Gas transfer within a multi-stage packed column oxygen absorber: Model development and application

Aquacultural Engineering 9 (1990) 33-59

Gas Transfer within a Multi-Stage Packed Column Oxygen Absorber: Model Development and Application Bamaby J. Watten National Fishery Research and Development Laboratory, US Fish and Wildlife Service, RD No. 4, Box 63, Wellsboro, Pennsylvania 16901, USA

& Claude E. Boyd Department of Fisheries and Allied Aquacultures, Auburn University, Alabama 36830, USA (Received 20 October 1989; accepted 30 January 1990)

ABS TRA CT A packed column oxygen absorber was developed in which oxygenflow is directed, in serial reuse, through parallel packed column stages receiving equal portions of the liquid being treated. The relative performance of the absorber was established using a computer simulation program employing finite difference-mass transfer calculations. The program was calibrated using packing specific mass transfer coefficients derived from pilot scale test data. A separate series of tests served to verify model assumptions and performance predictions. Simulation data indicated multi-stage operation can substantially reduce the column height required to achieve a selected oxygen absorption efficiency (AE); for example, the column height required to achieve an A E of 76"5% with an inlet volumetric oxygen-water ratio of 0"008 (column packing, 3.81 cm plastic ACTIFIL®; water temperature, 20"C; influent dissolved oxygen, 9.08 mgflitre; operating pressure (absolute), 760 mm Hg) was 0"27 m using a lO-stage system versus 1"39 m using a single-stage absorber. Reductions in column height achieved were related to oxygen and water feed rates, number of stages employed, mass transfer characteristics of the column packing used, and concentrations of dissolved gases in the liquid being treated. 33

34

B. J. Watten, C. E. Boyd

NOTATION AE BP

c, C,,.t

c,

DC DO DOin

DOout DN G/L Kd

K, Kp n oX

PW

QL Q,,x

R2 T

TE VP

x, Z a

#

Oxygen absorption efficiency (%) Bunsen coefficient of gas species i (litres gas/fit.re water at 760 mm Hg pressure) Atmospheric pressure (rnm Hg) Concentration of gas species i (mg/litre) Packed column influent concentration (mg/litre) Packed column effluent concentration (mg/litre) Dissolved gas saturation concentration (mg/litre) Dissolved carbon dioxide concentration (mg/litre) Dissolved oxygen concentration (mg/litre) Packed column influent DO (mg/litre) Packed column effluent DO (mg/litre) Dissolved nitrogen concentration (mg/litre) Overall mass transfer coefficient (dimensionless) Volumetric gas-liquid ratio (m3 h-m gas/m3 h- ~liquid) Distribution plate mass transfer coefficient (n- ~) Ratio of molecular weight to volume for gas species i (mg/ml) Specific packing mass transfer coefficient (m- ~) Number of distribution plates within the column (dimensionless) Mass density of oxygen (kg/m 3) Power used in pumping (kW) System water flow rate (m3/h) Volumetric flow rate of oxygen (m3/h) Coefficient of determination (dimensionless) Temperature (°C) Oxygen transfer efficiency (kg/kW h) Vapor pressure of water (ram Hg) Gas phase mole fraction of the Rh gas (dimensionless) Packing depth (m) Field water ( G)20/clean water ( G)20 (dimensionless) Field water CJclean water C~(dimensionless) ( G )20,g~, species//(G )2o,oxygen(dimensionless)

INTRODUCTION Packed column oxygen absorbers provide a relatively simple and efficient means of contacting oxygen with water to add dissolved oxygen

Gas transfer within a packed column oxygen absorber

35

(DO), or remove dissolved nitrogen (DN) (Speece, 1981; Colt & Watten, 1988). Accordingly, appfication of such systems in aquaculture has received considerable attention (Boerson, 1985; Severson et al., 1986; Dwyer et al. (in press) Westers et al. (in press)). Selection of the appropriate column design protocol must be based on reaction kinetics as well as on the expected contacting pattern within the reactor (Levenspiel, 1979). Nirmalakhandan et al. (1988)developed design recommendations for the ideal countercurrent case in which commercial oxygen and water proceed through the reactor, without longitudinal mixing, toward their respective discharge ends. Over the range of operating conditions considered, optimum bed depth, hydraulic loading, and volumetric gas-liquid ratios lie between 2.8 and 3.0 m, 50 and 120 kg/m2 s and 0.015 and 0"035, respectively. In the present study, a multi-component gas transfer model was developed for the non-ideal contacting pattern in which the gas phase within the reactor is homogeneous. Steady-state gas composition profiles established by Watten & Boyd (1990) indicated gas phase mixing within the column is extensive under the high packing irrigation rates and low gas flow operating conditions used by necessity in aquacultural applications. The homogeneous gas phase acts to restrict significant gas absorption-desorption to those regions within the upper portion of the column. Thus, column performance should be reduced to a level below that predicted with the algorithms developed by Nirmalakhandan et al. (1988). With extensive mixing in the gas phase, a higher mean gradient between gas phase and liquid phase concentrations can be provided by directing gas serially through packed column stages receiving equal portions of the liquid being treated (Warren, 1989). Figure 1 provides a flow schematic for a four-stage packed column system. In operation, the water being treated is divided into four portions, each directed through an independent packed bed reactor. Oxygen metered into the system is first directed through the single column on the left. Oxygen transfer and DN desorption in this segment is high because the effective volumetric gas-liquid ratio is relatively large. Movement of gas into and out of the liquid acts to lower the mole fraction of oxygen in the gas phase to something less than that in the feed gas. The extent of gas phase contamination, under steady-state conditions, is determined by water and gas flow rates, inlet dissolved gas concentrations, temperature, pressure, and the mass-transfer characteristics of the equipment selected (Speece et al., 1983; Warren & Beck, 1985; Warren et al. (in press)). The contaminated off-gas flow, normally vented to the atmosphere, is first directed through the second packed bed reactor. As the off-gas is exposed to untreated water, the dissolved gas deficit (C~-Ci) is re-established, allowing

36

B. J. Watten, C. E. Boyd Water Supply

O~ GAS Inlet

I ~dmn __

Pt ~d Co

nn

Off GAS

v Exit Oxygenated Effluent

Fig. 1.

