Arnwspheric Enuironment Vol. I I, pp. 775-781. Pergamon Press 1917. Printed in Great Britain.

COMPARATIVE PERFORMANCE EVALUATION OF CURRENT DESIGN EVAPORATIVE COOLING

TOWER DRIFT ELIMINATORS

JOSEPH CHAN and MIC~L W. GOLAY

~p~tment of Nuclear Engineering, M~sachusetts Institute of Technolo~, Cambridge, MA 02139, U.S.A.

(First received 10 January 1977)

Abstract-An analysis of the performance of standard industrial evaporative cooling tower drift elimina- tors using both numerical simulation methods and experimental techniques is reported. The simulation methods make use of the computer code SOLA as a subroutine of the computer code DRIFT to calculate the two-dimensional laminar flow velocity field and pressure loss in a drift eliminator geo- metry. This information is then used in the main program to obtain the eliminator collection efficiency by performing trajectory calculations for droplets of a given size by a fourth order Runge-Kutta numerical method. The experimental technique makes use of laser light scattering techniques for measurement of the droplet size spectra both at the inlet and outlet of the eliminator. From these measured spectra, the collection efficiency as a function of droplet size can be deduced. The results are found to be in good agreement with calculated collection efficiencies using no-slip boundary condi- tions. The pressure loss data for the eliminators are measured by an electronic manometer. The agree- ments between the measured and calculated pressure loss are good. The results show that both particle collection efficiency and pressure loss increase as the eliminator geometry becomes more complex, and as the flow rate through the eliminator increases. In ascending order of collection efficiency the eliminators tested are ranked as follows: sinus-shaned eliminator, three-segment eliminator and zig-zag

NOMENCLATURE

A(d) = droplet size distribution matrix d = droplet diameter (m) D = transverse displacement (m)

Fd = drag force on particle (N) H = eliminator length (m) K = proportion~ity constant

M(v) = voltage distribution matrix md = droplet mass (kg)

AP = Pressure drop R = droplet radius (m)

N(u,d) = voltage response matrix to droplets of dia- meter d

t = time (s) Q, = peak voltage (V) V, = air velocity (m/s) V, = droplet velocity (m/s).

1. INTRODUCTION

Drift, the current of water droplets rn~cha~~Iy entrained in the cooling tower exhaust flow, is reduced by requiring the exhaust flow to pass through drift eliminators installed in the cooling towers. Such separators operate by passing the two-phase flow stream through a curved duct, with the dense water droplets becoming trapped on the duct walls due to centrifugal acceleration. The performance of a drift eliminator can be determined in terms of its size- dependent collection efficiency and the pressure loss across the eliminator. For environmental protection, the collection efficiency should be high; for economi- cat tower ~rforrn~~ the pressure loss should be

low, and a compromise between these goals is required in an actual design.

Recently attempts by Roffman et al. (1973), Foster et al. (1974) and Yao and Schrock (1975) have been made to study the eliminator performance using theoretical analyses. All of these studies assumed either uniform or potential flow within the eliminator. Foster et al. (1974) tried unsuccessfully to compensate for the effects of turbulent wake regions by an effec- tive eliminator boundary defined from experimental flow visualization photography. Efforts in describing the turbulent wake region better are currently in pro- gress (Foster, 1976). Yao and Schrock (1975) also pre- dicted the pressure drop across the eliminator using the calculation method of Lieblein and Roudebush (1956).

Experimental evaluations of drift eliminator per- formance have mostly been performed by measuring the drift distribution at the exhaust side of the eli- minator which is installed in a particular cooling tower or in a simulated cooling tower facility. In most cases only the drift rate (defined as the drift mass current divided by the recirculating water flowrate in the tower) was measured. The drop size-dependent collection efficiency of the tower generally is not measured. Many methods have been developed to measure drift (ASME, 1975) since the tist attempt by Chilton (1952).

In this paper, evaluations of performance have been made for the three commercially available eliminator geometries which are shown in Fig. 1.

775

A.E. I I /R--H

716 JOSEPH CHAN and MICHAEL W. GOLAV

(a) (b) (c)

Fig. 1. Drift eliminator geometries. (A) Sinus-shaped eli- minator; (B) three-segment type eliminator; (C) zig-zag

eliminator.

