Desalination and Water Treatment 318 (2024) 100383

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Desalination and Water Treatment
journal homepage: www.sciencedirect.com/journal/desalination-and-water-treatment/

Experimental evaluation of liquid circulation flow rate induced by an airinjection-type aerator]]]]]]]]]]

Kazuhiro Itoha, , Norifumi Yoshidab, Hideki Hayashib, Shogo Taguchia⁎ab

University of Hyogo, 2167 Shosha, Himeji, Hyogo 671–2280, Japan
Aience Co., Ltd., 21–7 Edobori-1chome, Nishi-ku, Osaka 550–0002, JapanA R T I C L E

I N F O

Keywords:
Wastewater treatment
AeratorLiquid circulation flow
DiffuserA B S T R A C T

Despite previous research efforts, there is a lack of experimental data on liquid circulation flow rates, particu­
larly in real-scale facilities. The study aims to fill this gap by proposing a momentum transfer model using agas–liquid separation flow to estimate inner flow in a real-scale air-injection-type aerator. The liquid circulation
flow rate was evaluated using velocity measurements. The air supply pressure and dissolved oxygen (DO) were
determined. Experiments were performed at depths of 0.90, 1.30, and 1.75 m. The effect of the inner diffuser wasinvestigated at each depth. The inner diffuser reduced the circulation flow rate but increased the vertical upward
flow velocity. For the deepest condition (1.75 m), the inner diffuser reduced the air supply pressure and in­creased the overall oxygen transfer coefficient. Momentum analysis indicated a suppression effect of the air
velocity at the aerator exit owing to the inner diffuser. This indicated that an increase in the gas holdup owing to
the deceleration of the air velocity at the aerator exit reduced the air supply pressure owing to the reduction in
downstream pressure. This study provides valuable insights into aeration system optimization, consideringfactors such as energy efficiency and oxygen transfer effectiveness.

1. Introduction
Sufficient air supply and agitation are crucial in biological waste­water treatment. Aerators are essential equipment in aerobic waste­water treatment plants [1]. Aeration systems are also essential to otherindustrial processes, such as airlift reactors [2] and bubble curtains [3],significantly influencing performance. The use of aerators is expanding
to environmental purification and protection technologies [4] and themass cultivation of microalgae [5,6].
However, the energy consumption associated with aeration is typi­cally high, estimated to account for 15–20% of the total operating costsof aerobic fermentation [2]. Extensive effort has been made to improveoxygen transfer efficiency in aerator development. A large interfacial
area and low bubble-rising velocity positively impact oxygen transfer
efficiency. Therefore, membrane units have been employed rather thancoarse bubble-type diffusers in wastewater treatment plants [7].
Insufficient oxygen supply arises in scenarios involving high organic nu­
trient and large biochemical oxygen demand (BOD) loads, such as in foodwastewater. Siddiqui et al. [8] indicated that inadequate aeration results in
low treatment efficiency and increased costs. Sediment decomposition owing
to eutrophication accelerates hypoxia near the bottom of the tank. Anaerobic
conditions generate hydrogen sulfide (H2S) and malodorous gas [9–11].

⁎

To overcome these challenges, jet aerators have been developed
to ensure sufficient oxygen supply. Agitation in a deep tank can be
expected through momentum transfer from the strong air–liquid twophase flow to the wastewater [12–16]. Compressed air or pumped
liquid is vigorously mixed in the aerator and discharged into thetank. To develop an air-injection-type aerator, Iran et al. [17] eval­uated the induced liquid flow through particle image velocimetry(PIV). Choi et al. [16] simulated agitation and fluid flow throughcomputer fluid dynamics modeling. Burris and Little [18] andMcGinnis and Little [19] compared the experimental oxygen transfer
efficiency of an aerator with that of an analytical mass balance
model. Recently, interest has been focused on the liquid circulation
flow rate and gas holdup induced by air-injection-type aerators. An
attempt to experimentally evaluate these values has been reported[20] for the airlift loop reactors, which has the downcomers. How­ever, experimental data on liquid circulation flow rates remain lim­
ited, particularly in real-scale facilities for aerobic water treatment
tanks because of the difficulty of evaluation.
Attempts have been made to solve the momentum equation topredict the liquid circulation flow rate during airlift reactor develop­
ment. Liquid flow rate was determined from the motion of bubbles,
defined by buoyancy and fluid drag after being released from theCorresponding author
E-mail address: itoh@eng.u-hyogo.ac.jp (K. Itoh).

https://doi.org/10.1016/j.dwt.2024.100383
Received 6 March 2024; Received in revised form 30 April 2024; Accepted 2 May 2024
1944-3986/© 2024 The Author(s). Published by Elsevier Inc. This is an open access article under the CC BY license (http://creativecommons.org/licenses/by/4.0/).

