A THEORETICAL AND EXPERIMENTAL STUDY OF AIRLIFT PUMPING�

AND AERATION WITH REFERENCE TO AQUACULTURAL APPLICATIONS�

A Thesis

Presented to the Faculty of the Graduate School

of Cornell University

in Partial Fulfillment of the Requirements for the Degree of

Doctor of Philosophy

by�

Douglas Joseph Reinemann�

August 1987�

A Theoretical and Experimental study of Airlift Pumping

and Aeration with Reference to Aquacultural Applications

Douglas Joseph Reinemann, Ph.D.

Cornell University 1987

A theoretical and experimental study was conducted

pertaining to the pumping and aeration properties of the

airlift pump and its application in intensive

aquaculture facilities. The results and discussion of a

study of the effects of tUbe diameter on vertical slug

flow, specifically as it relates to 3 - 25 mm airlift

pump performance, are presented in Chapter One: Theory

of Small Diameter Airlift Pumps. The theory previously

presented by Nicklin (1963) is extended into this range

of tUbe diameters by taking into account the effects of

surface tension on bubble rise velocity. Differences

are noted between the rise velocity of a single gas slug

and a train of gas slugs in small vertical tubes.

Comparisons are made between experimental results and

theoretical predictions.

The results and discussion of a study of the flow

dynamics of a 38 mm diameter airlift pump, are presented

in Chapter Two: Hydrodynamics of the Airlift Pump in

Bubble and BUbbly-Slug Flow. Experimental flow patterns

ranged from dispersed bubble flow to bubbly-slug flow.

The effects of initial bubble size and water quality on

flow dynamics and flow pattern transition are examined.

Experimental data are compared with previous two phase

flow models and a new prediction equation is developed

for the bubbly-slug flow regime.

The results of an experimental study of the oxygen

transfer properties of a 38 mm diameter airlift pump are

presented in Chapter Three: Oxygen Transfer in Airlift

Pumping. The effects of varying initial bubble size,

flow rate, flow pattern, and water quality on oxygen

transfer are examined. A model to predict oxygen

transfer in airlift pumping is presented.

The results of an energy and cost analysis of

salrnonid production in water reuse systems is presented

in Chapter Four: Energy and Cost Analysis of Salmonid

Production in Water Reuse Systems. Various options to

increase system efficiency, including the use of

airlifts for pumping and aeration, are considered. The

energy inputs for salmonid production in water reuse

systems are compared with land based animal protein

production, other forms of aquaculture and traditional

fishing.

Douglas Joseph Reinemann 1987 ALL RIGHTS RESERVED

Biographical Sketch

Douglas Joseph Reinemann was born on January 25,

1958 in Frankfurt, west Germany, the second of five

children to Dr. John M. and Mrs. Joan Hug Reinemann.

In 1961 his father finished his tenure with the u.S.

Army and the family moved to Milwaukee, Wisconsin for

two years and then to Sheboygan, Wisconsin. Douglas

graduated from Sheboygan North High School in June of

1976, and enrolled in the University of Wisconsin­

Madison, where he received a B.S. in Agricultural

Engineering in December of 1980. He then worked as a

volunteer at the st. Francis Mission on the Rosebud

sioux Indian reservation for nine months. It was during

this time that he gained an understanding of and

appreciation for Lakota thought and philosophy. He

returned to the University of Wisconsin to complete an

M.S. in Agricultural Engineering in august of 1983.

Upon completion of his M.S. he returned to the Rosebud

for a period of one year to continue his studies of

Lakota and to work with the Wanekiya Cooperative, a

group which was formed during his first visit there. In

August of 1984, he was wed to Mary Kay Hauck, of

Missoula, Montana, who had been a teacher on the Rosebud

Reservation, at the st. Francis Indian School.

iii

Douglas entered in a Doctoral Program in

Agricultural Engineering at Cornell University in

September of 1984, where he has studied aquaculture and

water management. During his tenure at Cornell, the

couples first son, Joseph John, was born. The couple is

currently expecting their second child.

iv

Dedication

For my wife, my children, and all my relatives.

Mitakuye Oyasin

v

Acknowledgments

I would like to express my appreciation to the

members of my committee, M.B. Timmons, J.Y. Parlange,

and D. Pimentel. It has been an honor and a privilege

to work with them. I would also like to acknowledge the

teachers and fellow graduate students at Cornell who

have made my time here interesting and enjoyable:

especially Rakesh Gupta, Marc Parlange, and Dr. Zelman

Warhaft.

vi

Table of Contents

Chapter One Thcnry ~f ~ma]l oiamatQr Airlift Pumps 1

l.nt:ICoduul~on 1 Theory . . . . . . 3 Experimental Procedure 11 Results and Discussion 12 Conclusion 17

Chapter Two Hydrodynamics of the Airlift Pump

in Bubble and Bubbly-Slug Flow. 27 Introduction . . . 27 Theory . . . . . . . . . . . . . . . 29 Experimental Procedure 35 Results and Discussion . 37 Conclusion . . . . . . 42

Chapter Three Oxygen Transfer in Airlift Pumping 52

Introduction . . . .. . 52 Experimental Procedure . . . . 56 Results and Discussion . . . 58 Conclusion . . . . . . . 60

Chapter Four Energy and Cost Analysis of Salmonid Production

In Water Reuse Systems . 66

Comparison with Other Forms of Protein

Appendix A

Appendix B

Introduction . . . . . . . .. . 66 Salmonid Production in Closed Systems . 71

Production . . . . . . . 81 Conclusion . . . . . . . . . 82

Energy and Cost Analysis Details 88

Thermal Model Details 91

References . 93

vii

Table 1.1

Table 2.1

Table 2.3

Table 2.2

Table 3.1

Table 4.1

Table 4.2

Table 4.3

Table 4.4

Table 4.5

Table 4.6

List of Tables

Summary of Airlift Equations 18

Nomenclature and Definitions 43

Summary of Airlift Equations 44

Water Quality Parameters 44

Water Quality Parameters 61

US Fishery Products Supply 84

consumption of Selected Protein Products in the US . . . . . .. ... 84

Energy and Cost Analysis 85

sensitivity Analysis 86

Energy Inputs for Various Protein Production Systems . . . . 87

Protein Production and Land Area 83

viii

List of Figures

Figure 1.1. Typical Airlift Pump. · · · · · · · 19 Figure 1. 2. Experimental Apparatus. 20· · · · · · Figure 1. 3. Velocity Profile Coefficient vs.

Reynolds Number. · · · · · · · 21 Figure 1. 4. Efficiency vs. Gas Flow, 3.18 mm Tube. 22 Figure 1. 5. Efficiency vs. Gas Flow, 6.35 mm Tube. 23· Figure 1. 6. Efficiency vs. Gas Flow, 9.53 mm Tube. 24 Figure 1. 7. Theoretical Efficiency vs. Gas Flow. 25 Figure 1. 8. optimum Flow Characteristics vs. Tube

Diameter. . · · · · · · · · · · 26

Figure 2.1. Typical Airlift Pump. 45· · ·· Figure 2 . 2 . Flow Patterns. 46· · · · ·· · Figure 2 . 3 . Local liquid slug gas void ratio, bubble and liquid velocities in bUbbly-slug flow. 47· · · · · ·· · Figure 2.4. Experimental Apparatus. 48· ·· · · Figure 2.5. Bubble Flow Data. Experimental vs. Predicted Average Gas Velocities. 49

Figure 2.6. Bubbly-Slug Flow Data. Average Gas Velocity vs. Average Mixture Velocity. . 50· · · · · · · · · · · Figure 2.7. 38 mm Diameter Tube Flow Pattern Map. 51·

Figure 3.1. Typical Airlift Pump. 62· · · · · · · Figure 3 .2. Flow Patterns in Airlift Pump operation. 63· · · · · · · Figure 3 . 3 • Experimental Apparatus. 64· ·

Figure 3.4. Oxygen Transfer Coefficient vs. Reynolds Number 65· · · ·· · · · · ·

Figure 4.1. Thermal Model Detail. 92· · · · ·· ·

ix

Chapter One

Theory of Small Diameter Airlift Pumps

Abstract: The results and discussion of a study of the

effects of tube diameter on vertical slug flow,

specifically as it relates to 3 - 25 mm airlift pump

performance are presented. The theory previously

presented by Nicklin (1963), is extended into this range

of tube diameters by taking into account the effects of

surface tension on bubble rise velocity. Differences

are noted between the rise velocity of a single gas slug

and a train of gas slugs in small vertical tubes.

Comparisons are made between experimental observations

and theoretical predictions.

Introduction

A typical airlift pump configuration is illustrated

in figure 1.1. A gas, usually air, is injected at the

base of a submerged riser tube. As a result of the gas

bubbles suspended in the fluid, the average density of

the two-phase mixture in the tube is less than that of

the surrounding fluid. The resulting buoyant force

causes a pumping action.

The slug flow regime is most widely encountered in

airlift pump operation and is characterized by bubbles

large enough to nearly span the riser tUbe. The length

of the bubbles ranges from roughly the diameter of the

1

2

tube, to several times this value. The space botWG8n

the bubbles is mostly liquid filled and is referred to

as a liquid slug (Govier and Aziz, 1972). The large gas

bubble is also referred to as a gas slug or Taylor

bubble.

Extensive experimental and theoretical work has

been done on airlift pumps in the slug flow regime

(Apazidis, 1985; Clark and Dabolt, 1986; Hjalmars, 1973;

Higson, 1960; Husain and Spedding, 1976; Jeelani et al.,

1979; Nicklin, 1963; Richardson and Higson, 1962;

Sekoguchi et al., 1981; Slotboom, 1957; Stenning and

Martin, 1968). These studies have been confined to

air/water systems in tubes with diameter greater than 20

mm in which the effects of surface tension are small and

have therefore been neglected.

As tube diameter is decreased below 20 mm, the

effects of surface tension on the dynamics of vertical

slug flow become increasingly important (Bendiksen,

1985, Nickens, and Yannitell, 1987; Tung and Parlange,

1976; White and Beardmore, 1962; Zukoski, 1966). It has

been speculated that increased efficiency might be

obtained by using small diameter tubes at low flow rates

(Nicklin, 1963). Neither a satisfactory theory, nor

conclusive experimental evidence, however, has as yet

been presented for small diameter airlift operation.

The objective of this study is to examine the effects of

tube diameter on the hydrodynamics of the airlift pump

3

in the range of tube diameters in which surface tension

effects are significant.

Theory

In previous work, the rise velocity of a gas slug

in a vertical tube relative to a moving liquid slug has

been described by (Bendiksen, 1985; Collins et al.,

1978; Griffith and Wallis, 1961; Nicklin et al., 1962):

[ 1 . 1 ]

where

Vt = rise velocity of Taylor bubble (mjs)

Vts = rise velocity of Taylor bubble

in still fluid (mjs)

Co = liquid slug velocity profile coefficient

Vm = mean velocity of the liquid slug (mjs)

given by:

[ 1 . 2 ]

where

QI = vOlumetric liquid flow rate (m3js)

Qg = vOlumetric gas flow rate (m3 js)

A = tube cross sectional area (m2 )

Following the analysis used by Nicklin (1963), the

velocity of the Taylor bubble is set equal to the

average linear veIOC~~y of the gas in the riser tube:

Vt = ~ [1.3]

where

E = gas void ratio

4

It is convenient to express the vOlumetric liquid

and gas flows and bubble velocity in dimensionless form

as Froude numbers defined by:

QgQl' = ----1' Qg' = ----1' Vts'

1 [1. 4 ]

A (g D) 2" A (g D) 2" (g D) 2"

where

Ql' = Dimensionless vOlumetric liquid flow

Qg , = Dimensionless vOlumetric gas flow

,Vts = Dimensionless bubble rise velocity

in still fluid

D = tUbe diameter (m)

g = acceleration due to gravity (mj s 2)

From [l.lJ - [1.4J, the gas void fraction in the riser

tube can be expressed as:

[1. 5 J E = Co (Ql I + Qg') + Vts I

The submergence ratio is a parameter commonly found

in airlift analysis and is defined as:

[1. 6 J

where

~ = submergence ratio

Zl = lift height (m) (See figure 1.1)

Zs = length of tUbe submerged (m)

The submergence ratio is equal to the average

pressure gradient along the riser tube which is made up

of components due to the weight of the two phase mixture

and frictional losses. Performing a static pressure

5

balance on a vertical tube which is submerged in fluid

(see figure 1.1), it follows that:

[1. 7]

where

p = fluid density (kgjm3 )

This assumes that the weight of the gas is negligible

relative to the weight of the liquid. If the fluid in

the tUbe is moving, an additional pressure drop due to

frictional losses must be added to the right hand side

of [1.7]. The single phase frictional pressure drop can

be calculated based upon the mean slug velocity as:

Ps = f [1. 8]

where

Ps = single phase frictional pressure drop

(Njm2 )

f = friction factor (Giles, 1962) 0.316

= [1. 9]Re 0.25

Re = [1. 10]v

v = kinematic fluid viscosity (m2js)

The single phase frictional loss must then be mUltiplied

by (I-E), the fraction of the tube occupied by the

liquid slugs, to obtain the total frictional pressure

drop in the riser tube. The frictional effects in the

liquid film around the gas bubble have been shown to be

6

small compared to those in the liquid slug and arQ

therefore neglected (Nakoryakov et al., 1986) .

Including the frictional effects in the pressure

balance results in:

Dividing both sides by [p g (Zs+Zl)J and rearranging

gives:

ex = (I-E) (1 + f/2 (Ql' + Qg') 2) [1. 12 J

Thus, for a given tube diameter, imposing the gas flow

rate and the submergence ratio, the liquid flow rate may

be determined using the system of equations summarized

in Table 1. 1.

It is usual to define the efficiency of the airlift

pump as the net work done in lifting the liquid, divided

by the work done by the isothermal expansion of the air

(NiCklin, 1963):

n = efficiency

Pa = atmospheric pressure (N/m2 )

Po = pressure at base of riser tube (N/m2 )

Nicklin (1963) introduced the concept of point

efficiency which is accurate in describing total airlift

efficiency to within 1% for submergence lengths of up to

5 meters:

7

n = [1. 14 ]Qg' ex:.

