are dynamically similar if both satisfy the dimensionless form of the equation of motion, Equation (1).
We may write each of the forces of Equation (1) as:
where F p net pressure x area 5
F specific weight x volume g
F v F.
l shear stress x area a 3 µ ~u/~y x area surf ace tension x length density x volume x acceleration a 1, a 2, etc. = proportionality factors L
representative linear dimension
~p net pressure y
specific weight
µ dynamic viscosity surface tension p -
density u
representative velocity Substituting the above terms in Equation (1) and making it dimensionless by dividing by the inertial force, we obtain
-2 E
+
-2 F
+
~1 R
+
ct4 as
-2 w
1 (2)
6 where u
Inertia Force E
Euler number; Pressure Force lf':.p/p u
Froude number; Inertia Force F
Gravity Force v'gL u L Reynolds number; Inertia Force R
µ/p Viscous Force w
u Weber number; Inertia Force
/a/pL Surface Tension Force Since the proportionality factors, a., are the same in model and prototype, l
complete dynamic similarity is achieved if all the dimensionless groups, E, F, R, and W, have the same values in model and prototype.
In practice, this is difficult to achieve.
For example, to have the values of F and R the same requires either a 1:1 "model" or a fluid of very low kinematic viscosity in the reduced scale model.
Hence, the accepted approach is to select the predominant forces and design the model according to the appro-priate dimensionless group.
The influence of the other forces would be secondary and are called scale effects (2, 3).
Froude Scaling Models involving a free surface are constructed and operated using Froude similarity since the flow process is controlled by gravity and inertia forces.
The Froude number, representing the ratio of inertia to gravita-tional force, F
u/v'g'S
( 3)
~
- ~
lo.*.
where 7
u average velocity in the pipe g
gravitational acceleration s = submergence was, therefore, made equal in model and prototype.
F r F /F m P 1
( 4) where m, p, and r denote model, prototype, and ratio between model and prototype, respectively.
In modeling of an intake sump to study the formation of vortices, it is important to select a reasonably large geometric scale to achieve large Reynolds numbers and to reproduce the curved flow pattern in the vicinity of the intake (4). A geometric scale of L L /L = 1/2.8 was chosen for r
m p
the model, where L refers to length.
At sufficiently high Reynolds num-ber, an asymptotic behavior of energy loss coefficients with Reynolds number is usually observed (2).
Hence, with F
= 1, the basic Froudian r
scaling criterion, the Euler numbers, E, will be equal in model and pro-totype.
This implies that flow patterns and loss coefficients are equal in model and prototype at sufficiently high Reynolds numbers.
From Equa-tions (3) and (4), using s r
were:
u r Qr L r t r 2
L, the velocity, discharge, and time scales r
L 0.5 1/12:8 r
(5)
L 2.5 1/(2.8) 2 "5 u r r
(6)
L 0.5 1/12:8 r
(7)
8 Similarity of Vortex Motion The fluid motions involving vortex formation in the sumps of lpw head pump intakes have been studied by several investigators (1, 4, 5, 6).
Anwar (4) has shown by principles of dimensional analysis that the dynamic simi-larity of fluid motion in an intake is governed by the dimensionless para-meters given by where Q
d 4Q u
2 '
u 8 d hgs
_Q_
d
, and v s 2s discharge through the outlet tangential velocity at a radius equal to that of outlet pipe diameter of the outlet pipe Surface tension effects were neglected in his analysis, being negligible for weak vortices.
The influence of viscous effects was defined by the parameter Q/(v s), known as a radial Reynolds number, RR.
For similarity between the dimensions of a vortex of strengths up to and including a narrow air-core type, it was shown that the influence of RR becomes negligible if Q/(v s) was greater than 3 x 104 (4).
As strong air-core type vortices, if present in the model, would have to be eli-minated by modified sump design, the main concern for interpretation of prototype performance based on the model performance would be on the similarity of weaker vortices, such as surface dimples and dye-cores.
For the prototype of the present study, the values of RR for the oper-ating temperature ranges of 70° to 165°F, and using the submergence to 5
5 the first vortex suppression grid, ranged from 2.1 x 10 to 5.6 x 10.
The value of RR for the model was always greater than 3.2 x 104 for water temperatures of 40°F and above.
Thus, viscous forces would have
9 only a secondary role in the present study.
If so, dynamic similarity is obtained by equalizing the parameters 4Q/u8d 2, u/l2gs, and d/2s in model and-prototype.
A Froudian model would satisfy this condition.
Viscous and surface tension forces could influence the formation and strength of vortices (1, 5).
The relative magnitude of these forces on the fluid inertia force is reflected in the Reynolds and Weber numbers, respectively, which are defined as:
R u d/v (8) w u
(cr/ps)l/2 (9)
It was important for this study to ascertain any deviations in similitude attributable to viscous and surface tension forces in the interpretation of model results to prototype conditions.
Surface tension effects were considered negligible inasmuch as strong vortices were unacceptable, and the. free surface was essentially flat for all final tests.
Moreov_er, an investigation using liquids of the same viscosity but different surface tension coefficients (cr = 4.9 x 10-3 lb/ft to 1.6 x 10-3 lb/ft) showed practically no effect of surface tension forces on the vortex flow (1).
The vortex severity, S, is therefore mainly a function of the Froude num-ber, but could also be influenced by the Reynolds number.
S
==
S (F, R)
(10)
For consistent observations, it is convenient to clas_sify the free surface vortices from a swirl to an air core type vortex, as shown in Figure 4.
To compensate for any possible excessive viscous energy dissipation and consequently less intense model vortex, various investigators have pro-posed increasing the model flow and, therefore, the velocity, keeping the submergence constant.
Operating the model at the prototype inlet
10 velocity (pipe velocity) is believed by some researchers to *achieve the desired results (1).
This is often referred to as Equal Velocity Rule, and is considered to give conservative predictions of prototype perfor-mance.
The test procedure for the present study incorporated testing the model at prototype pipe velocities.
ARL Vortex Activity Projection Technique ARL has conducted an extensive research program to assure that the conclu-sions regarding vortex activity in the model are valid for the prototype.
A technique of extrapolating model vortex activity to prototype Reynolds numbers (17) by using elevated model water temperatures and varying model flow velocity (Froude ratio) has been applied to several studies (7, 12, 18, 19). Figure 5 illustrates the method used to investigate scale effects and predict vortex types in the prototype based on model results (7).