K. Itoh, N. Yoshida, H. Hayashi et al.

Desalination and Water Treatment 318 (2024) 100383

sparger [3]. However, applying these results to apply to air-injectiontype aerators is challenging because momentum transfer, denoting theacceleration of liquid, primarily occurs inside the device.
Airlift pumps are similar in terms of vertical two-phase flow ex­change momentum in the device. However, their air flow rate is gen­erally lower than that of jet aerators, and their riser tube is thinner andlonger. Thus, the flow pattern and gas holdup, based on the airlift pumptheory, do not align with the range of the aerator [21].
The novelty of this study is the experimental evaluation of the liquid
circulation flow rate induced by a full-scale air-injected aerator in theaeration tank. This study proposes a momentum transfer model using
gas–liquid separation flow to estimate the inner flow in a real-scale airinjection-type aerator. The influence of the inner components (i.e., thediffuser) on the velocity and gas holdup was investigated by sub­stituting the liquid circulation flow rate evaluated using the three-di­mensional velocity into the model. Finally, the tested aerator was ap­
plied in actual wastewater treatment facilities as a case study.

Fig. 1. Experimental facility.

2. Nomenclature
A
C
D
d
dPd
G
g
H
Hd
J
K
KLa
L
I
M
P
Pm
Q
Re
S1
T
u
Xtt

µ
i
w
l

1
2
g
l
r
s
z

(Hitachi Co., Ltd., VB-022-G, maximum air flow rate of 1.2 m3/min)placed on the ground.
Fig. 2 shows the dimensions of the aerator. Air was released verti­
cally upward from a nozzle with an inner diameter of d = 23 mm. The
nozzle was housed in a 134 mm diameter (D ) and 465 mm length (L )outer cylinder. The nozzle and outer cylinder were made of stainlesssteel. The distance between the nozzle and aerator exits was l= 315 mm. The gap between the aerator inlet and bottom of the tank
was G = 50 mm. The injected air and drawn water flowed upward
while colliding with the diffusers along the outer cylinder. Three sets of
two types of diffusers, as shown in Fig. 2(b), were alternately inserted,
totaling six units. Although these diffusers cannot rotate, their blades
were designed to create a swirling flow and enhance agitation in thetank [22].
Table 1 lists the experimental conditions, including the distancefrom the aerator exit to the liquid surface, called the liftoff height Hd .
Three taped water depths were tested, referred to as CASE 1, 2, and 3.
Two types of tests were performed for each case, with (A) and without(B) the inner diffusers, to investigate the influence of the inner diffusers.
Air flow rate was measured using an ultrasonic flow meter (Aichi
Tokei Denki Co., ltd., TRX50, measurement accuracy ±2%) installed on
the horizontal pipe on the ground. Because the data Qg , N Nm3/min was
obtained under normal flow conditions, the actual flow rate Qg m3/min
was converted using the following equation:

Cross-sectional area [m2]Oxygen concentration [mg/L]
Outer cylinder diameter [m]Nozzle diameter [m]Pressure loss at diffuser
Gap between floor and inlet of aerator [m]Gravitational acceleration [m/s2]
Depth of liquid [m]Liftoff height [m] (depth from water surface to aerator exit)
Interfacial area density [1/m]Correction coefficient of outlet pressure [-]Overall oxygen transfer coefficient [1/h]Length of aerator [m]Vertical distance from the nozzle to the aerator exit [m]Momentum [kg m/s]
Pressure [Pa]Measured air supply pressure [Pa]Flow rate [L/min]
Reynolds number [-]Drag coefficient [-]Temperature [°C]
Velocity [m/s]
Lockhart–Martinelli parameter
Greek symbols
Gas holdup [-]Aperture ratio [-]Liquid film thickness [m]Friction coefficient [-]Viscosity [Pa s]
Density [kg/m3]
Interfacial shear stress [N/m2]
Wall shear stress [N/m2]
Two-phase frictional multiplier correlation
Mass fraction of air
Subscripts
Section at the exit of the air nozzle
Section at the aerator exit
Gas phase
Liquid phase
Radial coordinate
saturatedvertical coordinate
Angular coordinate