Two important effects become significant when

airlift tube diameter is below about 20 mm. The first

is increased importance of surface tension. The second

is decreased Reynolds number. The effects of surface

tension can be characterized by the inverse E6tvos

number or surface tension number, ~, defined as:

a ~ = ------: [1. 15]

2 p g D

where

~ = surface tension number

a = surface tension (N/m)

White and Beardmore (1962) have defined a

dimensionless parameter which relates only to the

properties of the fluid and expresses the relative

importance of viscosity to surface tension:

y = -- [1. 16]

where

Y fluid viscosity/surface tension parameter

~ = fluid viscosity (kg/m s)

Experimental results have shown that when this parameter

is below 10-8 (which is the case for an air/water

system) viscosity does not influence bubble rise

velocity in still fluid (White and Beardmore, 1962).

8

Theoretical and experimental analyses of th0 rise

velocity of a single gas slug in still fluid have shown

that when both surface tension and viscous effects are

negligible, the bubble Froude number in still fluid (B)

assumes a constant value of about 0.35 (Bendiksen, 1984;

Collins et al., 1978; Davies and Taylor, 1950; Higson,

1960; Nakoryakov et al., 1986; Nickens and Yannitell,

1987; Nicklin et al., 1962; White and Beardmore, 1962;

Zukoski, 1966). This is the value which has been used

in previous airlift analysis (Nicklin, 1963; Clark and

Dabolt, 1986).

The bubble Froude number in still fluid is

influenced by surface tension when the surface tension

parameter is above about 0.02 (Bendiksen, 1985;

Bendiksen, 1984; Nickens and Yannitell, 1987; Tung and

Parlange, 1976; Zukoski, 1966). This corresponds to a

tube diameter less than about 20 mm in an air/water

system. As the tube diameter is decreased below this

value the bubble Froude number decreases. When the

surface tension number is above about 0.2 the bubble

will not rise in still fluid and the value of the bubble

froude number is zero. This corresponds to a tUbe

diameter of about 6 mm in an air/water system. When the

effects of viscosity can be neglected, as is the case

for an air/water system, the bubble Froude number in

still fluid can be expressed as a function of the

9

surface tension parameter alone (Nickens and YannitQll,

1987; White and Beardmore, 1962):

Vts' = 0.352 (1 - 3.18 ~ - 14.77 ~2) [1.17J

Correction can also be made on B for other gas/liquid

systems when viscous effects are significant (Nickens

and Yannitell, 1987; White and Beardmore, 1962).

Theoretical analyses of bubble rise velocity have

applied potential flow theory at the bubble tip,

expressing the stream function of the flow in terms of a

Bessel function series of the first kind and first order

(Bendiksen, 1985; Nickens and Yannitell, 1987; Tung and

Parlange, 1976). This treatment of the hydrodynamics

only at the bubble tip has been justified by several

experimental studies in which air/water bubble dynamics

have been shown to be independent of bubble length

(Nicklin et al., 1962; Griffith and Wallis, 1961).

The effects of surface tension are accounted for in the

application of the boundary condition of constant gas

pressure along the bubble surface. As the radius of

curvature of the bubble is reduced, surface tension acts

to increase the pressure at the gas/liquid interface.

This changes the flow dynamics at the bubble surface and

hence the bubble rise velocity.

Nicklin et aI, (1962), have shown that a value of

1.2 for Co is suitable when the liquid slug Reynolds

number is above 8000. For airlift pumps with diameter

greater than 20 mm, the Reynolds number is usually above

10

8000. The Reynolds number can be considerably less than

8000 when airlift diameter is less than 20 mm, however.

An increase in the velocity profile coefficient has

been observed for Reynolds numbers below 8000

(Bendiksen, 1985; Nicklin et al., 1962). The limiting

value of the velocity profile coefficient has been found

to be about 2 for Reynolds numbers approaching zero.

This rise in Co has also been predicted theoretically

when a laminar velocity profile was imposed in the

liquid ahead of the gas slug (Bendiksen, 1985, Collins

et al., 1978).

Bubble rise velocity as expressed in [1.1J can thus

be interpreted as its rise velocity in still fluid plus

the velocity of the fluid encountered at its tip. The

velocity profile coefficient is then the ratio of the

liquid velocity at the tube axis to the average velocity

of the liquid slug. The limiting values of Co (1.2 for

high Reynolds numbers and 2.0 for low Reynolds numbers)

reflect either turbulent or laminar velocity profiles in

the liquid slug.

Neglecting frictional effects, the efficiency of

the airlift from [1.5J, [1.10J and [1.12J is:

[1. 18 J n = Co (Ql '+ Qg') + Vts' - Qg'

Decreasing tube diameter in the range where surface

tension effects are significant will decrease the value

of the bubble froude number, Vts'. This will increase

11

efficiency. Previous experimental work has shown that 3

reduction in the liquid slug Reynolds number will

increase Co if the transition to a laminar velocity

profile occurs in the liquid slugs. This will reduce

efficiency. Thus, two opposing effects are predicted.

An experiment was performed to determine the relative

importance of the two effects.

Experimental Procedure

The test apparatus is illustrated in figure 1.2.

The reservoir and return sections were glass tube with a

38 mm inside diameter. The riser tubes were 1.80 m in

length and ranged in inside diameter from 3.18 mm to

19.1 mm. Volumetric air and water flow rates, bubble

rise velocity, submergence, and lift height were

measured after the flow stabilized for each trial.

Air and water flows were determined by means of

pressure drop measurements across calibrated orifices.

Bubble rise velocities in both still and moving liquid

were determined by timing a bubble over a known travel

distance. The flow was allowed to develop for a

distance of 0.8 meters before bubble velocity

measurements were started. Slug flow developed within 1

to 5 diameters of the entrance for all of the riser

tubes and flow rates tested.

The static head at the pressure tap immediately

before the riser tube was used as a reference level in

determining lift height and submergence (see figure

12

1.2). This same pressure was used as the air inlot

pressure. By using this pressure as a reference, all

losses in the water return line, air supply line and

across the orifices were separated from the riser tube

measurements. The resulting experimentally measured

flow variables are therefore as close as possible to

measuring the conditions of the riser tube alone.

Submergence ratios were varied by changing the

amount of fluid in the reservoir. Air was injected into

the system by means of a small diaphragm type compres­

sor. Air flow rate was controlled by a valve between

the compressor and the air flow measurement orifice.

The velocity profile coefficient was determined

using [l.lJ, and [1.2J with measured flow rates and

bubble rise velocities. The experimental efficiency was

determined using [1.12J with measured values of liquid

flow, gas flow and submergence ratio.

Results and Discussion

For all of the tube sizes tested, the bubble rise

velocity in still fluid corresponded very closely to the

prediction equation used and results reported by

previous workers (White and Beardmore, 1962; 1976;

Zukoski, 1966). Experimental results for tubes with

3.18, 6.35, and 9.53 mm diameters, showed the velocity

profile coefficient scattered closely about 1.2 with no

increasing trend for Reynolds number decreasing to as

low as 500 (See figure 1.3). This differs from earlier

13

results in which the velocity profile coefficient

increased for Reynolds numbers below 8000 (Bendiksen,

1984; Nicklin et al., 1962). The experiment was

repeated using a 19.1 mm diameter tube to determine

whether surface tension effects influenced this

phenomenon. For this tube size surface tension effects

were negligible as in previous studies. The results

again showed no increasing trend in the velocity profile

coefficient for low Reynolds numbers.

There are two major differences noted between the

previous experiments (Bendiksen, 1984; Nicklin et al.,

1962)and the present one:

1. In the previous experiments the motion of a

single gas slug moving through a moving stream

of liquid was studied. In the present

experiment the gas was introduced continuously

resulting in a series of gas slugs moving

through a series of liquid slugs.

2. The previous experiments used a pump to

regulate the liquid flow whereas in the

present experiment, liquid motion was the

result solely of buoyancy.

When a single gas slug is placed in a stream of

liquid whose motion is pump driven, the velocity profile

in the liquid ahead of the gas slug is a result of

single phase pipe flow. When a series of gas and liquid

slugs rise concurrently, the velocity profile in the

14

liquid slugs is a result of two phase slug flow

dynamics.

The results of the present experiment show that the

liquid slugs have a turbulent velocity profile for

Reynolds numbers as low as 500. Observation of the

motion of very small gas bubbles suspended in the liquid

slug showed erratic behavior further confirming the

presence of turbulence in the liquid slug at low

Reynolds numbers. A laminar velocity profile in the

liquid ahead of the gas slug was observed at low

Reynolds numbers in previous experiments (Bendiksen,

1985; Nicklin et al., 1962). It is believed that this

difference is the result of vorticity generated in the

liquid film surrounding the gas slugs and in their wake

when a series of gas slugs rise concurrently with a

series of liquid slugs. A value of 1.2 was used for the

velocity profile coefficient in all subsequent

theoretical airlift calculations since a turbulent

velocity profile was observed in the liquid slug for the

range of flow conditions tested.

The experimentally determined efficiencies versus

sUbmergence ratio and gas flow are shown in figures 1.4

through 1.6. Theoretical efficiencies for lines of

constant submergence are also shown. The agreement

between theory and experiment is good except when the

gas flow rate is low and the submergence ratio is below

0.7. This region signifies the approach of flow

15

oscillations which are not considered in the theoretical

model.

other workers have observed flow oscillations in

large diameter airlift operation (Apazidis, 1985;

Higson, 1960; Hjalmars, 1973; Sekoguchi et al., 1981;

Wallis and Heasley, 1961;). Oscillations have been

reported to both decrease air-lift efficiency, (Higson,

1960; Richardson and Higson, 1962) and increase

efficiency (Sekoguchi et al., 1981). Measurements taken

in the present study, in the region approaching

oscillatory behavior show efficiencies higher than those

predicted by theory for the tUbe sizes tested in this

regime.

It is instructive to examine the situation in which

no frictional losses are included in theoretical

predictions. This is an excellent approximation to

actual performance at low flow rates when frictional

losses are small. Efficiencies will drop increasingly

below the frictionless case as flow rates increase, (see

figure 1.7). For small tUbes (less than 6 mm diameter)

the bubble froude number in still fluid (Vts) is zero

and the efficiency is constant with respect to gas flow

and increases with increasing sUbmergence ratio in the

frictionless case.

For large tubes (greater than 20 mm diameter), The

bubble Froude number is equal to 0.35, its upper limit,

and frictionless efficiency depends on both submergence

16

and flow rate. Negative values of efficiency occur at

low flow rates, indicating a situation in which work is

done by the expanding gas and no useful work is being

performed pumping the fluid.

For tubes in the intermediate size range (6 mm to

20 mm), the bubble Froude number falls between its upper

and lower limits. Efficiencies fall between the positive

values encountered with small tubes and the negative

values for large tubes as flow rate decreases.

Frictional losses are negligible at low gas flow

rates. Frictional losses increase faster for higher

submergence ratios as gas flow increases. This causes

the characteristic crossing of the constant submergence

ratio efficiency curves (see figure 1.7).

A summary of the optimal flow characteristics of

the airlift pump versus tube diameter is presented in

figure 1.8. Nicklin (1963) concluded that optimal pump

efficiency and submergence ratio were insensitive to

tube diameter. This is indeed the case for air/water

systems when tUbe diameters are above 20 mm and surface

tension effects are negligible. As tube diameters are

decreased below this value, the effects of surface

tension act to increase optimal airlift efficiency and

submergence ratio confirming Nicklin's (1963)

speculations. The maximum attainable theoretical

airlift efficiency is 83% and occurs for tube with

17

diameter less than 6 mm in the limit of zero gas flow

and 100% submergence.

Conclusion

A difference has been observed between single

bubble and bubble train slug flow in air-water systems

at low Reynolds numbers. When a single gas slug rises

in a moving liquid stream the velocity profile

coefficient approaches a value of 2.0 for low Reynolds

number flow in air/water systems. This indicates a

laminar velocity profile in the liquid ahead of the gas

slug. When a series of gas slugs rise concurrently with

a series of liquid slugs, the velocity profile

coefficient remains at a value of 1.2 for Reynolds

numbers as low as 500. This indicates turbulent flow in

the liquid slugs. It is believed that this difference

is the result of vorticity generated in the liquid film

surrounding the gas slugs and in their wake.

It has been shown that including this effect and

the effects of surface tension on bubble rise velocity

allows the airlift pump theory previously described by

Nicklin (1963) to be extended to lower tube diameters of

from 3 mm to 20 mm. It has also been shown that airlift

efficiency and optimal submergence ratio increase in

this range of tube diameters. The theory described here

can be used with confidence to design small diameter

airlift pumps.

Summary Table 1.1

of Airlift Equations

18

E =

Q1'

f

Qg ,

(I-E) (1+f/2 (Ql '+Qg') 2)ex = 1.2(Ql'+Qg')+Vts'

Q1 Qg ZS = Qg

, = ex = 1 1

A (g D) "2 A (g D) "2 Zl + Zs

Vts , = 0.352 (1 - 3.18 2: - 14.77 2: 2 )

0.316 D(Q1 + Qg ) a

= Re = 2: = ReO. 25 l/ A p g D2

19

...-AIR INPUT

Figure 1.1. Typical Airlift Pump.

LIQUID PRESSURE TAPS

/"" LIFT, ~

I •

AIR FLOW ORIFICE ~ LIQUID FLOW ORIFICE

MANOMETER.

COMPRESSOR

Figure 1.2. Experimental Apparatus.

21

1.4

* #.

, '"' 0 0 'V

l-z w 0 ii: 11. w 0

1.3­

1.2

0

," o * # 'i ...

• 0 0,* 0 J t ott I: ,1 ti ",

TO Q t

+ # ,! #

4­

+ q.0

oro + 0

+ t t t

+

to

4­

J..

+ t + +

+

0

w 1.1 0 t .J ii: 0 0:: a. • I

~ 1.0 * D = 3.18 mm

0 0 .J + D = 6.35 mm w >

0.9 - o D = 9.53 mm

• D = 19.1 mm

0.8 I T r T I I r -T , 0 2 4 6 8 10

(Thousands) REYNOLDS NUMBER

Figure 1.3. Velocity Profile Coefficient vs. Reynolds

Number.

22

1.0

0.9 3.18 mm TUBE

0.8 71

NUMBERS ON GRAPH BODY INDICATE PERCENT

SUBMERGENCE OF EXPERIMENTAL POINTS.