The ordinate, F, is the ratio of model to prototype Froude number, while the r
abscissa is the inlet pipe Reynolds number, R. Assume the model to operate at flow less than Froude scaling (Fr< 1) at point a 1.
By increasing the discharge in the model while keeping the same submergence and temperature, Fr and Rare increased corresponding to a point, aN, where a vortex of type N was first observed.
The model Reynolds number can also be changed by varying the kinematic viscosity with temperature changes, and similar tests performed to locate bN, another point on the locus of type N vorti-ces.
Extrapolation *of the line of constant vortex strength of type N can be made to a prototype Reynolds number at the proper Froude number (F r
1), point pN.
The locus could represent any expedient measure of vortex severity.
Any scale effects due to viscous forces would be evaluated and taken into account by such a projection procedure.
The high temperature-high flow tests were used in the similar fashion for projecting the inlet loss coefficients (from the pressure gradient measurements) and the swirl severities (from vortimeter readings) over a wide range of Reynolds and Froude numbers.
11 Experience has shown that incoherent swirling flow is even less dependent on Reynolds number than a coherent vortex core.
Eliminating the tendency for coherent vortices axiomatically removes possible scale effects.
In reactor sumps, the design criteria eliminate the possibility of coherent vortex cores in an acceptable design.
Figure 6 shows the results of one typical recirculation sump model (18). As can be seen from the data, which are for the final design with vortex suppressor grids, there are no mea-surable changes in vortex strength with Reynolds number.
Therefore, it is concluded that no scale effects will be present in the final design.
Dynamic Similarity of Flow Through Screens In addition to providing protection from debris, screens tend to suppress non-uniformities of the approach flow. The aspects of flow through screens of concern in a model study are:
(1) energy loss of the fluid passing through the screen; (2) modification of velocity profile and the deflec-tion of streamlines at the screen; and (3) production of turbulence.
As all these factors could affect vortex formation in a sump with approach flow directed through screens, a proper modeling of screen parameters is important.
The loss of energy across the screen occurs at a rate proportional to the drop in pressure, and this loss dictates the effectiveness of the screen in altering velocity profiles.
The pressure drop across the screen is analogous to the drag induced by a row of cylinders in a flow field and could be expressed in terms of a pressure-drop coefficient K (or alter-nately a drag coefficient), defined as (8),
K lip liH (11) 1/2 pp u2 2
u /2g
12 where
~p drop in pressure across the screen U
mean velocity of approach flow p
density of the fluid
~H head loss across the screen g
acceleration due to gravity From the available literature on the topic (8, 9, 10), it may be seen that where R s S'
Pattern K
f(R, S', Pattern) s screen Reynolds number, U d~v, d being the wire diameter of the screen w
solidity ratio, equal to the ratio of closed area to total area of screen geometry of the wire screen (12)
If the solidity ratio and the wire mesh pattern are the same in the model and prototype screens, the corresponding values of K would only be a func-tion of the screen Reynolds number.
This is analogous to the coefficient of drag in the case of the circular cylinder.
It is known that K becomes practically independent of R at values of R greater than about 1000 (8, s
s 11).
However, for models with low approach flow velocity and with fine wire screens, it is necessary to ascertain the influence of R on K for s
both the model and prototype screens before selecting screens for the model which are to scale changes in velocity distribution.
Velocity modification equations relating the upstream velocity profile and downstream velocity profile have been derived based on different theories (8).
Most of these indicate a linear relationship between the upstream velocity profile and downstream velocity profile, shape and solidity ratio
13 of screen, and value of K.
If the wire.shape and solidity ratios are the same in the model and prototype screens, it is possible to select a suit-able wire diameter to keep the values of K approximately the same for the model and prototype screens at the corresponding Reynolds number ranges.
Identical velocity modifications would be produced by the respective screens if the loss coefficients were identical.
The pressure loss coefficient to Reynolds number relationship of fine screens have been investigated at ARL (12).
Based on the similarity of pressure loss and velocity modifications, an appropriate model screen was chosen, which in this case was the same as the prototype fine screen.
14 MODEL DESCRIPTION The model was constructed to a geometric scale of 1:2.8 with boundaries as indicated in Figure 7.
Model boundaries were chosen at approximate loca-tions where the flow pattern would be controlled in the prototype and based on previous studies (7, 12, 18, 19), which indicated little effect of far-field flow patterns once screen blockage is considered.
Screen blockage has consistently generated the most severe vortices and swirl in the numerous past ECCS sump studies at ARL.
The model was located in an existing elevat-ed tank and inflow was provided from a sump beneath the model by a vertical mixed flow pump.
Water level was controlled by an adjustable weir.
Flow straighteners at the model boundaries provided a uniform initial velocity distribution with relatively low turbulence levels.
Portions of the proto-type structure with outside dimensions greater than 3 inches, such as pipes, columns, conduit supports, and a stairway, in the immediate vicinity of the sump and below the water surface were modeled to the geometric scale.
Portions of the inner and outer drain trenches were modeled.
An analytic model of both trenches was used to predict the flowrates and lOsses within the trenches such that an equivalent head loss could be included at the end of the modeled trenches, forcing the total flow in the modeled trenches to be correct.
For ease of construction, the total area of the holes in the drain trench covers was modeled with 75% fewer holes of larger scaled dia-meter.
A separate tank was constructed for the inner drain trench so that the water level could be independently controlled at the level calculated by considering the head losses for the flow paths from inside the crane wall to the annulus in which the RHR sump is located.
The model was constructed using a combination of wood, steel, and clear acrylic, which allowed observation of flow patterns. The 17 inch ID verti-cal suction pipes were modeled for 8 pipe diameters, had access ports for vortimeter installation, and had 6 sets of piezometers for pressure grade-line measurement.
The clear acrylic sump is shown in Photograph 1 with the vertical RHR suction lines and piezometers.
ASME standard orifice flowmeters were provided to measure flow in each suction line.
Ir
15 As previous ARL basic studies (20) have shown~ geometrically scaled gratings are as effective in suppressing model vortices as are the prototype gratings in suppressing prototype vortices, therefore, all gratings employed in this study were geometrically scaled. Screens used in the model were of prototype dimensions; as previously discussed.
