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Institut National Polytechnique de Toulouse (INP Toulouse)
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Hayder Mohammed ISSA
24 Octobre 2013
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CHARACTERIZATION AND IMPROVEMENT OF A SURFACE AERATOR

FOR WATER TREATMENT
DPMF EPDUPSBMF et discipline ou spcialit
ED MEGEP : Gnie des procds et de l'Environnement
6OJUEFSFDIFSDIF
Laboratoire de Gnie Chimique (LGC), Toulouse
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Martine POUX, Ingnieur de Recherches (HDR), INP-ENSIACET, Toulouse
Catherine XUEREB, Directrice de Recherches CNRS-LGC/INPT, Toulouse
Jury :
Cathy CASTELAIN, Directrice de Recherches CNRS-LTN/INP, Nantes, Examinateur
Michel SARDIN, Professeur, ENSIC/INPL, Nancy, Rapporteur et Examinateur
Denis BOUYER, Professeur, Universit de Montpellier II, Rapporteur et Examinateur
Jean-Pierre GRASA, Prsident-Directeur Gnral , Biotrade, Toulouse, Examinateur
Remerciements et Ddicace
Je voudrais tout dabord remercier mes directrices de thse Catherine Xuereb et
Martine Poux, pour la confiance, le temps et la patience quelles mont accord au
long de ces trois annes, ainsi que pour leurs conseils aviss qu'elles ont port au
projet.
Jadresse ma reconnaissance M. Michel Sardin, Professeur INPL Nancy et M.
Denis Bouyer, Professeur lUniversit de Montpelliers II, qui ont accept de juger
ce travail et den tre les rapporteurs.
Je remercie galement Mme. Cathy Castelain, professeur lUniversit de Nantes
davoir accept de prsider le jury de thse.
Je remercie M. Jean-Pierre Grasa, PDG Biotrade pour son intrt ce travail, pour
son aide et ses conseils aviss et pour accepter de juger ce travail.
Mes remerciements sadressent Jolle Aubin et Karine Loubiere pour leur soutien
quelles mont fourni tout au long de ce travail.
Je souhaite remercier lquipe technique du LGC pour leur aide, tout particulirement,
Jacques Labadie, Lahcen Farhi et Alain Muller, pour leur disponibilit, et pour
leur aide dans la mise au point exprimentale. Mes remerciements s'adressent
galement aussi au service administratif du LGC particulirement Danile
Bouscary pour sa disponibilit. Je remercie tous mes collgues du laboratoire LGC.
Je ddie ce travail la mmoire de ma chre mre. Je ddie ce travail aussi ma
chre pouse Aseel, merci pour ton soutien et mes enfants Abdullah, Danya
et Mohammed, qui mattendent avec impatience.
Enfin, je tiens remercier tous ceux qui, de prs ou de loin, ont contribu la
ralisation de ce projet.
II
Abstract:
A new surface aeration system for water and wastewater treatment has been studied.
Its uniqueness lies in its ability to operate in two modes: aeration or simply blending
(mixing) by just reversing the direction of rotation. An experimental plant has enabled
to focus on mass transfer performance and hydrodynamics. The flow pattern and the
velocity field measurements inside the agitated tank were performed by both the Laser
Doppler Velocimetry (LDV) and the Particle Image Velocimetry (PIV) techniques for
the single phase (Mixing) mode and for the two phases (Aeration) mode. The oxygen
mass transfer occurs both in the water bulk and in the spray above water surface and
has been independently investigated. Different configurations and operational
conditions were tested during the experimental part in order to interpret phenomenon
effect of the draft tube and RTP propeller, rotational speed, turbine blades
submergence and else on the flow field and the oxygen mass transfer in the agitated
system that produced mainly by a cone shape turbine. The experimental part dealing
with hydrodynamics and flow field shows that the down-pumping operation mode
with the draft tube has the most convenient results in the mixing mode with respect to
turbulent flow field and mixing time. Whilst for the up-pumping aeration mode the
hydrodynamics experimental results show the whole system configuration is the most
convenient with regarded to mean velocities, turbulent flow intensity and mixing
time. For the oxygen mass transfer experimental part, it is found that the highest
standard liquid bulk aeration efficiency is achieved (SAEb = 2.65 kgO2 kw-1h-1) when
the whole system configuration is used. The highest standard aeration efficiency at 20
o
C for the water spray zone is accomplished ((Esp)20 = 51.3 %) with the whole system
configuration. Several correlations models have been derived for the oxygen mass
transfer in water bulk and spray zones, power consumption and mixing time, on the
basis of experimental results. They can be used as tools to estimate these parameters
for geometrical and dynamical similar systems at industrial scales.
Keywords: Surface aeration, Agitated tank, Mass Transfer, Oxygen mass transfer,
Hydrodynamics, Multiphase Flow, LDV, PIV, Dimensional analysis, Modelling
III
Rsum :
Un nouveau systme daration de surface pour le traitement des eaux uses a t
tudi. Sa spcificit rside dans sa capacit fonctionner selon deux modes : aration
ou simple brassage, en modifiant uniquement le sens de rotation du systme. Un
pilote a permis de cibler le travail sur ltude exprimentale du transfert de matire et
de lhydrodynamique.
Les champs d'coulement et les mesures de vitesse l'intrieur de la cuve agite ont
t raliss par vlocimtrie laser effet Doppler (LDV) et par vlocimtrie par
images des particules (PIV) pour le mode monophasique (brassage) et pour le mode
diphasique (aration). Le transfert d'oxygne se produit la fois dans la cuve et dans
le spray au-dessus de la surface de l'eau. Il a t tudi dans les deux zones.
Diffrentes configurations et conditions opratoires ont t testes
afin de
comprendre les phnomnes dinteraction : tube de guidage, hlice complmentaire
RTP, vitesse de rotation, niveau de submersion des pales de la turbine. La partie
exprimentale sur lhydrodynamique et les champs d'coulement montre que le mode
de fonctionnement en pompage vers le bas (brassage) avec tube de guidage procure
les meilleurs rsultats en termes de mlange si on se rfre aux champs d'coulement
et la mesure du temps de mlange. Pour le mode de fonctionnement en pompage
vers le haut (aration), les rsultats exprimentaux montrent que la configuration du
systme complet est la plus efficace si on considre le transfert doxygne, les vitesses
moyennes, l'intensit de l'coulement turbulent et le temps de mlange. Il est constat
que la meilleure efficacit d'aration standard est atteinte (SAEb = 2.65 kgO2kw-1h-1)
lorsque le systme complet est utilis. L'efficacit d'aration standard 20C la plus
leve au niveau du spray d'eau est obtenue ((ESP)20 = 51,3%) avec la configuration du
systme complet.
Plusieurs modles sont proposs pour calculer le transfert d'oxygne dans la cuve et
dans le spray, la consommation nergique et le temps de mlange. Ces relations
permettent dvaluer linfluence des diffrents paramtres gomtriques et de
fonctionnement dans des systmes similaires une chelle industrielle.
Mots-cls: Aration de surface, Cuve agite, Transfert de matire, Transfert
d'oxygne, Hydrodynamique, Ecoulement multiphasique, LDV, PIV, Analyse

dimensionnelle, Modlisation
IV
Table of Contents
Introduction and Outlines
Chapter 1: Surface Aeration Process for Water Treatment
1.1. Presentation of Different Aeration Technologies in Water Treatment
1.1.1. Diffused Aeration
10
I. Porous Diffusers
10
I.1. Plate Diffusers
10
I.2. Panel Diffuser
11
I.3. Tube Diffuser
11
I.4. Dome Diffusers
11
I.5. Disc Diffuser
11
II. Non Porous Diffusers
11
II.1. Fixed Orifice Diffusers
11
II.2. Valved Orifice Diffusers
11
II.3. Static Tube Diffusers
11
1.1.2. Submerged Aeration
12
I. Submerged Turbine Aerators
13
II. Jets Aerators
14
1.1.3. Aeration with High-Purity Oxygen
14
1.1.4. Aspirating Aeration
15
1.1.5. Surface Aeration
16
1.2. Surface Aeration for Water Treatment Processes
16
1.2.1. Types of Surface Aerators
16
I. Low Speed Surface Aerators
16
I.1. Low Speed Vertical Flow Aerators
16
I.2. Low Speed Horizontal Flow Aerators
18
V
II. High Speed Surface Aerators
18
1.2.2. Principals and Characterization
19
I. Principals
19
II. Surface Aerator Characterizations
23
II. Dissolved Oxygen Concentration Gradient Calculation Methodology
24
IV. Surface Aeration Oxygen Mass Transfer
27
IV.1. Operational Condition Effects
28
A. Rotational Speed Effect
28
B. Number of Impellers Effect
29
C. Liquid Level Effect
29
D. Clearance and Submergence Effect
30
IV.2. Geometry Effect
31
A. Tank Geometry
31
B. Baffles Effect
32
C. Draft Tube Effect
33
D. Surface Aerators Geometry
34
D.1. Surface Aerator Diameter
36
V. Impeller Position in the Treatment Tank
36
IV. Temperature Effect
37
V. Hydrodynamics
38
V.1. Flow Patterns
38
A. Flow Patterns Characterization
38
B. Vortex Formation
40
V.2. Air Bubble Size Distribution and Hold-up
40
V.3. Mixing Time
41
A. Mixing Time Characterization
41
B. Mixing Time Modeling
42
VI
V.4. Air Bubbles Entrainment
43
V.5. Surface Aeration Power Consumption
43
A. Operational Condition Effect
44
B. Power Consumption Relation with Oxygen Mass Transfer
44
VI. Contact Time between Water Droplets and Atmospheric Air
45
VII. Environmental Effects
46
1.3. Conclusions
46
Chapter 2: Experimental Setup and Calculation Methods
51
2.1. Introduction
51
2.2. Experimental Installation and System Description
51
2.3. The Measurement of Power Consumption
54
2.4. Hydrodynamics and Mean Velocity Measurements Techniques
54
2.4.1. Laser Doppler Velocimetry (LDV)
54
I. LDV Apparatus Description
54
II. Tracer Particles Seeding
55
III. Measurement Principals
55
IV. Signal Post-Processing
58
2.3.2. Particle Image Velocimetry (PIV)
58
I. Theory
58
59
III. PIV Principles
59
IV. Scattered Light
60
V. Laser Source
61
VI. Recording Techniques
61
VII. Image Analysis Method
62
VIII. The Cross-Correlation Calculation
62
VII
2.4.3. Other Flow Related Measurements Parameters
63
I. Mixing Time (tm)
63
II. The Pumping Number (NQp)
64
III. Circulation Number (NQc) and Flowrate
65
IV. Agitation Index (Ig) and Flow Quantification
66
2.5. Mass Transfer Experimental Setup and Calculation Methods
67
2.5.1. Oxygen Probe Description
67
2.5.2. Mass Transfer Coefficient
69
I. Liquid Bulk Oxygen Mass Transfer Zone
69
I.1. Introduction
69
I.2. Bulk Zone Oxygen Mass Transfer Calculation Methodology
71
I.3. Testing Different Probe Positions in the Vessel
71
I.4. Repeatability of Experimental Results
72
I.5. De-oxygenation and Re-oxygenation Processes
73
I.6. Oxygen Probe Response Time Measurement Verification
73
I.7. Determination Model of the Bulk Zone Oxygen Mass Transfer Coefficient
75
I.8. Measurement Procedure for the Bulk Zone Oxygen Mass Transfer Coefficient
76
I.9. Temperature Correction for the Oxygen Mass Transfer Coefficient
77
I.10. Oxygen Transfer Rate for the Bulk Mass Transfer Zone (OTRb)
77
I.11. Standard Oxygen Transfer Rate for the Bulk Mass Transfer Zone (SOTRb)
77
I.12. Standard Aeration Efficiency for the Bulk Mass Transfer Zone (SAEb)
78
II. Spray Oxygen Mass transfer Zone
78
II.1. Introduction
78
II.2. Spray Zone Oxygen Mass transfer Coefficient Calculation Methodology
79
II.3. Determination Model for the Spray Zone Oxygen Mass Transfer Coefficient
79
II.4. Temperature Correction for the Spray Mass Transfer Zone
81
II.5. Oxygen Transfer Rate in the Spray Mass Transfer Zone (OTRsp)
82
VIII
II.6. Spray Zone Mass Transfer Coefficient (klad) Measurement Procedure
82
2.6. Water Droplets Flight Time (tf)
83
2.7. Water Droplets Velocity and Volumetric Flow Rate
84
2.7. Conclusions
87
Chapter 3: Oxygen Mass Transfer in the Surface Mode
91
3.1. Water Bulk Mass Transfer Zone
91
3.1.1. Introduction
91
3.1.2. The Experimental Results
92
I. Effect of Geometrical Configuration
92
II. Effect of Impellers Rotational Speed
93
III. Turbine Blades Submergence Effect
95
IV. Effect of the Spacing between the Impellers
97
V. Power Consumption Measurements
98
VI. Standard Aeration Efficiency (SAEb) and Standard Oxygen Transfer Rate (SOTRb) for the
Water Bulk Zone
101
3.1.3. The Modeling
106
I. Mass Transfer
106
II. Power Consumption
108
3.2. Spray Mass Transfer Zone
110
3.2.1. Introduction
110
110
I. Impellers Rotation Speed Effect
110
I.1. Aeration Efficiency for the Water Spray Zone (Esp)
111
I.2. Spray Zone Mass Transfer Coefficient (klad)
115
I.3. Surface Aeration Water Spray Discharge Velocity and Volumetric Flow Rate
117
I.4. Spray Zone Oxygen Transfer Rate (OTRsp)
118
IX
I.5. Contribution Percentage of the Spray and Bulk Zones in the Overall Mass Transfer
Operation
119
II. The Effect of Turbine Blades Submergence
121
II.1. Water Spray Velocity and Volumetric Flow Rate
121
II.2. Spray Zone Aeration Efficiency (Esp)
122
II.3. Spray Zone Mass Transfer Coefficient (klad)
126
II.4. Spray Zone Oxygen Transfer Rate (OTRsp)
127
II.5. Contribution Percentage of the Spray and Bulk Zones in the Overall Mass Transfer
Operation
129
III. Effect of Propeller and the Draft Tube
130
III.1. Spray Zone Aeration Efficiency (Esp)
130
III.2. Spray Zone Mass Transfer Coefficient (klad)
133
III.3. Water Spray Velocity and Volumetric Flow Rate
134
III.4. Spray Zone Oxygen Transfer Rate (OTRsp)
135
IV. Comparing the OTRsp for the Whole System and Turbine Alone Configurations
136
3.2.3. The Modeling
137
3.2.4. Conclusions
142
Chapter 4: Hydrodynamics in the Single Phase Non-Aerated

Agitated Tank for Up and Down Pumping Directions Modes
147
4.1. Experimental Aspects
147
4.2. Mean Velocity Field and Flow Pattern
150
4.2.1. Down-Pumping Condition
150
I. Propeller and Draft Tube Configuration
150
II. Propeller Alone Configuration
158
4.2.2. Up-Pumping Mode
162
162
165
X
4.3. Power Consumption
168
4.4. Pumping Capacity
170
4.5. Agitation Index and Liquid Volume Quantification for the Down-Pumping
Mode with Draft Tube Configuration
171
4.6. Mixing Time
172
4.6.1. The Effect of RTP Propeller Rotational Speed
166
4.7. Conclusions
178
Chapter 5: Hydrodynamics and Flow Pattern in Aerated Agitated

Tank
183
5.1. Introduction
183
5.2. Flow Pattern and Mean Velocity Field in the Aerated Tank
183
5.3. The Effect of the Propeller and Draft Tube
185
5.3.1. Flow Pattern and Mean Velocity Field in the Aeration Tank
185
5.4. The Effect of the Draft Tube
186
186
5.5. Turbine Pumping Number and System Circulation Number
199
5.6. Agitation Index and Liquid Quantification for the Whole System
Configuration
199
5.7. The Mixing Time
200
5.7.1. Impellers Rotational Speed Effect
203
5.7.2. Effect of Propeller and Draft tube Presence
204
5.7.3. Effect of the Spacing between Two Agitators
205
5.7.4. The Effect of the Turbine Blades Submergence
206
5.7.5. Mixing Time Modelling
207
5.8. The Power Consumption in the Aerated Mode
209
5.9. Conclusions
211
XI
General Conclusions and Prospective
215
Appendix I: A Review for Biological Wastewater Treatment with

Activated Sludge
223
I.1. Industrial and Domestic Wastewater Treatment
223
I.1.1. Preliminary Treatment Process
225
I.1.2. Primary Treatment Process
225
I.1.3. Secondary Waste Water Treatment
226
I.1.4. Tertiary Wastewater Treatment
226
I.2. Activated Sludge Process
226
I.2.1. Nutrient Removal
228
A. Nitrogen Removal
228
B. Phosphorus Removal
229
I.3. Aeration Process in the Activated Sludge Treatment
231
I.4. Additional Treatments
232
I.4.1. Sludge Treatments
232
I.4.2. Odor treatment
233
Appendix II: The Derivation of Spray Flowrate for the Surface

Aeration System
237
List of Symbols
241
List of Figures
247
List of Tables
257
References
259
XII
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Introduction
Clean water is growingly demanded in the different fields of human activities, for
example water is more and more being used in the industry (Roubaty and Boeglin,
2007). On the other hand the continuous diminution of existing water resources made
water and wastewater treatment to become a truly developing and problematic
question. One of the ways used to maintain clean water resources for the diverse
industrial or urban demands, is wastewater treatment.
The implementation of the surface aeration process in the water and wastewater
treatment is established as an effective treatment for various wastewater types
especially in activated sludge biological and aerobic water treatment processes. This
technology has an important capacity of delivering the needed oxygen to the aerobic
micro-organisms for respiration and ensures efficient mixed condition for the entire
treatment tank through maintaining the microbial flocs in continuous state of agitated
suspension by accompanied mixing in order to achieve maximum contact surface area
between the flocs and wastewater (Gary, 2004). Surface aeration has various desirable
characteristics such minimum sludge residual is produced for the used activated
sludge process as a continuous operation of recycling the used sludge is implemented
for wastewater treatment plant (Nair et al., 2008; Ramalho, 1977).
Taking in to consideration the capacity of now used surface aerators in the water
treatment field with respect to the accomplished aeration efficiency, energy
consumption and the complicated maintenance as described in the related works.
It is expected from this work to provide the necessary investigations to prove the
flexibility and capability of the purposed innovative surface aerator to work in two
ways, aeration and mixing by simply reversing the sense of the rotation and acting on
a clutching system. The novel investigated surface aerator (FR Patent Demand, 2012)
is found useful and promising after the new method for combination between the
delivering the necessary oxygen into the water treatment tank by the up-pumping
aeration mode and to achieve an efficient mixing for the treatment tank constituents
by the down-pumping mixing mode.
Implementing the new surface aeration technology enables easy maintenance and
more energy saving. In addition the operation may be performed with minimum cost.
This research can be regarded as an approach that may open new opportunities to
enhance the aeration efficiency with optimized operation condition.
During the last decades, the surface aeration is considered as an effective oxidation
mean among the existed wastewater treatments tools. It consists in the dispersion of
the waste water into droplets and there projection through the atmospheric air, where
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a large contacting area is accomplished between the two phases the continuous
atmospheric air and the dispersed water droplets allowing much higher quantity of
oxygen transferred towards the droplets. Then these enriched oxygen droplets are
remixed and well distributed inside the tank. The large interfacial area generated
between the water droplets and the atmospheric air by the surface aeration leads to
high oxygen transfer. The oxygen mass transfer from the gas phase (atmospheric air)
to dispersed water droplets is only limited by the ability of the aerator to provide the
highest water volumetric rate that is exposed to air, (Mueller et al., 2002). Generally
the high speed surface aerators work with the rotational speed range of (1800 to 3600
rpm) depending of the specific conditions applied, while the low speed surface aerator
rotates between 40 to around 100 rpm; these speeds are varied depending upon the
power level of motors utilized. Normally the low speed aerators are fabricated for life
expectancy of 11.4 years before major maintenance or replacement (Stukenberg,
1984).
Mixing and agitation operations are considered as essential factors during the surface
aeration process, where the wastewater treatment process regarded as an effective
treatment according to the mixing condition occurred (Albal et al., 1983).
Both of the concerned mixing and aeration processes are governed by many
constrains that limit their performance, among of them is the mixing time beside the
power consumption, as its always desirable to achieve homogenization condition in
shorter time and lowest consumed energy to optimize power consumption and
reaching the sought contact condition between air bubbles and treated water in the
liquid bulk and its included microbial flocs. Usually the mixing time is considered as
a criterion of flow pattern in aerated and non-aerated conditions (Hadjiev et al., 2006).
There are two main aspects to evaluate a surface aeration system for water treatment
process, the oxygen mass transfer and the hydrodynamics investigations. Oxygen
mass transfer investigation is crucial to figure out the ability of the surface aeration
system and to attain a successful transfer of the oxygen from the atmospheric air to
the water inside the treatment tank or lagoon. Usually the oxygen mass transfer
capability of a surface aeration system is identified by determination two main
characteristic parameters; the standard aeration efficiency SAE, and the standard
oxygen transfer rate STOR. These parameters depend mainly on the achieved
volumetric oxygen mass transfer coefficient in the liquid phase kla during the
operation at standard condition (Sardeing et al., 2005).
For the surface aeration, most of the oxygen transfer is occurred in the generated
spray. The same characteristic parameters that the ones obtained for the liquid bulk
are determined in the spray zone such as the spray standard aeration efficiency
(SAE)sp and the spray standard oxygen transfer rate (SOTR)sp. These parameters are
calculated by determination the spray zone volumetric oxygen mass transfer
coefficient klad (Huang et al., 2009; McWhirter et al., 1995).
-4-

The other aspect of the surface aeration investigation is the hydrodynamic, where
basically for each agitated tank the hydrodynamic is studied to figure out the flow
pattern, the velocity field, turbulence intensity and the pumping capacity of the
implemented impellers. During the past three decades the main measurement means
that applied in this field are the Laser Doppler Velocimetry, LDV and Picture Image
Velocimetry, PIV, because these techniques are non-intrusive and dont interfere the
flow during the measurement (Adrian, 1991). Both the LDV and PIV allow us to
study the flow inside the agitated tank. The essential requirement for these flow
measurement techniques is that the measurement have to be carried out in a
transparent vessel and -if cylindrical- placed in a larger one filled with the same liquid
of the investigation to avoid laser beam diffraction (Aubin et al., 2001).
Outlines
The main objective of this work is to characterize the performance of the surface
aeration system in aeration mode and its blending capacity in the mixing mode. This
will be fulfilled by identifying the affecting parameters on the oxygen mass transfer
process that developed within the liquid bulk inside the tank and in the spray at the
water surface. These parameters are the operation conditions and the geometrical
configurations such as; the rotational speed, mixing time, impellers configurations and
the power consumption.
The other main objective of this work is to acquire the flow behavior for the two
phase (gas-liquid) condition (Aerated mode) and for the single phase condition
(Mixing mode) with related power consumption and impellers configurations.
This thesis consists in five chapters,

Chapter one presents a brief description for aeration types in the water treatment
process. In this chapter the literature review on the surface aeration according to the
most important characteristic parameters, such as the oxygen mass transfer, the power
consumption, the geometrical configuration, flow patterns and aerator types is
presented.
Chapter two includes the description for the pilot and the experimental techniques and
implemented apparatus in the hydrodynamic and the oxygen mass transfer
investigations, such as the LDV, PIV and dissolved oxygen concentration probes. In
this chapter the calculation methods and models applied for the oxygen mass transfer,
power consumption and fluid flow measurements are explained in details.
In chapter three the analysis is focused on the oxygen mass transfer operation
accomplished in the tested surface aeration system by means of dissolved oxygen
concentration measurement both in water bulk and water spray. The effect of the
-5-

impellers rotation speed and turbine blades submergence and presence of the RTP
propeller and the draft tube are investigated. Models for the oxygen mass transfer
dimensionless parameters in the water bulk and spray with effecting parameters will
be built.
Chapter four is dedicated to the hydrodynamics of the single phase stirred tank
(Mixing mode). In this case the aim of operation is to achieve an efficient mixing
condition of the water bulk without further aeration. The experimental runs of the
generated flow by the RTP propeller are carried out using the LDV and PIV for both
up-pumping and down-pumping modes. The effect of the presence of a draft tube will
be tested for these modes. The flow characterization is made through mixing time,
circulation number power number, and agitation index. A model is derived for the
mixing time for the down-pumping mixing mode correlating the influencing factors
with the dimensionless mixing time.
Chapter five involves the hydrodynamics investigations of the up-pumping flow in the
air-water system (Aerated mode). The flow patterns and velocity fields generated by
the turbine in the aerated tank will be characterized by the agitation index, mixing
time, pumping number and circulation number. The effect of different impellers
configuration and the draft tube effect have been tested. A model is derived for the
mixing time for the up-pumping aerated mode correlating the influencing factors with
the dimensionless mixing time.
-6-
-7-
-8-
Chapter One: Surface Aeration Process for Water Treatment
Chapter One
Surface Aeration Processes for Water Treatment
1.1. Presentation of Different Aeration Technologies in Water

Treatment
In the actual field of the water treatment, the oxygen mass transfer to the biologically
active microorganism masses is considered as an essential part of water treatment.
This transfer is improved with implementing the activated sludge in the treatment
tank. Different types of aeration systems have been employed in the water and
wastewater treatment field, the choice is made depending on the location and specific
treatment requirements.
The aeration process for water treatment is a way to achieve higher mass transfer rate
between the oxygen and the water by increasing the interfacial area between them.
Usually the agitation plays an important role in this process by keeping the
homogeneity of the liquid phase (water) with its included biomasses and creating an
acceptable dispersion of the oxygen gas phase. In this process, the size of oxygen
bubbles and the interfacial area are highly dependent on the condition and degree of
mixing (Ju and Sundarajan, 1992). The agitation has also another important role that
is to keep the bubbles as long time as possible inside the tank in order to prolong the
bubbles residence time.
Water and wastewater treatment by the aeration can be basically classified into four
main systems; (i) Diffused Aeration: the aeration is accomplished by various types of
aerators to diffuse the air into the treatment tanks and without implementing agitation
tools or pure oxygen gas sources. (ii) Mechanical and Submerged Agitators Aeration:
the aeration is achieved with the presence of one or more different types of the
agitators in the treatment tank beside the injected air sources. (iii) Surface Aeration:
the aeration is achieved by entraining the atmospheric air into the water bulk by one
or multiple impellers located in the treatment tank without injection of air or oxygen
gas. (iv) Pure Oxygen Aeration: this system is similar to the submerged aeration
except that the pure oxygen is injected throw the water instead of air.
In order to improve these aeration processes many studies were made with various
operation variables such as: the gas hold-up, bubble size distribution, the flow pattern
for gas and water, the circulation time (mixing and aeration time), the properties of
operating mediums (air and water), the transition between dispersion and flooding
states, rotational speed of the aerators, power consumption, geometrical
-8-

configurations (i.e. number of impellers, types of impellers, geometric ratios) and
oxygen mass transfer rate and coefficient in the water.
1.1.1. Diffused Aeration

The diffused aeration is defined as the injection of air or oxygen enriched air under
pressure below the water surface. Beside the gas injection, additional operations are
used like mechanical pumping or mixing and various devices are implied in these
additional operations such as jet aerators or sparged turbine aerators.
The diffused aeration was the first effort in the wastewater treatment with activated
sludge (Stukenberg et al., 1977). The mass transfer occurred for this type of aeration
is in fact consisting of two zones: the gas bubbling dispersion transfer inside the
treatment tank and the turbulent liquid surface transfer at the water surface. The
diffused aeration contains many configurations and operational variations such as the
bubble size formed, diffuser placement, tank circulation, gas flow rate and oxygen
transfer efficiency. The bubbles formed by the diffuser aerator vary from large
bubbles of diameter dB > 6 mm, to medium size bubble of diameter dB, range of 4 6
mm, or fine bubbles diameter less than 4mm.
Many types of diffusers are used in the diffused aeration but generally they are
classified into two major types, depending on the size and distribution of the gas
bubbles preferred (Mueller et al., 2002):
I. Porous Diffusers
They also in turn have many types as follows:
I.1. Plate Diffusers: These are usually having a form of (30 cm) square surface and
(25-38 mm) thick; most are constructed of ceramic media or made of porous plastic
media of (30 cm x 61 cm) surface area. Air is introduced below the plates through a
plenum (See Fig. 1.2).
Figure (1.1): Plate diffuser aerator,

(Permox H ceramic plate diffuser, (Supratec Co. Ltd.)
- 01 -
I.2. Panel Diffuser: These usually employ a plastic membrane which is stretched over
(122 cm) wide placed on base material of reinforced cement compound or fiber
reinforced plastic.
I.3. Tube Diffuser: These are constructed from stainless steel or a durable plastics,
they have generally (51-61 cm) long with (6.4 7.7 cm) diameter (Figure 1.2,a).
I.4. Dome Diffusers: They are usually used with the dimensions (18 cm) in diameter
and (38mm) high, the used medium is usually is ceramic materials.
I.5. Disc Diffuser: they are relatively flat but they differ in size, shape, method of
attachment and kinds of diffuser materials. Generally they have configuration of (1851 cm) diameter (Figure 2.13, b).
(a)
(b)
Figure (1. 2) : (a) Tube diffuser aerator, (b) Disc diffuser aerator, (Gemgate GmbH)
II. Non Porous Diffusers

These types are classified into:
II.1. Fixed Orifice Diffusers: They vary from a very simple form as an opening in
pipes to especially configured opening in a number of housing shapes. They employ
holes that usually range from (4.76 - 9.5mm).
II.2. Valved Orifice Diffusers: Their opening is designed in a way that provides
adjustment of the number or the size of air discharge openings. The air flow ranges
from (9.4 to 18.8 m3/h) and diameter hole is about (7.6 cm).
- 00 -

II.3. Static Tube Diffusers: They consist of stationary vertical tube placed over air
header that delivers bubbles of air through drilled holes, the tube diameters are
normally about (0.3 0.45 m), the average flow ranges are between (15.7 70.7
m3/h) (See Figure 1.3).
Figure (1.3): Static tube diffusers (Process Engineering s.r.l)
For the diffused aeration system with deep tanks between (4-6m), a combination of
aeration systems is used to the water treatment plants in order to improve the transfer
of oxygen. Usually one of applied means is using turbo-compressors to increase the
flowrate of injected air.
1.1.2. Submerged Aeration

The aeration takes place with the presence of an aerator inside the water treatment
tank or basin near the bottom. There are many types of submerged aerators such as jet
aerators or jet turbines. Mostly the turbines are joined with a sparger placed beneath.
Generally for these systems the compressed air is injected at lowest point in the
treatment tank below the aerators by using blowers. Then the aerators disperse the air
bubbles, which are in turn distributed in the water during their rising upward to the
surface. Low volatile organic compounds (VOC) are released to the atmosphere by
submerged aeration (Schultz, 2005). To achieve a satisfactory mass transfer between
the oxygen and the water for the submerged aeration systems, the interested
characters of the hydrodynamic of flow regimes occurring inside the tank and the
interfacial area between the gas and the liquid and the mass transfer coefficient, kla
are always taken in account (Tatterson, 1991).
There are many probabilities of operation conditions, when more than one impeller is
used the lower impeller always has very important effect over the operation. In the
agitated aerated systems, gas cavities can be developed behind the impeller blades (Lu
and Ju, 1989), which it may reduce the power consumption by the impellers or it also
- 01 -

may affect the flow characteristics in the tank. The cavity might be formed due to
many reasons; the most important one is the insufficient gas flow rate when it is
injected separately. There are three kinds of cavities the swirl cavity, adherent cavity
and grand cavities (Xuereb et al., 2006).
I. Submerged Turbine Aerators