SOUD UNES INDICATE THEORETICAL

UNES OF CONSTANT SUBMERGENCE.

0.7

0.6>­t)

Z w

0.50 Ii: IL w

0.4

0.3

0.2

0.1

0.0

0.0

Figure 1.4.

57

~THEORY

sr. THEORY + t 71r. THEORY

1.0 2.0 3.0 4.0

DIMENSIONLESS GAS FLOW (Qg')

Efficiency vs. Gas Flow, 3.18 mm Tube.

23

1.0

86 620.9 6.35 mm TUBE79

620.8 80r. THEORY

¥ 78 62

0.7 t-~-60r. THEOR~

0.6>-

"0 82 z 81 III

0.50 iL 93IL III

0.4

0.3

NUUBERS ON GRAPH BODY INDICATE PERCENT 90r. THEORY

0.2 SUBMERGENCE OF EXPERIMENTAL POINTS.

SOLID LINES INDICATE THEORETICAL

0.1 UNES OF CONSTANT SUBMERGENCE.

0.0 +------.---"""'T"""---.------...........----r-------,-------j

0.0 0.2 0.4 0.6

DIUENSIONLESS GAS FLOW (09')

Figure 1.5. Efficiency vs. Gas Flow, 6.35 mm Tube.

24

1.0

NUMBERS ON GRAPH BODY INDICATE PERCENT

0.9 9.53 mm TUBE SUBMERGENCE OF EXPERIMENTAl.. POINTS.

SOUD UNES INDICATE THEORETICAL

0.8 UNES OF CONSTANT SUBMERGENCE.

0.7

0.6 701. THEORY 48>­0 ..

71z w

0.50 iL lL w

0.4

0.3

0.2

0.1

0.0

0.0 0.2 0.4 0.6 0.8

DIMENSIONLESS GAS FLOW (Og')

Figure 1.6. Efficiency vs. Gas Flow, 9.53 mm Tube.

25

Figure 1.7. Theoretical Efficiency vs. Gas Flow.

26

1.1

1.0 SUBMERGENCE RAno (0)

0.9

0.8

0.7 DIMENSIONLESS UQUID FLOW (QI')

0: w W 0.6 ~ ( 0: (

0.5 tl EFFICIENCY (n)

0.4

0.3

0.2 DIMENSIONLESS GAS FLOW (09')

0.1

0.0

2 4 6 8 10 12 14 16 18 20

DIAMETER (mm)

Figure 1.8. optimum Flow Characteristics vs. Tube

Diameter.

Chapter Two

Hydrodynamics of the Airlift Pump

in Bubble and Bubbly-Slug Flow

Abstract: The results and discussion of a study of the

flow dynamics of a 38 mm diameter airlift pump are

presented. Flow patterns ranged from dispersed bubble

flow to bubbly-slug flow. The effects of initial bubble

size and water quality on flow dynamics and flow pattern

transition are examined. Experimental data are compared

with previous two phase flow models and a new prediction

equation is presented for the bubbly-slug flow regime.

Introduction

A typical airlift pump configuration is illustrated

in figure 2.1. A gas, usually air, is injected at the

base of a submerged riser tube. As a result of the gas

bubbles suspended in the fluid, the average density of

the two-phase mixture in the tube is less than that of

the surrounding fluid. The resulting buoyant force

causes a pumping action.

Extensive experimental and theoretical work has

been done on the airlift pump (Castro et al., 1975;

Clark and Dabolt, 1986; Kouremenos and Staicos, 1985;

Murray, 1980; Nicklin, 1963; Reinemann et al., 1987;

Richardson and Higson, 1962; Slotboom, 1957; Stenning,

27

28

and Martin, 1968; Todoroki et al., 1973). These studies

have been confined, however, to the slug flow regime.

In slug flow, the gas phase is contained in large

bubbles which nearly span the tUbe and range in length

from the tube diameter to several times this value.

These are referred to as gas slugs or Taylor bubbles.

The liquid filling the space between the Taylor bubbles

is referred to as the liquid slug. The liquid between

the Taylor bubbles and the tube wall is referred to as

the liquid film (see figure 2.2).

Other flow patterns are possible in vertical

gas/liquid flow. In bubble flow, the bubble diameter is

much smaller than the tube diameter and the bubbles are

distributed over the pipe cross section. Bubbles remain

close to their initial size, and there is little

interaction between bubbles (see figure 2.2). The

bubble flow pattern has been largely neglected in the

studies of the airlift pump because it has been assumed

to be absent in the useful operating regime. Clark et

ale (1985) presented a theoretical treatment of the

airlift in the bubble flow regime but offers no

experimental verification and does not clearly define

the bubble flow operating regime.

An intermediate regime referred to as bUbbly-slug

flow has been observed in several studies (Akagawa and

Sakaguchi, 1966; Fernandes et al., 1983; Mao and

Duckler, 1985; Nakoryakov and Kashinsky, 1981;

29

Nakoryakov et a1., 19B6~ serizawa et a1., 1975; Shiplay,

1984). In bUbbly-slug flow, small bubbles are found in

the liquid slug (see figure 2.2). The presence of these

bubbles is due to the region of extreme turbulence

encountered at the tail of the Taylor bubble. Small

bubbles are broken off of the Taylor bubble and

dispersed in the liquid slug.

In previous studies of bubble and bubbly-slug flow

dynamics, liquid motion has been pump driven (Akagawa

and Sakaguchi, 1966; Clark and Flemmer, 1985; Fernandes

et al., 1983). In airlift operation the sole driving

force is that developed by buoyancy. As a result of

this difference, many of the previous two phase flow

studies have been conducted at flow velocities much

higher than those encountered in airlift operation. The

objective of this study is to determine the

hydrodynamics of bubble and bubbly slug flow in the

range of gas concentrations and liquid velocities

generally encountered in airlift pump operation.

Theory

Bubble Flow: The drift flux model developed from the

kinetic theory of gasses by Zuber and Findlay (1965) is

widely used to describe two phase flows. The velocity

of the gas phase at a point is taken relative to the

volumetric flux density of the two phase mixture at that

point:

v ' = _J' + V [2.1]-g -gJ.1

30

The nomenclature and definitions used in this paper are

listed in table 2.1. The volumetric flux density does

not, in general, correspond to the velocity of either

phase but is used to represent the average velocity of

the two phase mixture. The average velocity of the gas

phase is obtained by taking a weighted average of [2.1J

over the tube cross section:

<s. Yg '> <s. !:I'> <s. Ygj'> Vg ' = E =---- + [2.2J

E E

Average gas velocity in bubble flow is generally

expressed in the following form:

[2.3J

with the distribution parameter, Cb, defined as:

<E J'>Cb = --=""'""=-:-­ [2.4 J

E <!:II>

The distribution parameter takes into account the

variation across the tube diameter of both the

vOlumetric flux density and the gas concentration. If

the gas concentration is higher than its average in

regions of higher than average flux the parameter will

be greater than one. Conversely if the gas concentration

is higher than its average in regions of lower than

average flux the parameter will be less than one. The

value of the distribution parameter (Cb), has been found

to range between 0.9 and 1.6 (Clark and Flemmer, 1985:

Govier and Aziz, 1972; Nicklin, 1962: Zuber and Findlay,

1965). Values of Cb less than 1 have been found when

31

flow velocities are less than about 1 m/s. Cb is

generally above 1 when flow velocities are above 1 m/s.

The last term on the right hand side of [2.2J is

generally expressed as the rise velocity of a bubble in

still fluid (Vbs'), multiplied by a correction term, (1

- ~ E), to account for reduced bubble velocity due to

the presence of other bubbles (Zuber and Findlay, 1965).

When average flow velocities are above about 1 meter per

second, this effect is often neglected (~ = 0). For

flow velocities less than 1 meter per second the value

of ~ has been found to fall between 0 and 2 (Govier

and Aziz, 1972; Nicklin, 1962; Zuber and Findlay, 1965).

Slug Flow: In slug flow, the average velocity of the

liquid in the liquid slug region can be shown to be

equal to the average volumetric flux density by

continuity considerations (see figure 2 • 2) • The rise

velocity of a Taylor bubble is taken as the rise

velocity of a Taylor bubble in still liquid plus the

velocity of the liquid encountered at the bubble tip and

is set equal to the average gas velocity (Nicklin et

al., 1962):

Vg ' = Cs <~'> + Vts' [2.5 J

The distribution parameter for slug flow, Cs ,

represents the ratio of the liquid centerline velocity

to the average liquid velocity in the slug. Several

workers have found that the distribution parameter

assumes a value of about 1.2 when the Reynolds number is

32

above about 8000 (Bendiksen, 1985; Govipr nnct ~~i~_

1972; Nicklin et al., 1962). This is approximately

equal to the theoretical ratio of centerline to average

velocity in turbulent pipe flow.

Bubbly-Slug Flow: The results of detailed local gas

concentration and liquid and bubble velocity

measurements in bubbly slug flow are summarized in

figure 2.3 (Akagawa, 1964); Akagawa and Sakaguchi, 1966;

Nakoryakov and Kashinsky, 1981; Nakoryakov et al., 1986;

Serizawa et al., 1975). The maximum gas concentration

has been found to occur at a distance of 3 to 5 mm from

the tube wall for tube diameters of 15 to 80 mm. This

corresponds approximately to the average diameter of the

small bubbles in the liquid slug.

The velocity of the small bubbles in the liquid

slug at the tube center is slightly higher than the

velocity of the Taylor bubble. The region just ahead of

the tip of the Taylor bubble can, therefore, be expected

to be essentially free of small bubbles. The small

bubbles near the tube wall move with a velocity equal to

or slightly less than that of the Taylor bubbles. These

bubbles either migrate into the liquid slug and rejoin

the preceding Taylor bubble, or stay near the tube wall

and combine with the following Taylor bubble.

Because of the small bubbles in the liquid slug,

the effective tube area available for liquid flow is

less than the total tube cross section. The velocity of

33

the liquid at the tip of the Taylor bubbles in bUbbly­

slug flow is, therefore, different from that found in a

slug flow free of small bubbles. The drift flux model

can be used to take into account the effect of the small

bubbles on the average liquid velocity in the liquid

slug. Equating the volumetric fluxes entering and

leaving the control volume abcd yields (see figure 2.2):

[2.6J

Retaining the physical interpretation that the rise

velocity of a Taylor bubble is equal to the velocity of

the liquid encountered at its tip plus its rise velocity

in still liquid, the velocity of the Taylor bubble can

be expressed as:

<JI> - <Ygs' ~s>J Vt' = Cbs + Vts' [ 2 • 7 J[ 1-E s

The coefficient Cbs corresponds to the ratio of the

liquid centerline velocity to its average value in the

liquid slug. Assuming that the average velocity of the

gas in the liquid slug is approximately equal to the

velocity of the Taylor bubble, the average gas velocity

in bUbbly-slug flow can be expressed as:

Cbs<J'> + Vts' (1-Es)Vg ' = [2.8J1 + (Cbs-1)Es

In the experimental work of Fernandes et al. (1983), it

was found that the average gas concentration in the

liquid slug, ES' in bubbly-slug flow was about 0.27 and

did not vary significantly when flow parameters were

34

changed. The liquid velocity profile in bUbbly-slug

flow is similar to those found in turbulent single phase

flows. The distribution parameter can therefore be

expected to be about 1.2. The denominator of [2.8] is

then expected to be about 1.05.

The submergence ratio is a parameter commonly found

in airlift analysis and is defined as:

Q = [2.9]

The submergence ratio is equal to the average pressure

gradient along the riser tUbe which is made up of

components due to the weight of the two phase mixture

and frictional losses.

In bubble flow, the frictional loss is generally

taken as the product of the single phase frictional

losses based upon the mean liquid velocity and a two

phase correction factor. Clark (1985) suggests (1 +

1.8E) for the two phase correction factor. In slug flow

the frictional effects in the film around the gas slug

are generally neglected and the single phase frictional

component is multiplied by the fraction of the tube

filled with liquid (I-E). In bubbly slug flow, the

single phase frictional loss must also be multiplied by

the two phase correction factor, (1 + 1.8Es), due to the

bubbles in the liquid slug. Assuming that the average

gas concentration in the liquid slug is about 0.27, the

correction factor is equal to about 1.5.

35

The single phase frictional pressure gradi8nt can

be written as:

F = f/2 <J:,>2 [2.10J

Where the friction factor, f, is obtained from

(Giles, 1962):

f = 0.316 ReO. 25 [2.11J

Thus for bubble flow:

cr = (1 - E) + (1 + 1.8E)F [2.12J

and for bubbly slug flow:

cr = (I-E) (1+1.5F) [2.13J

For a given tube diameter, imposing the gas flow rate

and the submergence ratio, the liquid flow rate may be

determined using the system of equations summarized in

table 2.2.

Experimental Procedure

The experimental apparatus consisted of a circular

loop of glass tubing (38 mm ID) with a 34 liter

reservoir (See figure 2.4). The riser tube was 2.35 m

in length. Volumetric air and water flow rates,

sUbmergence, and lift height were measured after the

flow stabilized for each trial. Calibrated orifices

were used to measure air and water flow rates. The gas

concentration in the riser tube was determined from the

measured gas and liquid flows, and submergence ratio and

equation [2.12J or [2.13J, depending on the flow

pattern.

36

Air was injected into the system by a rotary vanQ

type compressor. Air flow rate was controlled by a

valve between the compressor and the air flow

measurement orifice. Two different diffusers were used;

1) An aquarium air-stone which generated bubbles of 1

to 3 mm diameter, and 2) a 6 mm diameter tube which

generated bubbles of 10 to 15 mm diameter. The air­

stone produced the bubble flow pattern when the gas

concentration was low and the bubbly slug pattern when

the gas concentration was above some critical value,

which depended on the water type. The 6 mm tube

produced the bubbly slug pattern for both water types

and at all gas concentrations. Note that it was not

possible to produce the slug flow pattern, i.e., without

small bubbles, for any of the test conditions.

The static head at the pressure tap immediately

before the riser tube was used as a reference level in

determining lift height and submergence (See figure

2.4). By using this pressure as a reference, frictional

losses in the liquid return lines and across the flow

measurement orifice were separated from the riser-tube

flow measurements.