Qg =101.33 273.15 + T··QgN ,
Pm + P0
273.15(1)

where Pm is the air supply pressure measured using a pressure gauge(SMC Co., Ltd., ISE80, accuracy ±2.5%) installed on the horizontal pipeon the ground, P0 is the base pressure (101.33 kPa), and T is the airtemperature during the measurement.
Three-dimensional velocity was measured using an electromagnetic
flow meter (JFE Advantec Co. Ltd, ACM3-RS, accuracy ±0.005 m/s).
The measurement point was positioned 40 mm above the bottom and at
a 200 mm radial distance (rm ) from the center of the aerator, as in­dicated by the black circle in Fig. 1.
In this study, the measurements of flow velocity and of dissolved
oxygen (DO) were performed separately. The procedures of each were
described below. At first, the procedure for velocity measurement was
as follows: (i) The inverter-controlled blower was started from zero to
the maximum frequency (60 Hz); (ii) Velocity was measured at the
maximum air supply; (iii) The inverter frequency was reduced in 5 Hz
or 10 Hz intervals, and velocity data were recorded after a waiting
period of 3 min to achieve a steady state; (iv) These measurements were
continued until the inverter frequency at which air supply was nolonger possible. The instantaneous velocity values (ul, r , ul, , ul, z ) wereobtained at a 20 Hz sampling frequency over a 15 min duration. The

3. Materials and methods
3.1. ExperimentFig. 1 shows the layout of the experimental facility. An aerator(Aience Co., Ltd., Aquablaster AL-750) was mounted on the bottom ofan underground tank (2 m wide, 2 m deep, and 16 m long). The tank
was kept open to the atmosphere. Air was supplied using a blower
2

K. Itoh, N. Yoshida, H. Hayashi et al.

Desalination and Water Treatment 318 (2024) 100383

Fig. 2. (a) Dimensions of aerator and (b) photographs of inner diffuser.

3.2. Analysis model

Table 1
Experimental condition.

CASE 1 A
CASE 1B
CASE 2 A
CASE 2B
CASE 3 A
CASE 3BH [m]diffuserHd [mm]

DO-1
[m]

DO-2
[m]

0.90
0.90
1.30
1.30
1.75
1.75

Included
Not included
Included
Not included
Included
Not included

385
385
785
785
1235
1235

0.00
0.00
0.00
0.00
0.75
0.00

0.60
0.60
1.00
1.00
1.50
1.50

The injected air was assumed to rise to the center of the outer cy­
linder in a continuous phase because large bubbles were observed near
the outlet of the aerator (see supplementary video). A separated flow
model in an annular air–liquid two-phase flow was employed as shownin Fig. 3. The mass and momentum balance in the steady state between
Sections 1 and 2 without a phase exchange is described as follows:

A 2 g ug , z 2

circular liquid flow rate was calculated using the following equation:

Ql = 2 ·rm· G·ul, r ,(2)

A 1 g ug , z1 = 0,A (12 ) l ul, z 2

A (1

dMg =

Ag dP

g gA g dz(5)1 ) l ul, z 2 = 0,

2 ri i dz

Ag dPd,(6)(7)

substituting the averaged value in the radial direction calculated for
16384 (=214) data points into ul, r .
The DO measurements were performed on a different day from thevelocity measurements to reduce the oxygen level in the tank. Two DO
sensors (Iijima Electronics Co. Ltd., ID-150, accuracy ±0.1 mg/l) were
installed near the side wall of the tank at the depths shown in Table 1.
To measure the DO, the inverter frequency was raised from 0 Hz to50 Hz rapidly. Time-series changes in the DO were recorded without
waiting time. During data acquisition, the inverter frequency was fixedat 50 Hz. The overall oxygen transfer coefficient was calculated using
the following equation:C

KL a =

ln Cs

C1
C2

t2

t1

s

,

(3)

where Cs is the saturation oxygen concentration at system temperature,
and C1 and C2 are the DO values at times t1 and t2 , respectively. The
volume-weighted average values of the outputs of the two DO sensors
were substituted into C1 and C2 . The standard oxygen transfer coeffi­
cient at 20 °C was converted using the following equation:

KLa (20) = KLa· T(20 T ) ,(4)

where T = 19.7–21.7 °C, and T = 1.024.
For the value of velocity, liquid circulation flow rate and pressure, thestandard deviation of fluctuations within the measurement time interval
were displayed as an error bar in Figs. 4 and 5. For KLa (20) , the influences ofaccuracy of DO sensor (±0.1 mg/l) were shown by error bars in Fig. 7.