These turbines are submerged deeply inside the water tank or basin and consist of an
open-bladed turbine mounted on a vertical shaft driven by a gear motor assembly with
air sparger located under the turbine (As shown in Figure 1.4).
The submerged turbine develops both radial and axial flow. The oxygen transfer is
achieved within the turbulent flow created by the impeller crossed by the bubbles
discharged from the sparger holes. These aerators are generally implemented in deep
tanks aeration.
In the aeration with submerged turbines are mostly equipped with air injection supply
at the bottom of the tank, the most usual type is the disc turbines because they may act
as a second distributer. They collect an important proportion the rising bubbles from
its source before they reach the surface of the liquid and then they redirect them
toward the medium of the dispersion. Nowadays there are other types used to achieve
same objectives such as inclined blade-disc type turbine which admits for higher gas
flow rates at the system. The usual forms for the air distributers or spargers are as
punched ring type or cross-type. They have very effective influence on the dispersion
because they control the process and the flow pattern of the air. The air bubbles are
delivered from the holes on the upper part of the distributers. The best performances
are obtained when the number of holes is limited and has a small diameter. There are
two modes of operation for the air outlet from distributers; first is direct mode, where
all the air will go toward the distributer and then it is swept toward dispersion medium
in the tank; second, the indirect mode where a part of the outlet gas will stay around
the distributer; if the circulation of water is sufficient and the size of bubbles is small,
a part of it is withdrawn by the water flow from upper part of the impeller. The
indirect mode is occurred when larger size and the diameter of the distributor is used,
when ratio sparger to the impeller diameters (Ds/Dtur) is equal to 1.2 (Tatterson, 1991;
Xuereb et al., 2006).
- 02 -
Figure (1.4): Submerged turbine aerator, (ARS-ARS/S Radial submersible aerator,

(Caprari S. p. A.)
II. Jets Aerators

These aerators combine a liquid pumping with gas pumping to result in a plume of
liquid and entrained air bubbles. They are always positioned at the base near the wall
of the basins; the waste water is re-circulated and introduced with gas (that pumped
through separate header) within mixing chamber (Figure 1.5) (Engineers and
Federation, 1988).
Figure (1.5): Hydro Jet Aerator, Plaquette Aerodyn, (Biotrade Co.)
1.1.3. Aeration with High-Purity Oxygen

High purity oxygen aeration is implemented when an increase oxygen mass transfer
rate is highly needed. A 100% pure source of oxygen gas phase is used instead of air
supply. It could be carried out in covered and non-covered aeration. The pure oxygen
is supplied into the water by distributers positioned inside the tank (See Figure 1.6).
Usually mixing impellers are employed to enhance aeration potentials.
- 03 -
Figure (1.6): Flow diagram for uncovered pure oxygen aeration (Mueller et al., 2002).
1.1.4. Aspirating Aerators

They primarily consist of rotating hollow shaft attached to the motor shaft. These
aerators draw the atmospheric air into mixing chamber, where the wastewater will
contact the air and then the air-water mixture is discharged into the treatment tank
(See Figure 1.17). The submerged end of the rotating shaft is consist of propeller
fixed under the water mounted on a shaft, the rotation speed of the propeller is high
about (1800-3600 rpm) to ensure a drop in the pressure over the diffusing surface,
where the pressure is lowered around it and the air was entrained and mixed with
water and then enter the tank as fine bubbles then thoroughly dispersed though the
tank. The advantages for these aerators are; they create less noise than others, easy to
handle and portable, the projection of water drops not needed that is preferable with
limited size of basins but these types are lower efficiency of oxygen rate and more
complicated with mechanical point of view. There are two configuration of this
aerator, the first uses a tube mounted at an angle in the water with a motor and intake
the air above the water surface and the propeller is located below the surface, the
second type has submersible pump supplemented with a vertical air intake tube open
to the atmosphere. These types of aerators are manufactured in a variety of sizes from
0.37 to over 11 kwatt, the angle of the shaft with water surface made by supported
float can be adjusted the control depth of the shaft operation (Boyd and Martinson,
1984; Kumar et al., 2010a).
- 04 -
Figure (1.7): horizontal flow aspirating aerator, Aspirator, (AIRE-O2 Aeration

Industries International).
1.1.5. Surface Aeration

This system will be discussed in detail in next section.
1.2. Surface Aeration for Water Treatment Processes

1.2.1. Types of Surface Aerators
There are many types of surface aerators that are implemented in water and
wastewater treatment. They are primarily classified into several major groups
(Cumby, 1987a; Sardeing et al., 2005; Stenstorm and Rosso, 2008)
I. Low Speed Surface Aerators
Low speed surface aerators are divided into two major categories:I.1. Low Speed Vertical Flow Aerators
The aerators in this category are the older types that generate an upward axial flow
inside the tank and then the water is projected laterally in the air. They essentially
consist in blades fixed under a tray or directly to the shaft of agitation. These blades
are usually immersed in the water (See Figure 1.8).
This category contains some disadvantages that occur during the operation such as the
emission of aerosols or unwanted smells or it is source of noise but all these can be
overcame by installing a cover on the system unit. Usually the peripheral speed or the
blades tips speed is about (4-5 m/s) but this may change depending on the types of
motors used, like using powerful motors (75 KW) or using smaller motors. The
volumetric power consumption may vary between (30-80 W/m3). The surface aerators
- 05 -

of down-ward axial flow are developed recently, where in this type floated agitators
used, where the gas arrive point is fixed on the jacket of the aerator, the non-dissolved
gas was captured by this jacket and then recycled toward the water.
There is low speed aerator with downward flow, where instead of propelling water
droplets throw the air, the air is pumped into the tank near the impeller position and
then dispersed downward (See Figure 1.9). The up-ward flow category contains some
disadvantages that occur during the operation such as the emission of aerosols or
unwanted smells or it is source of noise same us upward type.
Figure (1.8): Low speed vertical flow aerator (Up-ward flow) (Praxair Technology).
(a)
(b)
Figure (1.9): Low speed vertical flow aerator (Down-ward flow), (a) Turboxal (Aire
Liquide), (b) Praxair (Praxair Technology).
- 06 -

I.2. Low Speed Horizontal Flow Aerators
These surface aerators are similar in action with vertical axis type. They are also
called horizontal rotors. Their shapes are like horizontal cylinder with blades, steel
angles, curvilinear or flat steel blades, plastic bars, or plastic discs. They are
submerged in the wastewater at one-half diameter fixed on its surface. These aerators
are usually used in oxidation ditches or in large rectangular treatment tanks. Their
aeration work is occurred while they span the channel or the tank. These rotors spray
the water up and down streams, with imparting a velocity to the water as the blades
rise out of the water. The oxygen is transferred when the droplets contact the
atmospheric air (See Fig. 1.10). The volumetric dissipated power is about (30W/m3).
For the rotating diameters of (0.7m), the peripheral speed at blade tips is about (4m/s).
Figure (1.10): Low speed horizontal flow aerator (Twin mini rotor aeration,
(Botjheng Water Ltd.).
II. High Speed Surface Aerator

These aerators are usually used with electrical motors of rotational speed ranged (7501500 rpm) without reducer; generally they contain a propeller or other types of
impellers with small diameter placed inside. The two advantages of this type are their
moderate price and high flexibility to apply; on the other side they consume excessive
energy and weak ability of agitation (See Figure 1.11). The oxygen transfer capability
and power consumption of this type are close to the aspirating horizontal flow aerator.
- 07 -
Figure (1.11): High speed surface aerator, Aqua turbo (AER-AS), (AQUATURBO
SYTEMS inc.) .
1.2.2. Principals and Characterization

I. Principles
Surface aeration is a mechanical tool that used to entrain the atmospheric oxygen into
the water bulk by surface agitation. The use of surface aeration becomes very
important to overcome several difficulties in the aeration processes. Because of their
simplicity and reliability and competitive oxygen transfer rate, surface aeration is a
popular choice for biological water and wastewater treatment systems (Huang et al.,
2009). It is applied to lessen the economic cost by decreasing the power consumption
requirements compared to other types of aeration. The surface aeration is applied to
improve the mass transfer rate between air and water by achieving larger interfacial
contact between water and atmospheric air. The surface aerators are designed to
promote growth of the aerobic micro-organisms, which in turn they reduce the
biologically demanded oxygen (BOD) of the wastewater by increasing dissolving the
oxygen in the water by creating largest possible contact area. This area is represented
by several calculation parameters such as the standard oxygen transfer efficiency and
the overall transfer efficiency.
The surface aeration process achieved either due to the projection and propelling the
water into the atmospheric air then re-falling of these liquid droplets into the water
again or/and the entrainment of the atmospheric air into the water by the rotation
function of impellers placed inside the liquid phase. Mixing is essential with the
surface aeration to insure the dissolution and the distribution of the oxygen of the air
bubbles into the water for the falling droplets or the directly entrained air bubbles in
inside the tank. The surface aeration includes the refreshment of water surface. In the
water and wastewater treatment the rotation speed of surface aerator should be higher
than specific speed which differs from one case to another on depending its operation
- 08 -

and configuration to prevent sedimentation of the deposits (Roustan, 2003). Surface
aeration is applied to treat waters of needed rates of oxygen up to 80 mg/l h
(Stukenberg et al., 1977).
The principle performances of the surface aerators and other types of aerators are
delivering the oxygen to the aerobic micro-organisms at appropriate conditions (i.e.
the temperature, impeller rotation speed . etc.) and accomplishing a homogenous
distribution of oxygen by the accompanied mixing process of the treatment tank. The
characteristic parameters of surface aerators performance can be represented by
various parameters that are related with the operation condition and the applied
system configurations such as; the mass transfer coefficient of oxygen in clean water
at standard condition kla20. For example the values of kla20 between (3.5 - 10 h-1)
correspond to the standard capacity of oxygenation of (30 -90 g/m3.h) respectively
and for a power consumption between (20-60 W/m3) (Roustan, 2003; Roustan, 2005).
The optimum immersion of the surface aerator turbine is very important and varies
from 2-3 cm to 15 cm depends on the type of turbine used. The treated standard
specific wastewater properties may change due to the variation of several centimeters
of the immersion of aerator. Also the power draw and the oxygenation capacity also
change and each aerator turbine has its optimum rotational speed (Roustan, 2003).
When another impeller is employed with the main surface aeration turbine, this
additional impeller is usually positioned below the main surface aerator turbine inside
the water bulk. The lower impeller helps to disperse more of the air bubbles so the
flow will modified in order to draw the air bubbles or to delay their rising toward the
surface.
The basic concept of the surface aeration is entraining the atmospheric air into the
water bulk. This objective is accomplished either by:
(i) Propelling the water from the surface by a turbine through the atmospheric air to
create direct contact with the air and then entraining the air bubbles into the water
with droplets impingement at the water surface.
(ii) The second way of surface aeration is air entraining from the atmospheric air into
the water bulk by the surface vortices that generated due to the impeller (usually
axial) rotation; these impellers are positioned near the water surface, the entrained air
bubbles are dispersed by the impeller blades (See Figure 1.12).
In the surface aeration process, there are numerous types of turbines that applied to
entrain the air as bubbles into the water bulk, where during the last decades many
types of impellers are invented to achieve the process successfully. Some of these
impellers has equipped with auxiliary propeller to enhance the gas bubbles dispersion
inside the water treatment tank.
- 11 -
Figure (1.12): The surface aeration regimes applying air entrainment from free liquid
surface (A) Direct entraining of the atmospheric air, (B) Spray formation and
entraining the air with droplets impingement at the surface,(Patwardhan and Joshi,
1998).
Various geometries are developed for surface aeration turbines, where the design is
emphasized to project highest quantity of water droplets into the atmospheric air by
achieving the largest contact area between the two phases. The limiting factor for the
surface aeration turbines is always taking in account reducing the radial discharge
flow of the droplets toward water surface level. The circulation of the water and
comprised air bubbles in the treatment tanks is generally consist of a main loop in
entire the tank, many secondary loops can be generated depending on many system
specifications like the number and type of impellers, tank geometry and impeller
position in addition to many system characterizations such as air bubbles hold-up and
retention time and else. For instance when the impeller is placed in the water bulk, the
secondary loops may appear around the impellers, or they can be developed in the
upper or lower part of the tank. When surface aeration turbine is positioned at the
water surface, the water bulk is usually engaged with one main circulation loop.
Deeper treatment tanks are generally preferred to ensure the needed residence time of
the air bubbles but on the other hand these tanks need sophisticated tools to let air
bubble reach the bottom of the tanks (Jakobson, 2008; McCabe et al., 1985; Nagata,
1975; Roustan, 2003; Tatterson, 1994; Xuereb et al., 2006) .
The circulation or mixing time is considered as a measurement indicator of the
average water bulk motion that generated by the impeller in tank. Measurements of
circulation and mixing times are considered as an indicator to understand the scalar
transport in the tank (Edwards and Baker, 2001 ). The circulation time in the surface
aeration is generally associated with; tank overall flow rate, air entrainment flow and
impeller pumping capacity. For more complicated configurations with multiple
impellers, the mixing or circulation time behavior depends on the created circulation
resultant of these impellers, where its not evident always when the number of
impellers increased that leads to shorten the mixing time (Wang et al., 2010).
For continuous flow surface aeration systems that is the case with open channels, the
circulation time is related with; the impeller speed, bulk motion and convective
- 10 -

transport (impeller water pumping capacity). Only one circulation is known during
this process that is the overall circulation related with that actually is of two types of
circulation times that the jet and mechanical agitation mixing times (Tatterson, 1991).
It is important to mention that surface aeration efficiency is highly affected by the
ambient temperature, since the major part the aeration is achieved at water surface it
is normal to have different efficiencies in winter and summer seasons. Each of surface
aeration systems has its characteristics of operation condition, where the selection is
generally made according to the view of the cost consideration and aeration efficiency
(McWhirter and Hutter, 1989).
Table 1.1 illustrates the comparison of the standard aeration efficiencies for different
aeration systems (SAE is defined as the transferred oxygen mass rate to the liquid per
the power consumed at standard condition). This efficiency is dependent on the input
power, air injection flow rate, the aerator submergence operation condition and tank
or basin volume and geometry as it can be noticed in Table 1.2, the performance of
fine bubble diffused aeration is varied according to the type of treatment. The pure
oxygen aeration has higher standard aeration efficiency but this aeration system has
limited application because it required pure oxygen source which is expensive.
Table (1.1): The standard aeration efficiency (SAE) for various aerators types
SAE
(KgO2/kWh)
Aerator Type
High Oxygen Purity Aerator
(3.5-5.5)b
Submerged Jet Aerator
(2.1-2.55)h
Diffuser Aerator (Fine Bubble)

Horizontal Flow Surface Aerator
Slow Speed Surface Aerator
Submerged Turbines
(with Draft Tube)
High Speed Surface Aerator
1.55a, (1.5-2.1)b, 1.66d,

2.2f, 2.27g
1.50a,(1.9-2.2)b
(1.6-2.4)b
1.05a,(1.1-1.4)b, 1.81e
Submerged Turbines(Axial)
(1.0-1.6)b
Submerged Turbines(Radial)
(1.1-1.5)b
Diffuser Aerator (Medium Bubble)

Aspirating Aerator
Orifice Diffuser Aerator
a
2.50a
1.00a, 0.9e
(0.4-0.9)b, 0.42c, 1.6e
0.60a
b
c
d
(Duchene and Cotteux, 2002); (Mueller et al., 2002); (Kumar et al., 2010a); (Moulick and Mal, 2009);
e
f
g
h
(Cancino, 2004a); (Boyd, 1998); (Thakre et al., 2009);
(Taricska et al., 2009).
- 11 -
Table (1.2): The standard aeration efficiency of the fine bubble diffused aeration in
different water treatment basins (Duchene and Cotteux, 2002)
SAE
(KgO2/kWh)
Basin Type
Large Open Channels
3.41
Small Open Channels
1.95
Cylindrical Tank (Flat floor)
3.11
Cylindrical Tank (Grid Arraignment)
2.12
II. Surface Aeration Systems Characterizations

The attempt to reach the desirable oxygen mass transfer in the different surface
aeration systems depends on both the capacity of the oxygen dispersion in the water
bulk and the achieved interfacial contact area between water droplets and air at the
surface for the each method. The employed mixing impeller has very important effect
on the flow of the water in the tank. Mostly the geometry of the tank and the surface
aerator determine the limits between the surface aeration systems, moreover the
performance of the surface aerator is highly affected by the properties of operating
materials.
Usually the characterizing parameters for the surface aeration are mainly the achieved
oxygen mass transfer, the agitation extent and the power consumption. Some of these
general parameters are comprised in more detailed characterizing parameters such as
impeller pumping number and other parameters of Froude number and Reynolds
number.
Many modifications on the surface aerators were made to improve the performance.
Numerous trials have been performed to enhance the operation efficiency by either of
increasing the mass transfer rate kla or by reducing the power consumed in the
operation. To evaluate these mass transfer or mixing performances of the surface
aeration direct experimentations with either global determination of their values or
local methods determination at numerous points in the vessel were made, the second
method is considered effective because for example the size of bubbles are not
uniform along the vessel so the mass transfer coefficient will vary according to that.
The effect of influencing parameters can be determined with varying several
parameters such as impeller rotational speed, impeller and tank geometry and fluid
properties. Various models were derived to relate the important dimensionless
numbers such flow number (pumping number) NQp, (which contains flow rate effect,
impeller rotational speed), Froude number (Fr), (the ratio of inertial forces to
- 12 -

gravitational forces), the related gas measurements (circulated and dispersed) and the
vortex regions characteristics around the impeller if it exists, and other geometric
dimensionless factors for the aeration system. These dimensionless factors were
proposed to characterize the performance of the surface aerators. The power
consumption is usually presented as the dimensionless number of Power number,
(Np), which depends on the type of the aerator and also depends on the Reynolds
number of the system. Usually the power number is between 0.4 - 1 according to the
type of the surface aerator impeller implemented for consumed power per unit volume
that varied between 20-50 W/m3 (Heduit and Racault, 1983a).
The effect of gas hold-up (air flow in the tank) on the rotational speed can be
identified as more gas holdup exists that results in more difficulties to recirculate the
air bubbles in the tank due to longer circulation paths for liquid because of gas voids
presence and the separation of the gas from liquid in the upper region of the tank,
these effects can be accounted by using the gas flow rate measurements (Nienow,
1997).
The relation between the interfacial area, gas holdup and bubble diameter is
commonly determined by employing gas dispersion approaches, which is achieved
by using the equation of continuity and motion with bubbles population balance
including bubble size and concentration distribution, bubble coalescence and
dispersion mechanism (Tatterson, 1994).
Most likely with agitated aerated tanks systems two sets of experiments are done to
evaluate the system potentials, first the experimental run in a lab-scale or in a pilot
plant system to calculate power consumption, liquid hydrodynamics, flow patterns
and mass transfer coefficient. And the second is determining same parameters in the
same configurations with resort to CFD to specify the opportunities to gain better
efficiency of the system by relocating or changing the geometrical positions of the
system components (Tatterson, 1991).
III. Dissolved Oxygen Concentration Gradient Calculation Methodology

The dissolved oxygen concentration transfer principle for surface aeration process has
been studied and investigated by many papers taking into account the factors that
affect the transfer of oxygen from the air to wastewater and the contained activated
sludge.
To model the oxygen mass transfer toward the water direction there are many theories
that describe the oxygen gas concentration gradient such as the two film theory,
penetration model, film-penetration model, surface renewal-damped model and
turbulent diffusion model.
- 13 -

The two film model was found more simple and close to simulate the occurred
process in our case, where it presumes that two laminar films for gas and liquid exist
regardless of the turbulent condition (Taricska et al., 2009).
The oxygen gas transfer to water (liquid phase) is generally represented by two mass
coefficients of gas and liquid side, kga and kla respectively, with assuming that
Sherwood numbers kgL/Dif, klL/Dif are same for both sides, where kl and kg change as
a characteristic length, L and the diffusivity changes, Dif (Taricska et al., 2009).
For aeration system, the high diffusivity in gas film (oxygen) and low diffusivity in
liquid film (water) leads to assume the mass transfer with concentration gradient in
gas phase is negligible, so the kla become the most important coefficient in the
aeration process (Tatterson, 1991).
The principal model that commonly used for oxygen mass transfer in the aeration
operations depending on the mentioned assumptions is:
(1.1)
Where, C and Cs represent the concentrations at any time and at saturation state
respectively
(Stukenberg et al., 1977) investigated the effects of biomass presence parameter in
wastewater treatment on the final determination of the oxygen mass transfer
performance model that presented by equation 1.1; they modified this principal model
that commonly used for oxygen by adding the oxygen transfer correction factor , and
oxygen saturation correction factor , as shown in equation 1.2, where dO/dt
represents oxygen concentration change with time.
(1.2)
While, (Dudley, 1995) has derived a new oxygen mass transfer model that represent
the surface aeration process. He suggested that the previously used models couldnt
represent the true condition of oxygen distribution inside the entire aerated tank.
(Dudley, 1995) derived the following model:
dC/dt= kla ( Cs-C) r MLSS us(dC/dz)
(1.3)
Where, MLSS is the mixed liquor suspended solids (the water with containing
microorganisms); r is the specific respiration rate; z is the length of each stage and us
is the liquid velocity. The modified model takes in account the correction from water
conditions to mixed liquor. This general model proposes a relative improvement in
the oxygen mass transfer coefficient (kla), but it still didnt reach the required true
value depending on standard operation conditions. (Ju and Sundarajan, 1992) found in
their studies on the oxygen transfer in surface aerated bioreactors with containing
microorganisms that the presence of the biomass showed no effect on the oxygen
transfer rate to water because of their formed film adjacent to the gas-liquid film is
- 14 -

extremely small. So it resistance to oxygen transfer considered negligible and the
correction factor can be set to 1 when experimental investigations are made on the
oxygen mass transfer to simplify the calculation and to prevent over estimation of the
biomass effect on the calculated kla.
The previous oxygen mass transfer model that applied to calculate kla had been
modified by (McWhirter et al., 1995) for surface aerator. They had divided the mass
transfer process zone into two essential zones: the first is for the droplets projection
form aerator turbine blades tips till it impinges the liquid surface, where they derived
a new model for this purpose by considering the liquid droplets traverse in continuous
infinite atmospheric air (gas) phase, the overall oxygen concentration distribution
within the two zones was:
CL= (A Cd*+B CLS */A+B) + (CS -(A Cd*+B CLS */A+B)) exp (-(A+B) t)
(1.4)
Where, A = Q Emd /VL; B = klsas; Cd*and CLS* are the saturated concentrations in the
droplet and water respectively and Emd is Murphree efficiency for the operation.
(Oliveira and Franca, 1998) modified the previously developed model for oxygen
transfer by (McWhirter et al., 1995) to represent the surface aeration for the turbines
that placed in water sub-surface position. They tried to simplify the model by
applying boundary initial condition and other considerations and assumptions to
eliminate the low effecting parameters, so they found the following model for oxygen
mass transfer coefficient with certain conditions:
dy/ dz = - At/G kla (CL* - CL) K2
(1.5)
With applying Henry's low and rearranging the equation the found following model
was:
C*L = C*st (1+yo/yo) [[Pb-Pv+L g (Zs-Z) / 1-Pv ] y/y+1
(1.6)
C*st= 32 / 18H (1- Pv) (1+yo)/yo
(1.7)
Where C*st is the standard DO level; Pb is the barometric pressure; Pv is the vapor
pressure; G is the gas flow rate; K2 conversion factor; At is the tank cross-sectional
area; C*L represents the true bulk liquid DO level; yo is the oxygen concentration in
the bubbles at (z = 0); and H is Henry's law constant.(Oliveira and Franca, 1998)
tested these models with the previous models experimentally, where they found that
acceptable fitting between the experimental and theoretical results, where the
dissolved oxygen at equilibrium state is decreased with increasing the temperature,
the highest oxygen transfer is noticed with low temperature values.
(Stukenberg et al., 1977) have studied the probability of the calculation errors with
several types of aeration equipment including the surface aeration. They studied the
- 15 -

procedure conducted to determine the dissolved oxygen concentration in the tanks for
different test durations and various effects on the exact results values of the
concentration and they compared between the theoretical and experimental results for
the achieved saturated dissolved oxygen and the mass transfer coefficient. They made
a comparison made between two methods. They called them the direct and
conventional of mass transfer coefficients, where they found that the different is by
the way of calculation. They investigated practically the evaluating important
parameters of the aeration in details like the methods to determine the dissolved and
saturated and , was discussed with suggesting the most correct results can be
obtained.
IV. Surface Aeration Oxygen Mass Transfer

It is very hard to classify the investigations and studies that made for the surface
aeration, as the influence of the relevant parameters is very merged and blended. So it
is quite tricky to identify what is criterion for the classification among the affecting
parameters when describing mass transfer operation or hydrodynamics in surface
aeration. The guidelines for this classification in this literature depend on the main
axes those were followed by the achieved studies in this domain.
(Heduit and Racault, 1983b) made a general assessment of the oxygen mass transfer
coefficient and aeration efficiency for the various surface aerators types that operate
in field. Among their study they found that about 85% of 111 tested low speed surface
aerator aeration efficiencies were between 1.2-1.9 kgO2 / kWh, and the average
aeration efficiency for all low speed turbines were 1.49 kgO2/kWh as shown in the
Fig. 1.13.
Figure (1.13): The histogram distribution of measured aeration efficiency for (111)
low speed surface aerators in field, the average is 1.49 kgO2/kWh, (Heduit and
Racault, 1983b).
- 16 -

(Fan et al., 2010) considered the surface aeration mass transfer operation is consisting
of three zones, they have created CFD model for a system of high speed surface
aerator. They found the most effective one is the mass transfer during the water spray
in the air and they assumed the ratio of re-aeration or air entrainment by water
droplets when impinging the liquid surface is not effectual. (Fan et al., 2010)
employed a single phase three- dimensional CFD model for fluid flow simulation to
represent the flow and dissolved oxygen distribution inside the tank and they
depended on the experimental results of dissolved oxygen in two positions in the tank
and the overall mass transfer coefficient, where the difference between them was
quite little.
(Patil et al., 2004) derived a general correlation for the surface aerators process mass
transfer depending on the previous works in the same field. They related oxygen mass
transfer coefficient with the effect aerator geometry and the operation condition and
power consumption in one model, the range of volumetric power consumption range
for this model is 90 < P/V < 400 W/m3.
kla/N =7 *10-6NP 0.71 Fr 0.48 Re 0.82(h/D) -0.54(V/D3) -1.08
(1.8)
IV.1. Operational Condition Effects

(Zlokarnik, 1979) has related the mass transfer performance for different aerator types
with the operational parameters such Froude number and Reynolds number and other
geometrical factors with changing the numbers and forms of impellers blades. He
formulated a dimensionless formulation that combines all the surface aerator
efficiency term (E) with aeration number and Froude number with what he called
sorption number (Y) , which is a dimensionless number and represents the oxygen
transfer, the model was developed for the ratio h/D =1.0,
(1.9)
Where, Y=G/c d3 (v/g2)1/3; G represents the oxygen uptake rate.
A. Rotational Speed Effect

The status of surface aeration changes due to the increasing the rotational speed as
founded by (Albal et al., 1983) in their investigations with the effect of operational
conditions. At low speed the oxygen gas transfer occurs only by diffusion at the
oxygen-water interface. With increasing the speed the oxygen mass transfer rate is
developed by creation convective forces inside the liquid, where velocity of air
bubbles increased, with further rotation speed increasing that leads to higher air
bubbles entrapped into the water bulk. (Backhurst et al., 1988) found out same
- 17 -

relation between the rotation speed and surface aeration efficiency with pilot and full
scales.
A critical rotational speed of the surface aerator (starting with this speed the aeration
efficiency or the oxygen transfer begins increasing relatively) can be observed when
the effect of rotational speed and the geometric parameters on the oxygen transfer
efficiency were tested as presumed by. The bubbles were created when the water
droplets hitting the water surface, in consequence the liquid bulk circulation and the
air entrainment are forming. (Takase et al., 1984) derived a model to represent the
standard aeration efficiency for the critical rotation speed higher in square tank:
SAE =6.6*10-6(ND)-0.1(D/WT)0.4 (D/h+WT) 0.5
for D/WT = 0.24, h/D = 2.5
(1.10)
Where; WT is the square tank width.
B. Number of Impellers Effect

(Veljkovi and Skala, 1989) have investigated the number of impellers that implied
for surface aeration process. They reached to the conviction of utilizing two impeller
gives higher oxygen transfer rates than using one impeller for turbine impeller type
and same rotation speeds, where the position of upper impeller at the water surface
enhances the intensity of surface aeration.
While a system of surface aeration consists of three immersed impellers may have
better operation performance for the gas holdup and dispersion consideration inside
the liquid as proposed by (Li et al., 2009). They examined several groups of three
impellers systems for the flow pattern, mass transfer coefficient. They found that the
three impeller system had very important effects on the gas distribution inside the
tank. They also found from results the best impellers combination was Rushton disk
turbine RTD, Techmix 335 hydrofoil impeller up-flow TXU and half elliptical blade
disk turbine HEDT distributed from above to bottom respectively.
C. Liquid Level Effect

The controlling factor that influences more the oxygen transfer is the liquid level
beside the rotation speed as founded by (Thakre et al., 2009) from their experimental
results. They have developed a correlation model for the oxygen mass transfer
coefficient in the oxidation ditches by applying curved rotor aerator, where these
relevant parameters presented in the developed model.
kla = 0.000746[(N)1.768 (h/D)1.038 ()0.031]
(1.11)
Where: is the blade tip angle. This model is applied within these ranges of Re
*103(50-84) and S/D (0.17-0.25).
- 18 -

D. Clearance and Submergence Effect
OTR, (kgO2h-1)
(Backhurst et al., 1988) examined the effect of blade submergence on the oxygen
transfer rate for different impeller types. Their results showed that there is an
optimum submergence (starting with submergence an efficient aeration is noticed) for
all tested impellers as illustrated in the Fig. 1.14.
Figure (1.14): The relation between surface aeration impeller blades submergence and
oxygen transfer rate for different blades number, (where H is the liquid level in the
tank) (Backhurst et al., 1988).
The clearance of surface aerator impeller in the treatment tank is commonly defined
as the distance between the lowest point of surface aerator and the tank bottom. While
the submergence is defined as the distance the water surface level and the specific
point on surface aerator blade. (Patwardhan and Joshi, 1998) concluded that with
increasing the submergence of surface aerator impeller the intensity of surface
aeration, oxygen transfer rate and gas hold up are decreased. They explained that the
amount of energy reaching the liquid surface is decreased. That also agrees with the
results obtained by (Backhurst et al., 1988). It is better always to set the impellers in
closer position to the liquid surface to enhance the reached energy to the liquid
surface and increase the gas holdup. This persuasion was found by(Deshmukh and
Joshi, 2006) by testing three types of surface aerators impeller of PBTU, PBTD and
- 21 -

DT. In their experimental work (Deshmukh and Joshi, 2006) examined each type by
varying the rotational speed with submergence for each type of impeller.
For these three different types of impellers that applied for surface aerator; PBTD,
PBTU and DT impellers, (Patil et al., 2004) found that an optimum submergence
position for PBTD can be found. Where best mass transfer coefficient was reached
with submergence (S/D) that equals 0.2 and rotation speed of (2.5 1/s) for tank
diameter 1.5 m. (Patil et al., 2004) studied the positions for three types of the
impellers; pitch blade turbine up-flow PBTU, pitch blade turbine down-flow, PBTD
and disc turbine, DT, where for all effective operations the impellers were located
near the liquid surface. They found that generally with increasing the submergence
the number of eddies at gas-liquid contact area were decreased, also the maximum jet
size accomplished by the impeller blades location at just near in the water. The
maximum value of kla was at the submergence ratio of 0.12D that was accomplished
by PBTU type. The other types didnt have the same ability to project the liquid in the
air. (Patil et al., 2004) determined the effect of impeller clearance, where the tests
were conducted with keeping the submergence constant where for impeller diameter
that equals Tv/3 it was found that optimum kla found at the clearance of (1.98D), but
for the impeller diameter that equals Tv/5 the kla was noticed decreasing constantly
with increasing the clearance.
IV.2. Geometry Effect