Tap water and waste water from an aquaculture

facility were the liquids used. Water quality

parameters were determined by standard analytic

procedures at the Cornell University Agronomy lab (see

table 2.3).

37

Results and Discussion

Regression analysis of both the bubble and bUbbly­

slug flow data showed no significant difference in flow

dynamics between the two gas diffusers or the two water

types. Differences were noted, however, in the critical

gas void ratio for transition from bubble to bubbly-slug

flow for the two water types, as discussed below.

Regression of the combined bubble flow data yielded

the following equation (see figure 2.5):

Vg ' = 0.62 <J'> + 0.44 (1-1.4E) [2.14]

(coefficient of correlation, R2 = 0.57)

in agreement with the theoretical form of [2.3]. The

results of the bubble flow data regression are

significantly different than those suggested by Clark et

al., (1985) (Cb = 1.2, ~ = 0). Regression of the

bubble flow data in the form suggested by Clark et al.,

(1985), with no dependence on the gas concentration

resulted in considerably reduced accuracy (Cb = 0.32, ~

= 0, R2 = 0.21). Thus, including the correction on Vbs

due to the presence of other bubbles as suggested by

Nicklin et al., (1962), considerably improves the

accuracy of prediction.

The prediction equation used by Clark et al.,

(1985) was obtained from bubble flow data in which the

liquid motion was forced by a pump. He considered flow

velocities in the range of 1 to 5 m/s. In airlift pump

operation, buoyancy is the sole driving force. In order

38

to increase the buoyant driving force the gas

concentration of the two phase mixture must be

increased. If the gas concentration is increased beyond

a certain point, however, slugging occurs and flow

dynamics change. Thus, flow velocities are limited in

the airlift pump in the bubble flow regime. In the

present study average flow velocities ranged from 0.1 to

0.3 mls in the bubble flow regime.

When flow velocities are high, the rise velocity of

the bubble in still liquid, (Vbs) , becomes negligible in

relation to the average flow velocity. The accuracy of

the bubble flow model is therefore very insensitive to

the value chosen for the bubble rise velocity (Vbs) , and

depends mainly on the choice of the distribution

parameter Cb. When flow velocities are reduced,

however, the bubble rise velocity in still liquid

becomes significant and prediction accuracy depends upon

the proper choice of this value.

The distribution parameter found in this study is

considerably lower than that recommended by Clark et

al., (1985). It has been observed by several workers,

that in low velocity vertical bubble flow, the maximum

concentration of bubbles occurs near the tube wall

(Akagawa, 1964; Akagawa and Sakaguchi, 1966; Nakoryakov

and Kashinsky, 1981; Nakoryakov et al.,1986; Serizawa et

al., 1975). This is also a region of lower than average

flow velocity. It is therefore expected that the

39

distribution parameter (Co) should assume a value less

than one. As flow velocities increase the bubble

concentration profile assumes a parabolic shape with its

maximum at the pipe center. Thus for high speed flows

Co would be expected to be above 1, as observed by Clark

et al. (1985).

When the slug flow model is applied to bubbly slug

flow data, it predicts average gas velocities

consistently higher than the experimental values (see

figure 2.6). Regression analysis of the bubbly-slug

data showed no significant difference in flow dynamics

between water or diffuser types. Regression of the

combined bubbly-slug flow data yielded the following

equation:

v'g = 1.1 <J'> + 0.75 Vts' [2.15]

(R2 = 0.90)

which follows the form of [2.8]. The results of the

bubbly-slug flow data regression agree well with the

model presented above for the bubbly-slug flow pattern.

The bubbly slug model is also very close to the

empirical relationship used by Hills (1976), for average

gas velocity in buoyancy driven bUbbly-slug flow in a

150 mm diameter tUbe 10.5 meters long. Thus, the

equation presented here is valid for tube of larger

diameter and greater length.

It has been shown (Reinemann et al., 1987) that the

slug flow model works well for describing airlift pump

40

performance when the riser tube diameter is less than 20

mm and a correction is made for the effects of surface

tension on the rise velocity of the Taylor bubble in

still liquid (Vts). Increased surface tension also acts

to suppress bubble breakup for tubes in this size range.

Small bubbles are less easily shed from the Taylor

bubbles and true slug flow results. When the tube

diameter is greater than 20 mm, small bubbles are more

easily shed from the tail of the Taylor bubbles and

bubbly-slug low results.

It has been speculated by Nicklin (1963), that for

long airlifts, the bubble and bubbly-slug flow patterns

are encountered only in the developing flow region.

Based on the results of this study and other vertical

air/water slug flow studies, it is doubtful that true

slug flow will ever exist in the airlift pump with the

exception of small diameter (D < 20 mm) riser tubes.

Previous studies of the airlift pump have concluded

that slug flow would develop regardless of the method of

introducing the gas. It has been demonstrated in this

study that a region of stable bubble flow exists for

airlift pumps. The first requirement for bubble flow to

exist is that the gas is introduced as bubbles much

smaller than the riser tube diameter. The second

requirement is that the gas concentration is below some

critical value. Above this critical gas concentration,

bubble collision and coalescence occurs. If the bubbles

41

increase in size to form Taylor bubbles the transition

from bubbly to bubbly slug flow occurs.

Flow pattern maps are commonly presented to predict

two phase flow behavior. In most two phase flow

applications the air and liquid flows can be

independently determined. In airlift pump operation,

however, for a given air flow volume and submergence

ratio, the liquid flow is fixed. The possible

operational regime of the airlift pump is, therefore, a

subset of the conventional flow pattern map (see figure

2 .7) •

The upper limit of the airlift operational regime

has been determined for a 38 mm diameter airlift with

100% submergence. No useful pumping work is being

performed in this configuration since there is no lift

provided. There are some practical applications,

however, such as the mixing or aeration of liquids.

The right hand boundary of the bubble flow regime

is defined by the critical gas concentration for the

transition form bubble to bUbbly-slug flow. When tap

water was used, a transitional gas concentration of 0.25

provided an accurate criteria for transition for bubble

to slug flow. When the waste water was used the

transitional gas concentration increased to 0.35.

Small amounts of surface contaminants have been

shown to affect the stability of bubble surfaces (Keitel

and Onken, 1982). The following effects indicative of

42

surface contamination were observed when the waste water

was used: 1) The stable bubble size in the flow was

slightly lower than that in tap water. 2) The bubbles

were more spherical in shape than those in tap water. 3)

The transitional gas concentration was higher than in

tap water. These surface effects could explain the

large variation of transitional gas concentrations found

in the literature.

Conclusion

Equations have been presented to describe operation

of the airlift pump in bubble and bubbly-slug flow. The

effect of small bubbles dispersed in the liquid slug on

the average gas velocity in bubbly-slug flow has been

accurately predicted. The stable operating regimes for

bubble and bubbly-slug flow have been identified. The

transition from bubble flow to bubbly-slug flow was

found to be sensitive to gas/liquid surface

contamination. Surface contamination did not however

significantly influence flow dynamics. Due to bubble

breakup and dispersal in the liquid slug, it is doubtful

that true slug flow will ever exist in the normal

operating regime of airlifts with the exception of

narrow diameter tubes (D < 20 mm, for air/water

systems) .

43

Table 2.1 Nomenclature and Definitions

D = pipe diameter [LJ

7f D2 2A = pipe cross sectional area = -4- [L J

X Point Quantity * <x> = average over tube cross section = l/A f X dA

Qg = vOlumetric flow rate of gas [L3/TJ ** Ql = vOlumetric flow rate of liquid [L3/TJ

~ = vOlumetric flux density of the mixture [L/TJ

<~> = average vOlumetric flux density of the mixture Ql + Qg

= [L/TJA

~ = point volumetric gas concentration [OJ

E = average gas concentration in riser tube [OJ

~s point gas concentration in liquid slug [OJ

ES = average gas concentration in liquid slug [OJ

Yg = point velocity of gas [L/TJ *** Qg

Vg = average gas velocity in riser tube =-- [L/TJA E

Ygj = velocity of gas relative to vOlumetric flux density [L/TJ

Ygs = velocity of gas bubbles in liquid slug [L/TJ

Yls = velocity of liquid in liquid slug [L/TJ

Vt = Velocity of Taylor bubble [L/TJ

Vts = Velocity of Taylor bubble in still fluid [L/TJ

z 0.35 (g D)1/2 for tube diameters greater than 20mm

Vbs = rise velocity of small gas bubbles in still fluid [L/TJ z 0.25 mls in air water systems considered

Re = Reynolds number = <J> D [OJ v

v = kinematic viscosity of liquid [L2/TJ

~ = submergence ratio [oJ

Zl lift height [LJ

Zs submergence [LJ

44

F = frictional head loss gradient expressed in meters of

fluid (pressure) per meter of pipe length [OJ

f = friction factor [OJ

g = acceleration due to gravity [L/T2 J

Cb, Cs , Cbs = distribution parameter for bubble flow,

slug flow, and bubbly-slug flow [OJ

* Underlined quantities depend on position.

** Dimensionless vOlumetric flow rates are obtained by

dividing flow rate by A (g 0)1/2 [L3/TJ, and are

primed (e.g. Ql', Qg').

*** Dimensionless velocities are obtained by dividing

velocity by (g 0)1/2 [L/TJ, and are primed (e.g. Vts').

Table 2.2 Summary of Airlift Equations

Bubble Vg , = 0.6 <~'> + Vbs' (1-1.4E)

Flow ex = (I-E) + (1+1. 8E) F

Bubbly- Vg , = 1.1 <~'> + 0.75 Vts

, Slug Flow ex = (l-E) (1+1. SF)

For all Flow patterns

Qg + Ql f <~,>2 0.316 <~'> = F = f =

A(g D)' 2 Re 0.25

(Ql + Qg)D Zs 0.25 m/s ,Re = ex = Vbs'= Vts = 0.35 A v Zl+Zs (g D)! (0 > 0.02 m)

Table 2.3 Water Quality Parameters

Tap Water Waste Water pH 7.8 6.9 BOD (mg/l) 0.0 60.0 Suspended Solids 0.0 97.5 conductivity (~mho/cm) 366.0 1300.0 chloride (mq/l) 13.8 41. 0

45

+-AIR INPUT

Figure 2.1. Typical Airlift Pump.

b e f i=I-U::=-------L.I::::I-i

d c h g OJ 01 OJ 01

BUBBLE FLOW BUBBLY SLUG FLOW SLUG FLOW

Figure 2.2. Flow Patterns.

----------

47

1.0

0.9 ~ lOCAl BUBBlE VElOCITY IN UQU/D SlU~

0.8

0.7

0.6 ------­ ~ It: LOCAl.. UQUID VELOCITY IN UQUID SlUG (\IIs')" ­w I­W ~ 0.5 4: It: 4: ll. 0.4

0.3

0.2 LOCAl.. GAS CONCENTRATION IN UQUID SLUG (Es)

" 0.1

0.0

0.0 0.2 0.4 0.6 0.8 1.0

RADIAL POSITON (r/R)

Figure 2.3. Local liquid slug gas void ratio, bubble

and liquid velocities in bUbbly-slug flow.

48

LIQUID PRESSURE TAPS

/"-. LIFT, ZL

RISER TUBE

AIR FLOW ORIFICE L LIQUID FLOW ORIFICE

MANOMETER

COMPRESSOR

Figure 2.4. Experimental Apparatus.

---------

49

--,-.0.7 - ­b

B BUBBLE FLOW, TAP WATER

Bb b /0.6 /

b BUBBLE FLOW, WASTE WATER b

0.5

..J

~ z fw 0.4

~ B Il Vg' = 1.1 <J') + 0.4 !b rP Bw Il. B X w 0.3

(JI

>

0.2

Vg' =0.62 <J') t 0.44 (1-1.4E)

0.1

0.0 +----.------r---...,-----r------r-----.----~

0.0 0.2 0.4 0.6

Vg' PREDICTED

Figure 2.5. Bubble Flow Data.

Experimental vs. predicted Average Gas Velocities.

50

1.1

1.0

0.9

0.8

0.7

lJl >

0.6

0.5

0.4

0.3

0.2

0.1

S = SLUG FLOW, TAP WATER

s = SLUG FLOW, WASTE WATER

so = SLUG FLOW, TAP WATER, NO DIFFUSER

BUBBLY SLUG UOOEL Vrf = 1.1 <JI) + 0.75 Vts'

S'

0.0

0.0 0.2 0.4

<J')

0.6 0.8

Figure 2.6. Bubbly-Slug Flow Data.

Average Gas Velocity vs. Average Mixture Velocity.

51

o ~

-' II.

o J o J III III W -' Z o III Z W Z Q

10.0~-----

38.1 mm ruBE

100'; SUBMERGENCE • • • • • • • • d . • • • I ••• .;. • • 00

~ '+++ 000000

.. +++ 0 .. + 0

+ 001.00 + 0 + 0

+ 0 E = 0.2 -.+ 0 ~ E = 0.3

oBUBBLE FLOW + 0 BUBBLY-SLUG FLOW

+ 0

o +

o

o 0.1 O+-.___.__---.----,-....--__.____.-..----~__r____,.___._____r_____.-.,...___.____.-._____r___1

0.01 0.10 1.00

DIMENSIONLESS GAS FLOW (Og')

Figure 2.7. 38 mm Diameter Tube Flow Pattern Map.

Chapter Three

Oxygen Transfer in Airlift Pumping

Abstract: The results of an experimental study of the

oxygen transfer properties of a 38 mm diameter airlift

pump are presented. The effects of varying initial

bubble size, flow rate, flow pattern, and water quality

on oxygen transfer are examined. A model to predict

oxygen transfer in airlift pumping is presented.

Introduction

The airlift pump has been of practical use as a

pumping device for many decades (see figure 3.1). It

has been reported that the famous Roman water

distribution system used airlifts 2000 years ago.

Airlifts also found application in removing water from

mines in the late 1800's. The first recorded pumping

studies were performed at that time and have continued

to the present.

Airlifts have become popular in the aquaculture

industry and in waste water treatment plants where large

volumes of water must be both circulated and aerated.

Several studies have been done on the aeration

properties of the airlift (Nagy, 1979; Zielinski et al.,

1978) but no theory or predictive equation for oxygen

transfer rates were given.

52

53

The flow dynamics of the airlift pump have been

described previously (Reinemann et al., 1987).