Therefore, the model screen Reynolds number and percentage open area, which have been shown to control head loss and effects on flow patterns (21) approximated the prototype values closely.
Photograph 2 shows the model vertical grating and screens installed.
The horizontal gratings at elevations 77 ft and 71 ft 6 inches are shown in Photograph 3.
The effects of asymmetrical approach flow and grating and screen blockages, were studied using 50% blockage at appropriate locations.
As the sump is located outside the crane wall and is, therefore, protected from direct im-pingement of breakflow jets, simulation of air entrainment from breakflows was not considered.
16 INSTRUMENTATION AND OBSERVATION TECHNIQUES Flow Measurement Flowrates were measured by ASME standard orifice meters using air-water manometers for differential pressure measurement.
Pressure Gradelines Each pressure gradeline in the suction line was measured by a pair of piezo-meters at six locations in each pipe using air-water manometers with the sump water level as reference pressure.
The pressure gradeline was extrapolated to the entrance by a linear least squares (linear regression) curve fit of the pressure measurements.
The area average velocity was used to calculate the pipe velocity head, which was added to the extrapolated pressure grade-line.
The total head within the sump was determined from a pressure measure-ment and the velocity head at that location.
The pipe total head was sub-tracted from the sump total head to determine the inlet loss.
An entrance loss coefficient was calculated by:
where Pipe Swirl Lrn.
K 1
2 v mean K
loss coefficient
~H.
inlet head loss, ft
].
2g (13)
Average swirl in the suction pipes was measured by cross vane vortimeters.
Studies at ARL (22) have shown that a vortimeter'with vane diameter 75%
17 that of the pipe diameter best approximates the solid body rotation of the flow.
The rate of rotation of the vortimeter was determined by counting the number of blades passing a fixed point in one minute.
An average swirl angle was defined as the arctangent of the maximum tan-gential velocity divided by the axial velocity.
The maximum tangential velocity of the vortimeter is the circumferential path travelled by blade tip per unit time, TI D N, and the average swirl angle is defined by:
where N
D v mean Vortex Activity e
arc tan (TI D N) v mean revolutions per second rotameter diameter, ft mean axial velocity (14)
Vortex activity was recorded by observing vortex strength on a scale from l to 6 (see Figure 4), and by determining the percent of time that each strength persisted.
Vortex strength was identified by using dye injection and addition of "trash" consisting of a slightly buoyant ball of paper.
Persistence of a vortex type was recorded as a voltage level on a chart recorder.
Percentage of time of given vortex type was determined from the chart recording to define persistence.
Observation of Flow Patterns Visual aids, such as dye, were used to observe flow patterns. Photographic documentation was taken whenever appropriate.
18 TEST PROCEDURE Tests were conducted at the normal laboratory water temperature. The mo-del was filled to an appropriate level, and all piezometer and manometer lines were purged of air and zero flow differentials checked.
The re-quired flowrates were then set and the water level allowed to stabilize.
The water level was checked and adjustments made if required and flow-rates were re-checked and re-adjusted, if necessary.
A 15 minute mini-mum settling time was allowed prior to initiation of the data recording.
Fifteen minutes of vortex observations were recorded and the required physical parameters, such as depth, manometer deflections, and vortimeter readings, were recorded.
Entrance losses were determined with the vorti-meters removed from the suction lines.
19 TEST RESULTS Vortex Activity with Normal Water Level -
81 ft, 9 inches Vortex activity was determined for uniform approach flow and clean screens, 6 simulated non-uniform approach flow schemes (shown in Figures 8 through 10), and 12 screen blockage schemes (shown in Figures 11 through 13).
All schemes utilized 50 percent blockage.of the total flow area and in the screen blockage cases, both the bar rack and the screens were blocked simultaneously.
Initial tests were conducted to determine the sensitivity of vortex per-sistence to the inner drain trench flowrate.
The initial tests were con-ducted with the water level in the simulated central portion of the con-tainment building set to that calculated based on the losses in the flow path.
Further tests were conducted and the water level was then increas-ed by 0.15 ft over the calculated value, which corresponded to the maximum velocity head in the flow path.
Test results varied slightly, but were within the probable scatter of the measurements, therefore, all remaining tests at water surface elevation 81 ft 9 inches were conducted with the calculated water level in the inner drain trench.
At no time during the test program were air core vortices noted.
All dye core vortices were disorganized by the horizontal grid at elevation 77 ft 0 inches and no coherent vortices were seen below that elevation.
Data on vortex persistence were difficult to obtain with high accuracy, especially for tests at prototype velocity scale, due to higher turbu-lence levels than at Froude scaled flow.
Persistence results are some-what subjective and, therefore, a single observer was used throughout the test program to achieve consistent results.
Accurate persistence data for vortices entraining air bubbles were especially difficult to establish due to the small quantities of air entrained and their tran-sient nature.
20 Table 1 surrunarizes the vortex persistence for the normal water level of 81 ft 9 inches, both pumps in operation, and both Froude and prototype velocity scale ratios.
TABLE l Vortex Persistence (Percent)
Vortex Strength Froude Velocity Prototype Velocity Scaling Scaling Blockage 1
2 3
4 5
l 2
3 4
5 None l
23 77 88 12 Non-Uniform 2
l l
86 11 2
Approach 3
4 34 62 l
6 74 16 4
Flow 4
l 18 81 4
36 45 9
10 II 5
6 23 71 l
14 47 26 12 II 6
l 4
45 38 13 II 7
l 21 53 19 7
Rack and 8
0 100 0
100 Screen 9
0 100 0
100 Blockage 10 0
100 0
24 48 11 17 11 0
0 100 0
6 48 17 29 12 0
31 69 0
7 72 3
18 13 0
91 9
0 0
73 4
23 14 0
45 55 0
0 100 15 0
100 0
91 9
16 0
61 39 0
12 88 17 4
19 75 2
0 2
95 3
18 26 31 43 0
73 27 19 12 75 13 0
56 36 3
5 Froude Velocity Scale The uniform approach flow case with no screen blockage (scheme 1) had only surface dimples or surface swirls. The three non-uniform approach flow schemes had a combination of surface dimples and dye core vortices.
For Froude velo-city scale, only one screen blockage scheme, number 17, caused vortex strength greater than a dye core (type 3), and that was a type 4 trash pulling vortex I
21 with persistence of 2 percent.