Flow schematic for a multi-stage packed column oxygen absorption system.

additional oxygen to be absorbed from the gas mixture. The serial reuse of the off-gas is then repeated in later packed column segments, allowing the oxygen partial pressure in the gas phase to approach the oxygen tension in the influent. Watten (1989) hypothesized that the reduction in column height afforded by the use of the multi-stage configuration significantly reduces the head required to achieve an acceptable degree of inlet gas absorption. We tested this hypothesis following the development and calibration of a numerical model that simulates the performance of both single and multi-stage packed column systems. NUMERICAL MODEL DEVELOPMENT A computer program simulating gas transfer within an enclosed packed colunm oxygen-water contactor was developed on the basis of a series of finite difference mass-transfer calculations. The model developed accounts for ( 1 ) oxygen transfer from the gas to the liquid phase and (2) nitrogen and carbon dioxide transfer from the liquid to the gas phase. The effects of gases present in trace quantities were ignored as were reactions of dissolved carbon dioxide (DC) in the liquid phase. Further, the gas phase within the contact enclosure was assumed to be homogeneous.

Finite difference calculations

Multi-component finite difference calculations, used to establish steadystate performance, are similar to those used in earlier studies addressing

Gas transfer within a packed column oxygen absorber

37

gas transfer in oxygen-water contact systems (Speece & Orosco, 1970; Mueller et aL, 1973; Watten & Beck, 1985; Watten etaL, (in press)). The following variables are calculated at each time step in the program: (1) mass of oxygen, nitrogen and carbon dioxide in the gas phase; (2) saturation concentrations; (3) dissolved gas deficits; and (4) changes in dissolved gas concentrations. At the start of the calculation series, the Ideal Gas Law was used to relate gas phase volume, temperature, pressure and molar composition. Operating pressure was taken as the sum of local atmospheric pressure and column gauge pressure. Gas phase volume was calculated on the basis of total column volume and dry packing porosity. Equilibrium gas saturation concentrations required to establish dissolved gas deficits were calculated using Henry's law (Colt, 1984): C~--/~(1000 K, B~X~(BP- VP )/760.O)

(1)

Bunsen coefficients (B) and vapor pressures (VP) used in this determination were obtained from relationships developed by Weiss (1970, 1974) and ASHRAE (1972). Gas absorption-desorption across the column was predicted by using the transfer model put forth by Hackney & Colt (1982): In

C -

Ci. = ¢k(nKd + KpZ) ct

(2)

Equation (2) is based on the two-f'dm mass transfer theory of Lewis and Whitman (1924). The expression on the right of eqn (2) is defined as the overall system transfer coefficient G, a required program input. Values for input G are corrected for the effects of temperature by using the APHA (1985) correlation:

(G).r=(G)2o (1"024) r- 2,,

(3)

After effluent dissolved gas concentrations had been calculated, and given the influent gas flow, the molar flow rate of gas exiting the column was established by applying a materials balance on the gas phase. Predicted effluent dissolved gas concentrations were compared to previous estimates and checked for convergence. The calculation sequence was repeated if the difference between the previous and current DO concentration estimated exceeded 0.01 mg/litre. Before the finite difference calculations were continued, the working composition of the gas phase was adjusted for changes resulting from gas inflow, absorption, and desorption. The time step used in all simulation runs was 5 x 10- ~ h.

B. J. Watten, C. E. Boyd

38

After completing finite difference calculations, we identified exit conditions and calculated performance indicators. Performance indicators

The computer program established steady state effluent DO, DN and DC concentrations. Effluent dissolved gas parameters, such as individual gas tensions and total gas pressures, were determined by using the computation methods summarized by Colt (1984). In the calculation of transfer efficiency (TE), the mass of oxygen absorbed per unit of energy input, we assumed a combined pump and motor efficiency of 70%: TE--

QL(DOou, - DOi. ) 10 - 3 PW

(4)

Transfer economy (total cost/kg 0 2 absorbed), and oxygen absorption efficiency (AE) were also calculated. The AE represents the ratio of mass oxygen absorbed to mass oxygen applied, as defined in the following expression: AE=

QL(DO,,u,

-

DOi.) 10- s

Q,,xP,,x

(5)

MODEL CALIBRATION Materials and methods

Before we used the model, we undertook two series of tests to provide predictive equations for the overall transfer coefficient G. In the first test series, G values were established for the plastic 2.54 cm Tri-Pack®t, 3.81 cm and 5.08 cm Nor-Pac ® packing employed in the gas phase axial dispersion study of Watten and Boyd (1990). The G values were established at bed depths of 0.362, 0.699 and 1.041 m with packing irrigation rates of 32.0 and 62.2 kg/m 3 s. The G value analyses were replicated either 2 or 3 times providing a total of 24 observations. In the second test series, G values were developed for 3.81 cm ACTIFIL® plastic packing at bed tThe mentionof trade names or manufacturers does not implythe endorsement of commercial products by the US Government.