2. THEORETICAL STUDY

Despite the fact that many of the important par- ameters which affect the performance of eliminators cannot be accounted for easily, a theoretical model

remains a very useful tool in evaluating the relative quality of performance of different drift eliminators,

and in designing improved devices. A computer program, DRIFT, has been written to

simulate the performance of drift eliminators (Chan, 1976). The main program of the code performs drop- let trajectory and capture efficiency calculations, using

the velocity field calculated by the subroutine SOLA- SUR. In the original SOLASUR code version, free- slip boundary conditions are used at the rigid curved

boundaries. In this study optional use pf no-slip boundary conditions at the rigid boundaries has been

added to the code so that the mass-averaged loss of total pressure between the inlet phase and the outlet phase of the liminator can be evaluated.

Air flow

(a)

Figures 2 and 3 show the calculated exhaust flow velocity fields within the three eliminators investi- gated. In the velocity field plots the length of the line

segments are proportional to the magnitudes of the velocities at the calculational mesh points, and the directions of the lines represent the directions of the flow at the mesh points.

Figure 2 shows the air velocity distributions calcu- lated for the sinus-shaped eliminator. In the free-slip case (A), the velocity distribution is approximately uniform. At the mid-length of the eliminator the vel- ocity is seen to decrease as the radius of curvature increases. Also, at the high pressure sides of the eli- minator the velocity is slightly greater than that at

low pressure sides. It can also be observed at each transverse cross section that the maximum velocity always occurs at the wall. This is not true in the no- slip case (B), where the maximum velocities occur at short distances away from the high pressure walls and then approach a value of zero at the walls. At the mid-length of the eliminator the maximum velocity occurs almost at the center of the cross section. Note that the velocities shown at the upper and lower boundaries of the eliminator do not represent the velocities exactly at the walls but at short distances away from the walls.

Similar observations can be made regarding the velocity fields of the other two geometries. Figure 3 shows the results of the three-segment geometry. For these more complicated geometries it can be expected

-. -.\ Air flow --\\ -“XL\\ ,

(b)

Fig. 2. Velocity distribution of air flow in sinus-shaped eliminator. (A) Free-slip conditions at upper and lower boundaries; (B) no-slip conditions at upper and lower boundaries.

---_\ ----‘\ ---.\\\

.-- --XL\ _--_ .X\\\

----.\\\\

-----\\\\

\\\

,111

\\\

\\

Air flow

\ // \\ /// \\\_--____//// \\\_____.//// \\\.____.//// ,,,.,__*../// \\,.___.-.,/// \\\.-- _-.,,/

,,.___-.,/ \.___I

\____I

(a)

jr

/I_

,/,-

/,,I_

//,.a /,/,- /,/,- / ,/-_ //I //

_ _ . __ -- _- __- _-- _--

__.._. Air flow

_ . -.__... ,

(b)

Fig. 3. Velocity distribution of air flow in three-segment eliminator. (A) Free-slip conditions at uppet and lower boundaries; (B) no-slip conditions at upper and lower boundaries.

Droplet dia = 40 micron

Cooling tower drift eliminators 717

Droplet dia = IO0 micron

\ / \ / (a) (b)

Fig. 4. Droplet trajectory plots for three-segment type eliminator with droplets entering the eliminator at left of figures. (A) 40~ droplet size; (B) 100~ droplet size.

that recirculating eddies and turbulent wake regions may exist at the bends in the flows and in interior corners. These effects are not taken into account in the current calculation. The recirculating eddies in in- terior corners could be resolved by use of a finer cal- culational mesh. However, this was not done since the mesh size chosen for use is in a range for which the calculated droplet capture efficiencies are rela- tively insensitive to the choice of mesh size. In addi- tion the droplet capture dynamics will be affected relatively little by whether the circulation of an eddy region is described exactly or if the region is treated as being approximately stagnant-as occurs with in- adequate spatial resolution of the calculated flow field.

The flow structure of any turbulent wake flow regions cannot be resolved by using a finer mesh, but could be treated explicitly by means of a “turbulence model” calculation (for which several different com- puter programs are available). However, in view of the success of DRIFT in predicting the experimentally observed behavior of the drift eliminators examined in this work it was decided that use of a turbulence mode1 would not be required. Effectively, the error introduced into the capture efficiency prediction by failure to describe turbulent eddy regions accurately is relatively small, mainly because for the devices examined the trajectories of the droplets which are captured lie far from such wake regions.

