reinemann phd thesis
TRANSCRIPT
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.