Fig. 3. Control volume for analysis of vertical upward two-phase annular flow.
3

K. Itoh, N. Yoshida, H. Hayashi et al.

dMl =

Al dP

l gAl dz + 2

Desalination and Water Treatment 318 (2024) 100383

ri i dz

2 rw w dz

Al dPd,(8)

The pressure loss dPd induced by the inner diffuser was assumed tobe approximated by the pressure loss for a perforated plate.

where is the gas holdup, dP is the pressure gradient, ri is the distancefrom the center of the aerator to the air–liquid interface, rw (67 mm) is
the outer cylinder radius, and dPd is the pressure loss induced by the
inner diffuser. The momentum variations, dMg and dMl , are expressed
as follows:dMg = A 2 g ug2, z 2

A2
1 g ug , z1,22 ) l ul, z 2

dMl = A (1

S1 =

(9)
2
1 ) l ul, z1.

A (1

dPd =

l

2
l ul, z .

2

(11)

The friction loss coefficient l is calculated by substituting the liquid
velocity at the aerator outlet into the Blasius equation;ul, z 2·D

Re =

(12)

1/4 ,

l = 0.0791 Re

.

(13)

l

The two-phase frictional multiplier l is proposed for estimating thepressure drop. When the single-phase pressure drop for the liquid, andthe two-phase pressure drops are described as (dp / dz )l and (dp / dz )F
respectively, the following relationship is obtained:

dP
=
dz F

2
l

dP
.
dz l

(14)

is calculated using the Chisholm–Laird equation for turbulent
flow in a smooth pipe [23] as follows:2l2
l = 1 +

21
1
+ 2,
Xtt
Xtt

x 0.9

1
x

g

(15)

0.5

µg 0.1
µl

l

.

i

2 g

i =

g

(ug , zm

1 + 300

D

(17)

5
Reg

2
g

,

(18)

where ug , zm and ui are the average gas and interfacial velocities, re­spectively. In this study, these velocities are denoted by ug , z 2 and ul, z 2 ,respectively. The friction loss coefficient g in gas was evaluated using
Eqs. (12) and (13) by substituting ug , z 2 with ul, z 2 . The average thicknessof liquid film is calculated as = D (12 )/2 , assuming that the gas
phase has a cylindrical shape, and its gas holdup can be represented by
2 . In the absence of waves and entrainment at the gas–liquid interface,
the interfacial area density J , defined as the interfacial area per unit
volume, can be expressed as

J=

4

2

D

.

}2

,

(21)

Fig. 4(a) shows the liquid circulation flow rate Ql , calculated fromthe average velocity ul, r in the radial direction. The horizontal axis re­presents the measured air flow rate Qg at each tested inverter frequency.
The liquid circulation flow rate increased with the depth H .
Compared with Qg ∼ 1000 L/min, Ql = 518 L/min in Case 3B waslarger than Ql = 495 L/min in Case 2B or Ql = 359 L/min in Case 1B.
When the depth of the tank decreased, the influence of waves on theagitation of the liquid became apparent. The momentum transfer fromthe bubble to the liquid decreased owing to the short liftoff height Hd .
The inner diffuser reduced the liquid flow rate in all cases according
to the local pressure drop. However, the maximum values of Qg on thehorizontal axis (i.e., at the maximum inverter frequency) in Cases 2 A
and 3 A were greater than those in Cases 2B and 3B, in which the inner
diffuser was not installed. The strict mechanism of the change in Qg
could not be explained. Nonetheless, the influence of the inner diffuseron the air supply pressure and injection rate is discussed in Section 3.3
based on the results of the momentum analysis.
Fig. 4(b) shows the average velocities in ul, z . A vertical upward flow
is crucial for preventing sediment growth at the bottom of the aeration
tank. The values of ul, z increased when the inner diffuser was installed.
In particular, a strong vertical flow occurred for Qg > 1000 L/min in
Cases 2 A and 3 A. Because the flow is no longer parallel to the bottom,
the value of Ql may be underestimated under these conditions. For in­
stance, ul, r and ul, z were 0.055 and 0.029 m/s, respectively for Qg= 1356 L/min in Case 3 A. Consequently, the flow angle with respect tothe bottom was estimated to be tan 1 (0.029/0.055) = 28 deg.
The influence of the inner diffuser on the circumferential velocity
ul, , as shown in Fig. 4(c), is similar to that of Ql . Thus, the diffuser
reduced the velocity in the suction and circumferential directions andincreased the velocity in the vertical direction.