Fig. 5. Schematic diagram of experimental drift measure- ment facility.

It is seen in Figs. 7 and 8 that the strong depen- dence of the predicted velocity field upon the choice of boundary condition is reflected in the calculated droplet size-dependent capture efficiency.

The drift trajectory is calculated by solution of the droplet equation of motion using a fourth order Runge-Kutta numerical method. A detailed discus- sion is given by Chan and Golay (1976). An example of trajectory plots is given in Fig. 4. It can be seen that most of the smaller droplets escape from the eli- minators while most of the larger droplets are cap- tured This is because the larger droplets have rela- tively larger ratios of intertial force to drag force as they move in the flow than the smaller droplets. The locations at which the droplets are captured as shown in the trajectory plots provide useful information for drainage system design. It can also be observed from these figures as the geometry becomes more complex that the collection effectiveness increases.

From the trajectory calculations, the collection effi- ciency as a function of droplet size can be determined. The collection efficiency is defined as the ratio of the number of droplets captured with the eliminator to the number entering the eliminator.

3. EXPERIMENTAL STUDY

In order to perform comparative performance studies of cooling tower drift eliminators experimen- tally (and in order to validate the DRIFT model), a Drift Eliminator Facility has been constructed, which simulates a cooling tower fill-outlet environ- ment in which drift eliminators can be installed for testing. A schematic illustration of the facility is pres- ented in Fig. 5. It is a 0.8 x 0.8 m low-speed wind tunnel with plastic walls. Air speeds can be adjusted by means of a two-speed fan to be either 1.5 m/s or 2.5 m/s, respectively, simulating natural-draft or mechanical-draft cooling tower conditions. Droplets are injected into the flow by a spray head which con- sists of 20 full-cone center-jet nozzles (SPRACO Mode1 3B). The drift quantity and droplet size spec- trum can be controlled by me&s of the spray flowrate valve, located at the pump. The droplets produced by the nozzles lie mainly in the diameter range from

778 JOSEPH CHAN and MICHAEL W. GOLAY

5 to 2OOpm. The air is recirculated continuously to

ensure that water vapor saturation is obtained, in order to prevent droplet evaporation. The vertical test section in which the eliminators are installed is shown in Fig. 5. Eliminators can also be installed lying hori- zontally in the horizontal test section of the facility. However, in the work reported in this paper the eli- minators have all been installed in the vertical test

section.

The voltage pulse height spectrum recorded in the analyzer represents the probability distribution of scattered light pulse height signals. In order to obtain the drift droplet size spectrum a transformation of

The drift spectrum is measured by a technique de- veloped in this work, using laser light scattering. The arrangements of the droplet spectrum measuring in- strument is shown in Fig. 6. The laser used is a

50-mW helium-neon steady-state laser. A stopper is used to reduce the beam to approximately 1 mm dia- meter. A scattering volume is defined by the intersec- tion of the laser beam and a coliminated phototube detector acceptance cone. The acceptance cone is defined by the two apertures preceding the photomul- tiplier detector, which is placed at an angle of ap- proximately 30” to the laser beam. When a droplet passes through the scattering volume light is scattered and is detected by the photomultiplier tube which produces a voltage pulse. The height of the pulse which is proportional to the projected area of the droplet is recorded by a multichannel analyzer. The spectrum being recorded by the analyzer represents the size spectrum of the droplets passing through the scattering volume. The principle of this technique is similar to PILLS system (Shofner, 1973). Limitation of this technique due to fog-induced background sig- nals, as has been reported in field-measurements of drift in cooling towers, is nonexistent in this labora- tory work since only cold water is used in the experi- ments. Multi-particle scattering is avoided by con- trolling the size of the scattering volume and by adjusting the drift density appropriately. A single par- ticle is never counted more than once because a stea- dy-state laser is used rather than a pulsed laser (as in the PILLS system).