Gas transfer within a packed column oxygen absorber

39

depths of 0.305, 0-648, 1.003 and 1-346 m. At each selected depth, G was established using packing irrigation rates of 34-3, 42.2, 53-5, 61-9 and 73"4 kg/m 2 s. The physical properties of all packing types evaluated are summarized in Table 1. We established G values by using the test packed column system described by Watten and Boyd (1990). Commercial oxygen was introduced into the column at a point representing 100% of the bed depth being evaluated. The number of distribution plate holes used with test hydraulic loading rates less than 53.5 kg/m 2 s was 1055/m 2. The number of holes used with all other test loading rates was 2096/m-'. Off-gas vented from the tower through the off-gas demister chamber was monitored with a YSI model 2650 oxygen sensor. Gas flow was elevated to a point that resulted in less than a 2% change in oxygen composition (gas phase) across the packed bed. After we established steady-state column performance, we measured influent and effluent DO concentrations along with temperature, alpha (a) and atmospheric pressure using methods described by Watten and Boyd (1990). The G values were subsequently calculated using eqn (2) and then adjusted to 20"12 with eqn. (3). Given test G and corresponding a values, we used least squares linear regression to establish predictive equations for (G)20, corrected to clean water conditions, as a function of the independent variables of packing type, bed depth and hydraulic loading. Results

During the course of the G value analyses, the range of temperature, influent DO, and a were 24.5-30"C, 7.0-8.1 mg/litre and 0"81-1.03. Table 2 gives the regression coefficients defining ( G)20 for Tri-Pack ® and Nor-Pac e media as a function of bed depth at each of the two hydraulic loading rates tested. The response of ( G)20 to increases in bed depth was adequately described by the linear model chosen. Values of (G):o predicted at the Y intercept (zero bed depth) reflect the extent of the end effect phenomenon (Watten & Boyd, 1990). End-effect (G)20 values are influenced by hydraulic loading and the design of the liquid distribution plate, but generally range between 0"08 and 0.68 (Colt & Bouck, 1984). The y intercept (G)2 o values established here, averaging 0"289, lie within this range. Figure 2 summarizes the (G)2 0 data developed for the 3"81 cm ACTIFIL ® packing at hydraulic loading rates of 34.3, 42.2, 53"5, 61.9 and 73.4 kg/m 2 s. At each bed depth evaluated, (G)2 0 remained essentially independent of flow as indicated by an average line slope,

B.J. Watten, C. E. Boyd

40

TABLE 1 Physical Properties of the Plastic Packing Used to Calibrate the Simulation Program" Variable

Packing type

Geometric surface area (m2/m 3) Void space (%) Specific mass (kg]m3)

2"54 cm Tri-Pack ®

3.81 cm A CTIFIL ®

3"81 cm Nor-Pac ®

5.08 cm Nor-Pac ®

279 91 99.4

139 92 76

144 93 60.9

102 94 52.9

"As reported by the manufacturer.

TABLE 2 Least Squares Linear Regression Coefficients Defining (G)_,o as a Function of Bed Depth /m) at Each of Two Hydraulic Loading Rates (HL) Packing Ope

Y intercept

Slope

Coefficient of determblation, R '

2"54 cm Tri-Pack® HL-- 32.0 kg/m-" s HL= 61.2 kg/m-" s

0.327 0.277

1.655 1.589

0.979 0.986

3.81 cm Nor-Pac* HL= 32.0 kg/m 2 s HL= 61"2 kg/m 2 s

0.324 0"398

1.555 1.428

0.990 0.976

5"08 cm Nor-Pac® HL-- 32.0 kg/m 2 s HL~ 61.2 kg/m 2 s

0.162 0.243

1.855 1.285

0.988 0.944

determined by regression, of just 1-29% (range 0.09-2.34%). Thus, the predictive equation for the ACTIFIL® media was established by regressing pooled ACTIFIL® (G)20 values against bed depth. The resultant linear equation is: (G)20 = 0-357 + 1.349(Z) R2=0.965

(6)

The observed insensitivity of (G)20 to changes in hydraulic loading is in agreement with data presented by Hackney and Colt (1982) for this same type of packing.

41

Gas transfer within a packed column oxygen absorber 2.25 Z = 1.346m

0 0

2.00

0

0

0 0

0

Z - 1.O03m

o

rD

1.75

1.50

.2

Z=O.648rn

0

1.25

0

0

0

M

u

~.o1.00 rD 0,75

Z=O.305m 0 0

O,50

I 30

0.25 2O

[ 40

I 50

I 60

I 70

80

HYDRAULIC LOADING ( k g / m Z . s )

Fig. 2. Summary of (G)2. versus bed depth data established for 3.81 cm ACTIFIL® media at five hydraulicloading rates and four bed depths (Z).

MODEL VERIFICATION Following model calibration, two series of tests were undertaken to verify single-stage and multi-stage packed bed performance as predicted by the computer model. Materials and methods Single-stage packed column analyses Watten and Boyd (1990) identified gas phase axial dispersion during a series of 90 column runs in which, at each of three bed depths (0.362, 0-699 and 1.041 m), all combinations of the following independent variables were tested: influent volumetric oxygen-liquid ratio, 0.008, 0.016, 0.026, 0.040 and 0.080; hydraulic loading, 32"0 and 61.2 kg/m z s; and packing type, 2.54 cm Tri-Pack ®, 3-81 cm Nor-Pac e and 5"08 cm Nor-Pac ®. Each unique combination of operating conditions was replicated once providing a total of 180 sets of observations. Steady-state concentrations of dissolved gas in the effluent measured during the

42

B. J. Wauen, C. E. Bo~d

analyses were compared here with model predicted values established by using as model inputs certain recorded operating conditions, including influent DO, DN, temperature, barometric pressure and a factors. Media-specific (G)20 coefficients required by the program were obtained by using the predictive equations developed during model calibration (Table 2). Similar steady-state performance comparisons were established with operating conditions representing either elevated influent DC (10 observations) or DN concentrations (10 observations). These later performance tests were conducted using 5"08 cm Nor-Pac ® media at a hydraulic loading rate of 32-0 kg/m 2 s. The packed column and data collection methods used were the same as those described by Watten and Boyd (1990). Influent dissolved gas concentrations were elevated by metering carbon dioxide, or nitrogen gas, at a constant rate into the head tank water supply line. Steady-state influent and effluent DC concentrations were measured in triplicate by the sodium carbonate titrimetric procedure (Boyd, 1979).