Complicating the prediction of the hydrodynamics of the

airlift is the fact that there are two flow patterns

possible in airlifts when tube diameters are greater

than about 20 mm. When the initial bubble size is much

smaller than the tUbe diameter and the gas void ratio is

low, the bubble flow pattern results. Small bubbles are

distributed over the pipe cross section. Bubbles remain

close to their initial size, and there is little

interaction between bubbles (see figure 3.2). If the

gas void fraction is above some critical value,

coalescence occurs and bubble size increases.

Bubbles with average diameter greater than about

0.7 times the riser tUbe diameter are referred to as gas

slugs or Taylor bubbles. The presence of Taylor bubbles

only is referred to as the slug flow regime. The slug

flow regime has been found to occur only when the riser

tube diameter is below 20 mm (Reinemann et al., 1987).

In the bubbly-slug flow regime, small bubbles are

found suspended in the liquid slug between the Taylor

bubbles. The presence of these bubbles is due to the

region of extreme turbulence encountered at the tail of

the Taylor bubble. Small bubbles are broken off of the

tail of the Taylor bubble and dispersed in the liquid

slug. Experiments were performed in both the bubble

flow and bubbly slug flow regimes to determine if any

54

significant difference existed between the gas transfer

properties in these two regimes.

The equation most commonly used to describe gas

transfer in gas/liquid dispersions is (Barnhart, 1969:

Clark, 1985: Colt and Tchobanoglous, 1981: Nagel et al.,

1977) :

e - KIa t [3.1]

where

Ci = initial gas concentration (mg/l)

Cs saturation gas concentration (mg/l)

C = gas concentration at time t (mg/l)

KIa = gas transfer coefficient (s-l)

t = gas/liquid contact time (s)

The gas transfer coefficient is a function of the rate

of molecular diffusion of gas in the liquid, the

gas/liquid surface area per liquid volume and the degree

of turbulence in the flow. The gas transfer coefficient

is commonly given as a function of the pipe Reynolds

number for two phase flow (Clark, 1985; Kubota et al.,

1978: Lin et al., 1976: Nagel et al., 1977: Shilimkan

and stepanek, 1977). The Reynolds number is a measure

of the degree of turbulence encountered in the flow.

The degree of turbulence influences both the gas/liquid

surface area and the transfer rate across the surface.

55

The Reynolds number for gaG/liquid pipo flow iG

calculated as (Govier and Aziz, 1972):

Re=~ [ 3 • 2 ] v

where

Vm = average velocity of two phase mixture

given by:

[ 3 • 3 ]

Ql = vOlumetric liquid flow rate (m3/s)

Qg vOlumetric gas flow rate (m3/s)

A = pipe cross sectional area (m2 )

D = pipe diameter (m)

v = kinematic viscosity of liquid (m2/s)

studies have been done on mass transfer in two

phase gas/liquid pipe flow (Clark, 1985; Kubota et al.,

1978; Shilimkan and stepanek, 1977;). These studies do

not directly apply to airlift operation, since the flow

speeds encountered in airlifts (Vm < 1 m/s) , are much

lower than those encountered in these studies (1 to 10

m/s). It has been shown that the results of high speed

two phase flow studies cannot always be extrapolated

into the slower flow regimes of the airlift (Reinemann

et al., 1987). The hydrodynamics and gas/liquid surface

area of two phase flow can change considerably as flow

speed increases. Both of these factors influence the

gas transfer coefficient. The objective of this study

is to determine oxygen transfer coefficients for the

56

flow patterns and flow velocities encountered in airlift

pump operation.

~xperlmental ~rocedure

The test apparatus consisted of a circular loop of

38 mm glass tUbing with a liquid reservoir (See figure

3.3). The total volume of the system was approximately

34 liters. The circuit was closed to the atmosphere

except for a gas outlet at the top of the riser tube.

Air was injected at the base of the riser tube. An

aquarium air stone was used to produce bubbles from 1 to

3 mm in diameter. This is the bubble size associated

with commercial fine bubble aerators. The diffuser

produced the bubble flow pattern at low gas flow rates,

and bubbly-slug flow at higher gas flow rates. A single

6 mm glass tUbe was used to produce large gas bubbles

and the resulting bubbly-slug flow pattern at all gas

flow rates. Air and water flows were determined by

means of pressure drop measurements across calibrated

sharp edged orifices.

Tap water and waste water from an intensive water

reuse aquaCUlture system were the liquids used to

examine the effects of contaminants in the liquid.

Water quality parameters were determined by standard

analytic techniques at the Cornell University Agronomy

lab (see table 3.1).

Sodium sulfite with a cobalt catalyst was used to

remove all oxygen from the tap water. The aquaculture

57

waste water was allowed to stand for 5 day9 in a clo9gd

container to allow the resident BOD to reduce the

dissolved oxygen level. The pump/reservoir system was

filled with a measured quantity of deoxygenated water

and the pump was started. Dissolved oxygen measurements

were taken with a YSI meter at 30 second intervals

beginning from the initiation of a run until saturation

was reached. Liquid temperature was also recorded for

each run. A typical raw data set is shown in figure

3.4.

The system gas transfer coefficient was determined

by regressing the left hand side of [3.4] vs. time:

In l~: = ~j = - KIa' t [3.4]

where

Kla'= system oxygen transfer coefficient (s-l)

The system gas transfer coefficient relates to the

gas/liquid contact area per total system liquid volume.

The oxygen transfer coefficient for the airlift pump

alone relates to the gas/liquid contact area per liquid

volume in the airlift pump riser tube. The airlift pump

gas transfer coefficient is therefore determined as

follows:

KIa = KIa' ~ [3.5]Vr

58

where

Vs total liquid volume in the system

Vr = liquid volume in the riser tube

(see figure 3.5)

correction was also made to adjust all Kla values

to standard temperature conditions (20°C). The

temperature correction used was (Barnhart, 1969);

Kla (T) Kla(20)

= (j (T-20) [3.6]

where

T = temperature (OC)

Kla(T) = gas transfer coefficient at temperature T

Kla(20) = gas transfer coefficient at standard temperature (20°C)

(j = 1. 02

Results and Discussion

Regression analysis of the gas transfer coefficient

as a function of the pipe Reynolds number showed no

significant difference between water type, flow pattern

or diffuser type (See figure 3.6). The data was

therefore pooled. Regression of the entire data set

yielded the following empirical correlation for the

oxygen transfer coefficient (R2 = 0.92, std. err. of

estimate = 0.009) :

Kla(20) = 6.5xIO- 6 Re [3.7]

The gas/liquid surface area in the bubbly-slug

regime is less than that in the bubble flow regime.

59

This tends to reduce the gas transfer rate. The

increased turbulence at the gas slug tail, however,

tends to increase gas transfer. These two effects thus

compensate for one another and the gas transfer rate

remains unchanged for the two flow regimes.

While the degree of surface active substances

present in the waste water produced an observable change

in the flow characteristics, the gas transfer rate was

not significantly reduced. contaminants on the

gas/liquid surface tend to reduce the gas transfer rate.

The bubble size is slightly reduced, however, increasing

the gas/liquid surface area. These two effects combined

to yield no significant change in the gas transfer rate

between waste water and tap water.

When tap water was used, the bubbles in the

developed region were elliptical with mean diameter of

2-3 mm. Coalescence was observed for gas void fraction

above about 0.25. When the waste water was used the

bubbles were spherical with average diameter of 1-2 mm.

Coalescence was not observed for waste water until the

gas void fraction was over 0.35. This effect has been

documented previously and is caused by surface active

agents attaching to and stabilizing the bubble surface

(Keitel and Onken, 1982).

The oxygen transfer for the airlift pump operating

in bubble and bubbly-slug flow can be described by

[3.1J, [3.6J and [3.7J. The gas/liquid contact time can

60

be expressed as the average liquid velocity divided by

the length of the riser tube:

[ 3 • 8 J

where

Z = length of riser tube (m)

VI = average liquid velocity (m/s) given by:

VI = QI I A (I-E) [3.9J

QI = liquid vOlumetric flow rate (m 3/s)

A = pump cross sectional area (m2 )

E = gas vOlumetric void ratio

Conclusion

Oxygen transfer was not significantly affected by

flow pattern, initial bubble size, or the wastes present

in the water studied. Wastes in the water did however

influence the transition from bubble to slug flow. An

empirical correlation is presented relating the oxygen

transfer coefficient (Kla(20)), to the two phase pipe

Reynolds number. It is not advantageous to use a small

pore gas diffuser to increase gas transfer in airlift

pumps. Reducing the orifice size increases the pressure

drop across the diffuser and the gas transfer rate will

not increase.

61

Table 3.1� Water Quality Parameters�

Tap Water Waste Water

pH 7.8 6.9

BOD (mgjl) 0.0 60.0

Suspended Solids(mgjl) 0.0 97.5

conductivity(~mhojcm) 366.0 1300.0

chloride (mgjl) 13.8 41.0

62

.-AIR INPUT�

Figure 3.1. Typical Airlift Pump.

BUBBLE FLOW BUBBLY SLUG FLOW SLUG FLOW

Figure 3.2. Flm'l Patterns in Airl ift Pump Operat.i 0:1.

[jJ:C"''' LIQUID PRESSURE TAPS

Iii LIFT, ZL /'"

RISER~lrJ

>

TUBE ::1

AIR FLOW ORIFICE L LIQUID FLOW ORIFICE

MANOMETER ----I~

COMPRESSOR

Figure 3.3. Experimental Apparatus.

65

0.11

0.10 s

0.09 s

0.08 s

"r-I u W III w

0.07

0.06

b

s

s

'"' 0 l'l v 0 y

0.05

0.04 b = BUBBLE FLOW, WASTE WATER

0.003 b B = BUBBLE FLOW, TAP WATER

0.02 s = BUBBLY-SLUG FLOW. WASTE WATE

0.01 S = BUBBLY-SLUG FLOW. TAP WATER

0.00

0 2 4 6 8 (Thousands)

REYNOLDS NUMBER

10 12 14

Figure 3.4. Oxygen Transfer Coefficient vs.

Reynolds Number

Chapter Four�

Energy and Cost Analysis of Salmonid Production�

In Water Reuse Systems�

Abstract: The results of an energy and cost analysis of

salmonid production in water reuse systems are

presented. Various options to increase system

efficiency, including the use of airlifts for pumping

and aeration, are considered. The energy inputs for

salmonid production in water reuse systems are compared

with land based animal protein production, other forms

of aquaculture and traditional fishing.

Introduction

The culture of fish for human consumption has a

long history. Records of fish husbandry have been found

in Egypt and China dating back several thousand years

(Brown, 1983). Despite this long history, aquaculture

is still in its infancy as a commercial enterprise.

Although only twelve percent of all edible fish and

shell fish produced in the United States in 1984 was

produced by aquaculture, this represents about twice

that produced in 1980, and the potential for further

growth is great (Greer, 1987).

Aquaculture systems can be more efficient than many

land based animal protein production systems (Pimentel

and Hall, 1984). The primary reason for this is that

66

67

fish have much better feed conversion ratios than land

animals (Large, 1976). When aquaculture systems are

combined with other plant and/or animal production

systems, energy inputs may even approach those for wheat

or rice (Pimentel, 1980).

Interest in aquaculture has been spurred by

increased demand and a static or decreasing supply of

seafood products worldwide. Most studies indicate that

traditional ocean and fresh water fisheries are

approaching their maximum sustainable yield (ADNYS,

1984; Brown, 1983; Greer, 1987; Wheaton, 1977). If

trends in over-fishing and pollution continue, reduced

fishery harvests from natural sources are expected.

In the US, the harvest of fish from natural sources

does not meet demand. In 1986 the US imported 64% of

the seafood consumed by humans (see table 4.1). The

demand for fish products is increasing faster than the

growth of population. Consumption of fish increased 24%

in the past ten years (see table 4.2). Population

growth accounted for only 1/2 of this increase. The

remainder is due to changing diets and a demographic

shift to those who prefer fish (ADNYS, 1984). The shift

in US dietary trends is illustrated in table 4.2.

Increases in the consumption of beef and eggs were far

below those for poultry and fish. Because of the

situation regarding the supply and demand of fish

68

products there is great potential for the development of

aquaculture particularly in the us.

One area of considerable interest in the us is the

production of fresh fish for local markets. Salmonids

(trout and salmon) have been identified as high

potential aquaculture products (ADNYS, 1984). There is

high consumer acceptance of these species, especially in

the northern us where environmental conditions are also

most favorable for their production. Another attractive

feature of salmonid production is that there is more

information available than for many other species as

they have been cultured successfully in the U.S. for

over 100 years (Brown, 1983).

Currently the major commercial producers of fresh

water salmonids are located where abundant natural

ground water supplies are available. The major

concentration is found in Idaho along the Snake River.

Here, ground water flows by gravity out of a porous

aquifer which acts to keep water temperature and flow

rates relatively uniform throughout the year. These

outdoor culture systems provide no treatment of the

culture water and rely upon the constant input of fresh

water to maintain water quality in the culture

environment. This is referred to as a flow through

system. This type of production is obviously limited to

areas with rather unique natural assets.

69

Because of the limited supply of surface water, the

expense of pumping ground water, and environmental

regulations concerning the quality of waste water

discharge, it is likely that the future growth of

aquacultural production will require the use of systems

which provide treatment of culture water. Unlike

systems which require large quantities of surface or

ground water, systems which treat and reuse culture

water can be located near population centers. Water

conservation is doubly important near cities which tend

to be highly competitive with agricultural operations

for available water.

Location near population centers reduces the time

between processing and consumption as well as minimizing

refrigerated transportation distances. This is

especially important for fish and shell fish which are

highly perishable and quite sensitive to handling and

storage.

Another advantage of closed system aquaculture is

that it allows greater control over the culture

environment. Water temperature can be economically

regulated to improve feed conversion efficiency. The

risks of introducing disease and pollutants from the

environment are also reduced when the water and feed

entering the culture system can be monitored.

The advantages of water reuse systems must be

balanced against their higher costs. The more intensive

70

the culture technology the higher the start-up and

operating costs and energy requirements. The monitoring

and control of the culture environment requires a high

degree of management time and skill. While the

introduction of pathogens from the environment is

reduced, mortality can result from sUb-optimal water

quality conditions such as the accumulation of waste

metabolites or suspended particulate matter, low

dissolved oxygen concentration or excessive temperature.