For the screen blockages,cfour schemes (8, 9, 10, 15) caused only surface dimples (type 2), while one scheme (ll) had a continuous dye core vortex.
The remaining schemes had some combination of dye core vortices and surface dimples.
Prototype Velocity Tests Using prototype velocity necessarily exaggerated the losses through the ver-tical grating and screens by the scale ratio, 2.8.
In cases of clean screens and aligned blockages, this effect was small, but some staggered blockages re-stricted flow paths parallel to the screen face caused losses to be near 1 ft prototype.
This decrease in depth within the screens caused higher velocities and increased vortex activity.
The increase to prototype velocity generally caused an increase in vortex strength. For the case of clean screens, a vortex pulling trash was evident about 10 percent of the time with a dye core the remainder of the time.
Non-uniform approach flow cases (schemes 2 through 7), which caused dye core vortices (type 3) in Froude velocity scale tests (schemes 3 through 5 only),
caused predominantly type 3 vortices at prototype velocity scale, but type 5 air bubble entraining vortices were noted about 10 percent of the time, and trash pulling vortices (type 4) had persistences of from 10 to 40 percent.
For screen blockage schemes where only surface dimples (type 2) were evident in Froude velocity scale tests (schemes 8, 9, 15), little change generally occurred for the prototype velocity cases.
The exception was scheme 10, where air bubble entraining vortices (typ~ 5) were evident.
In some cases, the maxi-mum strength did not increase (14, 16, 18), but persistence increased.
Scheme 17 had a trash pulling vortex, type 4, of less than 3 percent persistence in both cases.
In the remainder of the cases, some vortices entraining air bub-bles were noted.
In schemes 10, 11, 12, and 13, air entraining vortices were evident approximately 20 percent of the time, but only in scheme 12 were rela-tively significant amounts of air noted in the lower portion of the sump.
Be-neath the horizontal grid at elevation 77 ft, the air was not carried downward by vortex action, but the vertical velocity was sufficient to carry the bubbles downward.
22 The cases of screen blockage which entrained air bubbles generally had one or more of the inner screen corners blocked and lateral approach flow along the face of the screen.
The type 5 vortices generally formed within the blocked corner and were small in size. A large central eddy of strength 2 rotated in the opposite direction and reinforced the corner vortices.
Although the amount of air entrained during strength 5 vortices was small in the model, it is difficult to derive accurate scaling laws for.. the per-centage of air entrained.
Vortex Activity with Varying Water Level and Number of Pumps A few blockage schemes were tested for one pump operation.
Table 2 shows the vortex persistence results with one pump operation and includes compar-able two pump tests.
Since the two pump inlets are separated by only a ver-tical grid, which provides little hinderence to flow, vertical flow at the top of the sump (elevation 78 ft 0 inches) essentially fills the entire cross-section. Therefore, operation of a single pump is very similar to de-creasing the overall flowrate.
Since the ratio between Froude velocity scale and prototype velocity scale is 1 to 1.67, results with prototype velocity scale for one pump should produce results similar to Froude velocity scale tests with two pumps operating.
Essentially no adverse conditions were found for Froude velocity scale tests and decreasing the overall flowrate would de-crease vortex activity.
Therefore, only prototype velocity scale was used for single pump operation.
TABLE 2 One Pump Operation Prototype Velocity Scaling Blockage Water Level 1
2 3
4 5
1 81'-9" 0
0 88 12 Two pumps 1
81'-9" 0
52 48 East pipe only 1
81'-9" 0
37 63 West pipe only
. 13 81'-9" 0
0 73 4
23 Two pumps 13 81'-9" 23 52 20 4
1 West pipe only
23 With clean screens (scheme 1), no trash pulling vortic.es were evident and dye core vortices existed only 50 percent of the time.
Using the most ad-verse screen blockages, determined from two pump tests, resulted in sub-stantial decreases in persistence related to the two pump operation.
Table 3 shows the effect using prototype velocity of increasing the water depth l ft to elevation 82 ft 9 inches, which generally decreased vortex activity.
Since the top screens were in place, it was difficult to ob-serve any possible vortex action within the screen, but no obvious bubble entraining vortices were noted and no air was drawn into the sump.
Vortex activity outside the vertical trashrack was generally of the dye core type (type 3), but minor persistences of trash pulling and air bubble entrain-ing vortices were noted.
Blockage 11 11 12 12 13 13 18 18 TABLE 3 Increased Water Depth Prototype Velocity Water Level l
2 3
4 81'-9" 0
6 48 17 82'-9" 0
100 81'-9" 0
7 72 3
82'-l" 4
84 8
0 81'-9" 0
0 73 4
82'-9" 0
100 81'-9" 0
73 27 82'-8" 0
100 Scaling 5
29 18 4
23 Decreasing the water level abqut 9 inches to 81 ft 0 inches tended also to decrease activity for Froude scaie flow as shown in.Table 4.
Proto-type velocity scaling was not used at water levels lower than normal since non-representative head losses were produced by the screens, which resulted in water levels within the sump substantially below that out-side* the sump.
In some cases (scheme 13), the water level inside the
- -1~
24 sump was about 1 ft lower than the outside level, which approximated the Froude velocity scale tests with lowered water level. Attempts at lowering the water level 2 ft with Froude velocity scaling resulted in very turbu-lent flow in the sump, which tended to entrain bubbles without vortex acti-vity and, in fact, did not allow coherent vortices to form.
This phenomena was accentuated with prototype velocity scaling.
TABLE 4 Decreased Water Depth Blockage Water Level 10 81'-9" 10 81'-0" 12 81'-9" 12 81'-0" 13 81'-9" 13 81'-l" Inlet and Screen Head Losses Froude Velocity Scaling 1
2 0
100 0
100 0
31 3
90 0
91 0
100 3
69 6
9 4
5 Measurements of the head losses through the vertical bar rack, screens, and pipe inlet were made for several screen blockage geometries.
The velocity head of the approach flow was neglected such that the measured water level outside the screens was assumed to be the initial total head.
Static head in the sump was measured at elevation 376 ft and the velocity head in the sump, calculated using the area average velocity, was added to the static head to determine the total head.
Bar rack and screen losses were determined by subtracting the sump total head from the water level outside the screens.