(16)ui )2 ,

2

4.1. Liquid circulation flow rate and velocity

x is the mass fraction of the gas phase, denoted as quality
x = g /{ g + l (1)} . Using the above equations, the wall friction
term in Eq. (8) was calculated as 2 rw w dz = 2 rw w·l· l2 .
Interfacial shear stress can be evaluated using Wallis’s equation [24]for turbulent flow, expressed asi =)3/8 + 1

4. Results and discussion

where Xtt is the Lockhart–Martinelli parameter for turbulent flow, ex­pressed as

Xtt ={0.707(1(20)

where is the aperture ratio [25]. In this study, these values are= 0.357 for the wing-type diffuser, as shown in the bottom of Fig. 2(b),and= 0.487 for the needle-type diffuser, as shown in the top of
Fig. 2(b). The values at Section 2 were used for the gas and liquid ve­
locity and gas holdup for evaluating dPd .
Finally, the pressures in Sections 1 and 2 were assessed. The pres­sure at nozzle exit P1 was calculated from Pm by considering the cu­
mulative friction loss along the pipe and the local pressure losses due tothe elbows and reducers. p2 represents the pressure head, expressed as
P2 = m gHd , where m = 2 g + (12 ) l is the mixed density. Because
the discharged bubbles would expand or spread before reaching theliquid surface of the tank, the actual pressure acting on the aerator exit
differs from p2 . Subsequently, a factor K was introduced into the mo­mentum balance analysis. In Eqs. (7) and (8), the pressure gradient was
assumed as dP = P1 K · P2 = P1 K · m gHd .
Because the gas holdup at nozzle exit was constant 1= 0.0295 in thisexperiment, four variables, ug2 , ul2 , 2 , and K , were considered in solving
Eqs. (5)–(8). Eqs. (5)–(21) were translated into a Microsoft Excel worksheet.
A solution was obtained using the generalized reduced gradient (GRG)
nonlinear solution method by substituting the measured Qg and Ql . Con­vergence was considered achieved when the residual of the difference be­
tween the right and left hand sides decreased to 1×10−6.

(10)Because only the liquid phase is in contact with the wall, the wall
shear stress is determined as follows:
w =

1 2
u S1,
2

(19)

Therefore, the interfacial friction term in Eqs. (7) and (8) was de­termined as 2 ri i dz = i·J · D2 · l/4 .
4

K. Itoh, N. Yoshida, H. Hayashi et al.

Desalination and Water Treatment 318 (2024) 100383

Fig. 4. Variation of Ql , ul, z and ul, with Qg .

Fig. 5. Pressure in air supply pipe.

5

K. Itoh, N. Yoshida, H. Hayashi et al.

Desalination and Water Treatment 318 (2024) 100383

Fig. 6. Variation of oxygen deficit.

The air supply pressure Pm shown in Fig. 5 in this experiment was
lower than that of the membrane-type aerator (for example, 30 kPa for
Qg =13 L/min on Rizzardi’s experiment [7]). The inner diffuser had
relatively little effect on Pm in Cases 1 and 2. However, in Case 3, thepressure decreased when an inner diffuser was installed. Based on the
results of the momentum analysis, as discussed in Section 3.3, one ofthe potential causes is a lower back pressure owing to a higher gas
holdup after discharge from the nozzle.