Multi-stage packed cohimn analyses As in the single-stage analyses just described, steady-state performance of a multi-stage packed bed absorber was established and then compared with model predicted performance. The multi-stage system evaluated is illustrated in Figs 3 and 4. The system consists of four sealed packed bed segments operated in parallel with regard to water flow and in series with regard to gas flow. Each 20-3 cm diameter column segment was packed with 0"77 m of 3.81 cm ACTIFIL®media. During multi-stage tests, the system received water at a constant rate from the head tank system used in the single column analyses (Watten and Boyd, 1990). Flow to each column was regulated with a 3.81 cm gate valve before it entered a sealed distribution plate well. Each distribution plate contained nine 9.5 mm diameter holes. Water exiting each packed bed segment was collected within an effluent well 26.7 cm in diameter, outfitted with a sample port. Water levels within each well were regulated with a 3"8 cm diameter discharge gate valve as indicated by a sight tube reading. Individual column water flow rates were established by removing the effluent collection manifold (Fig. 4), then measuring the net weight of water collected during a known time interval. We metered commercial oxygen into the system at a constant rate using a pressure regulator and rotameter (Watten & Boyd, 1990). Offgas exiting the first column segment was directed through subsequent columns via gas recycle lines 1.9 cm in diameter. Gas exiting the last column segment was vented to the atmosphere through a submerged side wall orifice. Changing the depth at which the. off-gas was released, by

43

transfer within a packed column oxygen absorber d

t.

t,

L

J

Fig. 3. Front view of the four-stage, 20.3-cm diameter, packed column oxygen absorber: (A) adjustable column support bar; (B) flange and flange plate for the sealed column; (C) 1.3-cmdiameter gas inlet nipple; (D) 26.7-cm diameter effluent well; (E) 6-4 mm gas sample port; (F) 3-8-cm drain and water level control valve; (G) l-3-cm sample port with sight tube; (H) off-gas recycle line; (I) water inlet port; (J) gas sample port.

adjustment of the overhead support structure or well water level, provided control of contact enclosure operating pressure. In several test runs, either influent D C or D N was elevated by metering gas into the head tank influent line. T h e equipment used to regulate gas flow and the procedure used to measure DC is the same as that described in the preceding section on single-column model verification. After steady-state performance was established, column pressures were determined with a U-tube manometer, which was attached to the gas sample port shown in Fig. 3. Influent and effluent dissolved gases

44

B. Z Watten, C. E. Boyd J_

3 4 . 3 cm

In

6

~.2cm

I

7.6cm

I 12.7cm

I

"1 '.

21.3cm x ~

.\ \

\

\ xl ~d

/

43.7cm

~jIII

E

I

. x~L_~_,_q 21.4

I I

cm

i

........

I ,.]

39.3cm

14.2 cm

!

6.3 cm

End view, minus the support assembly, of the four-stage 20.3-cm diameter packed column: (A) inlet flow control valve; (B) water distribution plate; (C) media support plate; (D) sight tube; (E) effluent drain and water control valve; ( F ) anti-siphon vent. Fig. 4.

were determined, as described by Watten and Boyd (1990), along with temperature, a, gas composition (%02) and atmospheric pressure. Observed DO and DN in individual segments were compared with model predicted values established using recorded operating conditions as model inputs. Table 3 provides a summary of the operating conditions evaluated. The number of columns operated in series was varied, as well as volumetric gas-liquid ratio, column pressure, and influent dissolved gas concentrations. Each data set was replicated once, providing a total of 130 observed versus model-predicted effluent condition comparisons. The (G)20 coefficient required in the analysis was obtained from the (G)20 versus bed depth relation developed for 3-81 cm ACTIFIL® packing (eqn (6)).

Gas transfer within a packed column oxygen absorber

45

TABLE 3 Summary of the Test Conditions Used to Establish Performance of a Multi-Stage Packed Column Absorber. Hydraulic Loading Ranged from 34-5 to 56.7 kg/m 2 s

No. column segments

Gas-liquid ratio

2

0"012 0"022 0-041 0"056

3

0"027 0-051 0-091

4

0"014 0-032 0"115

4

0"012 0"025 0"055 0"092

4

0"014 0"025 0"057

4

0"018 0"045

Low column pressure"

High column pressure h

High influent nitrogen

High influent carbon dioxide

"Column pressure range -- 0-6-22-3 cm H_,O gauge. hColumn pressure range = 35.7-48.2 cm H20 gauge.