It is so difficult to categorize the most important geometric parameters in surface
aeration for various techniques that implied in the surface aeration (Kumar et al.,
2010b), but generally it is found that there are frequent parameters that can affect
performance for the majority of the surface aeration as following:
A. Tank Geometry
The surface aeration for water treatment is usually performed in cylindrical shape
tanks or basin, which are the most usual among the used tanks, but the geometry of
these cylindrical tanks may vary between plate bottom shape to curved and conical
shape (with 150 - 300 degree angle depending of the existed activated sludge
properties).The volume of the tank is commonly related with the height of the liquid
in the tank (Jakobson, 2008). Square or rectangular shape are also used but in very
limited way and for especial uses. On the whole, the water treatment tank volume that
equals height of the liquid is considered as standard geometric ratio for design
considerations.
(Rao and Kumar, 2007b) have derived a model for circular shape aeration tank, where
the derived model relates the mass transfer coefficient of oxygen, kla, with circular
- 20 -

tank on the basis of theoretical power input to the system. The mass transfer
coefficient used for theoretical power per volume, for baffled tank is:
k = (3.26 exp[-0.56/X] +0.21 - 0.426exp[-0.47(X - 0.878)2])10-6 X
(1.12)
Where X represents the theoretical power/ liquid volume (Fr4/3/Re1/3) and k =

kla20(2/g)1/3, the model is developed for the geometric ranges; h/D =1.0, S/W=1.26
and W/D = 0.24.
The experimental results of (Rao and Kumar, 2007a) showed that the square tank
aeration was also effective in surface aeration process, where higher values of mass
transfer coefficient k are achieved in shorter duration (that doesn't agree with previous
woks where the square shape not preferred because the formation of dead angles
exist) but in the power requirement point of view the circular aeration tanks were
more effective for less amount of power was required to reach the value of mass
transfer coefficient with keeping the other conditions constant during the experiment.
For square tanks, (Rao and Kumar, 2007a) verified a correlation that was developed
earlier in the previous work (Rao, 1999), which represents the mass transfer and
power measurement on electrical measured basis. The general correlation found by
(Rao, 1999) was for the mass transfer parameter k (where k = kla20(2/g)1/3), as a
function of geometric and physical properties that referred as, X, the ratio of the
Froude Number Fr, to Reynolds No., Re, for a baffled tank:
k = [17.32 exp(-0.3/X 1.05)+3.68 -0.925 exp(-750 )X-0.057)2)]10-6 X
(1.13)
Where; X is (Fr4/3/Re1/3). This model is applicable within X range of (0.01 8.0) and
it is developed for the geometric and operational ranges; h/D =1.0, S/W=1.26 and
W/D = 0.24
(Fuchs et al., 1971) have studied the surface aeration performance by examining the
effect the volume of the aeration tank according to volumetric power provided to the
operation, where they tried to keep the mass transfer coefficient constant during the
tests. (Fuchs et al., 1971) founded that the oxygen mass transfer coefficient was
generally increased as the volume of aeration tank is decreased for large volume
tanks, a satisfactory results found for high levels of provided power per volume ratios
B. Baffles Effect
Baffles are commonly used in the water treatment tank, where they are fixed near the
walls of the tank to reduce or prevent the formation of vortex that are generated
because of the centrifugal force created by impeller rotation especially when
cylindrical vessels used and when the impellers are centrally positioned in the vessel
- 21 -

(Jakobson, 2008). In general the baffles are considered as vertical blades works to
divide the primary motion into axial and radial movements according to the model
used. The number of baffles that used may vary according the method of use, but
generally the numbers between two and four are advocated (to do their mission
perfectly) because one is not enough to prevent the formation the vortex. It is very
necessary to put them in symmetrical form and it is advised to choose the number of
the baffles same with the number of the blades of the used impeller for reducing the
symmetrically mechanical stress on the shaft. The width of the baffles usually chosen
as (T/10), are not touched (affixed) to the walls, the distance between the baffles and
the walls is usually (T/50), and the long of the baffles is generally exceed the surface
of the liquid and reaches tank bottom (Treybal, 1980; Xuereb et al., 2006).
(Lines, 2000) has used a dual impeller surface aeration system to find the effect of
three types of baffles on the aeration with changing impeller speed and liquid height.
With angle-blade turbine and six flat-bladed discs turbine, the highest mass transfer
was obtained with 4-half height wall baffles and the gas-liquid mass transfer
coefficient was decreased with increasing the liquid height.
(Rao and Kumar, 2007b) investigated the effect of the baffles in circulated shape
tanks for wastewater treatment aeration process. They tested the performance of the
system by calculating the kla for both ratios of actual and the theoretical power per
unit volume. They applied a simulation of oxygen transfer coefficient in the two
cases. They deduced that the baffled system is more efficient in the treatment process
but it also more power consumer so they recommended using un-baffled system for
long duration treatments, where the power consumption will be more important. The
baffled circular shape tank system can be used for short duration treatment or in rapid
aeration process, where power consumption was less importance. They derived two
models for un-baffled and baffled tanks. For un-baffled as shown in following
equation:
105k = 7.38 PV exp (0.189/PV ) + 0.33(PV )0.5
(1.14)
For baffled tank:

105 k = 3.95 PV exp (0.85/PV) + 0.15(PV )0.5
(1.15)
Where, PV = actual or measured power/ volume and k = kla20 (2/g)1/3. These models
are developed for the geometric and operational ranges; h/D =1.0, S/W=1.26 and W/D
= 0.24.
C. Draft Tube Effect

Draft Tubes are used some times with the surface aerators to centralize the return flow
to the impeller and to centralize the direction and velocity to the suction region
(inward) of the aerator turbine.
- 22 -

It is needed to modify the surface aeration systems by adding draft tube to prevent
vortex formation or to enhance flow patterns. (Kirke and El Gezawy, 1997) in their
investigations have tested the effect of draft tube presence on the impeller function.
They found out the flow profile is improved and axial velocity rate was higher with
using the draft tube.
(White and De Villiers, 1977) have investigated the relation between the presence of
the draft tube and its effect on the aerated agitated tank oxygen mass transfer
performance. They found that it has remarkable effect on the aeration number N A,
where they observed the increasing in applied pressure and reduction of hydrostatic
head of water was maintained by the existence of draft tube.
D. Surface Aerators Geometry

Many types of aerators are used in the surface aeration; the design characteristics
changes due to the required operations. There is wide diversity of the impellers that
can be implemented in the surface aeration according the needs for each specific case.
The axial and radial impellers are used with many shapes and forms depending on the
particular conditions of the operation and the aim of the process. The most used axial
impellers types are the pitched-blade propeller and hydrofoil propeller. There are
numerous types of turbines that used in surface aeration. The number and the shape of
blades can be changed to curved, pitched and inclined blades if there are needs for
special performance. The turbines can participate to generate homogenous flow inside
the tank and prevents the air bubble rise to the surface of the liquid and they are able
to form a shear due to the force gradient of the velocity which is essential for the
systems of gas-liquid, where it improves the oxygen mass transfer as a result of this
shear that localized in reduced zone and by this the turbulent intensity was increased
(Roustan, 2003; Tatterson, 1994; Xuereb et al., 2006). It is important to know that not
all the turbines can act as successful aerator for all cases, where there are specific
impellers types that are suitable to perform the surface aeration correctly for each case
, as (Roustan et al., 1975) figured out by comparing several types of aeration systems.
(Cancino, 2004a) made a comparison of the performance of several types of axial
flow surface aerator that conducted in (Cancino et al., 2004), where he tried to
compare between these configurations in depending on factors such (water splashed
flow/ power consumption ratio, Q/P) and mass transfer coefficient with changing the
aerators geometric configurations like blades shapes and their types with inlet and
outlet angles of the water flow. He found that not always the increasing of aeration
efficiency is accompanied with increasing mass transfer coefficient because other
important factors may affect the aeration like the types of blades, inlet and outlet
angle and pattern of water projection in the air (i.e. the droplets that well dispersed).
In his assessment he preferred to take the ratio (Q/P) than using Q, itself because the
behavior of aeration efficiency with water flow is not clear where it is affected by the
- 23 -

power consumption. The best global mass transfer coefficient at 10 oC yielded was
(3.249 1/h); best standard aeration was SAE (1.805 kg O2/kWh) for the flat propeller
type.
In depending upon of the work of (Cancino, 2004a; Cancino et al., 2004),(Cancino,
2004b) investigated the best configuration of aerators groups in order to develop a
model by applying the dimensional analysis. He has chosen the most important
geometrical parameters that affect the process beside other parameters to characterize
two different aerators, the flat and pitched blade (PB) types. The results and
application of correlation equation showed difference in accuracy between the types
of aerators used, where the equation was about 80.3% close to the experimental
results for the flat propeller type, while it was about 56.5% accuracy to the
experimental results for other types. The general model he for flat blade impellers:
AE = (Q/P)-o.o461 (D N)-0.9345 (Re)0.003778 (Fr)1.3696 (2)-0.598 (D/S)0.039
(1.16)
AE is the aeration efficiency (it is defined as the transferred oxygen mass rate to the
liquid per the power consumed at actual condition).
For pitched blade impellers
AE= (Q/P)-o.o2678 (D N)-5.7148(Re)-0.3388(Fr)4.8695(2)-0.3676(D/S)-0.1256(1)0.2
(1.17)
Where Q is the water flow splashed by the aerator; 1 is 1/Inlet angle of paddle (rad1);
2 is 1/Outlet angle of paddle (rad1). The model is developed for impeller blade
submergence (58% - 228%).
For paddle wheel surface aerators performance, (Moulick et al., 2002) established a
general model by finding a modified standard aeration efficiency termed SAE'. They
verified the model by performing several set of experiments with varying the
geometric configuration for each one with SAE' to reach the optimal geometric
parameter, where they varied the paddle width / impeller diameter ratio, liquid
volume/ impeller diameter, impeller pitch/ impeller diameter and horizontal projection
of bent length / impeller diameter and bent angle. The correlated general equation
was:
(1.18)
Where (SAE') = (SAE) (v/g2)1/3 (c)-1 . .N3. D2and X= Fr4/3/Re1/3. The ranges of this
model are; Re (2105 - 8105), Fr (0.05 0.25) and (S/D)paddle wheel (0.025-0.225).
(Zlokarnik, 1979) compared between different surface aerator types with respect to
the oxygen mass transfer. He found that the surface aeration is highly influenced by
several geometric ratios. The investigation he made was as a function of aerator type
and their geometrical specification with other operational parameters like Froude
- 24 -

number and Reynolds number for different aerator types with changing the numbers
and forms of impellers blades (See Figure 1.15).
(Patil et al., 2004) found out the surface aerator with the PBTD type was the best in
the general in comparison between the impellers for surface aerator; PBTD, PBTU
and DT impellers within the tested range of system configuration with respect to the
values of mass transfer coefficient achieved by these aerators.
(Deshmukh and Joshi, 2006) have tested the impeller geometry effects on the liquid
flow profile for the surface aeration system with impeller placed in water sub-surface
position. They tested different impeller types PBTU, PBTD and DT. The results
showed that the performance of PBTU was the worst between the impellers according
flow patterns in the tank with increasing the rotation speed, where the created flow
was unsatisfactory and didnt cover all parts of the tank in contrary to the rest two
impellers PBTD and DT.
Figure (1.15): Relation between sorption number (Y) with the Froude number (Fr) for
the conical shape turbine surface aerator (Zlokarnik, 1979).
(Backhurst et al., 1988) built a correlation for the OTR (The oxygen mass transfer rate
to the liquid during the aeration (kgO2/h)) with affecting factors like the geometric
parameters, the flow and water surface conditions in the surface aeration system:
OTR=10-3(DCs Dm) Re1.90 Fr0.15 (D/h)0.2 (S/h) (n/n8)0.20 (Tv/D)0.05
(1.19)
Where n is the number of turbine blades; Dm is the oxygen diffusivity coefficient in

water at 20C; and Cs is saturation concentration, the specification for the tested
surface aerator of this model are illustrated in Figure 1.14.
- 25 -

D.1. Surface Aerator Diameter
Surface aerator impeller diameter is usually related to the tank dimension, for
example a minimum turbine diameter should be corresponded to the tank diameter
(i.e. for circular shape tanks, Tv/5) and to the depth of the tank to minimize the risk of
dead zone existence and that occurs more in the lower part of the treatment tank. It is
important for radial surface aerators diameter to keep the D/Tv ratio in a particular
range in order to keep the energy consumption within accepted range (Oldshue, 1983;
Tatterson, 1994; Treybal, 1980; Xuereb et al., 2006).
(Patwardhan and Joshi, 1998) figured out after reviewing many papers of various
surface aeration processes of air entrainment from free surface that the surface
aeration intensity, the oxygen mass transfer coefficient and suspension ability of
biomass are increased with increasing the impeller diameter due to increasing surface
turbulence and high velocity performed inside the liquid bulk. They tested the effect
of impeller diameter on surface aeration efficiency for both laboratory and industrial
scales.
The effect of the aerator diameter has been tested by (Patil et al., 2004), the ratio of
(P/V) (consumed power per liquid volume) had considered as distinguishing factor for
different aerator impeller diameters, where for (P/V < 100 W/m3) the impeller
diameter that equals Tv/5 exhibit in some kind more higher mass transfer coefficient
but for range (P/V > 150 W/m3) the impeller diameter that equals Tv/3 showed higher
values of mass transfer coefficient.
IV. Temperature Effect

Since the solubility of the oxygen in the water is highly temperature dependent factor.
The temperature effect is very limiting factor when the calculations are made for the
oxygen mass transfer to the water in the surface aeration. The calculated oxygen
transfer coefficient should always be corrected to standard temperature condition to
allow a meaningful understanding and assessment of this coefficient which in turn
helps to characterize the surface aeration efficiency of the concerned process.
The oxygen transfer rate can increase with temperature increasing, as the viscosity of
biomass in treatment tank is reduced with temperature elevation so the surface
mobility increased as proposed by (Cumby, 1987a).
The temperature has contradictory effects on oxygen mass transfer, where the mass
transfer coefficient is decreased with temperature increasing in the droplets while it
slightly increased with increasing the temperature in liquid bulk as founded by (Chern
and Yang, 2004) in their investigation for the performance of aeration system by
spraying the water droplets through the atmospheric air and then re-impinging with
the surface liquid bulk. They studied the distribution of dissolved oxygen
concentration in the droplets and inside liquid bulk and the turbulent flow created by
- 26 -

the impinging of droplets with the surface of liquid. Beside the different effects on the
system liquid like the height of droplets, recirculation rates of droplets and liquid
depth; (Chern and Yang, 2004) examined the effect of air condition, like temperature
and humidity on the oxygen transfer. (Chern and Yang, 2004) derived a general
model of oxygen mass transfer coefficient in the water droplets including the
temperature effect. They verified the results experimentally and found the equation
(1.20) can predict the behavior of the system with identical conditions used into
experimental work:
(1.20)
The mass transfer operation in their work is considered as two divided zones as it
proposed by (McWhirter et al., 1995). Where (kla)s is the volumetric mass transfer
coefficient in the surface mass transfer zone; is the temperature correction factor;
and Tw is the bulk liquid temperature, k3-k6 are equation constants.
(Wichterle, 1994) figured out that the heat transfer is scale dependent, this effect is
more clearly noticed with surface aeration process when is compared with other
operation on the basis of same rotational speed is used especially for smaller
tanks.(Wichterle, 1994) has developed a generalized correlation for surface aeration
temperature effect by depending on the laminar boundary layer concepts. He related
the heat transfer process at wall tanks with other dimensionless numbers Reynolds
number, Prandtl number and other operational parameters by depending on with the
flow velocity and pattern for aerated agitated baffled tank.
Nu =A Re Pr 1/3 Vi1/4
(1.21)
Where, A is a constant that depends on tank geometry; Nu is Nusselt No.; Pr is

Prandtl No.; Vi is the viscosity ratio (bulk/wall) and is the heat transfer coefficient.
The model was developed within the ranges A (0.3-1.2) and Re (250-235103).
V. Hydrodynamics
V.1. Flow Patterns
A. Flow Patterns Characterization
The flow pattern and circulation inside of surface aeration in the wastewater tank has
a very crucial importance to ensure distribution and efficient mixing for the dissolved
oxygen, the homogenization and suspension of the mixed liquor that contains the
activated sludge (Roustan et al., 1984).
Most of the agitation and mixing means that used for the surface aeration are
operating with including impellers or turbines that are placed on a turning shaft. There
are many of these impellers that can be characterized according to the generated
pattern flow. These impellers accomplish two main flow types. The axial flow
- 27 -

promotes a circulation toward down and top of the tank creating self-looping patterns
of circulation. The propeller is the most known device used in this type of operation
(Leng et al., 2008). For the radial flow, there are many types of impellers that all have
the characteristics of pushing the water in the radial form in perpendicular way to the
rotation shaft. The turbines are the principle agitators (impeller) that are used for
radial movement turbines in general imply a horizontal disc or cone on which the
paddles are fixed but there are numerous forms of modified turbines (Leng et al.,
2008; Nagata, 1975; Tatterson, 1994; Xuereb et al., 2006).
For the surface aeration systems, the axial flow may be produces in up-ward or downward directions by the impeller and then the flow will be redirected to the impeller
intake region by the tank walls in closed circulation profile.
In general the discharge flow that moves and hits the vessel walls, entrains liquids
from other parts outside its circulation field. The axial flow may invert to radial flow
during its general axial flow.
The generated radial flow by impellers blades moves always toward the vessel wall
and then the it is redirected by the wall for both upper and lower directions to goes
back toward the impeller (Nagata, 1975).
In the turbulent regime the flow patterns of various agitator speeds usually are
considered similar for each type of impeller (radial or axial). It is assumed that the
absolute flow velocity increases in proportion to the impeller speed, the flow pattern
is very important as a function of various geometrical and operational parameters such
as impeller diameter, tank diameter and Reynolds number (Tatterson, 1991).
The flow field in the aerated wastewater tank is dominated by the energy sources of
aeration and mixers, where the flow development in aerated tank with the presence of
bubbles is controlled by the buoyancy effects (Gresch et al., 2011).
The flow pattern of the water droplets impingement at the water surface plane was
examined by (Okawa et al., 2008). They noticed the generated liquid column on the
water surface after the water droplets collision is depending on the angle and
conditions of collision were convenient. When the angle wasnt too large, the
breakup of the liquid column led to the production of secondary drops. The number of
secondary drops depends on different factors like droplet diameter, the surface
tension, density, viscosity and impact velocity, where they are referred by
dimensionless parameters that called (k-number) that may control the number and size
of secondary drops.
(McWhirter et al., 1995) illustrated that the flow parameters of surface aerators with
water projection include an important modification that due to a new occurring mass
transfer zones concept. They found out the most important factor in the process is the
total liquid discharge velocity that refers to all the water droplets that propelled by the
turbine, which is very dependable on the geometry of the turbine. The discharge
- 28 -

velocity contains three component, two of them are constant the radial VR and
tangential V but the third one the vertical axial velocity Voy is changing with time
and it has a maximum value at a certain point and after this point it decreases which is
the first velocity that gets out at the upper tip of blades.
B. Vortex Formation
The depth and the form of formed vortex has the predominate effectiveness on the
process of un-baffled surface aerator. (Rao et al., 2009) built a general correlation for
specific configuration with emphasizing on the vortex forms. In order to correlate the
relevant parameters of the process with depending on previous work of (McWhirter et
al., 1995) to describe the formation and performance of vortex especially, the role of
critical speed of the impeller and other effective parameters in the system.
(Rao et al., 2009) performed experimental runs to deduce the mass transfer
coefficient related with vortex formation and they found that the mass transfer
coefficient at standard condition increase sharply near or above the critical speed. The
general modified correlation for scale-up purposes was:
hv / D = 43.2 Fr(0.1 Ga0.18)(h Cu W/L)0.16
(1.22)
Where hv is the depth of the vortex; Cu is the turbine blade clearance (in this case it
represents the distance between the horizontal bottom of the tank and the top of the
blades); L represents blades length; Ga = Re2/Fr. The model was developed within the
ranges hv / D (0.003-0.2) and Fr (0.005-0.02).
V.2. Air Bubble Size Distribution and Hold-up

When suitable impeller diameter and rotational speed are employed, with proper
bubble size distribution and residence time an improved oxygen transfer coefficient
was achieved as founded by (Lee et al., 2001). They studied the bubble distribution
inside a surface aeration for water treatment tank with turnover conical turbine, as
shown in the Figure (1.16) the air bubble distribution delivered by the axial impeller
blade rotation is homogeneously distributed in the entire tank and reaches the bottom.
While (Barigou and Greaves, 1992) have performed various tests on the aerated tank
to measure local bubble size distribution with changing the position. They found out
that there is a variation in the bubble size distribution within the tank and this
variation was related to other parameters variations like the rotational speed and gas
flow inside liquid bulk. The size of bubbles was lower at high rotation speed (these
results agree with those performed by (Kawecki et al., 1967), the location of bubbles,
they observed that bubbles size was changed especially between the vertical plane
coinciding with baffles and the mid-plane between two baffles location.
- 31 -
(a) D/Tv =1/3
(b) D/Tv =1/5
Figure (1.16): Bubble distribution of surface aerator system of rotation speed N=110
rpm, liquid level h=0.66 m. (Lee et al., 2001).
(Deshmukh and Joshi, 2006) pointed out that the gas holdup plays an important role
on the velocity field in the impeller region for the surface aeration system, which in
turn is affected by the impeller design, submergence and rotational speed of the
impeller.
These observations were noticed by (Sun et al., 2006). They found that the air bubbles
holdup profile was non-uniform in the surface aerated agitated tank, the bubbles
holdup was high in two regions, first close the liquid surface and second in the
impeller region. The weakest gas holdup region was under the impeller region. The
air bubbles holdup was measured at different operation conditions; the impeller was
Rushton disc turbine with surface baffle used.
V.3. Mixing Time

A. Mixing Time Characterization
The mixing time is generally considered as an indicating factor for perfect
accomplished mixing potentials in surface aeration, where it is necessary to achieve
the homogenization and effective mixing of dissolved oxygen inside the treatment
tank with maintaining the desired operation conditions for suspension and distribution
of activated sludge and biomass particles. Mostly the mixing time in aerated
conditions is less than non-aerated conditions, (Guillard and Trgrdh, 2003) reached
to same persuasion by examining the mixing time for both aerated and non-aerated
condition for disc turbine impeller. The mixing time usually referred as dimensionless
number called mixing number (tmN), which is very dependent on the turbulence flow
regime, where is always has lower values with turbulent flow than other flow for gasliquid systems (Hadjiev et al., 2006).
- 30 -

The mixing time is always decreased with increasing the impeller rotational speed as
seen in the Fig. 1.17, when the system consists in two impellers with the upper
clearance is set at the liquid surface as figured out by (Kang et al., 2001) in their
studies the influence of rotation speed effect on the mixing time for surface aerator
system, where they tested the effect of rotation speed on the mixing time for different
geometric configurations.
Figure (1.17): The relation between the rotation speed and mixing time for surface
aeration dual impeller system, (Kang et al., 2001).
(Guillard and Trgrdh, 2003) found out the general approach is difficultly applicable
in actual sized tanks and they deduced that the mixing time under aerated condition is
longer than non-aerated and depends on the tracer injection position also. They used
tracers by pouring pulses of concentrated acid; the injection was in three different
positions in the tank then the concentration gradient determined by PH electrode.
B. Mixing Time Modeling

(Hadjiev et al., 2006) developed a model to correlate the dimensionless mixing
number with the most important geometrical parameters and liquid flow and water
surface condition for the aeration systems:
Ntm=[-13.981(S/D)2+27.972(S/D )-4.1327] (Sp/D)-0.54 Re 0.275 Fr0.275 Fg-0.04
(1.23)
Where, S is the upper impeller submergence, and Fg, is the modified aeration number,
which equals G/(Nd3nim). The model was developed within the ranges Sp/D =2.12, Re
(2.0103-14103), D/Tv= 0.16 and h/Tv=3.5.
For micro- mixing in surface aeration process, (Rao and Kumar, 2009) derived a
model for mixing time by depending on the basis of theoretical power per unit volume
to simulate the process time with X = Fr4/3Re1/3,the ranges for the model are same in
equations 1.14 and 1.15.
(1.24)
- 31 -

V.4. Air Bubbles Entrainment
The flow inside the water treatment tank is produced indirectly by the turbine rotation,
the splashed water droplets are plunged in the water surface in tank. These plunged
water droplets entrain the atmospheric air as bubbles and go downward inside the
tank. Air bubbles are formed at the water surface when the turbulent kinetic energy is
strong enough to break the stabilizing effects of surface tension and
gravity(Bhattacharya et al., 2007). (Chanson, 2003) defined the air entrainment by the
water spray plunging as the entrainment or entrapment of un-dissolved air bubbles
and air pockets that are carried away within the flowing fluid.
(Bi, 1993; Ohkawa et al., 1986; Qu et al., 2011) have reached the persuasion that the
penetration depth of the entrained air bubbles is highly controlled by the water spray
velocity or flowrate in depending on combined experimental and numerical study of
the flow. (Kusabiraki et al., 1990) have added the water velocity distribution at the
impact position as effecting factor on the air bubbles entrainment as a results of their
experimental studies on the gas entrainment rate in a jet aeration system. The
importance of the water spray velocity is explained by (Ohkawa et al., 1986) is to
overcome the buoyancy forces of the air bubbles produced by the water spray jet.
Usually an impeller is coupled with surface aerator turbine in order to produce
significant shearing stresses in the mixing zone which enables air bubbles entrainment
from the surface (Heim et al., 1995).
V.5. Surface Aeration Power Consumption

In the surface aeration system, with the occurred agitation in the treatment tank it is
very indispensible to identify the necessary energy needed in the process. The power
or energy consumption in the surface aeration is the energy that is delivered from the
shaft to the surface aerator turbine and then to the water. Or it is simply can be
defined as the energy that leaving the impeller and entering the liquid in the tank.
Because of many variables can affect the energy consumption, identifying a general
characterization for the power consumption in surface aeration is so complicated,
where the power consumption here may influenced by the whole system variables of
the surface aeration, the chemical and physical properties of water and included solid
contents and the operational conditions like the rotation speed, gravity and water level
in the tank and other parameters.
There are many methods to measure the power consumption in the agitated tanks such
as the widely used electrical and torque-meters gauges are used for large scale tanks
(Ascanio et al., 2004) considered the torque-meter as an accurate method.
The power consumption is related of surface aerator and tank geometries and
operation conditions. It is quite clear that it will be influenced when different number
of impeller used, but it is not always dependent on these impellers equally, the
- 32 -

position and the bubbles flow profile around each impeller has its different effect (Cui
et al., 1996).
A. Operational Condition Effect

(Takase et al., 1982) have investigated the power consumption of surface aeration
system with aerator impeller (6 blade disc turbine, DT) positioned at the water surface
for different conditions of treated liquid with clean water and with two different
activated sludge mixed liquids, they found out the different conditions didnt consume
similar energy with surface aeration process, but their impact on the power number
were almost same for same operation condition and experimental equipment (See
Figure 1.18).
Figure (1.18): The relation between the Reynolds number and power number for three
different liquid condition ns, clean water and two types of activated sludge mixed
liquid (Takase et al., 1982).
B. Power Consumption Relation with Oxygen Mass Transfer

When comparing two surface aerators with same diameter and rotational speed, the
one that consume more power has higher oxygen transfer efficiency (Ognean, 1997).
For each surface aeration process according the power consumption there will be
three characteristic affecting parameters: the active volume, that identified by the
volume created by the aerator within the energy transferred to the liquid, the energy
transfer area and the mass transfer area; for certain operation condition and active
volume an optimum oxygen mass transfer can be achieved as supposed by (Ognean,
1993b).
(Ognean, 1993b) built a model that correlates the power consumption for the surface
aeration and the relevant dimensionless parameters with the aeration modified
sorption number YN at the maximum oxygen transfer condition. This model can be
- 33 -

applied to compare the different surface aerator types, where subscript m represents
maximum condition; G is the mass-transfer rate (oxygen uptake).
(YN)m = (OTR/P)m (1/Pm1/2)(D3N3/2)m(g5/63/2 1/3 c-1)
(1.25)
(1.26)
( )
Where de is the geometric equivalent dimension (
), c is dissolved
oxygen concentration in the water (mg/l).

(Ognean, 1993b) proposed models of dimensionless numbers for various types of
surface aerators by correlating the oxygen transfer efficiency with power consumption
and other affecting parameters like geometric ratios of the system. For an optimum
condition the maximum values for these parameters are correlated in the model. For
vertical shaft aerator, the maximum aerator diameter used was always considered as
relevant factor when trying to reach maximum aeration efficiency (AE)m.
(AE)m =K1 P1/2m (D3N3/2)m
(1.27)
K1 is the equation constant

While for horizontal shaft aerator he proposed another model, where (Ognean, 1993a)
considered dc as a characteristic length of the horizontal surface aerator and depends
on the contact area between the aerator and the water.
(AE)m = K2 P1/2 m (N3/2 D3/2 dc 3/2)m
(1.28)
K2 is the equation constant. The models 1.25 to 1.28 were developed within the ranges
D / Tv =0.125, h/D= 3.0, Np (1.0-3.0) and Fr (0.05-4.0).
VI. Contact Time between Water Droplets and Atmospheric Air

(Cancino et al., 2004) modified a model for the water spray droplets contact time with
the atmospheric air in the surface aeration process within the development of
theoretical design of the axial flow surface aerator. The model relates the time, te, with
the height of water droplets, Hd, in air and gravitational acceleration:
te = 8Hd/g
(1.29)
(Cancino et al., 2004) deduced that it is required to increase the contact time of water
droplets in the air to enhance the mass transfer operation and this can be done in the
way of increasing the height of water spray that projected or by increasing the
dispersion of the water droplets (decrease the size of the droplets). (Cancino et al.,
2004) demonstrated that to achieve these objectives it is required to increase also the
projected water flow rate to power consumption ratio (Q/P). (Cancino, 2004a)
- 34 -

declared that according to design consideration of surface aerator turbine it should be
made with considering that largest amount of water should be projected with largest
distance for droplets flight pattern should match the contact time within the limit of
lowest power consumed with regard to the fact that inlet and outlet angles of water for
the blades were well chosen.
VII. Environmental Effects
The performance of projected water and wastewater droplets as a spray in the
atmospheric air and the concerning process can produce undesirable emitted aerosols
causing air pollution, further it can be a source of unwanted odors (Cumby, 1987b).
(Chern and Chou, 1999) have studied the surface aeration but in different way, they
covered in their study the other sides of mass transfer of volatile organic compounds
that stripped from waste water to the air during the treatment operation, they divided
the mass transfer area into two zone for both continuous and batch aerations in order
of developing a general model to estimate the emission mass transfer for (VOC) in the
air, for the liquid spray zone for VOC emission rate the following equation was
derived:
(2.30)
For the surface re-aeration zone the emission rate was:
(
(2.31)
Where VOCER and VOCERs represent VOC emission rates; Cvoc is the dissolved
VOC concentration in the bulk liquid; Hc and Hcs are Henrys law constant of VOC at
water and air wet-bulb temperatures respectively; and CG represents VOC
concentration in the air.
Water and wastewater treatment plants that implement surface aeration are well
known as sources for noise pollution, so it is recommended to build these plants away
from urban zones.
1.3. Conclusions
In this chapter a brief presentation was made for the present day aeration systems for
the water and wastewater treatment with introducing the types and efficiencies for
each system. The surface aeration for water treatment was described somehow in
detail.
It is found that; many modification essays are made on the surface aerators to improve
the performance. There are many trials have performed to enhance the operation
efficiency by either of increasing the mass transfer rate or by reducing the power
- 35 -

consumed in the operation. In order to achieve these objectives there are many studies
are made concerning with the various operation variables like hydrodynamic of the
system (gas hold-up, bubble size distribution), the flow pattern for gas and water, the
properties of working mediums (air and water), rotational speed of the aerators, the
circulation time (mixing and aeration time), power consumption, geometric
configurations (i.e. number of impellers, types of impellers, geometric ratios), oxygen
mass transfer rate in the water and scale up, modeling process for the aeration system.
Form the literature review it can be presumed that in general the limiting factors for
successful surface aeration can be resumed as; the turbulent regime should be ensured
in the entire water treatment tank that contains sufficient dissolved oxygen entrained
from atmospheric air. The generated water flow by axial or radial surface aerators
must be sufficient to reach all parts of the treatment tank. The aeration process
accomplished the axial or radial impellers must be accompanied by mixing and
agitation performance to ensure an efficient distribution of the dissolved oxygen,
these constrains can be overcame by an appropriate surface aerator with suitable
pumping capacity to handle efficiently and effectively the large quantities of water.
- 36 -
- 37 -
- 49 -
- 50 -
Chapter Two: Experimental Setup and Calculation Methods
Chapter Two
Experimental Setup and Calculation Methods
2.1. Introduction
The experimental design and setup for surface aeration runs depend on many
important influencing factors, that generally can be classified as the geometric
configuration of the system, the operation conditions, the theoretical assumption that
describes the occurred process during the experimentation such as the models of the
transferred oxygen mass in water, the flow pattern of the air-water flow inside the
tank and at water surface, the physical and chemical properties for both water and air
used. The economic factor is an important factor that should be taken into account.
The experimental work was performed to interpret the important aims: the energy or
power consumed by the surface aeration system, the oxygen mass transfer coefficient
in water bulk kla and in the water spray (droplets) klad, the impeller configurations and
operational influences on the oxygen mass transfer and power consumption. Where,
the experimental runs can be classified into two main categories, the mass transfer and
hydrodynamics.
2.2. Experimental Installation and System Description

The surface aeration system was built in the laboratory and is the perfect scale-down
representation of the industrial system (MOS, Biotrade)(FR Patent Demand, 2012);
the surface aeration and mixing are achieved by employing the turbine (conical
overturned shape with 15o pitched blades) placed at the water surface with mixing
assembly positioned lower the turbine (See Figure 2.1). The turbine body is made of
PVC; while their blades are made of stainless steel adhered to the turbine body. The
mixing assembly is composed of reversible twisted pitched RTP propeller of four
blades with 45o pitch angle at the propeller hub (Milton Roy Mixing HPM204D) fixed
inside a draft tube (See Figure 2.2) to enhance the performance of the system. Three
baffles are fixed inside the draft tube (See Table 2.1).
The experimental runs are carried out in a cylindrical flat bottom vessel with
dimensions of 600 mm height and 800 mm in-diameter placed in a larger cubic vessel.
All the vessels are made of fibre glass (Perspex) except the front face of cubical
vessel, which is made of glass for easy cleaning and to eliminate causing scratches
during the cleaning and in order to avoid the laser beam diffraction, for the crossing
laser beams for (Laser Doppler Velocimetry) LDV and (Particle Image Velocimetry)
PIV application. The reason for the second vessel is to provide a compatible medium
for the laser beams in both inside and outside the pilot plant. The schematic diagram
of the system for both aeration and mixing is shown in figure 2.3 a, b.
- 15 -
Figure (2.1): The schematic diagram of the experimental apparatus.
Figure (2.2): The turbine, propeller and draft tube.