Fig. 6. Schematic diagram of light scattering drift measure- ment instrumentation. L-Laser, S-Stopper, M-Mirror, V-Scattering volume, A-Aperture, C-Photomultiplier

tube, P-Amplifier, MCA-Multichannel analyzer.

the recorded data is required. If the matrix N (u,d) represents the voltage response function for droplets of diameter d, and if matrix A(d) is the actual size spectrum of the drift, then the measured spectrum M(u) is given as

M(v) = N&d) A(d) (1)

Thus, the actual size spectrum can be determined as

A(d) = N-‘(I.$) M(V) (2)

In order to find the voltage response function N(u,d), calibration of the drift measurement instru- ment is performed by using monodisperse water drop- lets of known sizes, generated by a Berglund-Liu Monodisperse Aerosol Generator (Berglund, 1973). This generator uses the vibrating orifice principle, and the droplet size generated is controlled by orifice size, liquid feed rate and vibrational frequency.

In calibrating the droplet measuring system the droplet generator is placed under the scattering volume by the airflow in the Drift Eliminator Facility, which disperses the stream further due to turbulence. The droplets generated for the current work range from 50 to 1OOnm in diameter.

Because the laser light intensity is not uniform across the beam cross-section (it actually has a Gaus- sian intensity distribution), the monodisperse water droplets passing through the beam at different loca- tions on the beam cross-section will emit pulses of different intensity. Thus, the pulse heights recorded in the multichannel analyzer will not yield a single sharp peak as would be expected from the beam of uniform intensity. Calibration is performed for several different sizes of monodisperse droplets and from examination of the voltages, up, of the distribution maxima correlated with droplet sizes, it is seen that the relationship is obtained

L’~ = Kd=. (3)

This relationship can also be obtained from Mie- theory calculations of the intensity of monochromatic

light scattered from water droplets, and such calcula- tions have been performed in designing the droplet

size spectrometer (Shofner, 1973). By using the calibration drop size signal data and

equation (3) the response function N (E,d) can be

determined. Then the actual experimental droplet size spectrum A(d) can be calculated from the measured voltage distribution via equation (2). A computer pro- gram has been written to perform this conversion. In this experiment the scattering volume is kept fixed in space while the eliminators are either placed below the volume or above it. By placing the eliminators above the scattering volume the spectrum measured represents the inlet droplet spectrum, and similarly the spectrum measured when the eliminators are below the scattering volume is the outlet spectrum. From these two spectra the collection efficiency as a function of size can be determined.

A sensitivity analysis has been performed with

Cooling tower drift eliminators 119

regard to the effect of the response function on the collection efficiency calculation and it is found that the resulting efficiency is in general not very sensitive to the response function parameters, with the most sensitive parameter being the calibration droplet size. A 10% error in the calibration droplet size determina- tion results in relatively small errors in the measured collection efficiency for large droplets, but for small droplets error in the resulting collection efficiency can be significant (e.g. 20% error in collection efficiency for 4Opm droplets).

- Calculated result

--- Experimental result

This measurement has been performed for three in- dustrial drift eliminators at two different air speeds. The air speed is measured by means of a hot wire anemometer. The accuracy of the speed measurement is approximately 10%.

30 40 50 60 70 80 90 100 Droplet dio, micron

Fig. 7. Predicted and measured droplet collection effi- ciency functions for commercial drift eliminators at 1.5 m/s air speed. A-Sinus eliminator, &three-segment elimina-

tor, C-zig-zag eliminator.

4. DISCUSSION OF RESULTS

In Table 1 are compared the theoretically calculated pressure losses across the three types of drift elimina- tors with the experimentally measured values at dif- ferent air speeds. The calculated and measured results are in reasonably good agreement, and are within the bounds of experimental error. It is clear from these data that as the geometry of the drift eliminator becomes more complex the pressure loss increases.

Since the drift collection efficiency also increases with increasing geometrical complexity (see Figs. 7 and 8) it is necessary to strike a compromise in designing an eliminator in order to achieve both low pressure loss and high collection efficiency.

However, the good agreement of the calculated and measured results for the other two, simpler geome- tries, demonstrates the value of the calculations in reasonably smooth geometries in predicting pressure loss and collection efficiency.

From inspection of Table 1 it is seen that the pres- sure loss increases approximately as the square of the air speed for all three eliminators. The values in brackets are the resistances to the air flow expressed in terms of velocity head mainly due to entrance and exit losses.