momentum transfer after exiting the aerator (i.e., between the liftoff
heights) was considered to be greater than that of other conditions.
Fig. 8 shows the obtained solutions corresponding to the measured (Qg ,Ql ) values.
Fig. 8(a) shows the predicted gas holdup 2 at the aerator exit. The
2 values were larger than the inlet gas holdup of 1=0.00295 and in­creased with Qg . The reduction in the air velocity from ug , z1 to ug , z 2
owing to the frictional and interfacial shear stress in the aerator in­creased the cross-sectional area of air phase A· 2 in Eq. (5) and the 2values.
The triangular symbols in Fig. 8(a) represent the predicted values of
air velocity ug , z 2 at the aerator exit. In the cases with the inner diffuser
installed, as shown by the white triangles, ug , z 2 became lower than theblack symbols. Because ug , z1 at the nozzle exit was nearly the same for asimilar Qg between white and black, the local pressure drop at the inner
diffuser enhanced the air-phase deceleration.
The air phase diameter at the aerator exit can be evaluated by D 2 .
The value of ug , z 2 and D 2 provide an outline of released bubble
property. However, it is difficult to address the tendency of KLa (20) fromthese values because the influences of the breakup or the dispersion of
air bubble after exiting the aerator was not included in this analysis.
The correction coefficients of pressure K were nearly 1 for Cases
2B and 3B. However, the K values for Cases 2 A and 3 A were lessthan 1. The large gas holdup reduced the pressure head K · m gHd .
The momentum balance substituting measured Qg and Ql induced
that the mixed density of gas-liquid fluid between the liftoff height
Hd would less than the value of m at the aerator exit for Case 2 A
and 3 A. The reason is considered to be the deceleration of air
bubble because of the breakup by the effect of diffuser. This may
have caused the lower air supply pressure in Case 3 A comparedwith that in Case 3 B, as shown in Fig. 5. Moreover, the reduction in
K · m gHd may have also increased the maximum Qg in Cases 2 A and
3 A, as shown in Fig. 4, owing to a reduced air nozzle back pressure.
These prospects will be supported by the pressure measurements at
the aerator exit in the future. And the present analytical results
include errors due to neglecting the momentum transfer between
the liftoff heights. Incorporating a model for this section would
reduce the error, especially for low Ql conditions.

4.2. Oxygen transfer
Fig. 6 presents the time-series data of oxygen deficits( ln (Cs C ) ). The inner diffuser affected the variation in ln (Cs C )after 0.005 hr−1 in Case 1 and 0.007 hr−1 in Case 3. Fig. 7 shows thevalues of KLa (20) , calculated by substituting the slope of ln (Cs C ) into
Eq. (3). Notably, the KLa (20) values of the aerator in this study werelarger than those of a venturi-type aerator (KLa (20) = 12.8 h−1 at Ql= 216 L/min) evaluated by Dong et al. [12].
In cases where the diffuser was not installed (i.e., Cases 1B, 2B, and 3B),
KLa (20) decreased with increasing depth. This trend is consistent with thatobserved by Wagner and Pöpel [26] and Gillot et al. [27]. In Cases 1 A, 2 A,
and 3 A, the inner diffuser mitigated the variation in KLa (20) with H . Thus,
the difference between Cases 1 A and 3 A was smaller than that between
Cases 1B and 3B. In particular, the value of KLa (20) in the deepest condition(Case 3 A) exceeded that of Case 3B. The change in agitation caused by the
inner diffuser affected the KLa (20) value.
4.3. Analysis of momentum balance
The solution to Eqs. (5)–(8) can be obtained except for the minimum
Ql conditions in Cases 1 A, 3 A, and 3 B. Under these conditions, the4.4. Case study
A cut vegetable manufacturing plant in Iwate Prefecture, Japan, hasa private wastewater treatment facility with seven 31.5 m3 primary
aeration tanks. An overview of these facilities is provided in the
Supplementary Information. The average flow is 300 m3/d (0.08 mgd).
A total of 35 units, five units in each tank, of the aerator in this study(AL-750, inner diffuser included) were installed. For preaeration, 31
small-sized aerators (AS-250, 68 mm in diameter and 176 mm inheight) were installed in the 101.3 m3 flow-control tank upstream of the
primary tanks. The facility also has a 40.3 m3 sedimentation tankdownstream of the primary tanks. The air flow rate is 3.75 m3/min for