Three baffles were used in order to prevent or lessen the tangential circulatory flow
created by the system, (the baffles have the same height of the vessel see Table 2.1).
A cone is positioned at the vessel flat bottom exactly beneath the draft tube to
improve the inlet flow by preventing secondary eddies formation and dead zones
under the propeller. As it is noticed the geometrical ratio for the propeller (dpr/Tv) =
0.15, where that is lower than the general applied ratio (dpr/Tv) for axial flow
impellers (0.2 - 0.7) in classical agitated tanks.
- 15 -

Table (2.1): Geometrical configuration details
Vessel
Turbine
Draft tube
Propeller
Cone
Baffles
Draft tube
baffles
Impellers
spacing
Water
height
Diameter
(m)
0.8 (Tv)
0.19 (D)
0.15 (df)
0.12 (dpr)
0.15
Clearance
Tv/3.13
Tv/10
Tv/5.13
Height
(m)
0.6
Width
(m)
Blade width
(m)
No .of
blades
0.024
12
0.018
Numbers
0.1
0.06
0.6
Tv/10
0.06
df /8
0.076
0.28
Figure (2.3): Schematic diagram of experimental apparatus; (a) Up pumping flow

(Aeration mode), (b) Down pumping flow (Mixing mode).
- 15 -

2.3. The Measurement of Power Consumption
The consumed power by the surface aeration system was calculated by torque
measurements, which is applied for both the surface aerator turbine and lower
propeller. This measurement is achieved by two steps. First, the torque is measured in
empty vessel to determine the frictional mechanical loss due to air resistance during
the impellers and shaft rotation. Second the torque is measured with the rotation in
filled vessel with testing liquid (water). The actual consumed power is the difference
between the calculated powers in both filled and empty vessel by applying the
following equation:
P = 2N (To-Toe)
(2.1)
Where, To, and Toe, are the measured torques in filled and empty vessel respectively
in (Nm). The torque meter used with a torque capture transducer all are made by
(HBM). A motor (LEROY SOMER), (LS90SL, rated power 1.1 kW, max. rated speed
1420 rpm) utilized to achieve the desired rotation speed.
2.4. Hydrodynamics and Mean Velocity Measurements Techniques

In order to determine the velocity profiles and the occurred flow pattern inside the
tank, the mean values of the different velocities are measured by both the LDV (Laser
Doppler Velocimetry) and PIV (Particle Image Velocimetry). With LDV technique it
is the yield from set of velocity measurements for one or more of velocity vector
components on the same time (Mavros et al., 1998), while the PIV is used to study of
the instantaneous flow field for both single phase and multiphase (Aubin et al., 2004).
2.4.1. Laser Doppler Velocimetry (LDV)

I. LDV Apparatus Description
The LDV (Laser Doppler Velocimetry) is non-intrusive technique. The laser source
consists of two beams (514.5 mm green- 488 mm blue wavelengths) Dantec fibre
flow device. The measurement of the occurred velocity depends on the reflected
scattering beams by added tracer particles to the water. The light laser source is a 4 W
Argon ion of laser stability 2017 (Spectra Physics) with a focal length of 600 mm
(See Figure 2.4). The measurement that performed by the LDV is usually made in
regular plans inside the fluid bulk. The LDV measurements are carried out with a
Dantec fiber flow system is operating in back-scattering mode. The Doppler signal is
transmitted to a 58N20 flow velocity analyser (Dantec) for the signal processing. The
cooling is achieved by water medium that is circulated with pumping pressure of (5
bar). The employed software is BSA flow (version 4.5).
- 15 -
(4)
Figure (2.4): The Laser Doppler Velocimetry (LDV) testing apparatus, (1) Laser
source. (2) Traverse system, (3) Tested tank, (4) Flow velocity analyser.

In order to accomplish flow measurements with LDV the vessel is filled with
operating fluid (in our case is tap water) and seeded with small amount of tracer
particles of Iriodin 111 Rutile stain particles (Merck) ; the diameter of these particles,
dp is about 15 m.
III. Measurement Principals
The device operating concept depends on the reflected scattering beams by previously
added tracer particles to the water. The LDV device is equipped with a traverse
system for the purpose of moving the laser source to have the necessary flexibility of
reaching all the parts inside the tank easily; the coordinates of the traverse system
matches the actual lab coordinates.
Two laser beams are crossed and formed an interference fringe pattern, where this
area of interference is called the volume of measurement. This volume is consisting of
equally spaced interference fringe planes parallel to the bisector of the angle between
the two laser beams (See Figure 2.5). The model that indicates the set of parallel
fringe planes is equally spaced by a certain distance.
- 11 -

The equally distance between the fringes y, can be determined by the flowing
equation (Costes and Couderc, 1988):
y = o / 2 Sin (/2)
(2.2)
Where: o is the laser beam wave length.

When the tracer particles move through the region of interference fringes, they scatter
the light of the two laser beams, with modulated intensity corresponds to its
movement through the fringes. The phenomena is called the Doppler effect, where the
basic principle of the effect is the scattering of the two light beams by the particles,
which is propagating in a direction represented by a unit vectors with certain
frequency and wavelength to form an aggregate beam with a frequency different to
the two incident beams, the diffused light by the particles has a frequency relative to
the first light beams. The frequency of the modulated intensity at which the light is
scattered is called the Doppler frequency that is directly related to the components of
the particle velocity perpendicular to the interference fringes .
Figure (2.5): The LDV measuring volume fringe planes.
The quality of the LDV results depends on the size of the ellipsoidal measuring
volume formed at the beam intersection point and the number of the interference
fringes in this volume. The size is refined as much as possible in order to avoid the
presence of too many particles with different velocities and thus a random out of
phase signal. At the same time the number of interference fringes must kept
maximized in order to maximize the precision of the frequency of the signal (Aubin,
2001).
The dimensions of the measuring volume can be determined as (See Fig. 2.6):
- 15 -

2av = dp / sin(/2)
(2.3)
2bv = dp / cos(/2)
(2.4)
2cv = dp
(2.5)
Figure (2.6): The measuring volume.
The two velocity components vr and vz, are measured by changing the position of the
laser beams in the vessel. To measure the velocity component, the laser beams was
oriented such that they are in the same plane as the component and the bisector of the
angle formed by the two beams is perpendicular to it. Aligning the beams in a
horizontal plane allows the radial velocity component, vr to be measured and when in
a vertical plane the axial velocity component, vz can be determined. So, the measured
velocity components by the LDV laser beams are the axial and radial components
only. The other component of the flow velocity that the tangential is not measured,
especially the measurement concerned area applied by the LDV is the propeller
vicinity and the draft tube inlets, where the majority of occurred velocities are the
axial and the radial, where the laser beams should move in a direction perpendicular
to the optical axis (See Fig. 2.7).
IV. Signal Post-Processing
A Dantec burst spectrum analyzer (BSA) for time-resolved measurements is used as
the LDV signal processing, where the signal from the photomultiplier that contained
in the apparatus was filtered to remove the undesired low and high frequencies then
amplified by a variable gain. The number of samples required to achieve statistically
independent results depends on the flow field. It was found that at least 300 samples
- 15 -

were convenient for local turbulence intensity occurred in the testing area. The BSA
was also interfaced to a PC controlled by software installed in the PC. The BSA was
operated in Burst mode, that is, only one measurement was performed per detected
burst, where the frequency data are transformed to velocity data by the used software
(Lee and Yianneskis, 1998).
Figure (2.7): The LDV laser beams positions.
2.4.2. Particle Image Velocimetry (PIV):

I. Theory
The PIV technique is now widely used in the agitated tank flow field acquisitions
because of its convenience to obtain the full characteristics for the measured field (Li
et al., 2011). The PIV is applied to measure the instantaneous velocity field using the
images of tracer particles in the fluid flow (See Figure 2.8), where the PIV technique
provides a quantitative, instantaneous, whole-field visualisation and two dimensional
description of the flow (Udrea et al., 1997). The displacement of tracer particles is
measured through the analysis of the images that produced during the illumination of
these particles by pulsed sheets of light at precise time intervals and recorded on film
or a video camera array (Adrian, 1991). The average displacement is measured when
high image density of large particle concentration is applied (Keane and Adriane,
1992). The usual used PIV type is the two dimensional, where it can be applied to
acquire only both the radial and axial flow velocities components in the stirred tanks,
where the laser sheet is positioned vertically inside the vessel.
- 15 -
Figure (2.8): Schematic diagram of the PIV

The PIV experimental setup is generally consist of the addition of tracer particles to
the investigated flow stream. Tracer particles are small enough to ensure a good
tracking of fluid motion due to the difference in density between the fluid and the
tracer particles and in same time to ensure not they will not interact with the flow
being measured, they can frequently to be the order of 50-100 nm, their diameters are
then (1/10) to (1/5) the wavelength of green, green=532 nm. Seeding of the tracer
particles is made in order to achieve sufficient image contrast, where seeding can
easily be done by suspending solid particles into the fluid and mixing them to ensure a
homogenous distribution.
III. PIV Principles

After the addition of tracer particles, these particles are illuminated for detected plane
within the target flow for several times during short time interval. The displacement
of particles images between the light pulses determined by the evaluation of the
recording for the light scattered by these particles, which is recorded by high quality
lens of CDD cameras, where a post-processing then is applied to deal with these
collected data (Raffel et al., 2007). The main parts of PIV system are illustrated in
Figure 2.9.
IV. Scattered Light

The light scattered by the small tracer particles is a function of the ratio of the
refractive index of the particles to that of the surrounding medium, the particles size,
and their shape orientation, polarization and observation angle (Raffel et al., 2007).
- 15 -
Figure (2.9): The Particle Image Velocimetry (PIV) testing apparatus, (1) Laser
source, (2) Recording camera, (3) Tested tank.
The light is not blocked by the small particles but it spread in all directions, so
massive multi-scattering occurs when sufficient number of tracer particles exists
inside the light sheet, where the light scattered by more than one particle is imaged so
the recorded light by the lens is not only because of the direct illumination but also
because of the fraction of the light (See Figure 2.10).
Figure (2.10): Light scattering by the (10 m) glass particles in the water (Raffel et
al., 2007).
V. Laser Source
The laser is used in the PIV process is related to their ability to emit monochromatic
light with high energy density, where it easily can be bundled into light sheet for
illuminating and recording the tracer particles.
- 56 -

The laser device in the PIV consist of three main components: (a) the laser material,
which it in turn consists of an atomic or molecular gas, (b) The pump source; that
excites the laser material by the introduction of electro-magnetic or chemical energy,
(c) The mirror arrangement; i.e. the resonator allows an oscillation within the laser
materials (Keane and Adriane, 1992).
The used laser beam is usually Argon-ion laser (= 514nm, 488nm), which is a laser
gas, where this type of laser is characterized by the very high current is achieved for
the ionization process. For this type the laser mirrors that used is produce an emission
at several wavelengths, where specific wavelength can be selected in the laser
resonator, in such a manner that the individual wavelengths can be adjusted by
rotating the prism in the laser mirror. In general this type is used for the liquid fluid
mechanics.
The light sources are also used in the PIV flow investigation with laser sheet. For the
spectral output of the light source may suit for specific cameras. For the similarity in
the spectral sensitivity and that offer a repetition rate that matches the video rate.
Optical fiber bundles are used to link the light sources in order to achieve short pulse
separation times, where the generated light is simplified when the outputs of the fibers
are arranged inline. Generally three types of the lens are used to generate light source,
in which can be set in different combination of the various shapes used. The fiber
bundles can be used for the combination of two sources for shorter pulse separation
time (Hecht and Zajac, 2001).
VI. Recording Techniques

For each recorded frame of PIV, it represents a freeze of the flow in time. The
successive images average time can be applied to build flow maps that characterize
the impeller and the vessel configuration. For each frame, it can be considered as a
transient image of the flow, which exhibits the non-stationary phenomena, like flow
instabilities (Mavros, 2001).
In the PIV, the recording methods are achieved by either photographic or digital
recording. The recording process can be implemented in two ways: first the capture of
the illuminated flow with a single frame and then provide a single illuminated image
for each illumination pulse (single frame / single-exposure PIV) or for several
illumination pulses (single frame / multi-exposure PIV), the second way is the capture
of the illuminated flow on a multi frames and then provides a multi illuminated image
for each illumination pulse (multi-frame / single-exposure PIV). The (single frame /
multi-exposure PIV) recording was first used in conjunction with photography, as it
doesnt retain information on their temporal order of the illumination pulse, which
cause a directional ambiguity like what is called image shifting , while the second
way of recording preserves this temporal order of the particle motion(Raffel et al.,
2007).
- 55 -

The digital CDD (Charge coupled device) cameras in PIV recording are used because
they provide high repetition rates that needed for high pulse energy lasers. In CDD
cameras the individual pixels are grouped into a rectangular array to form a light
sensitive area, in which the array should to be read out sequentially in a two-step
process, that the accumulated charge shift vertically one raw at a time after exposing
CDD sensor into a masked off analog shift register on the lower edge of the sensors
active area.
VII. Image Analysis Method

In order to extract the needed information for target displacement from the recording,
the interrogation methods are applied on the selected images to tracking the individual
particles from exposure to exposure (Ag and Jimenez, 1987). The process is
implemented by image processing computer software. The processing step of the
images is depend on the dividing the complete image into the squared interrogation
areas (Escudie and Line, 2003).
The image evaluation by the experimental data with the results of numerical
calculation requires high density images or medium concentration of the images of the
tracer particles. The experimental applications for the image analysis are dependent
on the type of these experiments. For example for the air-liquid system, gas bubble
distribution, a careful arrangement of camera and light sources are needed to avoid the
over exposure of bubbles and the optical distortions from vessel wall (Laakkonen et
al., 2005).
VIII. The Cross-Correlation Calculation

The evaluation of the PIV recordings is made by the cross-correlating for two frames
of single exposure of the tracer particles. For each interrogation area with the crosscorrelation application, the calculation for most of the probable tracer displacement is
attained. With choosing the adequate time interval between the two successive
images, the instantaneous flow field inside the vessel can be calculated (Escudie and
Line, 2003).
(Raffel et al., 2007) demonstrated that the cross-correlation calculation can be
represented as assuming the displacement is constant Ddsp (See eq. 2.7) of all the
particles inside the interrogation volume, so the particle location during second
exposure during second exposure at time (t= t+t) is:
Xi = Xi + Ddsp
(2.7)
The particle image displacement is represented as:
- 55 -

(
(2.8)
Where M: is the magnification factor. The image intensity field for the time of the
second exposure:
N
I (x,) = Vo (Xj + Ddsp)(X-Xj-ddsp)

j=1
(2.9)
Where, Vo: is the volume of the interrogation during second exposure. So the crosscorrelation function of the two interrogation areas will be:
RII (s, , Ddsp)= 1/aI Vo (Xi) Vo(Xj+ Ddsp) aI (X-Xi) (X- Xj + S - ddsp)dX
(2.10)
The comparison between the PIV and LDV techniques showed good agreement
between them for the quantitative comparison of the radial and axial velocities (Myers
et al., 1997).
2.4.3. Other Flow Related Measurements Parameters

The mean velocity measurements that performed by the LDV and PIV are used to
calculate the flow rate or the pumping rate at the turbine and propeller vicinity or in
the entire vessel area. The calculated flow rate is necessary for example to compute
the pumping number and other flow related factors.
I. Mixing Time (tm)
The mixing time tm, experiments were performed for the system in aerated and nonaerated modes.
In aeration mode, the tested configurations consist in the turbine with RTP propeller
and the draft tube, the turbine, RTP propeller and the draft tube and for the turbine
alone in order to determine the mixing performance of each configuration and
elucidate its influence on the mixing time.
The mixing time in non-aerated vessel (the turbine is removed) was investigated for
both up and down pumping modes.
The injection position was kept for same position at the water surface for all tests to
eliminate the effect of injection variation on the results obtained.
- 55 -

The mixing time was determined by applying the colorization-decolourization
method. It consists in colouring the water with 10 ml iodine solution and then in
decolorizing it with 10 ml sodium thiosulfate solution with a slightly excessed volume
to insure completely decolorizing the iodine solution. The decolourization reaction
occurs progressively from yellowish brown colour to clear water (Brown et al., 2004;
Xuereb et al., 2006). The detailed mixing time experimental results are discussed in
separated chapter
II. The Pumping Number (NQp)

It is also called flow number. Each type of impellers has its own pumping number. It
is defined as a dimensionless number refers to the pumping capacity of the impeller
by relating the flow discharged by the impeller blades to its diameter and rotational
speed as presented in equation 2.11.
The liquid velocity that leaves the impeller blade tips could be tangential, axial and
radial and this depends on the geometry of the impeller such as the number of blades
and pitch angle for the blades (Bro et al., 2004). To compare between the
geometrically similar impellers, the pumping number is frequently used as a
characterization factor. Generally the axial velocity is generated by the propellers and
radial flow is generated by the turbines.
Impellers pumping number can be represented as:
Common types of propellers have the pumping numbers around 0.5. For ordinary flat
blade turbines and blade turbines are 1.3 and 0.87 respectively (Roustan, 2005).
The pumping number, NQp, for the propeller is obtained in depending upon the radial
and axial flow balance; the total inflow should equal the total outflow. The discharge
flow, Qp, is generated from propeller blades rotation is calculated from the measured
values of axial and radial velocities at known distance from propeller blades within a
determined volume calculated by equations 2.12a, 2.12b (See Figure 2.11). At
propeller vicinity the axial flow, Qpz, is calculated by measuring the axial velocities at
a distances of 1 mm above and below from the propeller edges, where these distances
are usually depend on the rotation position of the propeller. To calculate the radial
flow, Qpr the outer radial borders was taken at 1 mm from propeller edges side. The
formula of pumping rate calculation is depending on the velocity component occurred
within the controlled area as it explained with following equations (Sardeing et al.,
2003):
- 55 -

(2.12a)
(2.12b)
The Qp measuring volume
dr
VZ+
Vr
dz
-
VZ
Figure (2.11): The pumping number and pumping flowrate measuring volume.
To evaluate the impellers efficiencies, sometimes instead of using the pumping

number another dimensionless group was employed and called the discharge
efficiency. This efficiency is varied according to the impeller type and the flow profile
created by the impeller. The discharge efficiency relates with the power number to
pumping number of the impeller. When the ratio Np/NQp is large the impeller is
considered that has less efficiency and when the ratio Np/NQp is small, the impeller is
considered as it has more efficiency, which is called also the circulation efficiency
(Tatterson, 1991).
III. Circulation Number and Circulation Flowrate

The overall circulated volumetric flowrate in the vessel that generated by the
implemented impeller is characterized by the circulation flowrate Qc. The calculation
principle for the circulation flowrate is same as pumping the axial flowrate Qpz as
shown in the equations 2.12b, except that the limits of the integration are larger,
where the upper limit extends to the main circulation loop center in the vessel (Aubin
et al., 2001; Mishra et al., 1998).
- 51 -
(2.13)
The maximum occurred axial flowrate in the vessel is assumed to be equal to the
maximum radial flowrate (Jaworski et al., 1996). The circulation flowrate is
normalized by dividing by ND3 and known as circulation number.
(2.14)
Since the circulation number presents the overall flowrate entrained by the impeller in
the vessel, therefore the NQc is commonly greater than NQp (Jaworski et al., 1996).
IV. Agitation Index (Ig) and Flow Quantification

The agitation index was derived by (Mavros and Baudou, 1997) to evaluate and
compare the agitation quality of different impellers for the mixing process in stirred
vessels. The agitation index calculation depends on the measured mean velocities in
the r-z plane positioned on the posterior baffle (angle = 0o). The agitation index is a
dimensionless parameter as a percentage ratio of the total mean velocity to the
impeller tip velocity:
(2.15)
The 3-D flow inside the stirred vessel can represented by the corresponding 2-D
velocities for each volume grid that related to the vessel dimensions (See Figure
2.12), where the volume weighted average 2-D velocities can be used to compare the
agitation indices. These velocities are determined in assuming that each composite
mean local 2-D velocity (vij) corresponds to a 3-D cell liquid volume and related to
the grid coordinates (Garcia-Cortes et al., 2006; Mavros and Baudou, 1997). The
volume weighted velocity for the vessel is calculated as:
The cumulated volume summation (
) represents the vessel volume except
the volume swept by the impellers and the volume occupied by the cone and baffle
edges, the liquid volume calculations details were carried out as determined in
(Mavros and Baudou, 1997).
- 55 -
The 2-D composite mean local velocities was calculated for each grid point in
depending on the measured mean velocities in the r-z plane performed by the PIV, as
following (Garcia-Cortes et al., 2006; Mavros and Baudou, 1997; Mavros et al.,
1997):
(2.17)
Figure (2.12): The 3-D liquid volume cell and the vessel volume grids, (Garcia-Cortes
et al., 2006).
2.5. Mass Transfer Experimental Setup and Calculation Methods

2.5.1. Oxygen Probe Description
The employed oxygen probes in the experimental runs are of two types, the
polarographic probe and optical probe. The polarographic probe is Inpro 6000, series
O2 sensors, (Mettler-Toledo), type s-96 with DO-meter of Mettler-Toledo, transmitter
O2 type 4100. The function of the polarographic probe depends in its measurements
on the polarographic Clark electrode (i.e. electron transfer reaction principles). The
polarographic probe contains two electrodes, a cathode of platen and an anode made
of mercury immersed in the saline electrolyte of (PH=13). This probe requires a
voltage input from the meter to polarize the electrodes. The oxygen probe has a thin
membrane covering a layer of the electrolyte and the two metal electrodes (See Figure
2.13). The oxygen diffuses through the membrane at a rate proportional to its partial
pressure. The oxygen meter measures the current as the oxygen is reduced at the
cathode and more oxygen diffuses through the membrane, where the meter measures
- 55 -

the concentration of the dissolved oxygen by converting the current into concentration
units. The dissolved oxygen measurement is achieved in depending on the oxygen
partial pressure with correction for the affecting factors such as the temperature. The
concentration units are expressed in milligrams of oxygen per liter of water (mg/l).
Protective Cap
Cap Sleeve
the Membrane
the Cathode
the Anode
Figure (2.13): The schematic diagram of the polarographic probe.
The optical probe (HACH, LDO type) and the dissolved oxygen meter (HACH, type
HQ 10 (OENODEN)) were used. With this probe, the recalibration is not necessary
and the operation principle depends on the fluorescence properties of the electrode.
The amount of dissolved oxygen is determined from zero concentration to the
saturation. The probe includes blue LED, red LED, filter, lens, photo detector and
oxygen preamble foil. The concentration of the dissolved oxygen depends on the
delay by the oxygen molecules held back on the gas permeable foil in the probe cover
cap for the emitted light by blue LED inside the probe. The returning signals are
detected by a photo detector. This shifting or delay between the returning red light
and the blue excitation is measured to determine the dissolved oxygen quantity in
water (see fig. 2.14). The average percentage error of probes readings was calculated
by least squares best fit between experimental and theoretical readings. Average
percentages error were (3.49 %) and (4.8 %) for polarographic and optical probes
respectively.
LED
source
photodetector
cover
cap
foil
Figure (2.14): The schematic diagram of the optical probe.
- 55 -

2.5.2. Mass Transfer Coefficient
The mass transfer during the surface aeration process is occurred mostly at the water
droplets surface and inside the vessel. So the operation of the oxygen transfer is
assumed to be divided into two main essential mass transfer zones: first, the water
spray (droplets) zone that is generated at the water surface by the impeller rotation in
the encircling area around turbine. The second mass transfer zone is the water bulk
zone at the water subsurface near the vessel wall and in the remaining zone inside the
vessel (See Figure 2.15). The principals of oxygen mass transfer that occurs within
these two zones are clearly different due to the distinct way of gas-liquid contacting.
Mass transfer by surface aeration was previously identified only by the water bulk
zone, whereas the little attention has been paid for the droplet zone despite of oxygen
transfer is predominantly in the water droplets zone.
Figure (2.15): The schematic diagram of oxygen mass transfer zones.
I. Liquid Bulk Oxygen Mass Transfer Zone

I.1. Introduction
As the oxygen mass transfer is divided into two mass transfer operation zones of
water spray and bulk zones (See Figure 2.15), it will be two occurring oxygen transfer
coefficients in these two zones (McWhirter and Hutter, 1989). Our concern in this
section is the water bulk mass transfer coefficient kla. The zone will be discussed in
next section. The oxygen transfer in water bulk zone is related with many parameters,
such as water level, tank and impeller geometry, rotation speed, number of impellers
used and aeration rate and other operational conditions (Bandaiphet and Prasertsan,
- 55 -

2006; Nakanoh and Yoshida, 1980; Yatomi et al., 2008; Yoshida et al., 1960). The
oxygen transfer in water bulk occurs directly after the water droplets impinge the
water surface and then they plunge within water bulk; due to that air bubbles are
created and are entrained inside water bulk, in a manner that they are redistributed and
dispersed in whole tank.
The theory of oxygen transfer in water bulk is related to the mixing performance of
the implemented impellers. The mixing process that occurs within the tank and
around the propeller has been implemented widely in aeration processes in order to
enhance contact area between water and air phases (Van't Riet, 1979). As the
resistance to oxygen mass transfer is only considered in the water with neglecting all
other resistances in the system, the oxygen mass transfer coefficient in the water, kla
was regarded as an indicator for mass transfer rate beside the oxygen concentration
profile (Shluter and Decker, 1992). The available interfacial contact area for oxygen
mass transfer in aeration tank is essential for the rational design of a variety of gas
liquid equipment (Dehkordi and Savari, 2011). The interfacial contact area in the
liquid bulk during surface aeration depends essentially on the entrained air bubbles
size (El-Temtamy et al., 1984). The applied method to determine the volumetric mass
transfer coefficient kla in water phase is the dynamic method since the tested aeration
tank is batch system, wherein the initial concentration of dissolved oxygen in close to
zero. The water then is saturated by direct contact with air as a batch-wise manner
(El-Temtamy et al., 1984).
The water surface re-aeration oxygen mass transfer at the water surface is counted
within the water bulk zone, as the boundary of this zone begins directly where the
spray zone ends. The water surface re-aeration is actually not a part of water bulk
oxygen transfer, where at the water surface the atmospheric air is entrained into the
water as result of turbulent condition formed by the falling water droplets.
The re-aeration operation at water surface is more convenient to be included in the
spray mass transfer zone but it is so difficult to consider the surface re-aeration as
separate zone from bulk zone because of its boundaries are so hard to be identified
clearly as they are highly combined with water bulk transfer zone due to flow profile
and mixing effects take place in the tank.
The proposed model that represents the occurred oxygen transfer operation in all parts
(surface re-aeration and bulk inside) of liquid (water) bulk mass transfer zone is the
two-film theory (Atkinson et al., 1995).
The water bulk zone experiments were carried out with the turbine and the 4-bladed
reversible twisted pitched RTP propeller was mounted to the rotation shaft below the
turbine. The draft tube is installed around of the RTP propeller.
- 56 -

The water bulk zone experiments were conducted with the system configuration of
D/Tv= 0.238, C/Tv= 0.313, dpr/Tv= 0.15, Cpr/Tv= 0.2, h/Tv= 0.35, df/Tv= 0.188, but
these ratios were changed as several geometrical parameters are tested.
I.2. Bulk Zone Oxygen Mass Transfer Calculation Methodology

Since the whole resistance to the mass transfer is occurred in the liquid phase, the
oxygen mass transfer coefficient in the gas phase is neglected. It is assumed that no
other mass transfer occurs during the operation for the other constituents from the air
toward the water and also for the used nitrogen gas during the deoxygenating process.
The implemented calculation for the oxygen mass transfer, kla in the experimentation
was also based on several general assumptions: the vessel is efficiently mixed;
subsequently the kla values throughout the vessel can be represented as one value.
Also the saturated dissolved oxygen can be represented as one value, the other mass
transfer may occur that are related to other constituents from the air toward the water
as well as the nitrogen gas -during the deoxygenating process are neglected. Weak
heat gradient occurred during the aeration; in that manner the accompanied heat
transfer is ignored.
I.3. Testing Different Probe Positions in the Vessel

It is assumed that the liquid is well mixed and the variation in kla measured values
should not exist. In order to verify this, seven different positions for the oxygen
probes are tested. The measured kla for the seven different positions were close,
where the kla is changing about + 2.5% of the value 4.1 x 10-3 1/s in all tested points
of the tank (See Figure 2.16, a). The difference could be due to various affecting
factors like the accumulation of oxygen bubbles on the probes, the difference of time
lag for the two probes. The measurement of the dissolved oxygen is carried out by
both optical and polarographic probes placed in two different positions inside the
vessel (See Figure 2.16, b).
Because the measured kla for seven different positions is close and the difference
noticed is very small, the assumption of well mixed tank can be verified for the used
probes accuracy and for applied experimental conditions (Philichi and Stenstrom,
1989).
- 55 -

I.4. Repeatability of Experimental Results
The experimental results repeatability were tested by repeating each time two
experimental runs for same geometric configuration of same probe and for the two
used oxygen probes. The results were close for all the cases, as it can be noticed from
the figure 2.17 (the results obtained at a clearance of 0.10 m, from vessel bottom,
where the kla was about (2.4x 10-3 1/s) for both cases at N=2.08 rps).
Figure (2.16): (a) Experimental oxygen probes positions in the liquid volume,
(b) Experimental oxygen probes positions in the liquid volume, N=2.5 rps, (h/Tv) =
0.35.
- 55 -

3
polargraphic probe 1
optical probe 2
ln(Cs-Ct/Cs-Co)
2.5
2
y = 0.0024x + 0.1455
1.5
y = 0.0024x + 0.044
0.5
0
200
400
600
Time, (s)
800
1000
Figure (2.17): The repeatability of the mass transfer (kla) experimental results for both
oxygen probes, (N= 2.08 rps).
I.5. De-oxygenation and Re-oxygenation Processes

In order to perform the oxygen mass transfer coefficient experimental runs, it is
necessary to apply the de-oxygenation of the water, the implemented technique of
deoxygenating is accomplished by bubbling the nitrogen gas through the water inside
the vessel until it reaches the lowest possible dissolved oxygen concentration level
DO (< 0.5 mg/l) within reasonable time period (1/2 hr.). The operating fluid during all
experimental runs is tap water. To determine the oxygen mass transfer coefficient
during the surface aeration, direct measuring of DO is applied by the immerged
probes.
I.6. Oxygen Probe Response Time Measurement Verification