The importance of disease free stock is also greater

since stress induced by sub-optimal culture conditions

increases the danger of mortality due to these diseases.

Because of the risk of disease a large number of smaller

systems rather than single larger units may be desirable

(Muir, 1981; Timmons et al., 1987). This further

increases management requirements and reduces economies

of scale.

Aquaculture is still in its infancy as a commercial

enterprise. The base of information on appropriate

technologies, economic potentials, markets, and a host

of other considerations is minimal. Most experiments in

Europe with intensive aquaculture systems were economic

failures except where local conditions were favorable.

A major factor attributed to these failures is that

management problems were underestimated (Rosenthal,

1981) .

71

Intensive aquaculture practiced in reuse systems

has great potential for the production of fresh fish.

with the present state of aquacu1tura1 science, however,

the risks are high.

Salmonid Production in Closed Systems

An energy and production cost analysis has been

performed on a water reuse salmonid production system

based on recent experience gained at the Cornell

University aquaculture facility (Timmons et al., 1987).

A full scale water reuse system was used to grow brook

trout and Atlantic salmon to table size for a market

study conducted in cooperation with a local super

market.

The Cornell facility was designed to support a

maximum of 2 tonnes, live weight, of fish. The system

employs two 6.4 meter (21 foot) diameter rearing tanks

of 0.76 meter (2.5 foot) water depth, a 5.7 cubic meter

(200 cubic foot) gravel trickling filter for

nitrification, a 6.4 meter (21 foot) diameter settling

tank for the removal of suspended solids, and a number

of smaller units used to evaluate the performance of

various other water treatment methods. Flow rates and

filter sizing were determined from the design

information given by Speece (1973). The water

replacement rate was about 1% per day or about 570

liters per day. Water losses were those due to

evaporation and losses during solids removal.

72

The rearing and treatment tanks are above ground

with plywood walls and vinyl liners. The tank cost is

about 15.7 $/m3 (0.06 $/gal) which is considerably lower

than fiberglass tanks commonly used in fish culture.

The entire system is housed in a temperature controlled

dairy barn. The solids associated with this waste water

did not settle well due to their small size and

buoyancy. Active forms of particulate removal were

considered necessary (Timmons et al., 1987).

The system used a 1.1 kW (1.5 hp) submersible

centrifugal pump for water circulation with an identical

pump as backup in the event of pump failure or water

flow surges. The water circulation rate was

approximately 15 liters per second (250 gallons per

minute). A 0.75 kW (1 hp) blower was also used to

supply diffused aeration. The total steady power

consumption of the system for pumping and aeration was

therefore 1.85 kw (2.5 hp).

The system was stocked with 8 cm (3 inch)

fingerlings. A production cycle of 12 months produced

fish of 0.3 m (12 inch) which weighed 0.25 kg (1/2

pound) .

Because of the risks associated with the rapid

spread of fish diseases in intensive reuse systems,

particularly when the fish are subjected to

physiological stress, it is desirable to isolate the

culture system as much as possible from the environment

73

and from other fishes. Timmons et al., (1987) therefore

recommended that a production facility be divided into a

number of separate modules. Each module would consist

of a water treatment system serving one or more rearing

tanks. The size of a module would depend upon the

marketing volume and schedule along with the operational

skill of the plant manager.

The present management strategy at the Cornell

facility is that of a batch operation. A group of

fingerlings is introduced to the system and raised to

market size without the introduction of other fish into

the system. The system must be designed to handle the

final harvest weight of fish (carrying capacity) while

the total weight of fish in the system increases from

that of the fingerlings to the harvest size over the

growth cycle of the fish. The system therefore runs

under capacity for most of the growth cycle. In this

mode of operation the production to capacity ratio is

one (P:C = 1) • The production of the system is equal to

its carrying capacity over one growth cycle.

As management skill and system reliability increase

and the associated threat of disease is thereby reduced,

the system can be operated in such a manner so as to

keep the total weight of fish in the system at anyone

time more or less constant. Groups of fishes of varying

age and size would be served by one treatment system.

This necessitates the introduction of new fish into the

74

system during a production cycle, increasing the chance

of introducing disease. The production to capacity

ratio can thereby be increased making more efficient use

of the system but at higher risk.

The operating costs and energy inputs of a modular

aquaculture facility such as the Cornell system are

presented in table 4.3. More detailed information

regarding the determination of costs and energy

requirements are presented in Appendix A.

The costs and energy inputs have been calculated on the

basis of one metric tonne of system carrying capacity.

This was suggested as a reasonable size for one

production module by Timmons et al. (1987).

The heating and cooling requirements were

determined from a thermal model of an aquaculture

facility using Ithaca weather data. The enclosing

structure and tanks were assumed to be moderately

insulated (R = 1.9 w/m2/oc = 10 BTU/hr/ft2/oF). The

heating and cooling requirements represent the energy

required to maintain the culture water at 15°C. Details

of the thermal model are given in appendix B.

The system cost estimates presented here compare

favorably with those of a similar water reuse salmonid

production system (Meade, 1976; MacDonald et al., 1975)

when adjusted for inflation. The state of the art in

aquacultural science is in a period of rapid

development; it is therefore likely that costs of

75

production may vary considerably in the near future from

those quoted here. The costs of starting a system may

also depend highly on available assets (building,

natural water supply, existing tanks), size and

sophistication of the system. A sensitivity analysis

has therefore been conducted to determine the effects of

changing system costs and the production to capacity

ratio. The initial production to capacity ratio of an

aquaculture operation might be expected to be near one.

As management experience is increased and systems are

perfected the production to capacity ratio could be

increased to two or more. The results of the

sensitivity analysis are presented in table 4.4.

The base system (A) uses the production figures

estimated from the Cornell facility (Timmons et al.,

1987) and assumes a production to capacity ratio of one

(P:C=l). The effects of increasing the production to

capacity ratio using the base system are illustrated by

comparing systems A and B. Increasing the production to

capacity ratio from one to two decreases fish cost by 38

percent and energy inputs by 42 percent. Increasing the

production to capacity ratio had the largest effect on

fish price and energy inputs of all the variables

examined. This illustrates the importance of skilled

management in the economical and efficient operation of

aquacultural systems.

76

A high cost system (C) has been simulated by

doubling the capital cost of the culture system and

building. This might simulate a highly automated system

or one subject to severe site restrictions. A low cost

system (D) has been simulated by halving the capital

costs of a system. This might represent a situation in

which a building or other natural assets were already

available or new advances in system design were realized

to further reduce system costs.

The effect of varying the initial cost of the

system is illustrated by comparing systems B, C, and D.

Doubling and halving the base system cost estimates

results in a 20 percent change in production cost. The

energy cost of system construction represents from 1 to

4 percent of the energy budget. Changes in the initial

system energy inputs therefore have little effect on the

energy cost of fish production.

A vegetable feed system (E) has been considered to

examine the effects of substituting vegetable based feed

for the fish meal base currently used in the food of

fish such as salmonids and catfish. Ironically the

major constituent in commercial fish ration is fish meal

obtained from ocean harvests. Given this situation, the

aquaculture facility is a means of converting one type

of fish into another, presumably with higher market

value. Furthermore, the increased production of fish in

intensive aquaculture facilities would be dependant upon

77

increased ocean harvest. This has already been stated

as doubtful. In a system which uses fish meal based

feed this also represents the second largest energy

input. Substituting vegetable based feeds thus also

offers the advantage of reducing salmonid production

energy inputs. Fish cost is reduced 7% and energy

inputs are reduced 11 %.

Studies have been conducted in which soybean meal

has been substituted for all or part of the fish meal in

salmonid feed. Soybean meal contains about the proper

amino acid balance required by salmonids. These feeding

trials have been successful, although feed conversions

are not quite as high as with fish meal based feeds

(Hughes et al., 1983; Pitcher, 1977). A feed conversion

ratio of 1.6:1 was assumed for vegetable based feed.

This is about 20 percent higher than the 1.4:1 feed

conversion ratio attained in the Cornell facility with

fish meal based feed (Timmons et al., 1987). The effect

of sUbstituting vegetable based feed is illustrated by

comparing systems Band E. Fish cost is reduced 7% and

energy inputs are reduced 11 %.

Direct energy used for pumping, aeration, heating

and cooling represent 25% of the economic cost and 70%

of the energy cost of salmonid production in the base

system. This is typical of current intensive

aquacultural production facilities (Muir, 1981).

Reducing these direct energy inputs could therefore make

78

a significant impact on the efficiency of intensive

water reuse aquaculture systems.

The air-lift pump has qualities which make it

uniquely suited to aquacultural applications. The air­

lift has no moving parts in contact with the pumped

fluid. This allows it to handle abrasive or sediment

laden water, often encountered in aquaculture systems.

The air-lift combines both pumping and aeration

functions into one device. Based on research conducted

in conjunction with the Cornell aquaculture program, an

airlift pump has been designed to replace the

conventional centrifugal pump and diffused aeration

systems, thereby reducing capital costs and operating

energy requirements (Reinemann and Timmons, 1987).

Increased building and tank insulation would help

to reduce both heating and cooling loads. The large

thermal mass of the culture water will even out diurnal

fluctuations in system temperature. Several workers

have investigated the use of culture water as a thermal

storage system to regulate night time temperatures in

green houses (Zweig et al., 1981; VanToever and Mackay,

1981). Systems making use of solar gain could

considerably reduce winter heating requirements.

In-ground tanks have been shown to be successful in

maintaining system temperatures at tolerable levels in

summer months (Parker, 1981). This study was conducted

in Alabama where summer temperatures are much higher

79

than those found in the northern us. water temperature

0on 25 June was 21 C. While this temperature is

excessive for the culture of most trout species,

Atlantic salmon can tolerate this water temperature.

The use of in-ground tanks in northern climates could

considerably reduce or eliminate cooling requirements.

Disadvantages might include the necessity of stronger

tank walls and reduced access to fish.

Most studies of the effects of temperature on fish

have dealt with lethal effects. Much work needs to be

done to determine the effects of culture temperature on

fish growth rate, disease resistance, feed conversion,

and other parameters significant to the efficient

production of fish.

A conservation system (F) has been simulated which

would reflect the use of thermal conservation

strategies, vegetable based feed, and airlift pumps. The

feed conversion ratio is assumed to be 1.8:1 since some

temperature fluctuation is likely to be encountered.

Heating and cooling requirements are assumed to be zero.

The capital savings resulting from the use of an airlift

pump are assumed to be used in the thermal conservation

system. The effects of employing conservation

strategies is illustrated by comparing systems Band F.

Production costs are reduced 25% and energy inputs are

reduced 65%.

80

Cost estimates have also been made for a flow

through system (G). In such a system the pumping and

aeration and heating and cooling costs are zero since

the culture water is assumed to be gravity fed and the

water temperature unregulated. Capital costs are

assumed to be those of a low cost system since a

building and pumping and aeration equipment are not

required. The feed conversion ratio has been assumed to

be 1.8:1. This is a typical feed conversion ratio for

flow through and pond systems (Pitcher, 1977).

Improved feed conversion in reuse systems results

primarily from the possibility of increased management

of the system and the maintenance of a uniform water

temperature near its optimal value. In reuse systems,

fish are not sUbjected to high temperatures which cause

physiological stress or low temperatures which reduce

metabolism and growth rates. In water reuse systems,

control of water temperatures are economically possible.

Temperature regulation would be prohibitively expensive

in flow through systems which use large quantities of

surface or ground water. Flow through systems therefore

generally operate at the temperature of their water

source. Surface water temperatures vary considerably

through the year. Ground water temperatures tend to be

more stable, but additional energy is required for

pumping or sites are limited to those places in which

ground water is supplied by springs as is the case in

81

the Idaho salmonid producing region. The cost of

transporting the fish to market represent an average

delivery distance of 1600 km (1000 miles) since the

facility is subject to site limitations.

The cost and energy inputs for carefully designed

and properly managed salmonid production in water reuse

systems can be competitive with flow through systems.

Poorly designed and operated systems are not competitive

however.

Comparison with other Forms of Protein Production

The energy inputs for various protein production

systems including land based US agriculture, various

aquaculture systems and traditional fishing are listed

in table 4.5. The estimates made in this study of

energy inputs for salmonid production are comparable to

those made by other investigators. The energy inputs

for salmonid production in well designed and operated

water reuse systems can approach those for US beef and

pork production, flow through systems and deep sea

fishing (see table 4.5).

Aquaculture can exhibit widely disparate levels of

energy inputs. Some species of fish, such as carp and

milkfish, are able to survive in very turbid water with

low levels of dissolved oxygen and high temperatures.

These fish feed on aquatic plants, algae, or plankton.

Systems have been developed using these species to make

efficient use of waste nutrients from other animal or

82

fish production systems. These integrated systems have

very low energy demands and make very efficient use of

resources.

The incorporation of salmonids into other

production systems is more limited because of their high

water quality requirements. Some work has been done

with combined salmonid/hydroponic systems (Naegel, 1977;

VanToever and Mackay, 1981; Zweig et al., 1981). In

these studies, plants have been used successfully to

remove waste metabolites from culture water. Systems

such as these could further reduce intensive salmonid

production costs and energy inputs but would require

increased management skills.

The protein production efficiency per unit land

area is presented in Table 4.6 for various US land based

protein production systems as well as for salmonid

production using vegetable based feeds. Because of the

superior feed conversion efficiency of salmonids

compared to land animals, it is possible to produce more

protein per unit land area than with poultry, beef,

pork, milk, or egg production.

Fish breeding offers opportunities for increasing

the efficiency of aquacultural production.

Domestication of aquatic species has hardly begun. As

more species are sUbjected to selective breeding with

improvements in growth characteristics, aquaculture

83

would be expected to compete even more advantageously

with other protein production systems.

Conclusion

Based on current trends in fish and seafood

consumption, availability and prices, it seems likely

that the use of intensive water reuse aquaculture

systems will increase. since the harvest of fish from

natural sources appears to be near its sustainable

limit, it is also likely that the future price and

availability of fish and seafood will depend upon the

degree of advancement in the science of aquaculture.