The static head at the entrance to the RHR suction pipe was calculated by extrapolating the measured pressure gradeline along the suction pipe, deter-mined by a linear regression curve fit of the head at the six piezometers, to the pipe entrance.
The area average pipe velocity was added to the extra-polated value of static head to determine the total head.
This total head was subtracted from the total head in the sump to determine the pipe entrance
~*
25 loss.
An entrance loss coefficient was determined by dividing the entrance loss by the pipe velocity head, as defined by Equation (13).
The entrance loss coefficient was then comparable to published data.
To illustrate the techniques used, Figure 14 shows typical pressure measurements, the fitted linear curve, and the calculated total heads.
A computer program was de-veloped to plot the model data, fit the curve, and calculate prototype losses and inlet coefficients.
Table 5 sununarizes the loss data for several blockage schemes.
Due to low approach velocities, actual head losses were small and the variability of the data reflected the difficulty in resolving the small head differences.
Pipe entrance loss data showed no systematic variation with screen blockage configuration, as would be expected since, as will be shown below, little swirl was evident in any blockage scheme.
The average inlet loss coeffi-cient was 0.53, which corresponds to a prototype head loss of 0.39 ft. The measured inlet loss coefficient predicted from published data (23) for the same area contraction in pipe flow was 0.57.
TABLE 5 Inlet and Screen Head Losses Total Loss Screen Loss Iniet Loss Blockage (ft)
(ft)
(ft)
Coefficient 1
0.37 0.04 0.33 0.45 8
0.52 0.13 0.39 0.54 9
0.50 0.10 0.40 0.55 11 0.57 0.15 0.42 0.57 12 0.77 0.35 0.42 0.57 13 0.76 0.36 0.4:0 0.55 14 0.53 0.11 0.42 0.57 16 0.47 0.08 0.39 0.54 18 0.51 0.21 0.30 0.41 Average 0.39 0.53
26 Screen losses varied with blockage configuration due to the modification of the flow path between the bar rack and inner screen.
For 100 percent open area, head loss was about 0.04 ft.
For the 50 percent blockage cases with open areas of the bar racks and screen.' aligned, configurations 9, 11, 14, and 16, the head loss averaged about 0.11 ft. As the open areas'of the bar racks and screens were staggered in various ways, the flow paths from the outer rack to the inner screen became restricted since the flow had to move parallel to the face of the screen in locations with reduced cross-sectional areas.
This caused velocities to be significantly increased.and significant percentages of the attendent increased velocity head could not be fully re-gained.
Screen losses increased to about 0.36 ft in the worst cases, confi-gurations 12 and 13, which had only two narrow flow passages between the bar rack and screen. Measured total prototype head losses through the screens, including pipe inlet loss, varied from 0.37 ft for the clean screen case to 0.76 ft for the worst case of 50 percent blockage.
Total losses using the average pipe inlet loss were calculated to vary from 0.42 to 0.74 ft.
Swirl Angle Measurements Rotameter rotation rates, measured for all blockage schemes, were used to calculate swirl angles using Equation (14). Resulting swirl angles are list-ed in Table 6 for both velocity scales.
Generally, swirl angles were small with the average of all tests being about 0.6 degrees.
Comparison of the same blockage scheme for the two velocity scales generally showed good agree-ment, within the repeatability determined from repeated tests with the same velocity scale.
Variation in swirl angle between the two pipes was sometimes relatively large due to the existence of very local swirl patterns.
The maxi-mum measured swirl angle was 2.25 degrees for the case of blockage scheme 10.
Swirl angles somewhat greater than the maximum measured may result from single bends (24) and swirl angles resulting from combined bends could be three to four times greater (25, 26).
Swirl angle did not.correlate well with surface vortex action and the highest swirl angles were found for some of the block-age schemes with the weakest vortex activity.
27 TABLE 6 Swirl Angle (Degrees)
Froude Velocity Scaling Prototy:ee Velocity Scaling Blockage Pipe.l Pipe 2 Pipe 1 Pipe 2 1
.o. 79 0.14 0.42 0
2 0.42 0
3 0.29 0
4 0.70 0
0.46 0.17 5
0.71 0.13 0.21
.0. 08 6
0.16 0.21 0.09
- 0.09 7
0*.18 0.07 8
1.26 0.95 0.20 1.15 9
0.88 1.62 1.01 1.62 10 2.25 0.31
- 1. 76 1.04 11
- 1. 57 0.99 1.60 0.90 12 0.68 0.43 1.34 0.98 13 1.62 0
1.85 0.99 14 0.80 1.33 0.66
- 1. '18 15 0
0 0
0.07 16 0.48 0.53 0.31 0.52 17 0.24 0.24 0.47 o*
is 0.29 0.33 0.48 0.15 19 0
0 0.18 0.33 Final Design Test Results The model inner screens were constructed to an early set of drawings which did not include a horizontal anti-vortex grid that had actually been retro-fitted.
The grid, shown in Figure 15, was constructed of 3/16 inch x 1-1/4 inch bar on 1-3/16 inch centers.
The bottom of the grid was located at ele-vation 81 ft 5-1/2 inches.
Since the maximum screen head loss was measured to be 4-1/2 inches, a water level of 81 ft 9 inches outside the screens could result in a water level inside the screen below the grid level making the grid ineffective.
Since the blockage resulting in the greatest head loss also had the worst air en-trainment due to vortices in the model, the top of the grid was lowered to elevation 80 ft 7 inches.
28 The elevation of 80 ft 7 inches was chosen to assure a safety margin in the calculated minimum water level of 81 ft 9 inches.
The grid was designed to stop surface rotation, and will be effective if the water depth above the grid decreases.
Since prototype velocity resulted in more severe vortices than any other con-dition, i.e., Froude scale velocity and one pump operation, the final test series was conducted only with prototype velocity.
Table 7 shows the vortex persistence results with the anti-vortex grid.
The grid was effective in controlling vortex formation and maximum vortex strength was type 3.
The non-uniform approach flow resulted in the highest persistence for type 3 vor-tices, up to 17 percent.
The type 3 vortex observed was very small and was due to a local disturbance caused by the support angle frame and gussets at the center of the inner screen.
The low turbulence level of the clean screen case allowed the local swirl to organize.
The core of the vortex was broken up by the anti-vortex grid.
With the rack and screen 50% blocked, only con-figuration 18 was observed to have a type 3 vortex.
The vortex was observed in the same location as in the non-uniform approach cases.