The response time for each of the used probes (polarographic and optical) in the
experimentation is determined by depending upon both the experimental and
theoretical results of dissolved oxygen. The theoretical values of the dissolved oxygen
are determined by equation (3.11), (Merchuk et al., 1990; Mueller et al., 1967):
Ct= Cs (Cs Co) e-t/
(2.18)
While the experimental DO values are determined directly during the experimental
measurements of the probes response time through the direct changing from the zero
DO level solution (prepared by passing the Nitrogen gas through the water), then into
- 55 -

100% saturated solution (prepared by passing the air through into the water for
sufficient time), where the DO concentration were measured each 3 seconds until
reaching the saturation concentration level for the DO.
The , response time is determined by identifying the lowest difference value between
the experimental and the theoretical value that is calculated by equation (2.18) of
dissolved oxygen values which matches the highest value of the R-squared for the
least square method application, as it shown in the (Figure 2.18 and Figure 2.19). The
response time for the polarographic probe (Inpro 6000) is (9.2 sec) and for the optical
probe (HACH, LDO type) is (8.8 sec).
Ct , Experimental
10
8
R = 0.9973
6
4
2
0
4
6
8
Ct , Theoretical
10
Figure (3.18): Optical probe response time verification
Ct , (mg/l), Experimental
10
8
R = 0.9979
4
2
0
10
Ct, (mg/l), Theoretical
Figure (2.19): Polarographic probe response time verification.
- 55 -

I.7. Determination Model of the Bulk Zone Oxygen Mass Transfer Coefficient
(kla)
The relationship between the DO with the time during the surface aeration process
with taking in account the effect of probe response time can be represented as the
following equation (Arjunwadkar et al., 1998; Chern and Yang, 2003; Puthli et al.,
2005):
(2.19)
By applying the boundary condition at t=0, Cl =Co and at t=t, Cl =Ct:
(2.20a)
By rearranging equation (2.20a):
(2.20b)
On the other hand, the effect of probe response time on the dissolved oxygen
concentration measurement for aeration processes is identified according to the
following equation (Sardeing et al., 2005):
(2.21)
To evaluate the probe response time effect on the experimental results and the
calculated mass transfer coefficient kla, a comparison was made for same
experimental and calculated conditions. The dissolved oxygen concentrations in both
experimental and theoretical that calculated by equation (2.21) for the polarographic
and optical probes respectively are shown in the figure (2.20) and (2.21). The oxygen
mass transfer relation with time with considering the probe response time was implied
in the applied mass transfer equation that represents the experimental runs to take in
account the experimental errors that are produced due to that. The oxygen mass
transfer coefficient is calculated theoretically with applying equation 2.21, where the
effect of time lag constant for the oxygen probes are implied in our calculations of
oxygen mass transfer in water bulk zone kla.
- 51 -

10
Ct, (mg/l), (Theoretical)
8
R = 0.978
6
4
2
0
2
4
6
8
Ct, (mg/l), (Experimental)
10
Figure (2.20): The probe response time relationship verification for both theoretical
and experimental DO values for the optical probe.
Ct , (mg/l) , (Theoretical)
10
8
R = 0.9902
6
4
2
0
2
4
6
8
Ct , (mg/l), (Experimental)
10
and experimental DO values for polarographic probe.
I.8. Measurement Procedure for the Bulk Zone Oxygen Mass Transfer
Coefficient
At the beginning of the experimental runs, the water volume was de-oxygenated by
the bubbling of Nitrogen gas through it for at least 15 minutes until DO concentration
becomes below (0.5 mg/l). The DO measurements were conducted directly by the two
probes that situated in their positions as shown in the Figure (2.16), where the
dissolved oxygen was measured each 30 seconds till reaching the saturated level,
where run was terminated. The operation parameters and geometric configuration
- 55 -

were considered as variables according to the purpose of the experiments to identify
their influence.
I.9. Temperature Correction for the Oxygen Mass Transfer Coefficient

To calculate the necessary parameters for standard oxygen transfer rate and the
standard aeration efficiency its needed to correct the measured oxygen transfer
coefficient in the liquid to the standard conditions, where the following (Vant- Hoff,
Arrinous) equation was applied (Capela et al., 2004):
(2.22)
The volumetric coefficient of oxygen mass transfer is derived by depending on the all
the resistances in the water to oxygen transfer. The volumetric coefficient of oxygen
transfer, (kla)T is measured at operation temperature T, and corrected to standard
temperature here usually T=10 oC or 20 oC by the equation 2.22 (Roustan, 2003).
I.10. Oxygen Transfer Rate for the Bulk Mass Transfer Zone (OTRb)
The oxygen transfer rate (OTR), which is the transferred mass of oxygen rate at
experimental condition in the total volume of the filled vessel with water.
(2.23)
To have complete view for oxygen mass transfer for all surface aeration process and
to realize the contribution of droplets spray mass transfer zone in the oxygen mass
transfer operation, oxygen mass transfer rate in other side of water bulk re-aeration
mass transfer zone OTRb is needed to be calculated. OTRb was determined by
applying the following relation, which is modified to be agreeable for implemented
system depending upon mass transfer expression derived by (Chern and Yang, 2004):
(2.24)
I.11. Standard Oxygen Transfer Rate for the Bulk Mass Transfer Zone (SOTRb)
The standard oxygen transfer rate (SOTR)b is the transferred mass of oxygen rate at
standard condition in the total volume of tank filled with water. The SOTRb is the
standard criterions that employed to determine the surface aeration performance
- 55 -

beside the standard aeration efficiency SAEb (Cancino et al., 2004; Sardeing et al.,
2005).
(2.25)
I.12. Standard Aeration Efficiency for the Bulk Mass Transfer Zone (SAEb)
The standard aeration efficiency (SAE)b sometimes it called the overall transfer
efficiency which represents the mass of oxygen transferred to the water per the
consumed power (Cancino et al., 2004; Sardeing et al., 2005)
II. Spray Oxygen Mass transfer Zone

II.1. Introduction
The oxygen mass transfer coefficient at the water surface, which is identified as spray
(droplets) zone oxygen transfer coefficient klad is investigated in details in this study.
The spray zone begins with the creation of the water droplets thrown into the
atmospheric air rapidly as a projection (spray) at the water surface by the turbine
through the annular surrounding peripheral space until they impinge the water surface
at the end of their flight, where with this point the second zone of water bulk occurs
(See Fig. 2.15). The oxygen that transferred at this step mostly takes place at the wall
of these droplets, where for this kind of gas-liquid contacting is considered as a single
stage zone. The droplets are the dispersed liquid phase into the virtually infinite
continue gas phase with constant oxygen concentration in the atmospheric air. The
gas phase is considered as completely mixed of constant atmospheric air composition
with plug flow of dispersed liquid phase for water droplets (McWhirter et al., 1995).
It is assumed that any other possible secondary mass transfer can be occurred is
neglected such as the possibility of the Nitrogen gas transfer contained in the
atmospheric air toward water droplets. Since the mass transfer resistance is exist
merely at the liquid phase side, so the water mass transfer toward the gas phase is
neglected.
- 55 -

II.2. Spray Zone Oxygen Mass transfer Coefficient Calculation Methodology
The occurred operation within and around the water droplets during their flight is not
only the oxygen mass transfer, other parameters as well accompany the operation like
the heat transfer or the water evaporation. The heat transfer that combined with the
mass transfer has very important effect on the water droplets air surface temperature,
and it is not the same with the water bulk inside the vessel. The employed saturation
level here is considered as the wet bulb temperature of the air, C ds (Huang et al.,
2009). The basis used for this assumption is depending on the fact that the path of
water droplets in the atmospheric air follows its own conditions of constant air
composition, humidity and temperature which is completely different from those
existed in the water bulk. As a result the higher equilibrium oxygen concentration can
be reached theoretically for the droplet is the air wet-bulb temperature. The
constraining factor controls how deep the wet-bulb temperature goes inside the
droplet is depending on the droplet fight time, (which is too short in our case).
Consequently the wet-bulb temperature prevails only at water droplet surface, while
the inner droplet part temperature is highly effected by the relative high droplet
velocity in the air and high heat transfer caused by this velocity, it remains generally
at lower level than the surface temperature, while due to the inner motion the oxygen
mass transfer inside the droplets is enhanced (Sirivasan and Aiken, 1988).
II.3. Determination Model for the Spray Zone Oxygen Mass Transfer Coefficient
(klad)
The representative model for the spray zone that explains the oxygen mass transfer
during the travel of the droplets in the atmospheric air from turbine blades tips until
their impingement points at the water surface. The droplets trajectory are determined
by the following unsteady state oxygen mass balance equation (McWhirter et al.,
1995):
(2.27)
In equation (2.27) several assumptions are made to facilitate the explanation of the
oxygen transfer process for the intended zone. Each water droplet is assumed to be
completely mixed; the dissolved oxygen distribution is uniform within the droplets at
any time when they traverse the atmospheric air until their impingement points at
water surface with no back mixing might take place during this flight. The dissolved
oxygen concentrations at the limits of the droplet zone are related with other
presented zones, where at the zone inlet the concentration at any time is the same of
the well mixed liquid bulk for the dissolved oxygen concentration level inside the
- 55 -

vessel. The dissolved level at the outlet of the droplet zone (when the spray hits the
water surface) for the oxygen gas is considered as Cdt. The dissolved oxygen at outlet
of droplets zone, Cdt is the average of concentrations level for all the droplets at the
end of water droplets spray zone. If the droplets volume and the mass transfer
coefficient with the droplets oxygen saturated concentration are considered constant,
the equation (2.27) can be integrated and resolved as follows:
(2.28)
Equation (2.27) is a distinguishing equation that is built to explain especially the

droplet spray oxygen mass transfer zone; the oxygen behavior is well described
particularly for this step on the contrary for the equation (2.19), which is a general
equation that can be applied for the mass transfer zones but with the limited ability to
identify the actual mass transfer path that happens in each step. Equation (2.28) shows
that it is preferred to keep the droplets volume at the minimum during the surface
aeration process to reach possible highest dissolved oxygen level. For the same reason
the droplets should stay as long as possible in the atmospheric air before they fell
down into the liquid bulk and go inside the vessel.
In order to evaluate the efficiency for gas-liquid contacting systems, it is common to
apply a known indicating factor called (Murphree contacting efficiency Emd) for
various mass transfer operations (Baylar and Bagatur, 2000b; Chisti and JaureguiHaza, 2002; Sharma, 2007). In the surface aeration water spray zone and in other
similar processes, the implemented criterion for indicating the progressive oxygen
mass transfer that takes place is the Murphree contacting efficiency or it is called in
our case the spray mass transfer zone aeration efficiency, Esp (Toombes and Chanson,
2005; Vouk et al., 2005).
Equation (2.28) is rearranged to express the contacting efficiency is dependent on the
fraction between the actual and the theoretical oxygen transfer in the droplets spray:
Equation (2.29) shows that the contacting efficiency can be varied from its minimum
value (0.0) (no mass transfer occurred) to its maximum value of (1.0), which is an
ideal condition that can be reached at infinite time when equilibrium is achieved
between the water droplets and the atmospheric air (McWhirter et al., 1995). The
equilibrium level for the dissolved oxygen at the droplets outer surface depends on the
- 56 -

air humidity is less than the saturation level in the water bulk. So the saturation level
at this condition is better represented by the wet bulb temperature than the normal
dissolved oxygen saturation level in the water bulk; the presence of related heat
transfer takes place, as the air temperature here differs from the water bulk
temperature.
Rearranging equation (2.29) and re-writing it in the term of water spray zone surface
aeration efficiency gives us the following:
In the equation (2.30), the definition of Murphree contacting efficiency relates the
boundary levels for the actual and theoretical oxygen concentration gradients that take
place from water droplets propelling position to the impingement with water surface
location. In order to facilitate the calculation of the oxygen mass transfer results, the
previous equation (2.30) is rearranged in more appropriate correlation form (Baylar
and Bagatur, 2000a; McWhirter et al., 1995):
(2.31)
From equation (2.31) it can be estimated that the plot between the resulted oxygen
concentrations at the droplets impingement point with water surface at the
corresponding recorded water bulk dissolved oxygen gives a straight line with a slope
of (1-Esp), and the intercept with y-axis will be (Esp Cds).
Same as the surface aeration the similar efficiency calculation technique is applied in
other aeration processes like the aeration by weirs (Baylar et al., 2006; Kim and
Walters, 2001), where it is required to find a uniform basis for comparison purposes
between the obtained results, which can be achieved by normalizing the calculated
efficiencies to standard condition.
II.4. Temperature Correction for the Spray Mass Transfer Zone

The spray mass transfer zone aeration efficiency, Esp is corrected to the standard
temperature condition, mostly of 20 oC by applying, the widely applied temperature
correction relation is the Arrhenius type of water temperature correction. The
following relationship is developed by many previous work, (Baylar et al., 2010;
Gulliver et al., 1998; Khudenko, 1983; Nakasone, 1987; Wormleaton and Soufiani,
1998):
(2.32)
)
- 55 -
Where the (Esp)20 is the spray aeration efficiency at 20 oC, and Esp is spray transfer
aeration efficiency at the water bulk temperature, f is the exponent determined by the
following equation:
(2.33)
II.5. Oxygen Transfer Rate in the Spray Mass Transfer Zone (OTRsp)
It is necessary to take in account the relative humidity of the air when this approach of
OTRsp is applied if a remarkable change is occurred with the relative humidity during
the experimentation because the saturation level that used at the air wet-bulb
temperature is also depends on the air relative humidity.
Another important factor in the oxygen transfer in the water droplets spray zone is the
oxygen transfer rate OTRsp, which beside the standard aeration efficiency depends on
the water volumetric flow rate discharged by the turbine blades (Chern and Yang,
2004; McWhirter et al., 1995):
(2.34)
II.6. Spray Zone Mass Transfer Coefficient (klad) - Measurement Procedure

The measurements performed are similar to the mass transfer coefficient inside the
liquid bulk except that in this case it is needed to measure the dissolved oxygen at the
end of water droplets spray zone before they go inside the vessel, Cds. The
implemented measurement technique in this case for the given surface aeration
operation conditions and geometric configurations are depending on the dissolved
oxygen level in the droplets during their flight. In order to achieve the accurate
investigations it is needed to apply sampling for the droplets at nearest point above the
water surface prior to their clash with the surface and enter inside the bulk. The
available way for this sampling is collecting these droplets by narrow neck bottle
provide with rubber cup to prevent further contact with air after the appropriate
amount is collected (the collected water quantity was in the neck bottle for each
reading was approximately between 20 to 25 ml). Then an immediate measurement is
applied by an optical oxygen probe. The measurement is repeated periodically each
(60 seconds). At the same time, the dissolved oxygen level inside the water bulk is
measured also by the polarographic oxygen probe positioned inside the vessel; the
dissolved oxygen is registered with same time interval for each (60 seconds) (See
Figure 2.22).
- 55 -
Figure (2.22): The droplets zone mass transfer coefficient (klad) measurement.
2.6. Water Droplets Flight Time (tf)

In order to calculate water spray zone oxygen mass transfer coefficient klad, it is
necessary to be determined the droplets flight time. The droplets flight time is
calculated in function of the directly measured maximum droplets height Ym and the
height of apparent turbine blade above water surface H (See Figure 2.23), as it shown
in the following equation (McWhirter et al., 1995):
(2.35)
Equation (2.35) is derived principally from Newtons second law of motion, where
the initial vertical velocity of the water droplets has a changeable value during
droplets throwing operation (as it will be seen in next section). The droplets flight
time is affected by its trajectory, which in turn is affected by many parameters beside
rotation speed that act as lift force on droplets, like gravity, buoyancy and drag force
(Matsuura et al., 2003).
- 55 -
Figure (2.23): Schematic diagram illustrates the surface aeration water droplets spray
from turbine blades till the impingement point with some important relevant
dimensions.
2.7. Water Droplets Velocity and Volumetric Flow Rate

The water droplets are discharged with a considerably high velocity Vsp according to
turbine rotation speed used. The water droplets velocity is regarded as an essential
step for further droplets spray zone related calculations. Determining Vsp by directly
measuring turbine discharge velocity is somehow complicated. The alternative way is
applying the measurements to droplets flight path, where these projected droplets are
considered as free falling objects following physical laws that used in similar cases. In
order to simplify Vsp calculation, several assumptions are applied; (i) no friction
losses due to droplets flight in the atmospheric air, (ii) regardless to droplets
projection position by turbine blades above the static water surface, and (iii)
maximum droplets height during their trip is considered constant that equal to Ym that
corresponds to maximum spray radius Rm that fulfilled by droplets impingement at
the water surface. The total water spray droplets velocity can be feasibly represented
by one value, wherein in turn the three components of the spray droplets velocity are
constant. The spray droplets velocity is consisting of three components; vertical,
radial and tangential velocities, where their relationship can be represented by the
following equation (Huang et al., 2009; McWhirter et al., 1995):
(2.36)
- 55 -

Droplets vertical velocity propelled by the turbine in contrast to radial and tangential
velocities is determined by its initial velocity because at a certain point it begins to
change pursuant to acceleration of gravity; (Voy2) is determined by measuring the
spray maximum height (See Fig. 2.23), (Huang et al., 2009; McWhirter et al., 1995)
as follows:
(2.37)
The used straight tip blade turbine propels water droplets toward radial and tangential
directions in a horizontal plane, is parallel to the water surface, as it has been
described in Figure 2.24. To calculate these velocities for the given applied geometry
the following relationships (See equations 2.38 and 2.39) are employed (McWhirter et
al., 1995) with equation (2.35) for droplets flight time:
(2.38)
(2.39)
Figure (2.24): Droplets radial and tangential velocities propelled in horizontal plane
by turbine blades.
- 51 -

The determination of Vr and V separately is very complicated, because in general
they are calculated together as (Va). But Va in turn is also not easy to be measured or
identified, and it needs a sophisticated analysis that depends on the turbine geometric
configuration. (Huang et al., 2009) tried to simplify this analysis for straight blade tip
turbine. In order to have feasible value of oxygen transfer rate in the spray mass
transfer zone, (Huang et al., 2009) considered it is more logical to determine this
value at droplets impingement point with water free surface. The occurred velocities
at impingement point are essentially the radial velocity V'r and secondly the axial
vertical velocity V'y, when on the other hand the tangential is very weak and can be
ignored or actually it disappears. So the velocities at impingement position can be
calculated as follows:
(2.40)
(2.41)
At the impingement position the water droplets velocity is considered approximately

equal to the discharge velocity from turbine blades and the water spray droplets
direction is not different evidently from radial direction. For very short droplets flight
time up to (0.19 s), the energy consumed in the atmospheric air was neglected. (Vsp)
can be determined as:
(2.42)
When the droplets spray velocity is known, the volumetric flow rate of the droplets
that are propelled at turbine blades edges can be determined by applying the overall
conservation of energy to surface aeration turbine and the overall power consumed
through blades to project these droplets into atmospheric air. It is assumed the power
consumption is delivered to the free water surface as a kinetic energy, which
contributes to elevate water velocity up to discharge velocity leaving the blades as
droplets thrown in the air with power conversion is 100% effectively achieved (Baylar
et al., 2001). The energy balance relation gives an appropriate relation between the
spray volumetric flow rate and the power consumption (Baylar et al., 2001;
McWhirter et al., 1995). (Baylar et al., 2001) used in their approach the pumping
power relation in weir aeration process by water droplets creation. The relation was
applied within the boundaries between jet point and plunging point of droplets. In
surface aeration process the previous relation can be considered as a convenient
- 55 -

relation that employed in similar conditions, where the better way is applied the
energy conservation of converted mechanical energy to a kinematic energy within the
boundaries between the inlet and outlet points along turbine blades, where for 100%
energy conversion efficiency the general energy balance yields the following relation:
(2.43)
Where, the constant 1.38910-7 depends on the applied geometry of the surface aerated
equipment. For eliminating errors that may occur when whole system is used during
the energy balance calculation, the power consumption was measured only with
presence of the turbine to assure all the energy was converted to propel water droplets
into atmospheric air (i. e. the power consumption by the propeller with draft tube do
not have any relation with droplets generation outside the water surface but they
consume power only for water bulk mixing purposes inside the tank).
2.8. Conclusions
The procedures for measurements and calculation methods for the oxygen mass
transfer coefficients in both water bulk and water droplets spray zones, power
consumption, mixing time, water spray discharge flowrate, water droplets flight time
and other related parameters are represented in this chapter. These methods are
applied in the experimentation to propose the results that are showed in the next
chapters. The measurements of the flow field and mean velocity profiles in all parts of
the tank that achieved by Laser Doppler Velocimetry LDV and Particle Image
Velocity PIV are also have been described in this chapter.
- 55 -
- 55 -
- 89 -
- 90 -
Chapter Three: Oxygen Mass Transfer in the Surface Aeration
Chapter Three
The Oxygen Mass Transfer in the Aeration Mode
The surface aeration system has the capacity to operate in two different operation
modes;(i) Firstly, The aeration mode, where two phases (water and air bubble)
condition occurs, this mode is always exist when the turbine is employed. Mode,
where the single phase condition occurs, the mixing is accomplished by the RTP
propeller (no oxygen mass transfer takes place in the mixing mode).
The oxygen mass transfer in the surface aeration system is divided into two main
mass transfer operation zones of water bulk and spray zones according to the way of
contact between the water and the air (See Figure 2.14). The two occurring oxygen
transfer zones have two phases; the dispersed phase of the water droplets in the
continuous phase of atmospheric air. In water bulk mass transfer zone these two
phases are reversed, where the continuous phase is the water bulk and the dispersed
phase is the air bubbles. For these two zones there will be two different oxygen mass
transfer coefficients (McWhirter and Hutter, 1989). The water spray zone represents
the oxygen mass transfer that takes between the atmospheric air and water droplets in
the spray. While the water bulk zone represents the oxygen mass transfer that takes
place inside aeration tank between air bubbles and water bulk. It is so hard to identify
clearly the limits between the two zones. Theoretically the water bulk zone doesnt
begin directly when the water droplets impinge water surface where air bubbles are
created and entrained inside water bulk. Due to this impingement and the direct
contact with the atmospheric air a separate mass transfer zone occurs, where a reaeration takes place between the water surface and the atmospheric air in a manner
close to that happened in the water spray zone. This water surface aeration is actually
not a part of water bulk zone. But it is highly combined with water bulk transfer zone
due to flow profile and mixing effects that takes place in the tank zone, so practically
the effect of surface re-aeration is included in this water bulk zone.
3.1. Water Bulk Mass Transfer Zone

3.1.1. Introduction
The measurement of the dissolved oxygen deficit carried out by utilizing both of
optical and polarographic probes, they are placed in two different positions inside the
vessel (See chapter 2), as it is assumed before the liquid is efficiently mixed, in this
way the variation in kla measured values should not to be exist or to be very limited.
In order to verify that, seven different positions for the oxygen probes are tested, (the
measured kla for the seven different positions were close (the resulted kla was around
4.0x10-3 1/s, as illustrated in Figure 2.15), the difference noticed was very small, this
gap might be caused by the simple accumulation of oxygen fine bubbles on the probes
- 19 -

despite of the effective circulation occurs in the vessel or due to the difference of
saturated concentration levels between these positions.
I. Effect of Geometrical Configuration
Three geometrical configurations are examined beside the whole system
configuration; (i) the turbine and propeller (draft tube elimination), (ii) the turbine
alone (the draft tube and propeller elimination) and (iii) the turbine and the draft tube
(the propeller elimination) (See Figure 3.1). For all the modifications it was found that
the kla is relatively lower than that achieved by the whole system configuration
(Figure 3.2), which indicates that with propeller and/or draft tube has a contribution in
kla. It is obvious that the draft tube improves the propeller function, where the flow
pattern is affected by redirecting the flow from the lower propeller towards the
turbine. From Figure 3.2, it was found that neither the draft tube nor the RTP
propeller can improve kla individually, where their influences appear to be combined.
Figure 3.2, illustrates that the contributions of draft tube and propeller in the
enhancement of kla are negligible at low impellers speeds, their effect starting to
appear with impeller speed higher than 2.08 rps (Re = 75000, ND = 1.24 m/s), which
can be considered as critical rotation speed for the surface aeration as it will be
discussed the in next section.
Figure (3.1): Different tested geometric configurations, (a) Whole system, (b) Turbine
alone, (c) Turbine + Propeller, (d) Turbine + Draft tube.
- 19 -

8.5
Whole system
Turbine + Propeller
Turbine alone
Turbine + Draft tube
7.5
kla *10 -3, (s-1)
6.5
5.5
4.5
3.5
2.5
1.5
0.5
1.5
2.5
3.5
N, (s-1)
Figure (3.2): Impellers speed effect on the oxygen transfer coefficient with four
different geometrical configurations; (Whole System, without draft tube, turbine alone
and without propeller), D/Tv=0.24, C/Tv=0.31, h/D=1.47, Cpr/Tv=0.2, S/W=1.
II. Effect of Impellers Rotational Speed

The dissolved oxygen concentration profile showed in Figure 3.3 is for the water bulk
mass transfer zone with four rotation speed levels, the figure illustrates that with
increasing of impellers rotation speed N the dissolved oxygen concentration increases
till reaching its higher possible level near to its saturation dissolved oxygen in water.
With higher impeller rotation speeds, higher energy was consumed to achieve
efficient mixed condition of the dissolved oxygen in the water bulk, where the
dissolved oxygen reaches its equilibrium concentration in shorter time with higher N;
for example after (500 sec) from the beginning of experiment time at (N = 1.67 rps)
oxygen concentration level was reached (5.0 ml/g), while at (N= 3.33 rps) the oxygen
concentration level was reached (9.3 ml/g).
The oxygen mass transfer coefficient, kla value depends on the dissolved oxygen DO
range that implemented during the calculation in order to have the best fit of dissolved
oxygen deficit term (ln C*-Ct/C*-C0) versus experimentation time in the semi log
graph and to get rid of the nonlinear parts of the plot; usually the curved parts are
subject to truncations to reach this aim. As the calculated kla value is related with the
applied dissolved oxygen range during the measurement depending on the way that
the kla was calculated. Some authors have used linear regression of DO deficit with
time to calculate kla from equation 2.20; they have taken different dissolved oxygen
range (10%-70%) to (10%-90%) of the saturated oxygen concentration in water
- 19 -

(Boyd, 1998; Roustan, 2003), where this truncation was applied in order to find best
fit between actual and measured dissolved oxygen concentrations. In our calculations
we didnt use any truncations as we applied the least squares technique of nonlinear
regression for oxygen deficit (ASCE, 1993) for the equation 2.21, which it seems to
be more determinative for the experimentally reached mass transfer coefficient kla
values with the best accuracy.
The rotational speed of the turbine and the propeller was tested (with keeping other
parameters constant) on the oxygen transfer coefficient kla for the speed range of
(1.67 3.33 rps). As shown in the Figure 3.2, an increase of the rotation speed leads
to kla increasing, in other word the saturated oxygen condition was reached in shorter
time. This relation is explained as the impellers rotation speed was increased, the
amount of the water exposed to atmospheric air was increased consequently and the
interfacial contact area between the air and water droplets was enlarged, the upper
limit of the rotation speed for this relation 3.33 rps, after this level further increasing
in the rotation speed leads to the water droplets hit the vessel wall and some of them
splashed of the vessel, this situation is highly undesirable and considered as
inapplicable operation.
10
Saturated DO line
Dissolved oxygen conc., (mg l-1)
9
8
7
6
1.67 rps (Run1)

1.67 rps (Run2)
2.08 rps (Run1)
2.08 rps (Run2)

2.5 rps (Run1)
2.5 rps (Run2)
3.33 rps (Run1)

3.33 rps (Run2)
1
0
200
400
600
800
Time, (s)
1000
1200
1400
1600
Figure (3.3): The DO profile for different rotation speeds, h/Tv= 0.35, C/Tv=0.313,
Cpr/Tv=0.2, temperature =15 oC.
For the different geometrical configurations tested, the oxygen transfer coefficient kla
always shows a dependence on the impellers rotational speed, as it can be seen in the
- 19 -

Figure 3.2, this agrees with the results obtained by (Chisti and Jauregui-Haza, 2002).
Starting from N= 2.08 rps, it was visually observed that the formation of air bubbles
appears in the vessel due to water droplets plunging with water surface. So the
rotation speed 2.08 rps can be considered as a critical rotational speed.
III. Turbine Blades Submergence Effect

The effect of turbine blades submergence on the oxygen transfer coefficient was
examined (See Figure 3.4). The turbine blades submergence refers to the distance
from the water surface to the lower edge of turbine blade this effect is represented as
the dimensionless geometric ratio (S/W) as shown in Figure 3.4. This effect is also
can be considered as the effect of the water spray form and water droplets properties.
The turbine submergence (The distance from the tank bottom to the lower edge of the
turbine) is fixed during the experimentation. When the turbine blades submergence is
modified in the experimental runs, the liquid volume is modified too.
The effect of the water level inside the vessel that represented as the ratio h/D
represents here the effect of the liquid volume through its relation with flow condition
and circulation. It was found that the kLa is highly sensitive to the turbine blades
submergence and water level variation, this agrees with (Patil et al., 2004).
Five different turbine blades submergence are tested (0.17, 0.58, 1.0, 1.42, 1.83, and
2.25) that correspond the water levels h/D of (1.37, 1.42, 1.47, 1.53, 1.58 and 1.63)
respectively, where three conditions are tested; at S/W < 1 the turbine blades are
partially submerged (h/D lower than 1.47), at S/W >1 (h/D higher than 1.47) the
turbine blades are over submerged and at S/W=1 which corresponds to 100%
submergence of the blades (h/D = 1.47) here the water level is just at blades upper tip.
(a)
- 19 -

6
N=2.08
rps
S/W=1.42
S/W=1.00
S/W=0.58
kla*10-3 , (s-1)
1
0
1.30
6
5
kla*10-3 , (s-1)
1 (S/W), (-) 2
1.40
1.50
(h/D), (-)
1.60
1.70
1.5
2.5
N, (s-1)
(b)
(c)
Figure (3.4): (a) The limits of the tested turbine blades submergence, (b) Mass
transfer Coefficient kla relation with turbine blades submergence and water height in
the liquid bulk for three levels of rotational speed, (b) Mass transfer Coefficient kla
relation with the rotation speed for three levels of turbine submergence. D/Tv= 0.24,
C/Tv = 0.31, Cpr/Tv=0.2.
Figure 3.4 shows that higher S/W ratio leads to higher kla until the ratio S/W=1.75
after this point the kla values are decreased. This relation can be explained that with
higher (S/W) more water droplets are thrown into the air, consequently larger
interfacial area achieved between air and water and also the flight trip of these
droplets goes farther, so as a result the kla mount up, but at the lower impellers speed
1.67 rps, the kla stays relatively constant despite elevation the S/W ratio.
At higher impellers speeds, further increasing with S/W higher than 1.58 the k la goes
down, as illustrated in (Figure 3.4). The dropping down of kla with further increasing
of water level ratio is happened, when it is noticed with higher water levels the spray
begins to be deformed; the interfacial contact area is constricted. For the water levels
lower that 100% submergence (h/D = 1.47) of the aerator blades, the values of kla are
moderate. This relation is depending on the desired shape of droplets spray as well as
preserving interfacial area between air and water.
Figure 3.4, displays the influence of water level on the kla, as the higher (h/D) ratio
leads to higher kla produced, This relation can be explained that with higher (h/D),
more efficient fluid circulation loop is accomplished in the tank, so as a result the kla
is increased, but at the lower impellers speed 1.67 rps, the kla stays relatively constant
despite elevation the h/D ratio.
- 19 -

The kla reaches a maximum with the (h/D) ratio of 1.58, with further increasing with
(h/D) the kla goes down (2.8x10-3 1/s), as illustrated in Fig. 3.4. This behavior is usual
one, where the kla mount up with increasing of water level ratio turbine blade
submergence ratio, in consequence more of water droplets are thrown into the air and
also the flight trip of these droplets goes farther, where with S/W more than 1.75 (h/D
higher than 1.58) the kla goes down, as illustrated in Figure 3.4.
IV. Effect of the Spacing between the Impellers