Table 4.6� Protein Production and Land Area�

Total yield Protein yield kg/ha kg/ha

Salmonids 1860 267 Broilers 2000 186 Pork 490 35 Beef 60 6 Dairy 3270 114 Eggs 910 104 Soybeans 2600 885 Corn 7000 630 Alfalfa 11800 1840

84

Table 4.1� US Fishery Products Supply�

(Million Metric Tonnes)� From (USDA, 198~

1975 1985 change

Domestic catch 2.44 2.83 +16% human consumption 1. 26 1. 49 +34% feed & industrial 1.19 1. 34 +23%

Imports 2.81 3.99 +42% human consumption 2.10 2.70 +29% feed & industrial 0.71 1. 29 +82%

Total human cons. 3.36 4.19 +24% Total feed & industr. 1. 90 2.78 +46%

Table 4.2� Consumption of Selected Protein Products in the US�

(Million Metric Tonnes)� from (USDA 1986)�

1975 1985 change

Fish 3.4 4.2 +24% Beef 10.9 10.8 -1% Pork 5.2 6.7 +29% Broilers 5.0 8.5 +70% Milk 52.2 64.9 +24% Eggs 29.3 31.0 +6%

85

Table 4.3� Energy and Cost Analysis�

(See Appendix A for Details)�

Cost ($) Ener~y Inputs (10 kcal)

Initial Cost (per tonne of carrying capacity)� Building 8800 33.8� Culture System 4190 14.8� Land & site prep 50 0.5� Well and Pump 160 0.1� Heating & Cooling eq. 3000 0.4� Misc. Equipment 500 1.0�

Total Initial 16700 50.6

Fixed operating cost (per tonne of carrying capacity) Pumping & Aeration 1990 57.1 Heating & Cooling 220 6.3 Labor 1250 ---­Capital, Int. , Ins, . 2500 2.5

Maint. , Taxes

Variable operating cost (per tonne of )roduction) Feed 770 8.6 Stock 1100 5.2 Transportation 20 0.1

Total operating cost,P:C=l 7850 79.8

86

Table 4.4 sensitivity Analysis

Cost Energy $jkg ($jlb. ) 10 3 kcaljkg whole fish whole fish

A. Base System (P:C=l) 7.85 (3.57) 79.8 B. Base System (P:C=2) 4.89 (2.22) 46.9 C. High Cost (P:C=2) 6.14 (2.79) 48.1 D. Low Cost (P:C=2) 4.27 (1. 94) 46.2 E. Veg. Feed (P:C=2) 4.53 (2.06) 41.1 F. Conservation (P:C=2) 3.77 (1. 69) 17.6 G. Flow Through (P:C=2) 3.74 (1.70) 19.1

Operating cost $jkg ($jlb)

A B C D E F G

Feed 0.77 0.77 0.77 0.77 0.41 0.41 0.99 (0.35) (0.35) (0.35) (0.35) (0.19) (0.19) (0.45)

P&A 1. 99 1. 00 1. 00 1. 00 1. 00 0.29 0.00 (0.90) (0.45) (0.45) (0.45) (0.45) (0.13) (0.00)

H&C 0.22 0.11 0.11 0.11 0.11 0.00 0.00 (0.10) (0.05) (0.05) (0.05) (0.05) (0.00) (0.00)

Stock 1.10 1.10 1.10 1.10 1.10 1.10 1.10 (0.50) (0.50) (0.50) (0.50) (0.50) (0.50) (0.50)

Labor 1. 25 0.62 0.62 0.62 0.62 0.62 0.62 (0.57) (0.28) (0.28) (0.28) (0.28) (0.28) (0.28)

Trans. 0.02 0.04 0.04 0.04 0.04 0.04 0.40 (0.01) (0.02) (0.02) (0.02) (0.02) (0.02) (0. 18)

Capital 2.50 1. 25 2.50 0.63 1. 25 1. 25 0.63 (1. 14) (0.57) (1. 14) (0.29) (0.57) (0.57) (0.29)

Total 7.85 4.89 6.14 4.27 4.53 3.71 3.74 (3.57) (2.22) (2.79) (1. 94) (2.06) (1. 69) ·(1.70)

Energy Input (10 3 kcaljkg whole fish)

A B C D E F G

Feed 8.6 8.6 8.6 8.6 2.8 2.8 12.3 P & A 57.1 28.6 28.6 28.6 28.6 8.2 0.0 H & C 6.3 3.1 3.1 3.1 3.1 0.0 0.0 Stock 5.2 5.2 5.2 5.2 5.2 5.2 5.2 Trans. 0.1 0.1 0.1 0.1 0.1 0.1 1.0 Equip. 2.5 1.3 2.5 0.6 1.3 1.3 0.6

Total 79.8 46.9 48.1 46.2 41.1 17.6 19.1

P&A = pumping and aeration, H&C = heating and cooling,

87

Table 4.5� Energy Inputs for Various Protein Production Systems�

101 kcal I I kg protein

!Isalmonids in water reuse systems (1) 122-554

US Agriculture Pork (2) Beef (2) Eggs (2) Milk (2) Broilers (2)

171 100

71 47 39

Aquaculture Salmonids

Salmonids, flow through(l) Trout-Britain, ponds(3) Trout, raceway(4) Trout, automated hatchery(4)

Other Species Catfish US (2) Catfish-Thailand, ponds(3)

USA, ponds(3) Catfish, ponds(4) Milkfish-Taiwan, ponds(3)

Philippines, pens(3) Tilapia-Africa, ponds(3)

Thailand, ponds(3) Carp-Philippines, ponds(3)

Germany, ponds(3)

133 93

118 512

137 125 213 450

12 2 1

38 4

60

Fishing Herring-Inshore sea fishing(3) Cod-Deep sea fishing(3) Flounder, fishing(4)

6 74 94

(1) (2) (3) (4)

estimates from this study Pimentel and Hall, 1984 Edwardson, 1976 Rawitscher and Mayer, 1979

Appendix A� Energy and Cost Analysis Details�

tonne cc = metric tonne of carrying capacity tonne p = metric tonne of production

Capital costs Culture System:

Tanks: Combined rearing and treatment volume 73 m3jtonne� cc, 1.6:1 rearing to treatment volume, $20jm3 for wooden� tank (1)� 1.54XIO b kcaljm2 residential building, (2)� wooden tanks assumed to have similar energy cost as� residential building since construction materials and� techniques are similar� $1460jtonne cc 7.0xl0 6 kcal jtonne cc�

Pumps: 1 kw submersible pump per 2 tonne module,� $1200jpump, 1.3 service factor, 50 kg pump weight (1)� 20000 kcaljkg for electrical equipment (2~

$600jtonne cc 0.5xI0 kcaljtonne cc�

Blower: 0.75 kw (1 hp) blower for aeration system per� two tonne cc, $500jblower, 1.3 service factor, 20 kg� blower weight, (1)� 20000 kcaljkg for electrical equipment (2)� $250jtonne cc 0.2xl0 6 kcaljtonne cc�

Plumbing: 80 m of 150 mm PVC pipe and fittings per 2� tonne module (1)� 150 mm PVC pipe 96 kgjm, 28700 kcaljkg (2)� $500jtonne cc 5.7xl0 6 kcaljtonne cc�

Filter media: gravel, 3 m3jtonne cc, (1)� $40jm3 , $120jtonne cc�

Aeration system: 40 m of 59 mm PVC air lines and� fittings per 2 tonne module (1)� 59 mm PVC, 1 kgjm, 28700 kcaljkg (2)� $260jtonne cc 0.86Xl0 6 kcaljtonne cc�

Solids removal system: estimated� $lOOOjtonne cc 0.5XI0 6 kcaljtonne cc�

Total Culture System:� $4190jtonne cc 14.76xl0 6 kcal jtonne cc�

88

89

Appendix A Continued

Building:� Insulated pole building with insulated perimeter, 80� m2jtonne cc, building cost $110jm2 (1)� 422,000 kcal/m2 for farm buildings (2)� $8800jtonne cc 33.8xl0 6 kcaljtonne cc�

Land:� Near population center with site preparation, 0.01 Ha� per tonne cc, $4900jha (estimated)� $50/tonne cc 0.5 xl0 6 kcaljtonne cc�

Well and Pump:� $2500 for installation to serve 16 modules (estimated)� $160/tonne cc 0.03xl0 6 kcal/tonne cc�

Heating and Cooling System:� Heat pump, design heating load 14 kwjtonne cc, design� cooling load 3.9 kW/tonne cc, 20 kg system weight (1)� 20000 kcaljkg for electrical equipment (2)� $3000/tonne cc 0.4xl0 6 kcalj tonne cc�

Total Capital Cost $16700jtonne cc�

Total Initial Energy Input 49.49xl0 6 kcal/tonne cc�

Operating costs Feed: feed cost $550 per tonne ($0.25 per pound), 1.4 feed conversion (1) Energy input 6140 kcaljkg (3) $770/tonne p 8.60xl0 6 kcaljtonne p

Fingerling Stock: $0.25 per 8 cm fingerling to produce one 0.25 kg fish,xIO% mortality (1) energy input estimated from production of adult fish $1100jtonne p 5.2xl0 6 kcaljtonne p

Pumping and Aeration: 1.75 kw mechanical power, 1.3 service factor, 8760� hours, $O.10jkWhr (1)� 2863 kcaljkWhr for production of electrical energy (2)� $ 1990jtonne cc 57.1xl0 6 kcaljtonne cc�

airlift pump 0.5 kW mechanical power, 1.3 service factor, 8760 hours,� $O.lOjkWhr (1)� $ 570jtonne cc 16.3xl06 kcal/tonne cc�

90

Appendix A continued

Heating and cooling:� 2200 kWhr per tonne cc based on thermal model study,� 0.10 $/kWhr (1)� 2863 kcal/kWhr for production of electrical energy (2)� $220/tonne cc 6.3xl0 6 kcal/tonne cc�

Labor:� $20,000/yr labor to manage eight two tonne modules (1)� $1250/tonne cc�

Transportation:� 80 km average delivery distance, $0.25/km/tonne� transport cost (estimated)� 0.83 kcal/kg/km (2)� $20/tonne p 0.lxl0 6 kcal/ tonne p�

Cost of capital, insurance taxes, interest, maintenance 15% of total capital cost (4) Energy cost 5% of Initial based on 20 year life on equipment $2500/tonne cc 2.5xl0 6 kcal/tonne cc

Total Operating Costs $8074/tonne (P:C=l)

Total Energy Input 93.34xl0 6 kcal/tonne (P:C=l)

(1) Based on Cornell System (Timmons et al., 1987) (2) Pimentel, 1980 (3) Pitcher, 1977 (4) Muir, 1981

Appendix B

Thermal Model Details

The thermal environment of a water reuse system has

been modeled as shown in figure 4.1. Tank volume and

surface area and building volume and surface area were

determined per tonne of system carrying capacity. Waste

heat from the pumps and blowers was assumed to be used

to heat the culture water during the heating season and

to be exhausted to the environment during the cooling

season.

The culture tanks and treatment system tanks have

been assumed to have insulated walls (R-1= 1.9 w/m2/oc =

10 BTU/hr/ft2 /oF). This corresponds to 50 mm (2 inches)

of rigid polystyrene insulating material. The heat loss

from the water surface was taken from the ASHRAE

handbook for horizontal surfaces. The total UA value,

per tonne of carrying capacity, for the culture and

treatment tanks was 0.34 kW/oC.

The entire system is assumed to be housed in an

insulated building with the same R value as the tanks.

This corresponds to 100 mm (4 inches) of fiberglass

blanket insulation. The building was assumed to have

1.5 air changes per hour. The total UA value, per tonne

of carrying capacity, for the building was 0.25 kW/oC.

The culture water was assumed to be kept at a

constant� temperature. A design temperatures of 15°C was

91

92

simulated. Heat was assumed to be added to the culture

water by means of a heat pump. The primary concern in

maintaining culture temperature is to avoid lethal

effects. Low temperatures are generally not lethal to

salmonids, as long as the culture water does not freeze.

Metabolism rates and feed consumption would decline,

however. High temperatures, however have been shown to

cause physiological stress and to have lethal effects.

To avoid losses of fish to heat stress some sort of

cooling system is necessary. It is possible that proper

building and tank design combined with cooling from

subsurface soil could eliminate the need for a cooling

system in some climates. This warrants further

investigation.

1985 Ithaca average daily temperature data was used

to determine total yearly heating and cooling loads per

tonne of system carrying capacity. the results of this

simulation are given in Appendix A.

T air

Building Heat Loss T room

Tank Surface Heat Loss Tank Wallo ·/'/'.,......··,/'/'_...../v'./V'./"'/·· -.....-..< Heat

/ .-.. v"".......-./~~ --:-.//�

:.;.: :~-::::::T t k ~':;;':l::/·:+-~ Loss::;::: ::::::8:8:~~ a 11 AA2:%::::'':::;:'' :;.:

Figure 4.1. Thermal Model Detail.

References

ADNYS, 1984, Aquaculture Development in New York state, New York Sea Grant Institute, State University of New York and Cornell University Publication, 62 pp.

Akagawa, K., 1964, "Fluctuation of Void Ratio in Two­Phase Flow (1st Report, The Properties in a Vertical Upward Flow)," Bulletin of JSME, Vol. 7, No. 25, pp. 122-128.

Akagawa, K., and T. Sakaguchi, 1966, "Fluctuation of Void Ratio in Two-Phase Flow (2nd Report, Analysis of Flow Configuration Considering the Existence of Small Bubbles in Liquid Slugs) and (3rd Report, Absolute Velocities of Slugs and Small Bubbles, and Distribution of Small Bubbles in Liquid Slugs) ," Bulletin of JSME, Vol. 9, No. 33, pp. 104-120.

Apazidis, N., 1985, "Influence of Bubble Expansion and Relative Velocity on the Performance and Stability of an AirLift Pump," International Journal of Multiphase Flow, Vol. 11, No.4, pp. 459-475.

Barnhart, E.L., 1969, "Transfer Of Oxygen in Aqueous Solutions." Journal of the sanitary Engineering Division Proceedings of the American Society of civil Engineers, June 1969.

Bendiksen, K.H., 1985, "On The Motion of Long Bubbles in Vertical Tubes," International Journal of MUltiphase Flow, Vol. 11, No.6, pp. 797-812.

Bendiksen, K.H., 1984, "An Experimental Investigation of the Motion of Long Bubbles in Inclined Tubes," International Journal of MUltiphase Flow, Vol. 10, No.4, pp. 467-483.