Of the ten rack and screen blockages tested, only three had vortex strengths of type 2 or greater.
The higher velocities due to screen blockage created sufficient turbulence such that the very small local type 3 vortex of the clean screen cases was unable to form.
None Non-Uniform Approach Flow TABLE 7 Vortex Persistence _(Percent)
Horizontal Grating at Elevation 80 ft 7 inches Vortex Strength Blockage 1
2 3
4 1
6 90 4
2 22 71 7
3 26 69 5
4 4
96 5
7 86 7
6 19 81 7
11 72 17 5
Rack and Screen Blockage 29 TABLE 7 (continued)
Blockage 10 11 12 13 14 15 16 17 18 19 1
100 100 0
100 100 100 0
100 21 100 Vortex Strength 2
3 4
5 100 100 75 4
Blockage scheme 12 caused a very low water level within the screen only about 1 inch (elevation 80 ft 8 inches) above the grid.
The turbulence of flow through the screens caused bubble formation and these bubbles were car-ried into the sump by the high vertical velocity, not vortex action.
This water level was unrepresentative of that which would actually exist in the prototype due 1to exaggeration of screen and grate losses **with the prototype velocity scale when the water level outside the sump was at 81 ft 9 inches.
The water level outside the sump was, therefore, increased siowly until the bubble entrainment ceased.
A level within the sump of BO ft 11 inches was required to eliminate bubble entrainment, which is considerably. below the expected prototype level (about 81 ft 4 inches for outside level 81 ft 9 inches) with blockage scheme 12.
Swirl Angle Measurements Table 8 lists the swirl angle measurements with the horizontal grids in place.
Maximum swirl angle was 1.47 degrees which was less than without the grid.
Generally, the swirl angle decreased after the addition of the grid, but a few minor increases were noted.
30 TABLE 8 Swirl Angle (Degrees)
Horizontal Grating at Elevation 80 ft 7 inches Blockage Pipe 1 Pipe 2 1
0 0
2 0
0.11 3
0 0.04 4
0.11 0.19 5
0 0.17 6
0.04 0.08 7
0 0
10 1.10 0
11 0.24 0.34 12 1.46 0.84 13
- 1. 47 0.78 14 0.06 0.59 15 0.13 0.64 16 0.24 0.27 17 0.05 0.46 18 0.21 0.66 19 0.30 0.17 Combined Blockage and Non-Unifonn Approach To assure that non-uniform approach flow would not increase vortex activity with the screen blockages, several additional tests were conducted.
Screen blockage schemes resulting in vortex strengths of 2 or greater were used.
Two approach flow conditions were used as shown in Figure 16.
Table 9 com-pares the non-unifonn approach flow results to the previous results and in-dicates no significant changes to the vortex persistence.
TABLE 9 Vortex Persistence (Percent)
Non-Unifonn Approach and Screen Blockage Blockage 12 12A 12B Vortex Strength 1
2 0
100 100 100 3
4 5
31 TABLE 9 (continued)
Vortex Strength Blockage.
l 2
3 4
5 16 0
100 16A 17 82 l
l6B 24 75 l
18
- 21.
75 4
l8A 3
94 3
18B 13 86 l
NOTE:
A and B refer to non-uniform approach flow conditions shown in Figure 16.
32
SUMMARY
A 1:2.8 scale model of the containment building sump for the Salem Generat-ing Station, Unit 2, was constructed and tested. In the recirculation mode, residual heat removal pumps withdraw water from the sump after a postulated loss of coolant accident. A vertical bar rack and 1/8 inch mesh screen, both about 3 ft high, surround the sump to assure no debris is entrained into the pumping system.
The debris could block both the bar rack and screen, there-by producing adverse flow patterns in the sump. A wide range of possible ap-proach ~low distributions, bar rack and screen blockages, water levels, and pump operation combinations were tested to simulate possible undesirable flow patterns which could result in poor pump performance of the RHR pumps during the recirculation mode.
The model was operated with both Froude scale vela-city and prototype velocity.
Vortex activity was observed and the percentage of time a given strength vortex existed was recorded.
Head losses due to the bar rack, screens, and pump inlet and the flow rotation in the suction pipe were also measured.
Test results indicated that no air core vortices were formed and the vortex suppression grid at elevation 77 ft 0 inches was effective in breaking up co-herent swirling flow.
For Froude scale velocity, the maximum vortex strength was type 3 dye core vortices, except for one case of a trash pulling vortex with persistence of 2 percent.
Increasing the velocity to prototype velocity generally increased vortex strength and/or persistence of a given vortex type and four screen blockage schemes caused air bubble entraining vortices with persistences near 20 percent.
In three cases, the bubbles entrained were,
\\
\\
small in diameter and number and were difficult to discern in the sump.
In the remaining case, the bubbles were considerably larger, but total gas flux was small since the bubble buoyancy resulted in a decreased vertical velocity relative to the water.
The bubbles formed were sufficiently small to have a rise velocity less than the vertical velocity within the sump and, therefore, they were carried downward into the RHR lines.
Increasing the water depth by 1 ft and decreasing the water depth by 9 inches, and one pump operation de-creased vortex persistence and/or strength.
33 Average swirl angle in the suction pipe was 0.6 degrees and maximum value measured was 2.25 degrees.
Screen losses varied from 0.04 ft for a clean screen to 0.36 ft for the worst case of 50 percent screen blockage.
Losses for conditions where the blockage of the bar rack and the screen were align-ed averaged 0.11 ft.
The pipe inlet head loss averaged 0.38 ft.
Addition of an anti-vortex grid within the inner screen at elevation 80 ft 7 inches was found to be necessary to eliminate vortices greater than dye core vortices (type 3).
The grid functioned well for water levels about 5 inches below that which wo*uld exist for the most severe case of 50% block-age of both racks and screens.
Swirl angles were found to be approximately the same as those without the ogrid.
Combination of a non-uniform approach with the rack and screen blockage had approximately the same vortex activity as tests with only the rack and screen blockage.
34 REFERENCES
- 1.
Daggett, L.L., and Keulegan, G.H., "Similitude Conditions in Free Surface Vortex Formations," Journal of Hydraulics Division, ASCE, Vol. 100, pp. 1565-1581, November 1974.
- 2.