The lower RTP-U propeller in up-pumping mode (Aeration mode) was implemented
to assist the surface aerator turbine performance by redirecting the flow toward the
aerator intake in order to eliminate any shortage during the water projection and also
the other important propeller role is achieving the homogenous distribution of
dissolved oxygen in entire the vessel by ensuring the flow circulation moves through
all vessel parts especially the deep lower levels in the down-pumping mode (Mixing
mode), where with its presence the short circulations that may occur in the upper half
of the vessel is eliminated by forcing the flow to continue its circulation down to the
vessel deep bottom.
The obligatory position of the lower propeller inside the draft tube limits the choices
to vary the distance between the two impeller, where the performance of the propeller
is conjugated with that of the draft tube, as a result it is evident the propeller should
stay inside the draft tube, on the other hand the existence of draft tube baffles prevent
to push down the propeller to more lower positions (more geometric configuration are
tested see next section).
From Figure 3.5 it can be noticed that the relationship between kla and rotation speed,
N for two different spacing distances between the turbine and lower propeller; the
reference position that is implemented in the all experimental runs of 7.6mm with
lower position of 8.3 mm, where at lower rotation speed 1.67rps the lowest kla
achieved, which was about 1.3x10-3 s-1 for both configuration, while with elevating
the rotation speed to 2.5 rps, higher kla accomplished of (5.42 - 5.8)x10-3 s-1, when N
reaches its higher value 3.33 rps, kla was increased to around 7.1x10-3 s-1 to 7.89x10-3
s-1 for spacing 8.3 mm and 7.6 mm respectively.
From these results it is figured out with lower impellers speed (up to 2.5 rps), the
achieved kla was close for both the tested spacing 8.3 mm and 7.6 mm. For higher
impellers speed (higher than 2.5 rps) the kla is lower for the spacing of 8.3 mm that
that of 7.6 mm, and the difference becomes wider with impellers rotation speed
increasing (See Figure 3.5).
The slightly lowered kla at the lower tested position of the propeller Sp=8.3 mm,
compared with the standard reference position (upper position) Sp=7.6 mm for higher
rotation speed N >2.5 rps, comes from the fact that with lower position the inter- 19 -

action between the aerator and the propeller is weakened, since no other meaningful
reasons can be developed with this small replacement, in that way the spacing has an
effect on the oxygen transferred.
10
Spacing =7.6 mm
Spacing= 8.3 mm
kla *10 -3, (s-1)
1.2
1.7
2.2
2.7
N,
3.2
3.7
(s-1)
Figure (3.5): Spacing effect on the oxygen transfer coefficient, D/Tv= 0.24,
C/Tv = 0.31, Sp (Reference) =7.6 mm.
V. Power Consumption Measurements
In the aeration mode the power consumption was studied for various turbine blades
submergences (S/W), for the spacing variation between two impellers and for the
different geometrical configuration modifications (Figs. 3.6, 3.7 and 3.8).
It is observed that the power consumed is changing due to all these modifications; the
power consumed is highly relevant to each replacement or modification performed in
different manners. The difference in power consumption was very slight for the
spacing alteration and for the draft tube elimination, while it was remarkably affected
by the increasing of turbine blade submergence levels.
Figures 3.6 a and b illustrate that the power consumption in aeration mode is
dependent on both the Reynolds number (the diameter of the turbine D, was taken for
calculating Re) and water levels or turbine submergence. The power number
decreases with increasing the Reynolds number, as higher power is consumed with
higher impellers rotation speed (See Figure 3.6b). The power number increases when
the turbine blade submergence increases, wherein higher power consumed with
turbine blade submergence and higher water level. In the case of the water level is
higher than upper tip of turbine blades h/D =1.53 with blade submergence ratio
S/W=1.42, the turbine blades need more power to propel water droplets through the
air. The resulted power num er s for the three tur ine su m ergence levels are
- 19 -

somehow close at highest eynolds num er achieved, while the difference goes to e
more widely at lowest eynolds num er
6 104.
6
(S/W)=1.42
(S/W)=1.0
(S/W)=0.58
Np, (-)
0
20000
60000
100000
140000
180000
220000
Re, (-)
(a)
40.0
35.0
S/W=1.42
S/W= 1.00
30.0
S/W= 0.58
P, (watt)
25.0
20.0
15.0
10.0
5.0
0.0
1.0
2.0
3.0
N, (s-1)
4.0
5.0
6.0
(b)
Figure (3.6): (a) The relationship between power number and Reynolds No., for
different turbine blades submergence levels C/Tv = 0.31, (b) The relation between
power consumption and rotation speed, C/Tv = 0.31.
- 11 -

Figure 3.7, shows the power consumption in the water bulk as mentioned before, for
both tested spacing (7.6 - 8.3 mm), on the other hand both spacing configurations
showed close power consumption.
5
Position A (Spacing=7.6 mm)
Position B (Spacing =8.3 mm)
Np, (-)
1
40000
60000
80000
100000
120000
140000
Re, (-)
Figure (3.7): The spacing between two impellers effect on the power consumption.
Figure 3.8, shows that the power number Np relation with Re for geometrical
configuration effect on whole system, without draft tube and propeller and without
draft tube configurations, where the relation appears similar behavior with Figure 3.6.
It is important to explain that power consumption is less relevant to geometry
modification than the water level h elevation.
Figure 3.8 illustrates the effect of the draft tube and the RTP propeller on the power
number for the condition, where for all different configurations the water level is just
at the upper edge of turbine blades. It points to the power consumption was slightly
changed from whole system configuration for Reynolds number range (6104 to
18104). The power number profile was very similar for the draft tube elimination and
the turbine alone aeration mode. This refers that the presence of the draft tube and the
propeller doesnt contri ute to an increase of the power consumption.
Comparing the power consumption relation with mixing mode (down-pumping), the
power number in that case was relatively constant and it was around 0.6. In mixing
mode the RTP propeller operates only in liquid phase with low gas presence, therefore
the power number has close values for different Reynolds numbers, is for that power
number can be referred by one value.
- 911 -

4.5
Whole system (Aeration)
Turbine+Propeller(Aeration)
Turbine alone(Aeration)
3.5
Propeller+DT(Mixing)
Np, (-)
3
2.5
2
1.5
1
0.5
0
50000
100000
150000
200000
Re, (-)
Figure (3.8): The effect of various impellers configurations on the power

consumption.
VI. Standard Aeration Efficiency (SAEb) and Standard Oxygen Transfer Rate
(SOTRb) for the Water Bulk Zone
The standard oxygen efficiency of the water bulk mass transfer zone SAEb refers to
the relationship between the power consumption with achieved oxygen transfer rate
and calculated by equation 2.26, the SOTRb was calculated by equation 2.25. It is
noticed that SAEb was increased according to the elevation of rotation speed (See
Figure 3.9). The SAEb was affected by the increasing of impellers rotation speed for
the different tested geometrical configurations. For the whole system configuration,
when N was elevated from 1.67 rps to 3.33 rps the SAEb value was approximately
doubled this behavior was repeated for all tested geometrical configuration. Starting
from N=2.08 the whole system configuration shows competitive behavior with
turbine alone configuration till higher N >= 2.92 rps, where it looks clearly the whole
system achieves highest SAEb this behavior can explained that the contribution of
draft tube in redirecting the flow up-ward. As illustrated in Figure 3.9 the SAEb
relation with N wasnt similar for all tested configurations. At lower N levels it seems
that the propeller and draft tube have no contribution in the SAEb. With higher N
levels, the turbine and propeller performance become closer to turbine alone
configuration. But in general it can be notice that the draft tube presence looks
somehow necessary to assist the propeller in its work of pushing more gas liquid
dispersion to turbine intake region. The other advantage of draft tube is its power
consumption was quite low. On the other hand it was found that the SAEb didnt
changed effectively and proportionally with the turbine blades submergence or water
level variations. From Table 3.1 it was noticed that higher turbine blades submergence
or higher water level leads to higher accomplished SAEb till reaching the turbine
- 919 -

blades submergence ratio of S/W=1 After this ratio it is clear that the SAE b was
relatively kept its value for higher S/W ratios. SAEb was decreased gradually with
further increasing of the submergence of turbine blades by elevating the water level
till it reach lowest SAEb at S/W=2.25. The reason for this behavior is related with the
shape of water spray droplets projection, as it was observed that the spray form was
deformed at higher water level applied water level h/D = 1.63 (S/W=2.25) and the
SAEb value is lowered (See Table 3.1).
The achieved SAEb with the applied surface aeration system that shown in Tables 3.1
to 3.5 are considered successful and efficient when they are compared with other
surface aerators (See Tables 1.1 and 1.2), where the SAEb for slow speed surface
aerator that implement turbine is equal to 1.5 kgO2.kW-1h-1 (Duchene and Cotteux,
2002).
The standard oxygen transfer rate SOTRb was increased with increasing impellers
rotation speed or water level for the three examined configurations (See Tables from
3.1 to 3.6) and Figure (3.10). The obtained results of SOTRb can be explained as same
manner for kla results discussed before.
3
Whole system
Turbine + Propeller
SAEb, ( kg O2 kW-1 h-1)
Turbine alone
2.5
1.5
1.5
2.5
N, (s-1)
3.5
Figure (3.9): The relationship between the standard aeration efficiency SAEb for three
aeration mode configurations (Whole System, turbine + propeller and turbine alone)
with the impeller rotation speed, D/Tv=0.24, C/Tv=0.31, h/D=1.47, S/W=1.
- 919 -
Table (3.1): The SAEb and SOTRb for different water level ratios (h/D) and turbine
blade width to blades submergence ratio (S/W), D/T=0.24, C/T=0.31, Cpr/T=0.2,
N=2.803 rps, (Whole system Configuration).
Power
(watt)
2.932
5.418
7.552
9.148
9.698
10.025
h/D
(-)
1.37
1.42
1.47
1.53
1.58
1.63
S/W
(-)
0.17
0.58
1.00
1.42
1.83
2.25
(kla)T
(s-1)
9.80E-04
2.09E-03
2.84E-03
3.02E-03
3.10E-03
2.79E-03
(kla)10 C
(s-1)
8.38E-04
1.72E-03
2.52E-03
2.64E-03
2.74E-03
2.51E-03
SOTRb
(kgO2.h-1)
0.00448
0.00953
0.01451
0.01572
0.01686
0.01596
SAEb
(kgO2.kW-1h-1)
1.5265
1.7591
1.9208
1.7185
1.7386
1.5923
The SOTRb and SAEb were decreased when turbine blades submergence increased
S/W from 1.83 to 2.25 (water level ratio h/D was increased from 1.58 to 1.63), where
the lowest SOTRb took place (0.01596 kgO2/h). As turbine blades submergence was
increased to S/W= 2.25 (water level ratio was increased to h/D=1.63) that will cause
deformation of water spray shape created by turbine blades, so variation in contact
interfacial area was available.
The standard oxygen transfer rate SOTRb was increased with increasing impellers
rotation speed or water level for the three configurations examined that can be
explained as same manner for kla results discussed before (see Tables from 3.1 to
3.4). From Table 3.2, it can be observed that SOTRb is higher when higher water
height level used. Table 3.4 elucidate the relation between the experimental results of
SOTRb and with SAEb; where high SOTRb was reached with high surface aeration
efficiency SAEb was achieved, but this increasing is applicable for water level ratio up
to 1.58(h = 0.3 m).
Table (3.2) demonstrates that standard oxygen transfer rate for water bulk mass
transfer zone SOTRb is acted same as the standard aeration efficiency SAEb. SOTRb
was increased according to increasing of impellers rotation speed N, when N was
increased that will leads to more air bubbles entrapped into well mixed water bulk, so
larger contact interfacial area between the water bulk and air bubbles is provided. In
other hand water droplets propelled by turbine blades take longer travel time before
reaching impingement at water surface, this meaning larger contact area between
droplets and atmospheric air; as a consequence higher dissolved oxygen concentration
was accomplished.
- 919 -

Table (3.2): The SAEb and SOTRb for different impellers speed, D/T= 0.24, C/T=
0.31, Cpr/T= 0.2, h/D = 1.47 (Whole system Configuration, Aeration mode).
Power
(watt)
N
(rps)
(kla)T
(s-1)
(kla)10 C
(s-1)
SOTRb
(kgO2.h-1)
SAEb
(kgO2.kW-1h-1)
4.503
1.67
1.28E-03
1.1E-03
0.0063
1.4078
7.552
2.08
2.84E-03
2.52E-03
0.0145
1.9205
10.603
15.184
2.5
3.33
4.25E-03
7.89E-03
3.79E-03
7.01E-03
0.0218
0.0403
2.0570
2.6540
Tables (3.3, 3.4 and 3.5) are representing the SOTRb with different impellers rotation
speeds for three different configurations with the achieved oxygen mass transfer
coefficient at actual and standard conditions. From Figure 3.10 and from tables 3.3,
3.4 and 3.5 it can be noticed that the SOTRb is lowered when both the draft tube and
the propeller were removed or when the draft tube alone was removed from the
system. The explanation for this is similar to that mentioned earlier in this chapter for
the oxygen mass transfer. The same behavior was observed for spacing changing
between the two impellers and for the turbine alone implementation cases
respectively. The SOTRb was increased with impellers rotational speed (N)
increasing; this is logic as with higher impellers rotation speed higher oxygen mass
transfer coefficient is obtained.
Table (3.3): The SAEb and SOTRb for different impellers speed, D/Tv= 0.24, C/Tv=
0.31, Cpr/Tv= 0.2, h/D = 1.47 (Turbine and propeller configuration).
Power
(watt)
4.409
6.991
10.053
14.472
N
(rps)
1.67
2.08
2.5
3.33
(kla)T
(s-1)
1.23E-03
2.38E-03
3.77E-03
7.23E-03
(kla)10 C
(s-1)
1.10E-03
2.13E-03
3.35E-03
6.42E-03
- 919 -
SOTRb
(kgO2.h-1)
0.00631
0.01224
0.01926
0.03693
SAEb
(kgO2.kW-1h-1)
1.4318
1.7515
1.9154
2.5516

0.045
Whole system
0.04
Turbine + Propeller
Turbine alone
SOTRb, ( kg O2 h-1)
0.035
0.03
0.025
0.02
0.015
0.01
0.005
0
1.5
2.5
N,
3.5
(s-1)
Figure (3.10): The relation between the standard oxygen transfer rate SOTRb for three
aeration mode configurations (Whole System, turbine + propeller and turbine alone)
with the impeller rotation speed, D/Tv=0.24, C/Tv=0.31, h/D=1.47, S/W=1.
Table (3.4): The SAEb and SOTRb for different impellers speed, D/Tv= 0.24, C/Tv=
0.31, h/D = 1.47, (Turbine alone configuration).
Power
(watt)
4.472
6.82
9.943
14.263
N
(rps)
1.67
2.08
2.5
3.33
(kla)T
(s-1)
1.28E-03
2.37E-03
4.07E-03
6.93E-03
(kla)10 C
(s-1)
1.21E-03
2.22E-03
3.75E-03
6.38E-03
SOTRb
(kgO2.h-1)
0.00697
0.01275
0.02154
0.03668
SAEb
(kgO2.kW-1h-1)
1.5588
1.8700
2.1664
2.5715
Table (3.5): The effect the spacing between impellers on SOTRb and SAEb, D/Tv=
0.24, C/Tv= 0.31, h/D = 1.47, Sp=8.3 mm.
Power
(watt)
4.932
7.448
10.352
15.038
N
(rps)
1.67
2.08
2.5
3.33
(kla)T
(s-1)
1.40E-03
2.50E-03
4.00E-03
7.20E-03
(kla)10 C
(s-1)
1.25E-03
2.22E-03
3.52E-03
6.33E-03
- 919 -
SOTRb
(kgO2.h-1)
0.00720
0.01277
0.02024
0.03643
SAEb
(kgO2.kW-1h-1)
1.46
1.71
1.96
2.42

3.1.3. The Modeling
In order to determine and improve the oxygen mass transfer and power consumption
for surface aeration systems with larger scales in depending on the obtained
experimental results collected from the implemented water treatment tank model in
this work. The most affecting parameters during the surface aeration process were
correlated in two models that were derived for oxygen mass transfer and power
consumption in the bulk mass transfer zone. The influencing parameters such as
operational, property and geometrical parameters are considered as independent
variables, and are related with the dependent variable of mass transfer parameter or
power consumption.
It worth to mention that the number of the considered experimental points that applied
to derive the models for the transferred oxygen and power consumption are actually
the mean for the number of the results that obtained for the similar conditions at each
point.
I. Mass Transfer
From the previous studies that dealt with surface aeration dimensional analysis, it was
found that there are various trends to correlate different relevant variables. (Fuchs et
al., 1971) have correlated the mass transfer coefficient kla with the volumetric power
consumption (P/V) for tank volumes of (V > 200 liter). (Zlokarnik, 1979) proposed
two correlation models for both mass transfer and power consumption for various
types of surface aerator. The main distinction in the correlations made by (Zlokarnik,
1979) is the mass transfer parameter was represented as a dimensionless parameter
that called the sorption number Y, which represents both the transferred oxygen and
turbulence intensity (Y=klaV(/g2)1/3D3). (Zlokarnik, 1979) has related in his model
the oxygen mass transfer with the water surface flow characteristics. (Zeybek et al.,
1997) have applied the Box-Wilson method to optimize the values of mass transfer
correlation for both linear and non-linear models of kla in general application of
aerated agitated tanks. (Patil et al., 2004) performed various tests for different types of
surface aerators and they emphasized on the geometrical configuration that affecting
parameters such as liquid height and volume, tank diameter and impeller clearance.
The model was for optimum values of kla with respect to impeller submergence.
(Cancino, 2004; Chandrasekharan and Calderbank, 1981; Chisti and Jauregui-Haza,
2002; Deshmukh and Joshi, 2006; Kumar et al., 2010; Moulick et al., 2002; Rao,
1999) have attempted to correlate the oxygen transfer and / or the power consumption
for diverse affecting variables (for the aeration process), each one of them depending
on its implemented geometrical configuration and proposed relevant geometric,
material and process parameters.
- 919 -

The mass transfer model was developed in depending on the experimental results and
the previous performed studies. Normally in the surface aeration process there are
geometrical, operational and materials (physical and chemical properties) parameters
that govern the system performance. The operational parameters like the rotational
speed, the gravity of acceleration, the water level in the tank and power consumption
or many other parameters. The geometrical parameters are the turbine and tank
diameters. While the parameters of the chemical and physical properties are the water
viscosity and density. All these parameters may refer to the intensity of turbulence
inside the tank and at the water surface. The number of these parameters is different
according to each case, but generally they are the same essential parameters that work
for all the surface aeration cases. For example, the operational parameter that refers to
water wave action generated by the turbine or refers to flow regime. Always the effect
of the gravitational acceleration force, the rotational speed and the power consumed
are taken in account in the surface aeration with considering that the aerated fluid
properties; the water viscosity and density are crucial for any attempt to build a
surface aeration model. It is not restricted to use one normalized mass transfer
parameter, but it depends on the system implemented distinction, in addition to many
variables in relation with the geometrical configuration and operation condition. In
accordance to experimental results, it is found the most realistic independent variables
can affect the oxygen transfer are
(3.1)
The geometrical ratios were chosen upon the standard design ratios that generally
implied in the surface aeration process. By applying Buckingham theory the
equation (3.1) converted to the following relation:
(
(3.2)
Since the surface aeration process predominantly within the turbulent flow state, so
the Reynolds No. can be ignored, where for air-water submergence aerator the value
of Re is always ( > 104), so the Re, is considered irrelevant to the process objective
(Zlokarnik, 1979).
So equation (4.5) was reformed as:
(
(3.3)
Equation (3.3) was solved to determine the values of the constants by applying
multiple non-linear regressions. The following correlation was developed:
- 919 -

(
For surface aeration cases the constant (K1) which have the value 0.0322 in equation
3.4 it normally represents the effect of the system geometry, but it in our case in all
implemented geometrical configurations have the same effect because actually we
have kept the essential geometry for all the tested configurations, where the turbine
and tank geometries didnt changed during experimental runs.
From equation (3.4) it is noticeable that the water level has higher effect on the
dimensionless mass transfer parameter (kla/N) than the wave condition that
represented by Fr or the power consumption that represented by Np. This elevated
value for h/D parameters shows that the oxygen mass transfer is highly related with
the liquid volume and flow circulation condition in the tank
Thee mass transfer correlation (equation 3.4) is applicable for the three ranges, (P/V,
watt/m3) = (22 -100), (Fr) = (0.054 0.214), (h/D) = (1.37 - 1.58) with respect to the
applied geometry in the experimentation.
The comparison between the predicted values from applying above correlation
(equation 3.4) with the experimental results that obtained from 74 experimental runs
showed that the coefficient of determination of (0.985), which is accepted in the
frame of the error as shown in Figure 3.11a.
II. Power Consumption

The power consumption parameter in the derived model for the surface aeration is
related with the geometrical, material and dynamic parameters same as it was
explained with oxygen transfer correlation procedure. Furthermore most of the
dependent parameters are similar with that implied in the mass transfer correlation,
except that they are related directly with molecular diffusivity of oxygen in water.
During the past decades various attempts are developed to correlate the consumed
power in aeration applications that relied on each study condition of geometric and
operation specifications. (Sano and Usui, 1985) proposed several correlations for the
aerated tank, he mainly considered that the power is effected by geometrical
configuration, (Wu, 1995) regarded the Froude number as the main representative
independent variable for the power consumed by the aerator.
The attempt was conducted to drive a power consumption correlation with looking at
the experimental results as a prime source to have an idea about the affecting factors
on the consumed power. It was found that the power is a function of following
independent parameters can be represented as:
P = f (N, h, Tv, S, g, D, , )
(3.5)
- 919 -
By applying Buckingham theory of equation (3.5), the following correlation

developed:
(
(3.6)
Since the flow regime is always kept in turbulent region (Re > 104), the Re is
considered irrelevant to the process objective same as mass transfer model. Equation
3.6 was solved to determine the values of the constants by applying multiple nonlinear regressions, the following correlations developed:
(
The power correlation applicable for the ranges (22< P/V, watt/m3 <100), (1.37< h/D
<1.63) and (0.054<Fr<0.214) according to the applied geometry in the
experimentation. The comparison of predicted values from power consumption
correlation with experimental results values obtained from 58 experimental runs
found that the coefficient of determination of (0.972), (See Fig .3.11b).
2.5
2
Predicted (kla/N)*10-3
R = 0.985
1.5
0.5
0.5
1
1.5
2
-3
Experimental (kla/N)*10
(a)
- 911 -
2.5

6
Predicted Np
R = 0.972
4
3
2
1
0
Experimental Np
(b)
Figure (3.11): (a) The comparison between the kla/N, predicted by the correlation
model (Eq. 3.4) with the experimentally resulted kla/N, (b) The Comparison between
the (Np) values predicted by the Eq. 3.7 with the experimentally resulted (Np) values.
3.2. Spray Mass Transfer Zone

3.2.1. Introduction
The concept of dividing the mass transfer operation in the surface aeration process
into two zones above and under water surface was identified in many previous works
for the surface aeration and other types of aeration (Demoyer et al., 2003; McWhirter
and Hutter, 1989). As it was described in chapter two the oxygen transfer in the water
spray zone is related with the droplets form and condition that created by the aeration
turbine blades rotation at the water surface, where they propel the water droplets into
the atmospheric air with relatively high discharge velocity. The oxygen transfer
occurs with droplets creation till their clash on the water surface, wherein the second
mass transfer zone boundary begins.

I. Impellers Rotation Speed Effect
The effect of impellers rotation speed, N was tested for the whole system
configuration, D/T= 0.238, C/T= 0.313, d/T= 0.15, Cpr/T= 0.2, h/T= 0.35, df/T =
0.188. For the whole system configuration, the turbine and the RTP propeller are
mounted on the shaft and the RTP propeller is placed inside the draft tube. The
impellers rotation speed effect was tested with keeping the other operational
- 991 -

parameters constant, in order to interpret influence of N on the water spray droplets
oxygen mass transfer zone coefficient klad. The N was varied between (1.67 3.33
rps). The applied range represents the limits where the purpose to unequivocally
establish a correct and affective spray form of water droplets for the given operational
conditions and geometrical configuration, where at the impeller rotation speed lower
than 1.67 rps there is no spray creation, upper the limit of 3.33 rps the water spray
form begins to be deformed from its regular form and part of the droplets began to hit
vessel wall, wherein the benefit of droplets impingement with water surface was lost.
The repeatability of water spray zone aeration efficiency experiments was tested; two
experimental runs were repeated and then to identify the calculated spray zone
aeration efficiency relevant to these results (See Figures 3.13 to 3.17).
I.1. Aeration Efficiency for the Water Spray Zone

Figures 3.12 to 3.15 illustrate the dissolved oxygen concentration DO profile during
surface aeration in two different positions (same to that illustrated in Figure 2.15 b) in
the water bulk inside the tank and in the water spray for four impellers rotation speed
levels, where two experimental runs were performed for each level. It can be noticed
that the dissolved oxygen concentration DO in all positions is increased progressively
with the time, but in different manners. The DO concentration gradient in the droplets
was occurred in a pattern that is higher than that occurred in water bulk as the
saturated equilibrium DO concentration level in the droplets corresponds to the wetbulb temperature of atmospheric air, which is higher than the saturated equilibrium
DO of water bulk for the given operation conditions of water bulk temperature, air
temperature and atmospheric air relative humidity.
As it is shown from the Figures 3.12 to 3.16 the DO is gradually proceed to attain the
saturation level, where practically the DO isnt attained the 100% saturation due to the
various limitation that hinder the theoretical DO saturated equilibrium level to be
achieved. Among these limitations in the practical applications is the employed water
quality, where the DO saturation level is different for different water qualities. From
Figures 3.12 to 3.16 it was observed that for all impellers rotation speed levels
performed for same water level ratio h/D and other geometrical parameters. The DO
behavior is in the way of higher impellers rotation speed leads to shorter time needed
to reach its saturation level (highest DO concentration) as seen in the results.
- 999 -

12
10
DO (mg l-1)
8
C
1
CLt
Run1
Lt Run
Cdt Run1
CLt
Run2
C
2
Lt Run
Cds Run1
Cdt Run2
Cds Run2
200
400
600
800
Time (s)
1000
1200
1400
1600
Figure (3.12): Dissolved oxygen concentration profile with time of experiment in the
water bulk zone (Cdt), and the water spray zone (CLt) at N = 1.67 rps for two
experimental runs.
10
DO (mg l-1)
CLt
CLt Run1
Cdt Run1
CLt Run2
CLt
Run2
Cdt Run2
Cds Run1
Cds Run2
200
400
600
800
1000
1200
1400
Time (s)
experimental runs.
- 999 -
10
DO (mg l-1)
8
CLt
Run 1
C
Lt Run1
Cdt Run1
CLt
Run2
C
Lt Run2
Cdt Run2
Cds Run1
Cds Run2
100
200
300
400
500
Time (s)
600
700
800
900
experimental runs.
12
10
DO (mg l-1)
8
CLt
Run1
CLt Run1
Cdt Run1
Cds Run1
CLt
Run2
CLt Run1
Cdt Run2
2
0
Cds Run2
100
200
300
400
Time (s)
500
600
700
experimental runs.
- 999 -

10
10
Run 1
Run 1
Run2
y = 0.776x + 2.4913
R = 0.9884
Run 2
y = 0.791x + 2.3126
R = 0.9928
Cdt (mg l-1)
Cdt (mg l-1)
y = 0.772x + 2.2727
R = 0.996
y = 0.745x + 2.5914
R = 0.9909
Esp
Run 1= 0.224
Run 2= 0.209
2
0
CLt (mg l-1)
Esp
2
0
10
Run 1= 0.228
Run 2= 0.255
(a) N=1.67 rps
10
(b) N = 2.08 rps

12
12
Run 1
Run 2
Run 1
10
8
Cdt (mg l-1)
Cdt (mg l-1)
y = 0.6223x + 4.0444
R = 0.9912
y = 0.6689x + 3.0932
R = 0.9925
4
Esp
y = 0.518x + 5.0191
R = 0.9956
y = 0.529x + 4.4901
R = 0.9968
6
4
Esp
Run 1= 0.471
Run 2= 0.482
Run 1= 0.331
Run 2= 0.379
Run 2
10
CLt (mg l-1)
10
CLt (mg l-1)
10
CLt (mg l-1)
(c) N = 2.5 rps
(d) N = 3.33 rps
Figure (3.16): The linear regression correlation of water spray zone aeration
efficiency by plotting Cdt versus CLt for various rotation speeds.
It is noticed from Figures 3.12 to 3.15, the impellers rotation speed gradually
increasing from 1.67 rps to 2.08, 2.5 and 3.33 rps causes the increasing of oxygen
mass transfer coefficient klad in the spray zone. The increased oxygen mass transfer
coefficient klad means that the DO inside the droplets reaches its higher concentration
level (nearest to air wet-bulb saturated oxygen condition) in shorter time, but in the
same time the flight time for the water droplets from turbine blades to the water
- 999 -

surface is longer since these water droplets are thrown farther with increasing
impellers rotation speed.
Figure 3.16 illustrates the linear correlation obtained by plotting the dissolved oxygen
DO concentrations of the water droplets directly before they impinge water surface in
function of the DO concentration of the water bulk inside the tank. Applying the least
squares method of the DO concentrations best fit yields the water spray mass transfer
zone aeration efficiency Esp (The Esp was calculated by applying equation 2.31). When
the impellers rotation speed was increased from 1.67 to 3.33 rps; the water spray zone
aeration efficiency increased from 0.21 to 0.48. This particular behavior depends on
the fact that water droplets exposure time to the atmospheric air was prolonged and
consequently the amount of oxygen transferred from the atmospheric air to the water
droplets was increased too. Since the exposure time is same the flight time for the
droplets, so that higher impellers rotation speed results in longer flight time, but this
relation is applicable just up to 3.33 rps, after this limit desired shape of the water
droplets spray begins to be deformed, as consequence the interfacial contact area
between the air and water droplets is lessened.
Also Figure 3.16 shows that when the impellers rotation speed has been increased
from 1.67 rps to 2.08 rps, the spray zone aeration efficiency was slightly increased
where the average spray aeration efficiency was elevated from 0.215 to 0.245. With
further increasing of impellers rotational speed to 2.5 rps the spray aeration efficiency
was noticeably increased to its average value of 0.335. With continuation of impellers
rotation speed elevation to 3.33 rps it seems the corresponding related spray mass
transfer zone aeration efficiency was remained progressively increasing to its average
value 0.477. From these experimental results it was found that alteration of impellers
rotation speed has low effect on the water spray zone aeration efficiency up to 2.08
rps, while its effect on the water spray zone aeration efficiency was apparently
remarked after this limit.
I.2. Spray Zone Mass Transfer Coefficient (klad)

Right after droplets flight time was calculated for the four different rotation speed
levels and given operation condition, the klad was determined by applying equation
(2.30). The resulted klad, the droplets flight time and other related parameters for
different rotation speed are shown in Table 3.6.
Table 3.6 shows at the first glance that water spray droplets zone has oxygen mass
transfer coefficients klad much higher than of mass transfer coefficients in water bulk
mass transfer zone kla. The majority of oxygen mass transfer actually happens in the
spray zone (droplets) due to high interfacial area and turbulent conditions. This agrees
with (Yeh and Rochelle, 2003). In the surface aeration the bulk oxygen mass transfer
zone can be considered as a complementary or supplementary zone, where the
- 999 -

accompanied mixing or agitation process appears in order to accomplish distribution
and diffusion of dissolved oxygen in the entire tank. From Figure 3.17 it can be found
that the dependence of klad was increasing with impellers rotation speed N increasing
for the whole system configuration.
Table (3.6): The spray droplets mass transfer zone of klad, flight time tf and the related
measured parameters, (of 100% submerged turbine blades, H = zero), h/D = 1.47,
S/W = 1.
N
Tbulk
(oC)
(mg L-1)
(mm)
(mm)
Rm
Esp
(Esp)20
1.67
1.67
2.08
2.08
2.50
2.50
3.33
3.33
15.8
15.2
16.1
15.9
16.5
16.0
16.3
15.9
10.7
10.3
9.6
9.85
9.5
10.07
9.78
10.3
15
15
20
20
35
35
45
45
110
110
130
130
170
170
280
280
0.224
0.209
0.228
0.255
0.331
0.379
0.471
0.482
0.242
0.229
0.245
0.275
0.352
0.405
0.498
0.513
(rps)
Cds
Ym
tf
klad
0.11
0.11
0.13
0.13
0.17
0.17
0.19
0.19
2.305
2.131
2.000
2.264
2.365
2.802
3.351
3.462
(s)
(s-1)
4
3.5
klad, (s-1)
3
2.5
2
1.5
1
1.5
2.5
N,
3.5
(s-1)
Figure (3.17): The relation between the oxygen mass transfer coefficient klad and
impellers rotation speed and in the spray zone, S/W = 1.
- 999 -