Results

Single-stage packed column analyses During the course of the first series of single column analyses, water temperature, inlet DO and DN ranged between 21.3 and 32"8*(2, 5.4 and 9"5 mg/litre, and 11.6 and 15.3 mg/litre, respectively. Alpha waste correction factors averaged 1-00 with a range of 0.91-1"04. Figure 5 provides a comparison between observed and modelpredicted effluent DO concentrations for each of the three packing types and two hydraulic loading rates evaluated. Similar plots of effluent DN

46

B. J. Watten, C. E. Boyd 35 2.54cm Tri-Pack ® 0

3.81cm

AA

/

A

/

0

Nor - P a e ®

30

GO

5.08©ra Nor - P a c

1

O P, 25

Z;

,,..1 2O

[.. ¢~

15

10 I0

I 15

I 20

I 25

OBSERVED EFFLUENT DO

I 30

35

(ma/1)

Fig. 5. Comparison between observed and model-predicted effluent dissolved oxygen concentrations established by using three packing types during the single column model verification runs.

are given in Fig. 6. As shown, model predictions are in close agreement with pilot study data. Relative error of the effluent DO concentrations averaged only 8.63% (-7.91-20.4%) with 165 of the 180 predicted values being greater than observed concentrations. The mean relative errors associated with each packing type were similar, averaging 9.43% (-0.75-19-71%), 9"86%(-7-91-20.4%)and 6.61%(-5"61-15.42%) for the 2.54 cm Tri-Pack ®, 3.81 cm Nor-Pac e, and 5.08 cm Nor-Pac ® media. A trend towards higher relative error was observed in data pairs established using the hydraulic loading rate of 32.0 kg/m 2 min versus the alternative 61-2 kg/m 2 min, i.e the relative error of all low-flow test runs averaged 12.26% (4.47-20.4%) whereas the error of all high flow test runs averaged just 5.01% ( - 7.91-14.55%). A similar trend was present in the effluent DN data. Relative error of model predictions average 14.41% for the low flow and 8.24% for the high-flow runs. Relative errors associated with each of the three packing types were again similar, averaging 12.37% (-31.5-3.67%), for the 2.54 cm Tri-Pack ®, 11.34% (-25-11.06%), for the 3-81 cm Nor-Pack ® and 10"29% ( - 31.36-12.12%) for the 5.08 cm Nor-Pac ® media. The relative error of

47

Gas transfer within a packed column oxygen absorber 14

A

12

~0 v

5.08emNor-Pac

10

Z

®A J

F~ Z

0 0

l 2

I 4

I 6

I 8

[ 10

I 12

14

O B S E R V E D E F F L U E N T DN ( r a g / l )

Fig. 6. Comparisonbetweenobserved and model-predictedeffluent dissolved nitrogen concentrations established by using three packing types during the single column model verification runs.

all effluent DN estimates average 11-33% ( - 1"36-3"67%) with 164 of the 180 predicted values being less than observed concentrations. During elevated influent DN and DC test runs, water temperature ranged between 31-0 and 32.0°C, a averaged 1"00 (0.97-1.04), and DN was increased to levels representing 114-9-123.6% of air saturation concentrations. The addition of nitrogen gas to the supply pump discharge also acted to strip DO. Infhent DO concentrations then ranged between 4.4 and 5"9 mg/litre. Dissolved oxygen concentrations of 6"7-7.9 mg/litre were present during elevated DC test runs. Elevated DC concentrations were between 3.72 and 11.53 mg/litre. The relative error of model-predicted effluent DO concentrations averaged 10-57% ( - 7.15-15"6%) with high influent DC concentrations, and 9.65% (6"12-10.60%) with elevated influent DN. These values are similar to the mean relative error of 9.59% (4.47-15.52) established in the first series of runs in which the same packing type (5-08 cm Nor-Pac ®) and hydraulic loading rate (32.0 kg/m 2 s) were used. The relative error associated with model predicted effluent DN concentrations, however, were higher than in earlier rung, averaging 24.52%

48

B. J. Watten, C. E. Boyd

( - 34.14 to - 6-27%) for the high DN and 16.43% ( - 22-31 to - 4.79%) for the high DC test data. In these two test series, all model-predicted concentrations were lower than the concentrations observed.

Multi-stage packed column analyses Throughout the course of the multi-stage runs, water temperature ranged between 25.8 and 33"C. Alpha averaged 1.03 with a range of 0.95-1.17. The introduction of nitrogen gas into the influent line increased DN to levels representing 114.9-123.6% of air saturation concentrations, while reducing DO concentrations to 3"8-4.5 mg/litre. During all other runs, influent DO ranged from 7-2 to 8-5 mg/litre. Elevated DC concentrations varied between 9.1 and 9.7 mg/litre. The measured total head loss across the system was 1.00-1-34 m. As in the single column analyses, model predictions were in close agreement with observed effluent gas concentrations. The mean relative error averaged just 5.05% (range - 6-71-19.82%) for the oxygen data and 4.44% (range - 15.38-16.89%) for the nitrogen data. Figures 7-10 give predicted versus observed effluent DO data for each column of the multi-stage system. Inspection of these plots reveals that errors associated with model predictions under each of the four operating categories summarized in Table 3 are similar. Further, regarding DO predictions, there is a trend toward higher error with increasing column segment numbers, i.e mean relative error for column segments 1-4 were, respectively, 2.93% (N = 38), 4.88% (N-- 38), 5"53% (N-- 30) and 8"07% (N---24). The opposite trend was present in corresponding DN data; mean relative errors for columns 1-4 were (respectively) 5.72%, 5.14%, 3"16% and 2.91%. As a check of model calibration data, we calculated values of the mass transfer coefficient ( G)20 for each of the four column segments, in duplicate, using a selected set of recorded data. These data include temperatures, pressures, influent and effluent dissolved gas concentrations, a and column gas-phase compositions. The observed (G)20 values averaged 1.20, which represents 88-2% of the ( G)20 value predicted by using eqn (6). The predicted value was used in all simulation runs. The overestimate of ( G)2. by the regression equation, developed by using the 30.5 cm diameter column, may result from differences in multi-staged column geometry and distribution plate design. Use of an overestimate of (G)20 in the model comparison runs should increase the relative error above that observed in the single column analyses. The opposite response was observed here; the mean relative error established with the multi-stage

Gas transfer within a packed column oxygen absorber

. LOW PP~ESSURE

49

~ /

A HIGH DN ""

26

0

~,/

**

f .