Daily, J.W., and Harleman, D.R.F., Fluid Dynamics, Addison-Wesley Publishing Company, 1965.
- 3.
Rouse, H., Handbook of Hydraulics, John Wiley & Sons, 1950.
- 4.
Anwar, H.O., Weller, J.A., and Amphlett, M.B., "Similarity of Free-Vortex at Horizontal Intake," Journal of Hydraulic Research, IAHR 16, No. 2, 1978.
- 5.
Hattersley, R.T., "Hydraulic Design of Pump Intakes," Journal of Hydraulics Division, ASCE, pp. 233-249, March 1965.
- 6.
Reddy, Y.R., and Pickford, J., "Vortex Suppression in Stilling Pond Overflow," Journal of Hydraulics Division, ASCE, pp. 1685-1697, November 1974.
- 7.
Durgin, W.W., Neale, L.C., and Churchill, R.L., "Hydrodynamics of Vortex Suppression in the Reactor Building Sump Decay Heat Removal System," ARL Report No. 46-77/M202FF, February 1977.
- 8.
Baines, W.D., and Peterson, E.G., "An Investigation of Flow Through Screens," Trans. ASME, pp. 467-477, July 1951.
- 9.
Papworth, M., "The Effect of Screens on Flow Characteristics,"
British Hydromechanics Research Association, Report TN1198, November 1972.
- 10.
Weighardt, K.E.G., "On the Resistance of Screens," The Aeronautical Quarterly, Vol. IV, February 1953.
35
- 11.
Tennessee Valley Authority, "Flow.Through Screens," Report No.
87-8, May 1976.
- 12.
Padmanabhan, M., "Hydraulic Model Studies of the Reactor Containment Building Sump, North Anna Nuclear Power Station -
Unit l," ARL Report No. 123-77/M250CF, July 1977.
- 13.
Govier, G.W., and Aziz, K., "The Flow of Complex Mixtures in Pipes,"
Van Nostrand Reinhold, 1972.
- 14.
Chanishvili, A.G., "Air Entrainment and Vertical Downward Motion of Aerated Flows," IAHR, 8th Congress, Montreal, Canada.
- 15.
Muakami, M., Suehiro, H., Isaji, T., and Kajita, J., "Flow of En-trained Air in Centrifugal Pumps," 13th Congress, IAHR, Japan, August 31 -
September 5, 1969.
- 16.
Final Safety Analysis Report, J.M. Farley Nuclear Plant, Appendix 60, Nuclear Regulatory Commission, 1977.
- 17.
Durgin, W.W., and Hecker, G.E., "The Modeling of Vortices at Intake Structures," Joint Symposium of Design and Operation of Fluid Machinery, Colorado State Uniersity, June 1978.
- 18.
Padmanabhan, M., "Hydraulic Model Investigation of Vortexing and Swirl Within a Reactor Containment Recirculation Sump," Donald D. Cook Nuclear Power Station, ARL Report No. 108-78/Ml78FF.
- 19.
Padmanabhan, M., "Assessment of Flow Characteristics Within a Reactor Containment Recirculation Sump Using a* Scale Model,"
McGuire Nuclear Power Station, ARL Report No. 29-78/M208JF.
36
- 20.
Padmanabhan, M., "Selection and Scaling of.Horizontal Gratings* for Vortex Suppression," ARL Report No. 62-78, July 1978.
- 21.
Padmanabhan, M., and Vigander, S., "Pressure Drop _Due to Flow Through Fine Mesh Screens," Journal of the Hydraulics.Division, ASCE, HY8, August 1978.
- 22.
Durgin, W.W., and Lee, H.L., "The Performance of Cross-Vane Swirl Meters," -ASME Winter Annual Meeting, 1980.
- 23.
Miller, D.S., Internal Flow Systems, BHRA Fluid Engineering, 1978;
- 24.
Unpublished ARL Experimental Results.
- 25.
Padmanabhan, M., "Investigation of Flow Distribution and Swirl Due to a Combined Pipe Bend," McGuire Nuclear Power Station, ARL Report No.
12-79/M208MF, December 1978.
- 26.
Nystrom, J.B., "The Effects* of Combined Bends on the Velocity Distri-bution and Swirl at the Inlet to a Pump," ARL Report No. 122-80/MlOSAF, August 1980.
FIGURES
INNER DRAINAGE TRENCH NOTE:
WATER STOPS EL 78'-2" ARROWS INDICATE FLOW PATH MODEL BOUNDARY MODEL BOUNDARY FIGURE 1 PLAN OF REACTOR CONTAINMENT BUILDING AT EL 78'-0"
w
(.!)
<(
(.)
a:
w I-0 0
C9.
1'-0" w
(.!)
<(
(.)
a:
w z z 0 t.
C9
(>>
1 '-0" 14----- B'-6" l.S. OUTER CAGE-----
1 '-6" 5'-0" l.S. INNER CAGE-1--2'-0" CONC..
CURB
~
r...
VERTICAL SCREEN (1/B" x 1/8" (0.12) MESH)
ALL AROUND 1-1/4 x 3/16 AT 1-3/16 CENTERS VERTICAL GRID ALL AROUND (U.S. CAGE ONLY)
~
REMOVABLE HORIZONTAL GRID (1-1/4" x 3/16" AT 1-3/16" CENTERS)
B/EL 77'-0" (TYP.)
FIGURE 2 PLAN OF CONTAINMENT BUILDING SUMP AT ELEVATION 78'-8"
EL 81'-11" WELDED ANGLE FRAME (TYP.)
r SOLID TOP PLATE
'""""'="="'"'="'='=~"'"""""=""="'="=""'1 +--~-t-tt--t--- TOP OF SCA EE N EL 81 '-9"
. f ~~~-~9;~:~-:~,
- \\"'.~ -*. **.EL n*~o"* **,1--__._..~-------1*
- ._::.:.4 -..<:.:... :.. :.-.....,_*,,.Jl*:<
I i:
- _.-.*~**9*~.....
-~."--,
I ~1111-11'.
REMOVABLE HORIZONTAL GRIDS VERTICAL GRID 1-1/4" x 3/16" AT 1-3/16" CENTERS REMOVABLE*
HORIZONTAL GRIDS VERTICAL FINS
- ~~-~ ~:.
' I I II II l I l 1
I!! 11111 I ! 111 I! I 4" x 1 /4" AT ------J'-----'._
4" CENTERS AHR SUCTION PIPE 17" 1.D.