Brown, E.E., 1983, World Fish Farming: Cultivation and Economics. AVI Publishing Co., Westport, Connecticut, 516 pp.

Castro, W.E., P.B. Zielinski, and P.A. Sandifer, 1975, "Performance Characteristics of Airlift Pumps of Short Length and Small Diameter," Proceedings of the 6th annual meeting World Mariculture Society, J.W. Avault, R. Miller (eds.), World Mariculture Society, La State University, Baton Rouge, pp. 541­461.

93

94

Clark, N.N., and R.J. Dabolt, 1986, "A General Design Equation for Airlift Pumps Operating in Slug Flow," AIChE Journal, Vol. 32, No.1, pp. 56-64.

Clark, N.N., T.P. Meloy, and R.L.C. Flemmer, 1985, "Predicting the Lift of Air-Lift Pumps in the Bubble Flow Regime," Chemsa Vol. 11, No.1, PP 14-17, January 1985.

Clark, N.N., and R.L.C. Flemmer, 1985, "Predicting the Holdup in Two-Phase Bubble Upflow and Downflow using the Zuber and Findlay Drift-Flux Model," AIChE Journal, Vol. 31, No.3, PP 500-503, March, 1985.

Clark, N.N., 1985, "Gas-Liquid Contacting in vertical Two Phase Flow," Industrial Engineering Chemistry Process Design and Development, Vol. 24, No.2, pp. 231-236.

Collins, R., F.F. DeMoraes, J.F. Davidson, and D. Harrison, 1978, "The Motion of a Large Gas Bubble Rising Through Liquid Flowing in a Tube," Journal of Fluid Mechanics, Vol. 89, part 3, pp. 497-514.

Colt, J.E., and G. Tchobanoglous, 1981, "Design of Aeration Systems for Aquaculture," Proceedings of the Bio-Engineering Symposium for Fish Culture, Traverse City, Michigan, Oct. 1979, American Fisheries Society.

Davies, R.M., and G.I. Taylor, 1950, "On the Motion of Long Bubbles in Vertical Tubes," Proceedings of the Royal Society, Vol. 200 A, pp. 375-379.

Deckwer, W.D., R. Bruckhart, and G. Zoll, 1974, "Mixing and Mass Transfer in Tall Bubble Columns," Chemical Engineering Science, Vol. 29, pp. 2177-2188.

Edwardson, W., 1976, "Energy Demands of Aquaculture - a Worldwide Survey," Fish Farming International, December 1976, pp. 10-13.

Fernandes, R.C., R. Semiat, and A.E. DuckIer, 1983, "Hydrodynamic Model for Gas-Liquid Slug Flow in Vertical Tubes," AIChE Journal, Vol. 29, No.6, pp. 981-989.

Giles, R.V., 1962, Schaum's Outline of Theory and Problems of Fluid Mechanics and Hydraulics, McGraw­Hill Book Company, New York.

95

Govier, G.w., and K. Aziz, 1972, The Flow of Complex Mixtures in Pipes. Van Nostrand Reinold Co. New York, 792 pp.

Greer, W.R., 1987, "Hook Line and Sinker: American Appetite for Fish Spurs Aquaculture Boom," New York Times News service, March 25, 1987.

Griffith, P. and G.B. Wallis, 1961, "Two Phase vertical Slug Flow," Journal of Heat Transfer, Vol. 83, pp. 307-312.

Higson, D.J., 1960, The Flow of Gas-Liquid Mixtures in vertical Pipes. Thesis, Imperial College of Science and Technology, England.

Hills, J.H., 1976, "The Operation of a Bubble Column at High Throughputs-I. Gas Holdup Measurements," The Chemical Engineering Journal, Vol. 12, pp. 89-99.

Hjalmars, S., 1973, "The origin of Instability in Airlift Pumps," Journal of Applied Mechanics, June 1973, pp. 399-404.

Hughes, S.G., G.L. Rumsey, and M.C. Nesheim, 1983, "Dietary Requirements for Essential Branched-Chain Amino Acids by Lake Trout," Transaction of the American Fisheries Society, Vol. 112, No.6, pp. 812-817.

Husain, L.A., and P.L. Spedding, 1976, "The Theory of the Gas-Lift Pump," International Journal of Multiphase Flow, Vol. 3, pp. 83-87.

Jeelani, S.A.K., K.V. KasipatiRao, and G.R. Balasubramanian, 1979, "The Theory of the Gas-Lift Pump: A Rejoinder," International Journal of Mu­Itiphase Flow, Vol. 5, pp. 225-228.

Keitel, G., and U. Onken, 1982, "Inhibition of Bubble Coalescence by Solutes in Air/Water Dispersions," Chemical Engineering Science, Vol. 37, No. 11, pp. 1635-1638.

Kouremenos, D.A., and J. Staicos, 1985, "Performance of a Small Air-lift Pump," International Journal of Heat Fluid Flow. Vol. 6, No.3, pp. 217-222.

Kubota, H., Y, Hosono, and K. Fujie, 1978, "Characteristic Evaluations of ICI Airlift type Deep Shaft Aerator," Journal of Chemical Engineering of Japan, Vol. 11, No.4, pp. 319-325.

96

Large, R.V., 1976, "The Food Production Efficiency of Undomesticated Species of Animals," Food Production and consumption: The Efficiency of Human Food Chains and Nutrient Cycles, Ducham, A.N., J.G.W. Jones, and E.H. Roberts (eds.), North Holland PUblishing Company, pp. 199-214.

Lin,� C.H., B.S. Wang, C.S. Wu, H.Y. Fang T.F. Kuo, and C.Y. Hu, 1976, "Oxygen Transfer and Mixing in a Tower Cycling Fermentor," Biotechnolgy and Bioengineering, Vol. XVII, pp. 1557-1572.

MacDonald, C.R., T.L. Meade, and J.M. Gates, 1975, "A Production Cost analysis of Closed System Culture of Salmonids," University of Rhode Island Marine Technical Report, No. 41, 13 pp.

Mao,� Z.S. and A.E. DuckIer, 1985, "Rise Velocity of a Taylor Bubble in a Train of such Bubbles in a Flowing Liquid," Chemical Engineering science, Vol. 40, No. 11, pp. 2158-2160.

Meade, T.L., 1976, 'The Technology of Closed System Culture of Salmonids," University of Rhode Island Marine Technical Report, No. 36, 30 pp.

Mendelson, H.D, 1967, "The Prediction of Bubble Terminal Velocity from Wave Theory," AIChE Journal, March 1967, pp. 250-253.

Muir, J.F., 1981, "Management and Cost Implications in Recirculating Water Systems," Proceeding of the Bio-Engineergin Symposium for Fish Culture, Traverse City, Michigan, Oct. 1979, American Fisheries Society, pp. 116-127.

Murray, K.R, 1980, "The Design and Performance of Airlift Pumps in a Closed Marine Recirculation system," Proceedings of the World Symposium of Aguaculture in Heated Effluents and Recirculation Systems, 28-30 May, Vol 1.

Naegel, 0., H. Kurten, and B. Hegner, 1977, "Design of Gas/Liquid Reactors: Mass Transfer Area and Input of Energy," in Heat and Mass Transfer in Two-Phase Flow Chemical Systems.

Nagy, Z., 1979, "The Airlift Aerator and its Application in Sewage Treatment," Progressive Water Technology, Vol. 11, No.3, pp. 101-109.

97

Nakoryakov, V.E., and O.N. Kashinsky, 1982, "Local Characteristics of Upward Gas-Liquid Flow," International Journal of Multiphase Flow, Vol. 7, pp 63-81.

Nakoryakov, V.E., O.N. Kashinsky, and B.K. Kozmenko, 1986, "Experimental study of Gas-Liquid Slug Flow in a Small Diameter vertical Pipe," International Journal of Multiphase Flow, Vol. 12, No.3, pp. 337-355.

Nickens, H.V., and D.W. Yannitell, 1987, "The Effects of Surface Tension and Viscosity on the Rise Velocity of a Large Gas Bubble in a Closed, vertical Liquid­Filled Tube," International Journal of MUltiphase Flow, Vol. 13, No. I, pp. 57-69.

Nicklin, D.J., J.O Wilkes, and J.F. Davidson, 1962, "Two-Phase Flow in vertical Tubes," Transactions of the Institution of Chemical Engineers, Vol. 40, No. 2, pp.61-68.

Nicklin, D.J., 1963, "The Air-Lift Pump: Theory and Optimization," Transactions of the Institution of Chemical Engineers, Vol. 41, pp. 29-39.

Nicklin, D.J., 1962, "Two-Phase Bubble Flow," Chemical Engineering science, Vol. 17, pp. 693-702.

Parker, N.C., 1981, "An Air-Operated Fish Culture System with Water-Reuse and subsurface Silos," Proceedings of the Bio-Engineering Symposium for Fish Culture, Traverse City, Michigan, Oct. 1979, American Fisheries Society, pp. 181-187.

Pimentel, D., 1980, Handbook of Energy utilization in Agriculture. CRC Press, 475 pp.

Pimentel, D., and C. Hall, 1984, Food and Energy Resources, Academic Press, Orlando, Florida, 268 pp.

Pitcher, T.J., 1977, "An Energy BUdget for a Rainbow Trout Farm," Environmental Conservation, Vol. 4, NO.1, pp. 59-65.

Rawitscher, M., and J. Mayer, 1979, "Energy Requirements of Mechanized Aquaculture," Food Policy, August 1979, pp. 216-218.

98

Reinemann, D.J., and M.B. Timmons, 1987, An Interactive Program for the Design of Airlift Pumping and Aeration Systems, Department of Agricultural Engineering, Internal Report, Cornell University, Ithaca, NY.

Reinemann, D.J., J.Y. Parlange, and M.B. Timmons, 1987, "Theory of Small Diameter Airlift Pumps," International Journal of Multiphase Flow, (accepted for Publication)

Richardson, J.F. and D.J. Higson, 1962, "A Study of the Energy Losses Associated with the Operation of an Air-Lift Pump," Transactions of the Institution of Chemical Engineers, Vol. 40, pp. 169-182.

Rosenthal, H., 1981, "Recirculation Systems in Western Europe," Proceedings of the World Symposium on Aquaculture in Heated Effluent and Recirculation Systems, Stavanger (ed).

Sekoguchi, K., K. Matsumura, and T. Fukano, 1981, "Characteristics of Flow Fluctuation in Natural Circulation Air-Lift Pump," Bulletin JSME, Vol. 24, No. 197, pp. 1960-1966.

Serizawa, A., I. Katocoka, and I. Michiyoshi, 1975, "Turbulence Structures of Air-water Bubbly Flow," International Journal of MUltiphase Flow, Vol. 2, pp. 221-246 .

Shilimkan, R.V., and J.B. Stepanek, 1977, "Interfacial Area in Cocurrent Gas-Liquid Upward Flow in Tubes of Various Size," Chemical Engineering Science, Vol. 32, pp. 149-154.

shipley, D.G., 1984, "Two Phase Flow in Large Diameter Pipes," Chemical Engineering Science, Vol. 39, No. 1, pp. 163-165.

Slotboom, J.G., 1957, "The Behavior of a Gaslift Pump for Liquids," Transactions of the 9th Int. Congress of Applied Mechanics, Vol. II, pp. 371-383. (Brussels: the University)

Speece, R.E., 1973, "Trout Metabolism Characteristics and The Rational Design of Nitrification Facilities for Water Reuse in Hatcheries," Transactions of the American Fisheries society, No.2, pp. 323-334.

99

stenning, A.H., and C.B. Martin, 1968, "An Analytical and Experimental study of Air-lift Pump Performance," Journal of Engineering for Power Transmission ASME, Apr 1968, pp. 106-110.

Timmons, M.B., W.D. Youngs, P. Bowser, J. Regenstein, G. German, and C. Bisogni, 1987, "A Systems Approach to the Development of an Integrated Trout Industry for New York state," New York State Agriculture and Markets Report, Department of Agricultural Engineering, Cornell University, Ithaca, NY.

Todoroki, I., Y. Sato, and T. Honda, 1973, "Performance of Air-lift Pumps," Bulletin of JSME, Vol. 16, pp. 733-740.

Tung, K.W., and J.Y. Parlange, 1976, "Note on the Motion of Long Bubbles in Closed Tubes-Influence of Surface tension," Acta Mechanica, Vol. 24, pp. 313-317.

United states Department of Agriculture, 1986, Agricultural Statistics, united states Government Printing office, Washington D.C.

VanToever, W.V., and K.T. Mackay, 1981, "A Modular Recirculation Hatchery and Rearing System for Salmonids Utilizing Ecological Design Principles," Proceedings of the World Symposium on Aquaculture in Heated Effluent and Recirculation Systems, Stavanger (ed.).

Wallis, G.B., and J.H. Heasley, 1961, "Oscillations in Two-Phase Flow Systems," Journal of Heat Transfer, August 1961, pp 363-369.

Wheaton, F.W., 1977, Aquacultural Engineering. Wiley & sons, Inc. 618 pp.

White, E.T., and R.H. Beardmore, 1962, "The Velocity of Rise of Single cylindrical Air Bubbles Through Liquids Contained in vertical Tubes," Chemical Engineering science, Vol. 17, pp. 351-361.

Zielinski, P., Castro, W.E., and P.A. Sandifer, 1978, "Engineering Considerations in the Aquaculture of 'Macrobium Rosenbergi' In South Carolina," Transactions of the ASAE, Vol. 21, No.2, pp. 391 -394,398.

100

Zuber, N., and J.A. Findlay, 1965, "Average Volumetric Concentration in Two-Phase Flow Systems," Transactions of the ASME - Journal of Heat Transfer, November 1965, pp. 453-468.

Zukoski, E.E., 1966, "Influence of Viscosity, Surface Tension, and Inclination Angle on Motion of Long Bubbles in Closed Tubes. Journal of Fluid Mechanics, Vol. 20, pp. 821-832.

Zweig, R.D., J.R. Wolfe, J.H. Todd, D.E. Engstrom, and A.M. Doolittle, 1981, "Solar Aquaculture: An Ecological Approach to Human Food Production," Proceedings of the Bio-Engineering Symposium for Fish Culture, Traverse City, Michigan, Oct. 1979, American Fisheries Society.