HIGH DC

~

'

•

Q 24

**

e~ ~

22

!,o 18

16

I 18

16

I 20

I 22

I 24

I 26

I 28

30

OBSERVED DO (rag/l)

Fig. 7.

Comparison between observed and model-predicted effluent dissolved oxygen concentrations established with the first column of a multi-stage absorber.

26 24

-

* LOW PRESSURE o HIGH PRESSURE

22

/

* ~ . / / / *

/

A HIGH DH /

D HIGH DC

.-. 2O

q

~/~

1~1 12 lO 8 6 4 4

| 6

I O

I 10

I 12

I 14

I 16

I 16

I 20

I 22

I 24

26

OBSERVED DO (rag/l)

Fig. $. Comparison between observed and model-predicted effluent dissolved oxygen concentrations established with the second column of a multi-stage absorber.

50

B. J. Wauen, C. E. Boyd

28 26

~

LOW PRESSURE

/ /

O HIGH PRESSURE

24

/ O /

n HIGH DN

2 0 ~22 -.~

O

Z~t. /

O HIGHDC

.i ~ ~ .

18

~

~

O

14.

I0 8 6

r+/

6

l

I

I

l

I

Z

I

8

10

12

14

16

18

20

I

22

I

I

24

26

28

OBSERVED DO (mg/l)

Fig. 9.

Comparison between observed and model-predicted effluent dissolved oxygen concentrations established with the third column of a multi-stage absorber.

2,5 ,~ LOW PRF.,SSURE O HIGH PRESSURE 20

r, HIGH DN D HIGH DC

/

I,5 0

10

,'0

L 20

2,5

OBSERVED DO (rag/l)

Fig. 10. Comparison between observed and model-predicted effluent dissolved oxygen concentrations established with the fourth column of a multi-stage absorber.

Gas transfer within a packed column oxygen absorber

51

system averaged 5-05%, versus 8"63% in the single-stage system. Nevertheless, mean relative errors in both types of systems were considered acceptably low. MODEL APPLICATION After verification, the simulation program was used to identify the operating characteristics and relative performance of multi-stage packed column equipment. Unless otherwise noted, all simulation runs were performed with the following operating conditions: temperature, 20"C; a, 1.0; t , 1.0; atmospheric pressure, 760 mm Hg; column pressure, atmospheric; DO, 9"08 mg/litre; DN, 14-88 mg/litre; DC, 0 mg/litre; inlet gas composition, 100% oxygen; water flow rate, 0.38 ma/min; hydraulic loading, 64.8 kg/m 2 s; packing type, 3"81 cm ACTIFIL®; and relative (G)2 0 for nitrogen, 0.94. Total head loss for a given (G)20 was taken as the bed depth calculated using the ACTIFIL® regression correlation (eqn (6)), plus 0.1 m to account for the assumed presence of water in the distribution plate wells. Results of a series of runs used to establish the change in gas composition and relative gas flow within each stage of a 10-stage system are given in Fig. 11. Values of (G)~0 for each column were based on 0"5 m of packing. Data were generated by using volumetric oxygen-water ratios of 0.005, 0.01 and 0-02. The concurrent absorption-desorption processes that occurred within each packed bed segment reduced the oxygen purity of the gas phase, as well as off-gas flow. The extent of the changes increase with lower system oxygen-water ratios, e.g. with a gas liquid ratio of 0"02, the mole fraction of oxygen, X02, dropped from 1"00 to 0.69, whereas with a gas-liquid ratio of 0.005, X02 drops further to 0"22. Values of X02 below atmospheric levels (0.21) can be achieved when treating waters with DO below saturation concentrations. The changes in gas phase composition along the column segment series are reflected in the corresponding stage wise plots of effluent dissolved gas concentrations given in Fig. 12. The change in DO and DN across the column segment is reduced with lower values of X02. Results of a series of program runs performed to establish the sensitivity of multi-stage column performance to changes in bed depth, number of column segments, gas-liquid ratio, and influent DO concentrations are summarized in Figs. 13-15. Relative performance in these plots is indicated by the increase in oxygen absorption efficiency (AE) realized over that achieved with a single stage system. During runs summarized in Fig. 13, the volumetric gas-liquid ratio was held at 0.01 while the

B. J. Watten, C. E. Boyd

52 100

90

80 v

a¢ 0 r.i)

70

60

0 50

l

I

l

I

I

1.0

r~

0

0.8

•To

o.e

.J~

0.4

0.2

* 0.005 o 0,01 ,', 0,02 A

i

i

i

2

4

6

8

i

10

STAGE NUMBER Fig. 11. Predicted changein gas composition and relative gas flow within each stage of a 10-stage oxygen absorber at three operating volumetric gas-liquid ratios.

number of stages employed was varied from 1 to 12 at each of two bed depths. The two bed depths (0.28 m and 1.02 m) are indicated by (G)2 0 values of 0"6 and 1.6. In both series of runs, the change in AE increased at a decreasing rate with number of stages. The asymptotic percentage value approached is greater in the system with the shallow bed than the system employing the deep bed. Further, it is apparent that fewer column segments are required in the deep bed system to realize the benefit afforded through use of the multi-stage configuration. Figure 14 illustrates the effect of operating volumetric gas-liquid ratio on multi-stage system performance. In this test (G)20 was held at 0.6 while the number of stages was varied from 1 to 12 at each of two gasliquid ratios (0.01 and 0.03). The data plotted indicate that the increase in AE resulting from multi-stage operation will be greater when operating gas-liquid ratios are relatively low, e.g. with 12 stages, the increase in AE over single colurrm performance is 20.1% and 11.1%, respectively, for gas-liquid ratios of 0.01 and 0.03.