I.3. Surface Aeration Water Spray Discharge Velocity and Volumetric Flow Rate
For eliminating the errors of interpretation that may occur when the whole system was
used during the energy balance calculation (i. e. the lower propeller and draft tube
were placed under the turbine and their contribution in the power consumption is
mainly for mixing purposes inside the tank, where that doesnt have any relation with
the droplets generation outside the water surface), the power consumption was
measured only with presence of the turbine to assure all the energy was converted to
propel water droplets into atmospheric air. For the case of turbine alone it was noticed
that spray configuration didnt changed significantly, which means the presence of
propeller and draft tu e doesnt have noticea l e effect on the spray configuration. By
applying equation (2.43) the water spray flow rates were calculated for various N
levels as shown in Table 3.7.
To evaluate the accuracy of the equation 2.43, the calculated water spray flow rates by
this equation were compared with those obtained by the PIV technology at the same
turbine rotation speeds and it is found that the equation 2.43 can predict the flow rates
with an average error less than 2.0%
Table (3.7): Water spray droplets velocity and volumetric flow rate of given operation
and geometric configuration (whole system) for various rotation speed levels.
(rps)
(mm)
Ym
(mm)
Rm
(mm)
Rsp
Esp
1.67
1.67
2.08
2.08
2.50
2.50
3.33
3.33
15
15
20
20
35
35
45
45
85
85
115
115
175
175
270
270
180
180
210
210
270
270
365
365
0.224
0.209
0.228
0.255
0.331
0.379
0.471
0.482
tf
Vsp
(s)
(Watt)
(m s-1)
(L h-1)
0.111
0.111
0.128
0.128
0.169
0.169
0.19
0.19
4.472
4.472
6.820
6.820
9.943
9.943
14.263
14.263
1.71
1.71
1.76
1.76
1.80
1.80
2.14
2.14
11010
11010
15851
15851
22094
22094
22422
22422
From the first glance on the results illustrated in Table 3.7, it is obvious that the
droplets spray flow rate was increased when the turbine blades rotation speed was
increased, this agrees with (Takase et al., 1982), this relation can be easily explained
as the higher rotation speed leads to higher water droplets were projected through the
air by the turbine blades. The droplets velocities in the spray at the impingement
position at water surface is approximately considered equal along the water spray
from turbine blades tip to plunging point and the energy consumption within water
spray path was ignored since the droplets flight time is extremely short. With
elevating the turbine rotation speeds it was noticed that the spray velocity was
increased, this relation was achieved with higher rotation speed or in other word with
higher power consumed, so higher mechanical energy was converted to a kinetic
- 999 -

energy and delivered into water bulk, which in turn causes that the water droplets to
be discharged faster from turbine blade tip through the atmospheric air. It is worth
mentioning that actually the occurred droplets velocities are slightly different
according droplet position in the spray, where the droplet position in the spray
depends on its projection position from blade tip. For the droplets velocity
determination purposes, only the maximum spray height was considered in order to
simplify the calculation since the effect of this variation generally is limited (Huang et
al., 2009; McWhirter et al., 1995).
I.4. Spray Zone Oxygen Transfer Rate (OTRsp)

As mentioned before the fast water droplets in the spray travel with very short flight
time, where only the droplets surface has higher dissolved oxygen DO concentration,
while the DO concentration becomes lower for further deeper levels inside the
droplets. Despite the short flight time of droplets, a high mass transfer rate and may
be also a high heat transfer rate can take place during the operation, wherein the outer
surface layer temperature of the droplets becomes equal to the wet-bulb temperature.
The remaining important factors for evaluating purposes in the droplet spray mass
transfer zone is the overall oxygen mass transfer rate (OTRsp) that occurred along
water spray by contacting droplets surface with atmospheric air. After determination
of water spray discharge velocity Vsp and flow rate Q, OTRsp can be calculated easily
by applying the equation (2.34) (Chern and Yang, 2004; McWhirter et al., 1995).
The OTRsp values obtained by applying equation (2.34) are illustrated in Table 3.8; in
the view of the fact that higher oxygen transfer rate takes place with higher rotation
speed employed. The OTRsp results agree with those obtained for Esp, it is obvious
that OTRsp in the water droplets zone is high.
Table (3.8): Spray zone oxygen transfer rate OTRsp variation with impellers rotation
speed.
N
Esp
1.67
1.67
2.08
2.08
2.5
2.5
3.33
3.33
0.22
0.21
0.23
0.26
0.33
0.38
0.47
0.48
(rps)
tf
Vsp
OTRsp
(s)
(m s-1)
(L h-1)
(gO2 h-1)
0.111
0.111
0.128
0.128
0.169
0.169
0.19
0.19
1.71
1.71
1.76
1.76
1.80
1.80
2.14
2.14
11010
11010
15851
15851
22094
22094
22422
22422
25.649
23.011
33.972
39.005
68.012
80.973
99.050
107.642
- 999 -

Table 3.8 and Figure 3.18 show that the oxygen transfer rate in the spray zone
increases with increasing of impellers rotation speed N. Increasing N leads to the
water droplets take longer time in their travel before reaching impingement position at
water surface, in other word more time is available for the oxygen in atmospheric air
to transfer to the water droplets. A higher dissolved oxygen DO concentration in these
droplets is accomplished. The difference in the resulted OTRsp of each experimental
run for the same N may occur due to the variance in operation conditions such as the
air and water bulk temperatures effect.
120
OTRsp, (gO2 h-1)
100
80
60
40
20
0
1.5
N, (s-1)
2.5
3.5
Figure (3.18): The relation between the oxygen transfer rate OTRsp and impellers
rotation speed and, in the spray zone, S/W = 1.
I.5. Contribution Percentage of the Spray and Bulk Zones in the Overall Mass
Transfer Operation
To have complete view for the oxygen mass transfer for all surface aeration process
and to realize the contribution of spray mass transfer zone in the overall oxygen mass
transfer operation, the oxygen mass transfer rate in the water bulk mass transfer zone
OTRb is needed to be calculated. OTRb was determined by applying the equation
(2.24) (See Appendix II).
- 991 -

Table (3.9): The oxygen mass transfer rates for spray and bulk zones OTRb and OTRsp
with the percentage of contribution.
(rps)
(s)
tf
(gO2 h-1)
OTRsp
(gO2 h-1)
OTRr
OTRsp%
OTRb%
1.67
1.67
2.08
2.08
2.50
2.50
3.33
3.33
0.11
0.11
0.13
0.13
0.17
0.17
0.19
0.19
25.649
23.011
33.972
39.005
68.012
80.973
99.050
107.642
6.753
6.810
10.183
10.214
20.621
19.935
35.307
35.541
79.16
77.16
76.94
79.24
76.73
80.02
73.72
75.18
20.84
22.84
23.06
20.76
23.27
19.98
26.28
24.82
Results from Table 3.9 and Figure 3.19 show that for different impellers rotation
speed, N levels (1.67-3.33 rps); the spray zone oxygen mass transfer OTRsp has a
contribution range between (80.02% - 73.72%) of the overall oxygen mass transferred
during the operation, which is higher than the contribution of bulk mass transfer zone
in the overall oxygen transfer rate. At higher N, the percentage of contribution for the
oxygen mass transfer rate in the bulk zone OTRb in the overall oxygen transfer rate
stays the same although the OTRb was increased with N elevation, where more
turbulent state of water surface was occurred by water droplets at the impingement
and plunging position due to more mechanical energy was converted to kinetic
energy. As a consequence of this condition more air bubbles were generated and
entrained into the water bulk, beside that more fine air bubbles were diffused into the
tank (Ozkan et al., 2006). So the range of the oxygen mass transfer rate of the bulk
zone is improved. The improvement of the OTRb assists the bulk zone to keep its
contribution somehow similar within the tested range of N. It is also important to
mention that both the OTRb and OTRsp were calculated at the same water bulk
temperature and ambient temperature for each run.
Percentage in the overall OTR, (%)
100%
80%
60%
OTRb%
40%
OTRsp%
20%
0%
1.67
2.08
2.5
3.33
N, (s-1)
Figure (3.19): The influence of the impellers rotational speed on the spray zone
oxygen transfer rate OTRsp and in the bulk zone oxygen transfer rate OTRsp
contributions the overall transfer rate OTR, S/W = 1.
- 991 -

II. The Effect of Turbine Blades Submergence
II.1. Water Spray Velocity and Volumetric Flow Rate
Table 3.10 illustrates that the calculated water spray flow rate Q was increased as the
dimensionless turbine blades submergence ratio S/W was increased, where with
higher water level higher amount of water droplets are propelled into atmospheric air
by the turbine blades. As it was mentioned before, water spray droplets velocity at
impingement position of water surface is approximately considered equal to the water
spray discharge velocity by turbine blades tip according to physical laws for
projectiles.
The energy consumption within the water spray path is ignored since the droplets
flight time is extremely short. By elevating the water level it was noticed that the
spray velocity remained somehow around (1.8 m/s).
For high turbine blades submergence levels (S/W >1) it was found that even though
more power was consumed by turbine blades with increasing (S/W) the spray
velocities were relatively close. The excess power was consumed to convert the
needed mechanical energy into kinetic energy and delivered into water bulk in order
to overcome the excess in water amount that to be projected into atmospheric air with
achieved spray velocity.
In the case of turbine blades were partially submerged in water S/W= 0.17
(particularly for minimum water height used h/D=1.37), it was noticed at the given
rotation speed (N=2.5 rps) the spray velocity is somehow lower, the reason for that
may rely on the fact that a part of the mechanical energy didnt converted to kinetic
energy into water bulk , and has been wasted by air resistance due to direct contact of
apparent blade part above water surface, which in turn causes less discharged water
droplets from turbine blade tip through the atmospheric air.
Generally the water spray volumetric flow rate is increased with increasing turbine
blades submergence level, as more water droplets can be propelled when higher S/W
is implemented. The spray velocities and volumetric flow rates for the water droplets
were calculated by equations (2.42) and (2.43).
The values of power consumption were measured for the turbine alone condition at
various S/W levels to eliminate the error in calculation because power consumed by
the lower propeller and to insure that all the measured power are equivalent to the
converted mechanical energy from turbine blades to kinetic energy in water bulk.
- 999 -

Table (3.10): The water spray velocity and volumetric flow rate for the whole system
geometrical configuration with various turbine blades submergence and water levels,
N= 2.5 rps.
S/W
(-)
0.17
0.17
0.58
0.58
1.00
1.00
1.42
1.42
1.83
1.83
h/D
(-)
1.37
1.37
1.42
1,42
1.47
1.47
1.53
1.53
1.58
1.58
Ym
Rm
Rsp
(mm)
(mm)
(mm)
23
23
26
26
33
33
35
35
38
38
55
55
85
85
170
170
180
180
187
187
150
155
180
185
265
265
275
275
282
282
Esp
(-)
0.182
0.185
0.217
0.223
0.353
0.340
0.362
0.360
0.392
0.384
tf
Vsp
(s)
(Watt)
(m s-1)
(L h-1)
0.093
0.093
0.109
0.109
0.164
0.164
0.169
0.169
0.176
0.176
4.214
4.214
6.754
6.754
9.943
9.943
11.624
11.624
12.872
12.872
1.74
1.74
1.79
1.79
1.81
1.81
1.82
1.82
1.80
1.80
10021
10021
15176
15176
21850
21850
25264
25264
28602
28602
II.2. Spray Zone Aeration Efficiency (Esp)

The oxygen mass transfer in water spray zone is affected by various factors. Beside
impeller rotation speed N it is necessary to consider the effect of turbine blades
submergence or water level on the spray mass transfer zone characteristics. The
turbine blades submergence primarily affects the droplets volumetric flow rate,
whereas the amount of the projected droplets by turbine blades is changed due to the
difference of water free surface height. The operation condition and geometrical
configuration were kept constant, while the turbine blades submergence was tested as
a dimensionless geometrical ratio of S/W, where the submergence of the turbine
blades S has been normalized with turbine blade width W. The submergence of
turbine blades S is referring to the distance from the water surface to the lower edge
of turbine blade. The ratio S/W was changed for different water levels of (0.17, 0.58,
1.0, 1.42, 1.83, and 2.25) (See Figure 3.20) wherein N was kept constant at 2.5 rps.
The five tested levels of the submergence ratio S/W correspond to the water levels
h/D of (1.37, 1.42, 1.47, 1.53, 1.58 and 1.63) respectively. The submergences of
turbine blades can be referred as submergence percentage by considering the water
level at the turbine blades upper edge as 100% of submergence, the submergence
percentage variation in the experimentation were (167% - 33%). At S/W < 1 (h/D
lower than 1.47) the blades are partially submerged, at S/W >1 (h/D higher than 1.47)
the blades are over submerged and at S/W =1 (h/D = 1.47) the water level is just at
blades upper tip (See Table 3.11). All the possible conditions were covered from
excessively submerged, totally submerged to partially submerged turbine blades. The
repeatability of water level height experimental runs was tested also, where spray
zone aeration efficiency is calculated twice for each experimental run.
- 999 -

Figure 3.20 shows the DO concentration profile in the water spray with the
experiment time at different turbine blade submergence levels with constant rotation
speed. When the turbine blade submergence increases, the DO reaches its higher
possible level near to its equilibrium state within shorter time. As higher turbine blade
submergence was used, more water droplets were projected into the atmospheric air,
which leads to more contact area between water droplets and air was provided. In the
same time water droplets were propelled farther toward the periphery space around
the turbine, where these droplets travel in atmospheric air longer flight time before
they impinge water surface.
From Figure 3.20 it was found that there is no noticeable difference of dissolved
oxygen DO profile at high blades submergence level S/W= 1.42 and 1.83 (h/D =1.53
and 1.58). It was noticed that from the level S/W= 1.42 (h/D 1.53), the regular form of
water spray begins to be deformed by creation of instability of its configuration
during the turbine rotation. At low turbine submergence levels S/W < 1 (S/W= 0.17
and 0.58) it was also observed that the DO profiles are similar and the DO takes
longer time to reach equilibrium level, which means with low submergence of turbine
blades (33% - 67%) correspond to (h/D =1.37 and 1.42), the amount of water droplets
that propelled by turbine blades is lower and flight time is shorter. In other word less
contact time is achieved between water in droplets and atmospheric air, so longer time
is needed to reach equilibrium DO concentration.
DO, (mg l-1)
10
8
S/W=0.17
S/W=0.58
S/W=1.00
S/W=1.42
S/W=1.83
4
2
0
200
400
600
800
Time, (s)
1000
1200
1400
Figure (3.20) Effect of the turbine blade submergence on spray mass transfer zone
dissolved oxygen concentration at rotation speed (N=2.5 rps) and turbine clearance
(C/Tv = 0.35).
From Figure 3.20 it was found also that elevating S/W ratio to 1.42 has no
enhancement on DO behavior, on the contrary higher S/W levels causes deformation
of water spray, in consequence the achieved DO is lowered. In the other hand at lower
S/W ratio (1.37 and 1.42) it seems that DO profiles with time are close together.
- 999 -

10
Run 2
Cdt (mg l-1)
y = 0.8177x + 2.8226
R = 0.9892
y = 0.8155x + 2.8716
R = 0.9804
Esp
Run 1= 0.182
Run 2= 0.185
2
0
Run 1
Run 2
Cdt (mg l-1)
10
Run 1
4
6
CLt (mg l-1)
y = 0.7823x + 2.2402
R = 0.9928
y = 0.7776x + 2.2324
R = 0.9936
Esp
Run1= 0.217
Run2= 0.223
10
5
CLt (mg l-1)
(b) S/W = 0.58
(a) S/W = 0.17

10
Run 1
Run 2
10
y = 0.6468x + 3.4357
R = 0.9938
Cdt (mg l-1)
Cdt (mg l-1)
12
Run 1
y = 0.6599x + 3.4707
R = 0.9915
4
Esp
Run1= 0.353
Run2= 0.34
2
0
4
6
CLt (mg l-1)
Run 2
y = 0.6377x + 3.9956
R = 0.979
y = 0.6398x + 3.8801
R = 0.9877
Esp
Run1= 0.389
Run2= 0.36
2
0
10
10
5
CLt (mg l-1)
10
(d) S/W = 1.42
(c) S/W = 1.00

10
Run 1
Run 2
Cdt (mg l-1)
y = 0.6076x + 3.8876
R = 0.9907
y = 0.6156x + 3.8405
R = 0.9946
Esp
2
0
Run 1= 0.41
Run 2= 0.42
0
10
CLt (mg l-1)
(e) S/W = 1.83
efficiency by plotting (Cdt) versus (CLt) for various turbine blades submergence, N =
2.5 rps.
- 999 -

Figure 3.21 shows the obtained linear regression correlations of the spray zone
efficiency. The DO concentration levels of droplets before the impingement at water
surface and of the water bulk inside the tank were plotted for different turbine blades
submergence ratios S/W (from 0.17 to 1.83), by applying the least squares of the DO
concentrations best fit to identify (Esp).
The increasing in spray zone efficiency depends as mentioned before on the fact that
water droplets flight time in atmospheric air is longer, the amount of oxygen
transferred from the atmospheric air to the water droplets is increased. As a
consequence it can be deduced that water droplets aeration efficiency Esp depends on
turbine blades submergence S and the ratio S/W, so that leads to higher Esp was
accomplished with higher S/W.
From Figures 3.21and Figure 3.22 it was founded that when S/W was increased from
0.17 to 0.58, the average spray zone aeration efficiency Esp is increased from 0.183 to
0.22. But it is noticed that at S/W equals 1.47 the Esp was effectively increased to its
average value of 0.35. With further elevating of S/W from 1.00 to 1.42 and 1.83 it
was found that the achieved Esp was increased respectively to their average values
0.36 and 0.388. From these experimental results it seems that elevation of S/W has
remarkable effect on Esp for higher than 1.00, while S/W has a relative effect on (Esp)
especially for the levels lower than 1.00. These experimental results show that Esp is
effectively lowered when the submergence level of turbine blades is less than 100 %
(S/W < 1), where Esp was changed from 0.346 to 0.183. The Esp is relatively affected
by turbine blades submergence level higher than 100% (S/W >1), where the Esp was
changed from 0.358 to 0.392. From Figure 3.20 it can be noticed as (S/W) is
increased, the water spray zone aeration efficiency is increased from 0.182 to 0.392.
0.45
0.4
Esp, (-)
0.35
0.3
0.25
0.2
0.15
0.1
0.5
1
S/W, (-)
1.5
Figure (3.22): The relation between the spray zone aeration efficiency Esp and turbine
blade submergence ratio, N = 2.5 rps.
- 999 -
II.3. Spray Zone Mass Transfer Coefficient (klad)

Table 3.11 and Figure 3.23 show that the oxygen mass transfer coefficient in the
water spray zone klad as mentioned before is also much higher than the mass transfer
coefficients in water bulk mass transfer zone kla for different turbine blades
submergence (See section 3.1 in this chapter). Table 3.11 illustrates the dependence of
klad on turbine blades submergence for whole system configuration at given impellers
rotation speed (N=2.5 rps), where the klad was increased from 2.16 1/s to 2.827 1/s
when S/W was increased from 0.17 to 1.83. It is important here to elucidate that the
water spray aeration efficiency Esp is relevant to S/W elevation, the klad showed also
the same behavior as a result of enlarging the interfacial contact area and flight time
elongation. The water droplets flight time was calculated by equation (2.35). The
standard water spray aeration efficiency at 20 oC was calculated by the equation
(2.32).
Table (3.11): The calculated water spray mass transfer zone (klad), droplets flight
time(tf) and some related measured parameters and turbine blades submergence and
water level, at constant rotation speed (N =2.5 rps).
S/W
(-)
0.17
0.17
0.58
0.58
1.00
1.00
1.42
1.42
1.83
1.83
h/D
(-)
1.38
1.38
1.42
1.42
1.47
1.47
1.53
1.53
1.58
1.58
Tbulk
Cds
Ym
Rm
(oC)
(mg L-1)
(mm)
(mm)
17.7
17.8
18.4
18.3
16.8
16.7
16.7
16.5
15.2
15.1
10.73
10.9
9.45
9.94
9.5
10.07
10.5
10.2
9.94
10.0
23
23
26
26
33
33
35
35
38
38
60
60
90
90
170
170
180
180
187
187
- 999 -
Esp
(-)
0.182
0.185
0.217
0.223
0.353
0.340
0.362
0.360
0.392
0.384
(Esp)20
(-)
0.190
0.193
0.224
0.230
0.373
0.360
0.383
0.382
0.424
0.417
tf
klad
0.093
0.093
0.109
0.109
0.164
0.164
0.169
0.169
0.176
0.176
2.160
2.199
2.244
2.449
2.655
2.534
2.659
2.641
2.827
2.785
(s)
(s-1)

(S/W), (-)
0
0.5
1.5
2.5
2
1.35
1.4
1.45
1.5
1.55
1.6
klad , (s-1)
2.75
2.5
2.25
(h/D), (-)
Figure (3.23): The relation between the spray zone oxygen mass transfer coefficient,
klad with the turbine submergence and liquid level impellers rotation speed, N = 2.5
rps.
II.4. Spray Zone Oxygen Transfer Rate (OTRsp)

Table 3.12 and Figure 3.24 illustrate the oxygen transfer rate of the water spray zone
(OTRsp) calculated by equation (2.34), where higher oxygen transfer rate takes place
with higher blade submergence used. The experimental results of OTR sp and Esp
elucidate a direct relation with them; high OTRsp was accomplished as high spray
zone surface aeration efficiency was obtained. It was indicated in the Table 3.12 that a
higher oxygen transfer rate at droplet spray zone was reached (106.514 g/h) when the
highest turbine blades submergence was used (S/W =1.83), while a lowest OTRsp took
place (16.232 g/h) with lower blades submergence level used (S/W =0.17). When h/D
or S/W was increased that causes more water droplets to be propelled into
atmospheric air, so larger contact interfacial area is available and also the longer flight
time during their travel before reaching impingement position at water surface was
fulfilled, so higher DO concentration in droplets is accomplished.
- 999 -

Table (3.12): Water spray zone oxygen transfer rate (OTRsp) variation with different
turbine blades submergence and water levels, (N = 2.5 rps).
S/W
(-)
0.17
0.17
0.58
0.58
1.00
1.00
1.42
1.42
1.83
1.83
h/D
(-)
1.37
1.37
1.42
1.42
1.47
1.47
1.53
1.53
1.58
1.58
Esp
(-)
0.182
0.185
0.217
0.223
0.353
0.340
0.362
0.360
0.392
0.384
(s)
tf
(m s-1)
Vsp
(L h-1)
OTRsp
0.093
0.093
0.109
0.103
0.164
0.164
0.169
0.169
0.176
0.174
1.74
1.74
1.79
1.79
1.81
1.81
1.82
1.82
1.80
1.80
10021
10021
15176
15176
21850
21850
25264
25264
28602
28602
16.232
16.407
29.211
30.424
71.346
67.975
85.511
85.221
106.514
104.340
0.5
(g h-1)
(S/W), (-)
120
1.5
2.5
OTRsp , (gO2 h-1)
100
80
60
40
20
0
1.3
1.4
1.5
1.6
(h/D), (-)
Figure (3.24): The relation between spray zone oxygen transfer rate OTRsp with the
turbine blades submergence and liquid level, N = 2.5 rps.
- 999 -

II.5. Contribution Percentage of the Spray and Bulk Zones in the Overall Mass
Transfer Operation
Table 3.13 and Figure 3.25 show the results for the comparison between the oxygen
transfer rates for both bulk and spray mass transfer zones in order to have a general
understanding on the overall oxygen mass transfer occurred in the system and its
distribution in the zones. At low turbine blades submergence (S/W=0.17) or low
water level (h/D = 1.37) the resulted spray mass transfer zone OTR sp has a
contribution range between (60.1% - 61.12%), which is higher than the contribution
of occurred re-aeration in water bulk mass transfer zone OTRb in the overall oxygen
transfer rate.
At moderate turbine blade submergence (S/W=0.58) or the water height level
(h/D=1.42), the contribution percentage of the spray oxygen mass transfer rate OTRsp
in overall oxygen transfer rate is higher and reaches (66.56%). With elevated turbine
blade submergence level S/W >=1 (h/D =1.47 and higher), the contribution of OTRsp
was high and reaches (78.86%) but it stays close to this value even though the h/D
was increased, which means that further elevation in water level for totally immerged
tur ine lade s doesnt affect the contri ut ion of water spray mass transfer zone in the
overall oxygen mass transfer in the system.
The oxygen mass transfer rate for water spray (droplets) mass transfer zone OTRsp
and bulk (re-aeration) mass transfer zone OTRb are calculated according to equations
(2.34) and (2.24) respectively.
Percentage in the overall OTR, (%)
100%
80%
60%
OTRb%
40%
OTRsp%
20%
0%
0.17
0.58
1.42
1.83
S/W, (-)
Figure (3.25): The influence of the turbine submergence on the spray zone oxygen
transfer rate, OTRsp and in the bulk zone oxygen transfer rate OTRsp contributions the
overall transfer rate OTR, N = 2.5 rps.
- 991 -
Table (3.13): Oxygen mass transfer rate and percentage of contribution for both water
spray and bulk (re-aeration) zones, (N= 2.5 rps).
S/W
(-)
0.17
0.17
0.58
0.58
1.00
1.00
1.42
1.42
1.83
1.83
h/D
(-)
1.37
1.37
1.42
1.42
1.47
1.47
1.53
1.53
1.58
1.58
tf
OTRsp
OTRb
OTRsp%
OTRb%
0.093
0.093
0.109
0.103
0.164
0.164
0.169
0.169
0.176
0.174
16.232
16.407
29.211
30.424
71.346
67.975
85.511
85.221
106.514
104.340
10.778
10.437
14.673
16.607
19.319
20.578
23.540
22.841
29.969
30.157
60.10
61.12
66.56
64.69
78.69
76.76
78.41
78.86
78.04
77.58
39.90
38.88
33.44
35.31
21.31
23.24
21.59
21.14
21.96
22.42
(s)
(g h-1)
(g h-1)
(%)
(%)
III. Effect of Propeller and the Draft Tube

The function of the propeller and the draft tube that positioned below the turbine is
mainly to enhance the distribution of dissolved oxygen in the water bulk. The
investigation of the effect propeller and draft tube presence on the water spray mass
transfer characteristics was performed by the determination of oxygen mass transfer
coefficient, aeration efficiency, surface aeration water velocity, volumetric flow rate
and oxygen transfer rate for the water spray mass transfer zone with turbine alone
geometrical configuration in order to find out if there will be any preliminary
difference occurred. The way to recognize the relative effects consider in testing the
mentioned parameters for turbine various rotation speeds, N for constant water level
ratio h/D=1.47 and turbine blades submergence ratio S/W = 1.00. The other operation
condition and geometrical configuration are also kept constant.
III.1. Spray Zone Aeration Efficiency (Esp)

The dissolved oxygen DO concentration profile in the water spray during surface
aeration of turbine alone configuration at four rotation speed levels is shown in Figure
3.26. As it was found in the whole system configuration, with turbine rotation speed,
N increasing the DO increases till it reaches its saturation level at wet-bulb of
atmospheric air. With higher N applied, more water droplets were propelled into the
atmospheric air and at the same time the water droplets DO concentrations are close
to those obtained in the whole system configuration, where DO reaches the
equilibrium concentration in shorter time with higher N used.
- 991 -

10
DO (mg l-1)
8
6
N=1.67 rps
4
N=2.08 rps
N=2.5 rps
N=3.33 rps
250
500
750
1000
1250
1500
1750
2000
Time (s)
Figure (3.26): The effect of the turbine rotation speed on the spray mass transfer zone
dissolved oxygen concentration at h/D=1.47and C/T=0.35, (Turbine alone
configuration).
The spray zone aeration efficiency Esp for the turbine alone configuration was
calculated by the linear regression for the Cdt versus CLt plot. Aeration efficiency Esp
was determined for the four different rotation levels by applying best fit with the least
squares method. The results are shown in Figure 3.27.
As in the whole system configuration, with higher N, higher water spray zone aeration
efficiency Esp is achieved. Figure 3.26 shows that when N was increased from 1.67
rps to 2.08 rps; the Esp was increased from 0.183 to 0.206. At moderate N=2.5 rps, the
aeration efficiency Esp was noticeably elevated to 0.292. At relatively elevated
N=3.33 rps, the Esp was found to be 0.356. From Figures 3.27 and 3.28 it is found that
the Esp is remarkably lowered when the propeller and the draft tube are removed.
The Esp values are relatively lower than the whole system configuration aeration
efficiencies for all tested N levels, for N= 1.67rps, the Esp was lowered from 0.21 to
0.183, while for N= 3.33rps, the Esp was lowered from 0.471 to 0.356. From these
comparisons it can be deduced that the presence of the propeller and draft tube has a
relative influence on the water spray efficiency Esp especially at higher rotation
speeds.
- 999 -

10
10
y = 0.817x + 1.5911
R = 0.998
y = 0.7949x + 1.8793
R = 0.9959
Cdt (mg l-1)
Cdt (mg l-1)
4
Esp= 0.206
Esp= 0.183
2
0
CLt (mg
l-1)
10
(a) N=1.67 rps
(b) N = 2.08 rps
10
10
y = 0.7076x + 2.6796
R = 0.9919
6
4
y = 0.6449x + 3.1303
R = 0.9762
Cdt (mg l-1)
Cdt (mg l-1)
Esp = 0.292
2
0
10
CLt (mg l-1)
6
4
Esp= 0.356
10
CLt (mg l-1)
10
CLt (mg l-1)
(c) N = 2.5 rps
(d) N = 3.33 rps
Figure (3.27): The linear regression of the spray zone aeration efficiency of the plot
(Cdt) versus (CLt) for various rotation speeds, (Turbine alone configuration).
The performances of both the draft tube and the propeller are to enable the system to
obtain an efficient mixing condition of the DO in the water bulk mass transfer zone
inside the aeration vessel, which has also a relative effect in the enhancing of the
aeration efficiency Esp in the spray mass transfer zone through increasing the
dissolved oxygen DO in the liquid bulk. The water bulk here is considered in the same
time as the intake of water spray droplets.
- 999 -

For this configuration it can be said that each zone of the two existed spray and water
bulk mass transfer zones is regarded as complementary to other zone and each zone
cannot be evaluated separately from the other one.
0.6
Turbine alone
0.5
Whole system
Esp, (-)
0.4
0.3
0.2
0.1
0
1.5
2.5
N, (s-1)
3.5
Figure (3.28): The relation between the impellers speed N and the spray zone aeration
efficiency Esp for both turbine alone and whole system configurations, S/W = 1.0.
III.2. Spray Zone Mass Transfer Coefficient (klad)

The water spray zone oxygen mass transfer coefficients (klad) for turbine alone
configuration is illustrated in Table 3.14. The occurred klad values are slightly lower
than the klad values for the whole system configuration as shown in Figure 3.29.
4
Turbine alone
3.5
Whole system
klad, (s-1)
3
2.5
2
1.5
1
1.5
2.5
N, (s-1)
3.5
Figure (3.29): The relation between N and the oxygen mass transfer coefficient klad
for both turbine alone and whole system configurations, S/W = 1.0.
- 999 -

Table 3.14 and Figure 3.29 show that the klad was higher for the rotation speeds 2.5
and 3.33 rps (klad = 1.837 1/s) than of low rotation speed (1.67 and 2.08 rps), where
the klad was increased to 2.316 1/s. The klad showed same behavior as spray aeration
efficiency Esp, where their values are decreased when the draft tube and propeller are
removed. As a result of that the water bulk DO concentration is reduced inside the
water bulk. It was found that the accomplished water spray maintains its dimensions
for both whole system and turbine alone configurations. The water droplets flight time
is calculated by equation (2.35), while klad is calculated as mentioned in chapter 2.
Table (3.14): The spray mass transfer zone (klad), flight time (tf) and some related
measured parameters, S/W=1.00, h/D =1.47, (Turbine alone configuration).
N
Tbulk
(oC)
(mg L-1)
(mm)
(mm)
1.67
2.08
2.5
3.33
19.6
20.8
21.0
19.7
8.85
8.74
9.28
9.01
15
20
35
45
110
130
170
280
(rps)
Cds
Ym
Rm
Esp
(-)
0.183
0.206
0.292
0.356
(Esp)20
(-)
0.184
0.203
0.287
0.358
tf
klad
0.11
0.13
0.17
0.19
1.837
1.774
2.031
2.316
(s)
(s-1)
III.3. Water Spray Velocity and Volumetric Flow Rate