Figure 7 shows droplet collection efficiencies as functions of droplet size for the three eliminators at the low fan speed (1.5 m/s). The calculated results using both no-slip and free-slip wall conditions are compared with experimental results. It is seen that both the calculated and the experimentally measured droplet capture efficiency values agree well for the sinus-shaped eliminator, for droplets with diameters larger than 7Opm, with the no-slip boundary collec- tion providing the more accurate predictions. It is also seen that droplets smaller than 50 pm in diameter escape the eliminator easily.

The calculations for the zig-zag type eliminator have been unsuccessful because of the complex eli- minator geometry and because of the limitations of the calculation method in describing turbulence.

By comparing the calculated and measured collec- tion efficiency data for the three-segment eliminator, it is seen that this eliminator is more efficient in cap- turing droplets of any size than is the sinus-shaped eliminator. Also, the agreement between the no-slip prediction and the measurement is reasonably better

Type of eliminator

Sinus

Three-segment

Zig-zag

Table 1. Pressure drop across eliminator

Air speed (m/s)

1.5

2.5

1.5

2.4

1.5

2.3

Calculated Measured Accuracy of value (Torr) value (Torr) measurement

0.02868 0.03054 13% (2.82)* (3.00) 0.06546 0.06279 11% (2.32) (2.23)

0.03482 0.03734 12% (3.43) (3.68)

0.07663 0.08119 11% (2.72) (2.88)

0.07424 8% (7.31)

- 0.16079 6% (5.70)

* Values in brackets are expressed in units of velocity head (= AP/$p).

MO JOSEPH CHAN and MICHAEL W. GOLAY

than the agreement between the free-slip prediction and the measurement.

Data are also displayed for the zig-zag eliminator, and it is seen that the measured performance is simi- lar to that observed with the three-segment elimina- tor.

In Fig. 8 the measured collection efficiencies of the three eliminators are compared to the calculated effi- ciency results (using no-slip boundary conditions) for the high fan-speed case. The individual air speeds of the three eliminators are slightly different from one another because the eliminator contributes in each case, to the total pressure loss in the wind tunnel, and each eliminator has a different associated pres- sure loss. The air speeds in the eliminators range from 2.2 to 2.5 m/s. The collection efficiencies at these high air speeds are consistently greater than those observed at low air speeds. This is true for all three eliminators at all droplet sizes.

An increase in the collection efficiency for droplets of a given size as the air speed increases would be expected on aerodynamic grounds. The force main- taining the droplet in the flow stream is the aerody- namic drag force, which varies in proportion to the relative wind on the particle. This force acts to impart transverse momentum of the particle (thereby deflect- ing it from the eliminator boundary) during the time t that the particle is within the eliminator (where t = H/V,). Thus, the maximum transverse displace- ment D of a droplet of given size in the flow will vary as

Consequently, the net drag-induced displacement of a particle from its inlet trajectory decreases in propor- tion to l/V,, as the air speed in~e~~-r~ulting in easier droplet impaction with the eliminator and greater capture efficiencies at higher air speeds, since an incoming droplet is diverted less from its inlet tra- jectory.

- Colculoted result

---- Experimental result

0 I I I I I 1 1 30 40 50 60 70 80 90 100

Droplet dia I mrcro”

Fig. 8. Predicted and measured droplet collection efli- ciency functions for commercial drift eliminators at 2.5 m/s air speed. A-Sinus eliminator, B-three segment elimina-

tor, C-zig-zag eliminator.

The pressure drop across the eliminator is measured with a differential electronic manometer. Two pitot tubes which monitor static pressure are located at the inlet and outlet of the eliminator. The sensor is a Barocell pressure transducer. The differen- tial pressure across the experimental chamber is first adjusted to be zero with no eliminators being in- stalled and with air flowing. The pressure loss across the eliminators is then the differential pressure measured with the eliminators installed.