Gas transfer within a packed column oxygen absorber

53

35

30 Z

25

~

20

0 CO

10

0

i

5 ~

i

i

L

,

16

E 14 Z

* 0.005 oo.o

/J

~.---'~ -

0.02

0 Z ~

8

,-1 0

6

m 40

2

4

6

8

STAGE NUMBER Fig. 12.

Predicted stage-by-stage plots of effluent dissolved oxygen and nitrogen obtained at three volumetric gas-liquid ratios.

The influence of inlet DO on multi-stage column performance is demonstrated in Fig. 15. Again, ( G)2 o was held constant at 0.6 along with a selected gas-liquid ratio of 0.01. The number of stages was varied from 1 to 12 by using influent DO concentrations of either 0 or 9-08 mg/litre. Inspection of the figure reveals that relative performance of the multistage system is greater when concentrations of DO in the influent are high. A separate series of program runs were undertaken to demonstrate the reduction in oxygen flow that may be allowed, through multi-stage operation, for a given required effluent DO concentration. The value of ( G)20 was set at 0-9 which corresponds to a total head loss of 0.5 m. The DN in the influent was set at 120% of air saturation levels. Gas-liquid ratios were varied from 0-005 to 0.03. Plots of required gas-liquid ratios versus effluent DO were established for both a one-stage and 10-stage contact system (Fig. 16). Note that a considerable reduction in required gas flow has been achieved with the multi-stage configuration. The extent

54

B. J. Watten, C E. Boyd 22

(G~ *

20

0.6

o 1.6 18

16

z 14 co < CJ

12

z_

10

8

6

i

z

2

i

4

,

6

8

ltO

12

N U M B E R OF STAGES

Fig. i 3.

Predicted increase in oxygen utilization resulting from multi-stage operation at two bed depths.

25

~/L * 0.01

o 0.03 20

Z 5D ,< c~

I0

Z

0 0

l

2

L

4

I

6

I

8

llo

12

N U M B E R OF STAGES

Fig. 14.

Predicted increase in oxygen utilization resulting from multi-stage operation at two gas-liquid (G/L) ratios.

Gas transfer within a packed column oxygen absorber

55

22

DO ( m ~ / I ) * 20

-

18

-

9.08

o 0.00

v r.~ 16
_z 14

--

12

--

co r_)

z

10

8

l

6

I 4

2

I 6

I 8

t 10

12

N U M B E R OF STAGES

Fig. 15.

Predicted increase in oxygen absorption efficiency resulting from multi-stage operation at two influent oxygen concentrations.

00 3 5 No.

00 3 0 ~q

o:

Stages

*

1

o

10

f

~

/

0.025

C

E

0.020

v 0015 r.~ 0 010 C~

0005I 0000

12

114

L1

6

J

18

~

20

I

22

I

24

26

REQUIRED EFFLUENT DO (mg/1) Fig. 16.

Predicted gas-liquid ratio (G/L) required to achieve a given effluent dissolved oxygen concentration with a single stage and 10-stage absorber.

B. J. Warren, C. E. Boyd

56 100

9O

Z

o

Z r~ 0

70-

Z

60 0

NO

I

[

0.005

0.010

I

I

0.015

I

0.020

0.025

0.030

G/L (dimensionless)

Fig. 17.

Predicted effluent dissolved nitrogen concentration as a function of gas-liquid ratio for a single stage and I O-stage absorber.

1.6 1.4 v

1.2

No. Stages

o10

1.0 Z ,-.1 0

0.8 0.6

~

0.4 0.2

060

I

65

7tO

715

810

I

85

90

AE (Z)

Fig. 18.

Predicted column height required to achieve a given oxygen absorption efficiency for a single stage and 1O-stage absorber.

Gas transfer within a packed column oxygm~ absorber

57

12

No. S t , ~ u 10

*

1

o

10

8 J:

6 -

S-, 4

2

0

Fig. 19.

60

70

5

Av (x)

80

65

90

Predicted transfer efficiency required to achieve a given oxygen absorption efficiency for a single stage and 10-stage absorber.

of the savings realized is shown to increase with effluent DO concentrations. For example, a single-stage effluent DO concentration of 16 and 24 mg/litre, dictates gas flows 29% and 55%, respectively, above that predicted for the 10-stage configuration. Packed bed oxygen contact equipment often is used to reduce DN to or below saturation for controlling gas bubble disease. The extent of nitrogen stripping is regulated by adjusting oxygen flow or system operating pressures. Figure 17 shows effluent DN data, as per cent of air saturation concentrations, corresponding to the simulation runs used to develop Fig. 16. Inspection of Fig. 17 reveals that DN desorption across the 10-stage contactor exceeds that in the single stage system at a given volumetric gas-liquid ratio. The improvement in DN stripping increases with the operating gas-liquid ratio. The multi-stage configuration has been developed to reduce the head required for packed column operation to that which often exists (0-3-1-0 m) between the water supply and operating water level of raceway culture systems (Watten, 1989). Reducing the head requirement would provide a saving in energy costs as well as a reduction in risk of system failure by eliminating the need for pumping. In line with this objective, the total column height required for a selected AE has been established for both a one-stage and 10-stage absorber (Fig. 18). Simulation data were generated by using a gas-liquid ratio of 0.008. The (G)20