FLOW ~
(MINIMUM-WATER LEVEL)
EL 78'~9" EL 71 '-6" SECTION A-A.
(SEE FIGURE 2)
FIGURE 3 CROSS-SECTION OF CONTAINMENT BUILDING SUMP
--1
VORTEX TYPE 2
3 4
5 6
'\\l TRASH cb '
y
()
AIR BUBBLES
~
II INCOHERENT SURFACE SWIRL-SURFACE DIMPLE; COHERENT SWIRL AT SURFACE DYE CORE TO INTAKE; COHERENT SWIRL THROUGHOUT WATER COLUMN VORTEX PULLING FLOATING TRASH, BUT NOT Al R VORTEX PULLING AIR BUBBLES TO INTAKE FULL Al R CORE TO INTAKE FIGURE 4 CLASSIFICATION OF FREE SURFACE VORTICl;S
IF =
r 2
LOCUS, TYPE N VORTEX 7 I
....--+~+PROTOTYPE.RANGE~~---
IR FIGURE 5 ARL VORTEX ACTIVITY EXTRAPOLATION TECHNIQUE
~
0 j::
<(
a:
a:
w IJl
- J z w
c
- J 0 a:
- u.
2.0*-.----.---.~~~~~~~--.-~~~~..--~~-.-~---.~~~---.-~~-r-~~~~~~~~.......-~~~---.~~~
48°F 1.5, 1.0 0.5 (2,3)
NOTE: NUMBERS WITHIN BRACKETS REPRESENT MAXIMUM VORTEX TYPES CASE 2 WATER SURFACE EL 732.3' 1
- PROTOTYPE 70°F -
165°F 106 NOTE:
FROM REFERENCE 19.
REYNOLDS NUMBER, ud/V FIGURE 6 TYPICAL MODEL RESULTS FOR ARL VORTEX ACTIVITY EXTRAPOLATION TECHNIQUE
INNER DRAIN TRENCH TANK
....__FLOW..
FLOVli STRAIGHTENER FLOW CRANE.
WALL.
I I I I I I I I
~.RHR L__f:/SUMP NOTE:
VERTICAL GRATING AND SCREENS NOT SHOWN.
CONTAINMENT WALL FIGURE 7 PLAN OF MODEL LEVEL CONTROL WEIR
CONFIGURATION.NO. 2 BOTTOM 50% HORIZONTAL BLOCKAGE D
CONFIGURATION NO. 3 TOP 50% HORIZONTAL BLOCKAGE D
FIGURE 8 NON:--UNI FORM APPROACH FLOW, SCHEMES 2 AND 3
CONFIGURATION NO. 4 VERTICAL BLOCKAGE D
CONFIGURATION NO. 5 VERTICAL BLOCKAGE D
FIGURE 9 NON-UNIFORM APPROACH FLOW, SCHEMES 4 AND 5
--~
75%
50%
TOP CONFIGURATION NO. 6 VERTICAL BLOCKAGE D
CONFIGURATION NO. 7 HORIZONTAL BLOCKAGE 50%
TOP D
FIGURE 10 NON-UNIFORM APPROACH FLOW, SCHEMES 6 AND 7 p
(...
l*
CONFIGURATION 8
~~~~
.. _l r.i;;;~~~iliiiiiiiiiiiiiii~(.\\----:---11:~,rr 1'1"'"'11"'1'"11 Ii
- :---!. --~1~. ~=! b
-. ~ - :
"...!~'
--'--~ I I
11 l
11
~
~7=-=-:~ _: !
~3L---. -----
~
r--
1.. :.:.
- ~-I I
I
---~ I I
,1
-.. --. ~\\
~=~2-~L=ll r-------
.,iJ.'
1_,;*
CONFIGURATION 10 CONFIGURATION 9 CONFIGURATION 11
- /
FIGURE *11
-SCREEN BLOCKAGE, SCHEMES 8 THROUGH* 11*
CONFIGURATION 12 CONFIGURATION 14 CONFIGURATION 13 I I.
I *'
- I*
I I *:
I CONFIGURATION 15 FIGURE 12 SCREEN BLOCKAGE, SCHEMES 12 THROUGH 15
'r CONFIGURATION 16 (BOTTOM HALF)
I
,+,
.,.,_-tJ)
.1*
I CONFIGURATION 18 (BOTTOM HALF OUTSIDE)
(TOP HALF INSIDE)
CONFIGURATION 17 (TOP HALF)
CONFIGURATION 19 (OUTSIDE BOTTOM HALF)
(INSIDE TOP ALTERN_ATING BOTTOM HAl,.F)
FIGURE 13 SCREEN BLOCKAGE, SCHEMES 16 THROUGH 19 *
-I
- 1. 4"---------------------,
WATER LEVEL
- 1.
r.----+-- TOTAL HEAD IN SUMP
- 1. ee 1-ttl e. ee LL I
~ e. se LL.I
- c
- e. 4g TOTAL HEAD IN PIPE
- 4. ee 6.11
- 8. Bl DISTANCE PIPE DIAMETERS m BLOCKAGE CONFIGURATION 16 FIGURE 14 TYPICAL PRESSURE GRADELINE PLOT
fl r
- -<t_ -,__
2'-6" 7-1/8"
.C r-I
- 1~
I
-- ~~
L l~I
' 4-3/4" '1 1 G
r -rm 1'-2-7 16" -
2'-3'
~~_/
BAR 3/16" x 1-1/4 ON 1-3/16" CENTE 1'-3-3/8" LL 4"
-. w 7-1/2" EL 80'7" --*-
5'-0" TOP VIEW L----------
I
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.SECTION FIGURE.15.
DETAIL OF ANTl~*-VORTEX GRID
CONFIGURATION NO. 4 VERTICAL BLOCKAGE D
CONFIGURATION NO. 5 VERTICAL BLOCKAGE D
FIGURE 16 NON-UNIFORM APPROACH FLOW, SCHEMES 4 AND 5
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PHOTOGRAPHS
Photograph 1 Model Lower Sump with RHR Suction Pipes and Piezometers
Photograph 2 Construction of Model - Vertical Grids and Screens Installed
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Photograph 3 Interior of Model Sump with Horizontal Grids at Elevation 77 Feet and 71 Feet, 6 Inches
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