In Fig. 8 the agreement between the calculated (no- slip) and measured sinus-shard eliminator collection efficiency results is seen to be very good. IIowever. the theoretical-experimental agreement is relatively poor for the three-segment eliminator, due to diffi- culty in treating turbulent wake effects. In this case it is observed that the predicted collection efficiencies are signifi~tly greater than those observed experi- mentally. A similar trend-attributed to turbulence effects-has been reported in Foster et al. (1974). It is seen that the experimentally-determined collection efficiency for droplets of a given size increases as one progresses respectively from the sinus-shaped ehmina- tor to the three-segment to the zig-zag eliminator (as is also observed at lower air speeds). Calculated col- lection efficiency data for the last eliminator are not shown in Fig. 8 due to the inability of the calculations to describe wake turbulence and recirculation encountered in this case, for which Reynold’s number is in the range of 5ooO and the geometry is complex.

It is observed for all eliminators at all air speeds that the collection efficiency increases monotonically with droplet sizes. This occurs because, at a given air speed, the aerodynamic drag force varies approxi- mately in proportion to RZ (see equation 1) while the particle mass varies in proportion to R3 (and the time during which the particle is affected by drag is approximately constant). Thus, the net acceleration on a particle, and the resulting displacement vary ap- proximately as l/R, resulting in easier capture for larger droplets.

Water film ejixts. The droplet efficiency calcula- tions have ignored the possible effects of the water film on the eliminator surfaces. One class of effects arises from the drag of the exhaust flow on the water film, and the other class arises from the impacts of captured water droplets on the film. The presence of the water film modifies the exhaust flow boundary condition from being one of no-slip to being one of matching the air-water velocities at the gas-liquid in- terface. Except in the case of a thick film the effect of this boundary condition change upon the velocity distribution within the eliminator should be small. However, drag on the liquid film can lead to droplet generation by means of drawing the water film over the trailing edge of the eliminator, and by stripping droplets from the film surface through formation of Helmholtz instability waves.

With sufficiently thick films and with droplets im- pacting at sufficiently acute angles (Jayaratne, 1964)

Cooling tower drift eliminators 781

it has been observed that a droplet can rebound from the film, or “bounce” back into the exhaust flow. In this work the experiment was not set up to observe “bounce” effects. The water loading on the eliminator is so small that no continuous water film has been observed on any of the eliminators.

The work of Yao and Shrock (1975) indicates at an air speed of 2.0m/s in a smooth drift eliminator geometry that the minimum film thickness required for droplet stripping is 0.2 mm, and that substantially higher air speeds are. required for droplet generation via pickup of the peaks of Hehnholtz ins~bility waves for a film of this thickness. In the m~surements the water film on the eliminator was not measured; how- ever, visual observations were unable to detect either droplet stripping at the trailing edge of the eliminator or droplet generation in the interior. The absence of droplet stripping at the outlet is implicit evidence (Yao, 1975) for the absence of wave generated drop- lets in the interior also.

5. CONCLUSION

The performance of three commercially available cooling tower drift eliminators has been investigate both by using numerical simulation techniques and by laboratory-scale experimental testing. It is found that the agreements between these two techniques are best for smooth-geometry eliminators and at low air speeds. This suggests that the turbulence wake effects within eliminators are not negligible for complex geo- metries at high air speed, and should be taken into account in design calculations. However, the current computer program constructed to calculate the eli- minator collection efficiency and resistance to the air flow is seen to be valuable in predicting drift elimina- tor performance for cases in which laminar flow con- ditions obtain. It is seen that improved collection effi- ciency for droplets of a given size can be obtained by increasing the flow speed through the eliminator, in each case with an attendant increase in pressure loss. In ascending order of capture efficiency perform-

ance the eliminators tested are ranked as follows: sinus-shaped eliminator, three-segment eliminator, and zig-zag eliminator. The same ordering is obtained when the eliminators are ranked in ascending order of observed pressure loss.

Acknowledgemen&--This work was funded through the Waste Heat Management Program of the M.I.T. Energy Laboratory, supported jointly by the New England Electric System and Northeast Utilities.

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Roffman A. et al. (1973) The state of the art of saltwater cooling tower for steam electric generating piants. WASH-1244, U.S. Atomic Energy Commission, -- E.l-E.21.

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Yao S. C. and Schrock V. E. (1975) Aerodynamic design of cooling tower drift eliminators. Am. Sot. Mech. Engrs. Paper No. 15-WA/Pwr-5.