58

B. Z Watten, C. E. Boyd

values were varied from 0.5 to 2.1. The predicted shift of the 10-stage data to the right of the single stage plot resulted in a substantial reduction in required column heights. For example, the height required by the 10-stage system to achieve an AE of about 76.5% (0-27 m) was only 19-4% of that required by the single column absorber (1.39 m). In this example an AE value of about 85% is obtained with a column height of less than 0.7 m. Thus, it appears the multi-stage configuration is capable of reducing column heights to the desired range of 0.3-1.0 m . Figure 19 compares TE corresponding to the head loss and AE values presented in Fig. 18. The savings in column height provided by the 10-stage system are reflected here in higher TE ratings. With a required AE of 75%, for example, the TE for the 10-stage system is 9"5 kg/kW h--a 413% improvement over the corresponding TE of 2.3 kg/ kW h established for the single-stage system. Clearly, the benefits afforded by using the multi-stage configuration can be substantial, particularly in applications employing shallow beds, high influent DO concentrations, low gas-liquid ratios, and high required values of AE.

REFERENCES APHA (1985). Standard Methodsfor the Examination of Water and Wastewater, 16th edn. American Public Health Association, Washington, DC, 1268 pp. ASHRAE (1972). Handbook of Fundamentals. American Society of Heating, Refrigeration and Air Conditioning Engineers, New York. Boerson, G. (1985). Michigan's use of pure oxygen for gas supersaturation control. Paper presented at Annual Meeting, American Fisheries Society, Sun Valley,Idaho (unpublished). Boyd, C. E. (1979). Water quality in warmwater fish ponds. Ala. Agric. Exp. Stn., Auburn University, Alabama. Colt, J. (1984). Computation of dissolved gas concentrations in water as functions of temperature, salinity and pressure. American Fisheries Society, Bethesda, Maryland, Special Publication No. 14. Colt, J. & Bouck, G. R. (1984). Design of packed columns for degassing. Aquacult. Engng, 3, 251-73. Colt, J. & Watten, B. J. (1988). Applications of pure oxygen in fish culture. Aquacult. Engng, 7,397-441. Dwyer, W. E, Kindishi, G. A. & Smith, C. E. (In press). Evaluation of high and low pressure oxygen injection techniques. American Fisheries Society Bioengineering Symposium, Portland, OR. Hackney, G. E. & Colt, J. E. (1982). The performance and design of packed column aeration systems for aquaculture. Aquacult. Engng, 1,275-95. Levenspiel, O. (1979). The Chemical Reactor Omnibook. Oregon State University, Corvallis, Oregon. Lewis, W. K. & Whitman, W. C. (1924). Principles of gas adsorption. J. Ind. Engng Chem., 16, 1215-20.

Gas transferwithina packed column oxygenabsorber

59

Mueller, J. A., Mulligan, T. J. & Ditro, D. M. (1973). Gas transfer kinetics of pure oxygen systems. J. Environ. Engng Div., ASCE, 99, 269-82. Nirmalakhandan, N., Lee, Y. H. & Speece, R. E. (1988). Optimizing oxygen absorption and nitrogen desorption in packed towers. Aquacult. Engng, 7, 221-34. Severson, R. E, Stark, J. L. & Poole, L. M. (1986). Use of oxygen to commercially rear coho salmon. Papers on the Use of Supplemental Oxygen to Increase Hatchery Rearing Capacity in the Pacific Northwest, ed. G. R. Bouck. Bonneville Power Administration, Portland, OR, pp. 25-34. Speece, R. E. (1981). Management of dissolved oxygen and nitrogen in fish hatchery waters. In: Proceedings of the Bioengineering Symposium for Fish Culture, eds L. J. Allen & E. C. Kinney. American Fisheries Society, Bethesda, Maryland, pp. 53-62. Speece, R. E., Nirmalakhandan, N. & Lee, Y. (1988). Design for high purity oxygen absorption and nitrogen stripping for fish culture. Aquacult. Engng, 7, 201-10. Speece, R. E., Eheart, J. W. & Givler, C. A. (1983). U-tube aeration sensitivity to design parameters. J. Wat. Poll. Control Fed., 55, 1065-9. Speece, R. & Orosco, R. (1970). Design of U-tube aeration systems. J. San. Engng Div., ASCE, 96, 715-25. Warren, B. J. (1989). Multiple stage gas absorber. US Patent, Department of Commerce, Washington, DC, No. 4,880,445. Warren, B. J. & Beck, L. T. (1985). Modeling gas transfer in a U-tube oxygen absorption system: effects of off-gas recycling. Aquacult. Engng, 4, 271-97. Watten, B. J. & Boyd, C. E. (1989). Gas phase axial dispersion in a packed column oxygen absorber. Aquacult. Engng, 8, 421-34. Warren, B. J., Colt, J. & Boyd, C. E. (In press). Modeling the effects of dissolved nitrogen and carbon dioxide on the performance of pure oxygen absorption systems. American Fisheries Society Bioengineering Symposium, Portland, OR. Weiss, R. E (1970). The solubility of nitrogen, oxygen and argon in water and seawater. Deep-Sea Res., 17, 721-35. Weiss, R. E (1974). Carbon dioxide in water and seawater: the solubility of a non-ideal gas. Mar. Chem., 2, 203-15. Westers, H., Boerson, G. & Bennett, V. (In press). Design and operation of sealed columns to remove nitrogen and add oxygen. American Fisheries Society BioengineeringSymposium, Portland, OR.