Fig. 7. KLa (20) [1/h].
6

K. Itoh, N. Yoshida, H. Hayashi et al.

Desalination and Water Treatment 318 (2024) 100383

Fig. 8. Prediction ofeach aeration tank and 7.75 m3/min for the flow-control tank. Table 2
lists the wastewater characteristics, all of which meet the Japanese
effluent standards.
The next facility treats kitchen wastewater (60 m3/d (0.02 mgd)
from food and dish washing in the cafeteria of a measurement hardware
manufacturing factory in Kanagawa Prefecture, Japan. Twenty-five AL750 are installed in five aeration tanks, each with a capacity of 23.8 m3.
A small 6th aeration tank (10 m3), a sedimentation tank (7.8 m3), andan effluent flow-control tank (4.0 m3) are installed downstream of thefive aeration tanks. Two AL-750 units and five AS-250 units are in­stalled in the 6th aeration tank and effluent flow-control tanks, re­spectively. The total air flow rate supplied to the 1st to 3rd aeration
tanks is controlled at 13.5 m3/min. The total air flow rate supplied to
the 4th to 6th aeration tanks and effluent flow-control tank is 12.3 m3/min.
Table 3 summarizes the results. In the previous equipment, organic
fats and oils treatment was incomplete, generating foul odors. Afterinstalling the AL-750 units, the residues of the organic material werereduced, and odors were resolved. The sludge disposal interval was
drastically extended.

2 , ug,2 and K .

Table 3
Wastewater characteristics on kitchen of cafeteria of manufacturing factory.
Unit: [mg/l].

Influent
Effluent

BOD

COD

SS

n-hex

1000
40

500
40

1200
50

400
< 1.0

5. Conclusions
This study investigated the influence of the inner diffuser in­corporated in an air-injection-type aerator across three aeration tankdepths. The results limited for a range of the experimental conditions
and the model of aerator was concluded as below.
The inner diffuser reduced the liquid circulation flow rate up to49%. However, the vertical component of velocity increased when the
diffuser was installed. The maximum increase in vertical velocity wasachieved in Case 1 A, which was 75 times greater than that of Case 1B.
Under the deepest condition (Case 3 A), the inner diffuser reduced theair supply pressure up to 13%.

Table 2
Wastewater characteristics on the cut vegetable manufacturing plant. Unit: [mg/l].

Influent
Effluent
Effluent standards

BOD

COD

SS

n-hex

T-N

T-P

920
7.0
160

830
22
160

960
7
200

8.7
1
5

100
4.8
120

14
0.5
16

BOD = biochemical oxygen demand.
COD = chemical oxygen demand.
SS = suspended solids.
n-hex = normal hexane
T-N = total nitrogen
T-P = total phosphorous7

K. Itoh, N. Yoshida, H. Hayashi et al.

Desalination and Water Treatment 318 (2024) 100383

For oxygen transfer, the presence of an inner diffuser reduced thesensitivity of KLa (20) to depth. KLa (20) in Case 3 A was 1.4 times greaterthan that in Case 3 B, where the diffuser was not included.
The air velocity and pressure at the aerator exit were evaluated bysubstituting the measured Qg and Ql values into the momentum balance
equation for annular two-phase flow. The local pressure drop in the inner
diffuser reduced the air velocity and increased the gas holdup at the aerator
exit. A higher gas holdup may reduce the air supply pressure.
This study provides valuable insights into the design and optimi­
zation of air-injection-type aerators. The potential benefits include thereduction in liquid circulation flow rate, improved oxygen transfer ef­
ficiency, and energy optimization through careful consideration of
inner diffuser characteristics.

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Funding
This work was supported by an innovation grant from City of Osaka,
H28.
CRediT authorship contribution statement
Kazuhiro Itoh: Writing – original draft, Investigation, Data cura­tion, Conceptualization. Norifumi Yoshida: Supervision, Project ad­
ministration. Hideki Hayashi: Validation, Supervision. Shogo
Taguchi: Validation, Software.
Declaration of Generative AI and AI-assisted technologies in thewriting process
Statement: During the preparation of this manuscript, the authors
used Deep L for Japanese to English translation. After using this service,
the authors reviewed and edited the content as needed and take full
responsibility for the content of the publication.
Data availability
Data will be made available on request.
Declaration of Competing Interest
The authors declare that they have no known competing financial
interests or personal relationships that could have appeared to influ­
ence the work reported in this paper.
Acknowledgment
The authors express their sincere appreciation to Mr. Fuya
Yoshimura and Mr. Tomoaki Shimada of the University of Hyogo.
Appendix A. Supporting information
Supplementary data associated with this article can be found in the
online version at doi:10.1016/j.dwt.2024.100383.
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