The water spray flow rate Q as well as the water spray velocity Vsp for the turbine
alone configuration match those in whole system configuration, since both the Q and
Vsp depend in their calculations on the shape of water spray which it is identical in
both cases. Table 3.15 shows the Q and Vsp in the turbine alone configuration and
elucidates the relation between the water spray dimensional values with the generated
spray. Same as the whole system configuration both the spray velocity Vsp and spray
flow rate Q increase with rotational speed increasing. Equations (2.40), (2.41), (2.42)
and (2.43) are applied to calculate water spray discharge velocity Vsp by the turbine
blades tip according to physical laws for projectiles and water droplets spray flow rate
Q. It was found that when the turbine rotation speed was elevated from 1.67 rps to
3.33 rps the achieved spray flow rate Q was approximately doubled from 11010 g/h to
22422 g/h. The increasing in N higher than 2.5 rps didnt produce a noticea le rising
in the projected water droplets spray flowrate Q.
Table (3.15): The spray velocity and volumetric flow rate for various turbine rotation
speed (Turbine alone configuration).
N
Ym
Rm
Rsp
(rps)
(mm)
(mm)
(mm)
1.67
2.08
2.5
3.33
15
20
35
45
85
115
175
270
180
210
270
365
Esp
(-)
0.183
0.206
0.292
0.356
- 999 -
tf
Vsp
(s)
(Watt)
(m s-1)
(L h-1)
0.111
0.128
0.169
0.19
4.472
6.820
9.943
14.263
1.71
1.76
1.80
2.14
11010
15851
22094
22422
III.4. Spray Zone Oxygen Transfer Rate (OTRsp)

Table 3.16 and Figure 3.30 show that a higher oxygen transfer rate OTRsp was
produced with higher water rotation speed N. It was noticed in turbine alone
configuration the OTRsp and Esp water spray efficiency are related in direct relation,
where higher OTRsp is obtained as high Esp is achieved. From Table 3.15 it was
founded that at the highest employed rotation speed (N=3.33 rps), the maximum
oxygen transfer rate was reached (67.131g/h). The lowest OTRsp (16.622 g/h) was
occurred with lower N used (N=1.67 rps). Table 3.15 shows that as higher N causes
more water droplets to be projected into atmospheric air as consequence larger contact
interfacial area is available. The calculated OTRsp is by applying equation (2.34). The
important of the OTRsp calculation for this case is to compare this performance with
that obtained for whole system configuration in order to understand the effect of the
draft tube and the RTP propeller on the achieved OTRsp for the system as it is
explained in the next section IV.
120
Turbine alone
Whole system
OTRsp, (gO2 h-1)
100
80
60
40
20
0
1.5
N, (s-1)
2.5
3.5
Figure (3.30): The relation between spray zone oxygen transfer rate with the
rotational speed for the whole system and turbine alone configurations, S/W = 1.0.
- 999 -

Table (3.16): The Spray zone oxygen transfer rate OTRsp for different rotation speed
levels, (Turbine alone configuration).
N
(rps)
1.67
2.08
2.5
3.33
Esp
(-)
0.183
0.206
0.292
0.356
(s)
tf
(m s-1)
Vsp
(L h-1)
OTRsp
0.111
0.128
0.169
0.19
1.71
1.76
1.80
2.14
11010
15851
22094
22422
16.622
26.253
55.353
67.131
(g h-1)
IV. Comparing the OTRsp for the Whole System and Turbine Alone
Configurations
To find out the effect of the RTP propeller and draft tube presence on oxygen mass
transfer rate in spray mass transfer zone OTRsp, a comparison was made between the
calculated OTRsp in both whole system and turbine alone configurations. It is found
that OTRsp in whole system configuration has moderately higher values than the
turbine alone configuration as illustrated in table 3.17 and Figure 3.31. The
improvement in OTRsp when propeller and draft tube placed under the turbine is
generally has an excess percentage around (30 %). The experimental results showed
that OTRsp is always greatly depending on the aeration efficiency for various
conditions tested in this work as seen in Table 3.16. The excess in the oxygen mass
transfer rate for the spray mass transfer zone, OTRsp, when the RTP propeller and the
draft tube are positioned in their places is due to the improvement of the mass transfer
rate in the tank as a result of flow pattern improvement which is highly related to the
oxygen mass transfer in the spray zone since the initial concentration of the DO in the
spray zone is in the same time the final concentration of DO in the bulk zone.
Percentage in the OTRsp, (%)
100%
80%
60%
OTRsp% (Whole system)
OTRsp% (Turbine alone)
40%
20%
0%
0.17
0.58
N, (s-1)
1.42
Figure (3.31): The influence of the tank internals (Propeller and draft tube) on the
spray zone oxygen transfer rate, OTRsp contribution the overall transfer rate OTR,
S/W = 1.0.
- 999 -
Table (3.17): The Comparison for the spray zone oxygen transfer rates OTRsp
between the whole system and turbine alone configurations, S/W =1.00, h/D=1.47.
N
OTRsp
OTRsp
(rps)
(L h-1)
Turbine alone
Whole system
1.67
2.08
2.5
3.33
11010
15851
22094
22422
(g h-1)
16.622
26.253
55.353
67.131
OTRsp
(g h-1)
Excess percentage (%)
24.330
36.488
74.493
103.346
31.68
28.10
25.69
35.04
(Whole system)
3.2.3. The Modeling

In order to characterize the oxygen mass transfer rate achieved in the spray zone by
the turbine through propelling water droplets in the air and to determine the optimum
process condition. A model of the water spray mass transfer zone efficiency was
developed depending on applied works in the same field of investigation and the
obtained results from the experimental results. The model is built to interpret the
oxygen mass transfer operation to water droplets, where have considered the standard
water spray efficiency (Esp)20 as the dependent factor instead of (kla/N) that applied in
the model for bulk (re-aeration) mass transfer zone (See equation 3.4), (Kim and
Walters, 2001; Kucukali and Cokgor, 2009; Watson et al., 1998) have also considered
the (Esp)20 as the independent parameter in their models. Same as the water bulk mass
transfer zone, the most affecting independent parameters during the surface aeration
process are related with the dependent mass transfer parameter but here for the
parameter (Esp)20, these independent parameters represent the important objectives
achieved through the experimentation. A total of 47 experimental runs were
conducted for determining the standard water spray mass transfer zone efficiency of
surface aerators at 20 oC, (Esp)20. At the experimental run temperature, Esp spray zone
efficiency coefficient was calculated using the equation 2.30. The equations 2.32 and
2.33 are applied to convert the calculated spray mass transfer zone efficiency Esp to
standard spray mass transfer zone efficiency at 20oC (Esp)20.
The spray zone oxygen mass transfer depends on various relevant geometric, dynamic
(process) and materials (property or physical) parameters. The most important
parameters are shown in the following list:
- 999 -
Geometrical Parameter:
Turbine diameter
Turbine blades width
D
W
Material Parameters:
Density
Viscosity
Process Parameters:
Gravitational constant
Turbine speed
Turbine blades submergence
g
N
S
The spray zone standard efficiency varies noticeably with the turbine rotation speed.
The standard spray efficiency depends also on the submergence level of the turbine
blades, where the turbine blades submergence S/W showed also a remarkable effect
on the standard spray zone efficiency.
As a primary essay to correlate the standard efficiency of the spray mass transfer zone
with the influencing parameters in the same way that applied for the water bulk zone
as represented in the equation 3.1.
(Esp)20 = f (N, , , D, g, S, W)
(3.8)
Equation 3.8 is solved to determine the values of the constants by applying

Buckingham theory equation (3.8) was converted to the following correlation:
(
(3.9)
Solving equation 3.9 by multiple non-linear regressions, the following model is

obtained:
(
( )
Fr=0.54-0.215
(3.10)
S/W=0.17-1.83
Where Reynolds number Re was ignored because it is considered irrelevant to the
process objective (Esp )20 according to the multiple non-linear regressions, as the spray
flow is predominantly turbulent the value of Re is always > 104 (Zlokarnik, 1979).
The plot of the predicted values of spray zone standard efficiency calculated from
equation 3.9 with the experimental values showed a coefficient of determination of
(0.9095) as shown in Figure 3.32.
Equation 3.10 means the correspondences (Esp)20 (N)1.23 or (Esp)20 (S/W)0.355 are
achieved when the turbulent regime was prevailing (Re > 104) and within indicated
ranges of Fr and S/W.
- 999 -

0.6
2
R = 0.9
Predicted (Esp)20
0.5
0.4
0.3
0.2
0.1
0
0.1
0.2
0.3
0.4
Experimental (Esp)20
0.5
0.6
Figure (3.32): The Comparison between the (Esp)20, Predicted by the Correlation
model (Eq. 3.10) with the experimentally resulted (Esp)20.
From the earlier works (Baylar and Bagatur, 2006; Baylar et al., 2010; Nakasone,
1987; Tarshish et al., 2000; Wormleaton and Soufiani, 1998), they have found that the
form of the water spray standard efficiency correlation is more complicated than the
power low correlation applied in equation (3.10). For the same influencing parameters
as shown in equation (3.8), this was solved by applying multiple non-linear
regressions and least square method. The following model was derived:
( )
(3.11)
This means the target quantity (Esp)20 increases in a correspondence to (S/W)0.535

within same indicated ranges in equation 3.10. The plot of the predicted values of
spray zone standard efficiency calculated from equation 3.11 with the experimental
values showed a coefficient of determination of (0.9025) as shown in Figure 3.33. The
range of Re in equation is (60000-120000). While the range for Fr and S/W are same
in equation 3.10.
- 991 -

0.6
2
R = 0.9
Predicted (Esp)20
0.5
0.4
0.3
0.2
0.1
0
0.1
0.2
0.3
0.4
0.5
0.6
model (Eq. 3.11) with the experimentally resulted (Esp)20.
(Baylar and Bagatur, 2006; Nakasone, 1987; Wormleaton and Soufiani, 1998)
developed a model of weir type aeration, where they showed that spray discharge
flow rate and droplets height involves the effects on the standard spray zone
efficiency.
(Tarshish et al., 2000) have built a model for the achieved oxygen mass flow rate by
the surface aeration in the wastewater treatment tank, they showed the oxygen mass
flow parameter in the model is affected by power consumption and jet flow rates of
waste water and activated sludge. (Wormleaton and Tsang, 2000)studied the aeration
behavior in weirs with. For the spray zone, they derived the model of standard spray
aeration in terms of a spray Froude number and Reynolds number.
The effect of spray discharge flow rate and droplets height included in the previous
models are not considered separately here, but they are implicated in more general
dimensionless parameters.
In dependence on experimental results and earlier studies, it is found that standard
spray zone efficiency for the surface aeration, it is considered that the target quantity
of standard spray zone efficiency (Esp)20 for the given process conditions (the
discharge spray flow rate/spray radius, qsp and the maximum height of the spray Ym)
depends on the turbine blades submergence S, turbine blades width W, gravitational
constant g, and kinematic viscosity of the water . The following relevance list is:
[(Esp)20 ;qsp, Ym ;W, S; g, ]
(3.12)
- 991 -

As mentioned earlier the various relevant variables included in equation 3.12 are not
considered separately and are implicated in more general dimensionless parameters.
Thus, dimensionless parameters for standard spray zone efficiency coefficient can be
taken as function of the dimensionless parameters of Reynolds, Froude and power
numbers. The equation (3.12) converted to the following relation:
(
(3.13)
(3.14)
(3.15)
The spray Reynolds number is calculated by equation (3.14), the spray Froude
number is calculated by the equation (3.15). Equation (3.13) was solved to determine
the values of the constants by applying multiple non-linear regressions and least
square method. The following correlation was developed:
(
( )
(3.16)
The mass transfer model (equation 3.16) is applicable within the ranges, (Resp*10-3) =
(22.1 - 46.2), (S/W) = (0.17 - 1.83), (Frsp) = (0.044 1.80) and (water spray maximum
height Ym/D) = (0.08 0.24).
Equation 3.16 shows that the determined powers for the three parameters are
somehow close and they are proportionally related the spray zone standard efficiency.
This means the effect of these parameters have close effects on the objective
parameter (Esp)20. For example the equation 3.16 refers that the standard spray zone
efficiency (Esp)20 increases by the factor 1.3 when the turbine blade submergence ratio
(S/W) was altered from 0.17 to 0.58.
The standard error of estimation for the equation is 0.0215, which indicates a 98%
probability that the equation 3.16 will predict (Esp)20 within + 0.01 of its true value.
The correlation coefficient for this equation is R2 = 0.9568. The measured (Esp)20
values are compared with those predicted by the equation 3.16 are given in the Figure
3.34. An acceptable agreement between the measured values and the values computed
from the predictive equation is obtained.
- 999 -

From Figure 3.33 it can be found that the points with higher error compared with the
experimental values are the points of lower spray Reynolds number with the ratio of
S/W=1.
0.6
Predicted (Esp)20
0.5
0.4
0.3
0.2
0.1
0
0.1
0.2
0.3
0.4
0.5
0.6
model (Eq.3.16) with the experimentally resulted (Esp)20.
3.2.4. Conclusions
With the totality of experimental runs performed for the power consumption in water
bulk zone, it was found that the power consumed is related with rotation speed and
water height level (turbine blades submergence) variations more than other
parameters occurred in aeration mode.
For the given impellers rotational speed range, it was observed that the desirable
performance of the aeration mode and flow patterns in the bulk zone is achieved
starting with impellers rotation speed of 2.08 rps, the formation of air bubbles appears
in the vessel due to plunging water droplets with water surface, so the impellers
rotational speed 2.08 is considered as the critical speed to accomplish aeration
process.
At the given water level ratios and turbine blades submergence, the water bulk zone
oxygen mass transfer coefficient kla was increased with the rotation speed N. It was
observed that the kla was highly dependent on the impellers rotation speed, however
this relation is limited: the outer edge of the spray should not colloid the vessel wall
what can be observed with rotational speed N lower up to 3.33 rps.
An optimum water height level came across from the water bulk zone surface aeration
experimental testing and it depends on a satisfactory formation created by the aerator
turbine blades. Where, neither too low nor too high blade submergence is preferred.
- 999 -

The spacing between the impellers affects slightly the water bulk zone oxygen mass
transfer coefficient kla. This influence stays related with the obligatory of remaining
the lower propeller inside the draft tube. The reliance of the kla on the spacing is
conjugated with the interaction between the propeller and the daft tube.
Through the four tested different geometrical configurations of the water bulk zone,
the highest kla was achieved when all the aerator system parts exist together (the
turbine with both the draft tube and RTP propeller).
Upon the calculated water bulk zone standard aeration efficiency SAE b values, it
seems the SAEb is more sensitive to rotational speed variation, while on the other
hand the SAEb didnt show same response toward water height level variation.
For the whole system configuration it is found that at the spray mass transfer zone
both the oxygen transfer rate OTRsp and spray zone efficiency Esp are highly
dependent on the impellers rotation speed N, dimensionless turbine blade
submergence ratio S/W and dimensionless water level ratio h/D.
The majority of oxygen mass transfer is achieved in the spray zone due to high
interfacial area and turbulent mass transfer occurred in this zone
The discharge flowrate of the water spray (droplets) Q is increased when impeller
rotation speed or water level are increased for the given turbine blades submergence.
The OTRsp in whole system configuration is slightly higher than OTRsp for the turbine
alone configuration.
Two models are developed for water bulk zone oxygen transfer coefficient kla/N and
power consumption, applicable in specified range. A model is developed for surface
aeration spray zone, standard oxygen transfer efficiency coefficient; the model is
applicable for specified limits.
- 999 -
- 999 -
- 145 -
- 146 -
Chapter Four: Hydrodynamics in the Single Phase Non-Aerated Agitated

Tank for Up and Down Pumping Directions Modes
Chapter Four
Hydrodynamics in the Single Phase Non-Aerated Agitated
The investigated surface aeration system consists in two agitation tools; the turbine
positioned at liquid surface which propels droplets into atmospheric air, thus forming
the water spray. The second tool is the mixing assembly (RTP propeller and draft
tube) placed below the turbine. The proposed function of the mixing assembly is to
improve mixing capability of the surface aeration system by enhancing air bubbles
distribution inside the vessel and redirecting the flow toward the upper turbine. The
mode of mixing assembly employment is either individual or conjugated with turbine,
where mixing assembly operates separately from the turbine when its needed to
achieve well-mixed condition inside the vessel, when sufficient dissolved oxygen is
accomplished but without satisfying distribution of dissolved oxygen throughout the
liquid bulk.
In this chapter the turbulent flow regime in liquid phase generated by the reversible
twisted pithed, axial 4-bladed and 45o pitched blade propeller (RTP) (Milton Roy
Mixing HPM204D) is investigated to determine velocity field for different areas of
propeller vicinity, inside the draft tube and in the vessel, whereas the upper turbine is
not combined. Thus there is only the liquid phase and very few gas bubbles are
entrapped inside the liquid bulk can be neglected. Comparison between flow patterns
resulting from up-pumping and down-pumping operation modes of RTP propeller is
made. The up-pumping condition is performed ordinarily with surface aeration
system; so the propeller turns in same manner than the upper turbine and pushes the
fluid upward. The down-flow is applied mainly when propeller and draft tube are used
separately aiming that mixing objectives are to be fulfilled for the liquid bulk.
4.1. Experimental Aspects

Both PIV and LDV (details for the devices are shown in chapter two) are applied to
measure mean velocity fields. The pumping capacity of the propeller is calculated for
all tested flow conditions at different positions inside the draft tube. In order to
achieve LDV and PIV measurements the cylindrical tank is placed inside another
square tank with front face made of glass to eliminate laser distortion that may affect
the measurements. The outer tank is filled with water in order to avoid refraction and
deflection of laser beams at cylindrical and square tanks walls.
All flow patterns tests were conducted in the baffled vessel. At the beginning unbaffled vessel was examined to find out its compatibility for given experimental runs.
- 147 -

Un-baffled vessel seems not to be suitable because of oscillation occurred at water
surface resulting from agitation due to impellers rotation.
The LDV and PIV measurements were applied for the flat bottom cylindrical vessel
(Tv =0.8 m) provided with three transparent baffles their width are equal to (b =
Tv/10). A draft tube (df =0.15 m, Lf = 0.1 m) was positioned in the tank with clearance
(CDT = 0.1m) and was removed for some of the measurements. The four 45o pitched
and twisted blade propeller RTPwas studied at clearance (Cpr/Tv = 0.2 m) with
rotation speed was (N= 5 rps) to ensure staying always in turbulent regime for created
flow field (Re = 72000), where the LDV measurements were focused on the core of
the tank where draft tube is settled to measure the flow pattern, the mean velocity
field and the liquid flow rate created by the propeller inside draft tube in different
planes that cannot be reached by the PIV.
Figure (4.1): LDV implementation for measuring velocity in a (r-z) plane, (Plane B).
The LDV is a Dantec Fiberflow system with two green and blue laser beams,
wavelengths of 514.4 nm and 488nm respectively. The laser source in the LDV is 4W Argon with a focal length of 600 mm (Spectra Physics). The LDV measurements
were performed through an r-z plane (Plane B) at (90o) to posterior baffle and (30o) to
two lateral baffles (See Figure 4.1), where for each measurement, about (4000 to
5000) samples are applied with a maximum measurement time of 300 seconds. The
tracer particles that were used to seed the liquid are (Iriodin 111) (dp =15m).
The PIV were performed in order that flow patterns can be determined for all vessel
areas. the vessel dimensions are considered somehow large in comparison with same
kinds of experimentations, so it is preferred to conduct flow field measurements by
- 148 -

PIV for time and effort saving. PIV creates a laser sheet which allows us to inquire
flow and velocity profile at various spatial parts inside the vessel. LDV provides
detailed information for single points with good temporal resolution while with PIV
we can get detailed spatial information for different instants of time (Atkinson et al.,
2000).
The obtained PIV measurements detailed information depends on the number of
interrogation areas that divides the simultaneous images taken by used CDD camera
after seeding tested liquid bulk with (Rhodamine) red fluorescent tracer particles
(dp=15 m). In the experiments the laser sheet created by used Nd: YAG laser source
is placed in the middle of vessel, where r-z measurement plane was positioned
directly on posterior baffle (0o) and at (60o) from the two lateral baffles from front
face (see figure 4.2). (32%x32%) with (50% overlap) is applied for velocity
measurements. The number of image pairs was (300) and exposure time delay (t)
between each two sequent images was (1000 s) depending on operation condition
such as propeller tip velocity and image scale factor adapted during the acquisition.
Figure (4.2) PIV implementation for measuring velocity in a (r-z) plane, (Plane A)
For single (liquid) phase flow the used propeller rotation speed was kept at (5 rps) to
ensure reasonable vector displacement between two frames and since upper limit of
rotation speed due to turbine presence and consequence droplets impingement
position does not exist. The applied software for acquisition system is (DaVis 8.0.5
flow master), while image processing is (PIV imager pro x2M system).
- 149 -

4.2. Mean Velocity Field and Flow Pattern
In order to obtain a good knowledge of the flow profiles inside the vessel in mixing
mode, it is needed to examine both the up-pumping and down-pumping flow
generated by the axial 4-bladed, 45o pitched blade RTP propeller (See Figure 4.3)
with or without draft tube. The flow is discharged from propeller axially in up-ward
direction when the shaft is rotated in its normal direction, while it is discharged in
down-ward direction when the shaft is rotating in reverse mode. Since the propeller
blades are twisted and made in a symmetrical way, the propeller can be settled upside
down, the blades always push the fluid in same direction. The propeller is placed
inside a draft tube, the presence of draft tube having theoretically noticeable effects on
the generated flow (Aeschbach and Bourne, 1972; Bro et al., 2004; Kumaresan et al.,
2005; Tsui and Hu, 2008). A part of the discharge flow from the propeller goes in
radial direction and needs to be redirected toward axial direction. Few studies were
made for a down-pumping flow generated by an axial pitched blade propeller with
draft tube presence; on the contrary many studies were made for this configuration for
the airlift and bubble column reactors.
Figure (4.3): The axial 4-bladed, reversible 45o pitched and twisted blade propeller
RTP propeller (HPM204D, Milton Roy Mixing)
4.2.1. Down-Pumping Condition:

The time averaged velocity field in the taken r-z middle plane of the vessel at (30o)
from the two lateral baffles (Plane A) performed by the PIV (See Figure 4.2) is shown
in the Figure 7.4, where the flow is produced by the reversible twisted pitch blade
RTP-D propeller with draft tube (Lf = 0.1 m) in down-pumping direction mode.
- 150 -

Form Figure 4.4 it is noticed that one general circulation loop is formed in the entire
vessel, that begins just below propeller blades and goes downside in axial pattern and
then toward vessel wall in radial pattern along the vessel bottom, the circulation loop
continues in upward before it returns to impeller inward flow region. The flow returns
to impeller region and discharges axially due to draft tube. A part of the circulated
liquid downs axially to the vessel base in the region between the draft tube wall and
the middle of radial distance to the vessel wall without being entrained by the RTP-D
propeller, this is occurred because the low ratio of dpr /Tv = 0.15 as the low pressure
inflow region generated by RTP-D propeller blade rotation isnt significant enough to
entrain all the returned downward flow into this RTP-D propeller inflow region.
As shown in Figure 4.4, two secondary circulation loops are produced; one of them is
directly below the RTP-D propeller. This secondary loop has been also observed by
(Kumaresan et al., 2005) for hydrofoil impeller (HF4 type) with draft tube in downpumping mode, with (Cimp/TV = dimp/Tv=0.33) and length of draft tube (Lf =0.167 m).
(Kumaresan et al., 2005) found that the axial flow created by impeller (HF4 type) was
increased in the core of the vessel in presence of a draft tube.
The third circulation is formed beside the RTP-D propeller blades; this circulation is
actually the radial discharged flow from the RTP-D propeller redirected by the draft
tube upward to the RTP-D propeller inward flow volume.
The mean axial velocities are more intense at vessel core and are significantly smaller
with moving away from vessel core toward the radial direction except in the region
near the vessel wall, where higher axial velocities are achieved due to the interaction
with vessel wall. The low mean velocities near the water free surface should be taken
into account carefully because of the error that may occur due to the reflection from
the free surface. The distribution mean average velocities are presented more clearly
in the contour map (See Figure 4.5) of the r-z time averaged velocities in downpumping mode (Mixing mode).
The highest downward axial velocity below the blades of the RTP-D propeller is of
the order 0.27Vtip, whilst the highest downward velocity above the impeller is
0.22Vtip. In the inflow region above the RTP-D propeller the value of downward axial
velocities are of the order 0.1Vtip. In the secondary loop beside RTP-D propeller
blades inside the draft tube, the highest upward axial velocity in the vessel is found
0.34Vtip, this loop is actually in the discharge flow from the impeller in radial
direction but it has been redirected by the draft tube as an upward axial velocity.. In
the central part of the vessel exactly in the mid-planes between the draft tube and
vessel wall, predominate velocity is the downward axial velocity with the order of
0.02-0.03 Vtip; these values are of the order 0.02-0.04 Vtip in the upper part of the
vessel. For the RTP-D propeller the maximum radial velocity is found also in the
secondary loop beside impeller blades, this maximum radial velocity is of the order
0.22Vtip, which moves toward the impeller blades. In the intake flow region above the
- 151 -

RTP-D and the maximum radial velocity in this region is of the order 0.06Vtip. The
highest radial velocity in the primary loop is produced in the discharged flow from
RTP-D propeller inside the draft tube. The maximum radial velocity in this region is
of the order 0.2Vtip. The discharged flow below the impeller becomes mainly radial
when it approaches the vessel bottom. This flow pattern is achieved with aid of the
cone placed at the vessel bottom below the draft tube. The highest radial velocity in
this zone is of the order of 0.16Vtip. Radial velocities are clearly found in both the
upper and lower part of the vessel and have highest radial velocities of the order
0.12Vtip and 0.06Vtip respectively.
The observed low axial and radial values in the upper, central and near the vessel wall
regions indicates the weak mixing and flow pattern in these parts, the main cause for
this performance is the small ratio of dpr/Tv =0.2, where weak interaction between the
discharged liquid flow by the impeller and the liquid bulk in these regions as
mentioned earlier. Although the influence of applied impeller diameter to tank
diameter ratio causes a poor circulation especially in farer regions from the core of the
vessel, it is preferred to keep this ratio because the main purpose of mixing mode is
maintain an acceptable mixing condition for the already aerated and mixed tank
contents and the mixing mode should be accomplished with minimum power
consumption in order to prevent any additional cost for the treatment operation; this
cannot be done when higher dpr/Tv is used.
The values of axial velocities with those obtained by (Kumaresan et al., 2005) for
three pithed blades with the angle 35o and twist in the tip region, with 3 blades
hydrofoil impeller of (HF4 type) with two different draft tube lengths (Lf = 0.125 and
0.167 m) in down-pumping mode, with (Cimp/TV = dimp/Tv=0.33). The axial velocities
below the impeller were of the order 0.335 Vtip and 0.274 for the longer and shorter
draft tubes respectively, whilst they were of the order 0.22 Vtip and 0.18 Vtip for the
longer and shorter draft tubes respectively. Compared with the velocities produced by
the axial HPM204D-D impeller, the order of the axial velocity above the impeller the
RTP-D is approximately than HF4 for the longer draft tube and smaller than HF4 with
shorter draft tube. For the axial velocities in the discharge flow below the impeller,
the order of the axial velocity for RTP-D is approximately the same than with HF4 for
the longest draft tube and greater than HF4 with the shortest draft tube.
- 152 -
- 153 -
Figure (4.4): The (r - z) flow field map for the RTP-D propeller with draft tube configuration in the entire vessel.
- 154 -
- 155 -
Figure (4.5): The (r - z) velocities contour map for the RTP-D propeller with draft tube configuration in the entire vessel.
- 156 -

More mean velocity field measurements were performed in the core of the vessel as
this region was found more significant in terms of efficient circulation and mixing
according to the measurements made by the PIV. As the draft tube is baffled with
three half-length baffles the velocity field generated by the RTP-D propeller was
measured from different planes. This was performed by LDV in the r-z plane (Plane
B) at (90o) to posterior baffle and (30o) to two lateral baffles (See Figure 4.1).
Figure (4.6) shows the quantitative analysis of the velocities in the down-pumping
(Mixing mode) as a radial profile for the normalized axial velocity inside the draft
tube, above and below the impeller in the two different measurement planes (A at 0o
and B at 90o from the posterior baffle). Above the RTP-D propeller the highest axial
velocity are of the order 0.23Vtip and 0.25Vtip in the planes A and B respectively,
whilst below the PTBD impeller in the discharge region the highest axial velocities
are of the order 0.27Vtip and 0.16Vtip in the planes A and B respectively.
0.1
Above impeller (plane A)

Below impeller (plane A)
Above impeller (plane B)
Below impeller (plane B)
0.05
0
Vz/Vtip, (-)
-0.05
-0.1
-0.15
-0.2
-0.25
-0.3
0.2
0.4
0.6
0.8
r/R, (-)
Figure (4.6): Dimensionless axial velocity profile at different liquid levels (below
impeller z/h=0.55, above impeller z/h=0.62) for RTP-D with draft tube, Cpr /Tv =0.2,
N= 5 rps in two different measuring planes; plane A at 00 and plane B at 900 from the
posterior baffle respectively.
The axial velocity in the discharge region of the impeller looks higher in the
measurement plane A than those in plane B and also it has a different profile more
than in inward region above the impeller. This difference is due to the effect of the
- 157 -

draft tube baffle position, where in plane A the discharge flow is surrounded by two
close baffles at 45o each side, when on the other hand in the plane B only one of the
baffles is close at 45o and the other is at 90o. The effect of the draft tube baffles on the
discharge flow will be confirmed with repeating this quantitative analysis for
propeller alone configuration.

Figure 4.7 shows the time averaged velocity fields in the taken r-z middle plane in
same position that performed by PIV for the previous propeller and draft tube
configuration, where the flow is produced by the axial 4-bladed, 45o pitched twisted
blade propeller (RTP-D) alone in down-pumping mode. The flow pattern and mean
velocity profile measurements were focused on the vessel core as it was observed in
previous section for the propeller and draft tube configuration with down pumping;
this region is more active in terms of generated mean velocity profiles. The effect of
the propeller on the produced flow pattern was weaker near the walls as the
interaction between the tank wall and the propeller is quite low (dpr/Tv = 0.15).
The downward discharge flow by the RTP-D propeller is predominantly axial and the
secondary loop that was observed beside the impeller blades edges in the propeller
and draft tube configuration has disappeared here. This confirms that this secondary
loop was due to the presence of the draft tube.
The discharge flow shown in Figure 4.7 (a, and b) is pushed intensively downward in
a regular manner. This observation somehow disagrees with that noticed by
(Armenante and Chou, 1996; Tatterson et al., 1980) where (Tatterson et al., 1980)
observed that the 450 pitched blades impellers PBT (Chemineer HTD-2, dim/Tv = 0.35)
push the flow downward in a chaotic way of two different manners; a high speed jets
and low speed flow, (Tatterson et al., 1980) also remarked trailing vortices formed at
the blades tips, they assumed the created downward flow profiles by axial pitched
blades impellers is changing due to tank and impellers geometric ratios. The produced
discharge flow is similar with (Ranade and Joshi, 1989) observation The observed
discharge flow agrees with that by (Ranade and Joshi, 1989) for PBT impellers (30o
60o pitch angle, dim/Tv = 0.25 - 0.5), where they found that an intense axial flow
pushed downward especially at impeller vicinity with negligible radial flow at this
region, they reported that the velocity profile under the impeller was relatively
changed with impeller clearance variation, for the tested clearance range (C= h/2
h/6). (Armenante and Chou, 1996) found out that the down flow in impeller region
induced by 6 blades PTD (45o pitched angle) had a contrary flatter form to other
studies; (Armenante and Chou, 1996) imputed the non-convergence of down flow
under impeller blades to the geometry and PTD clearance applied (dim/Tv = 0.876,
Cim/T=0.33) . It is worth mentioning that the observations on the flow pattern created
- 158 -

by axial pitched impellers were made for dim/Tv ratio higher that performed in this
work dim/Tv=0.2, where the interaction between the flow generated by acting axial
pitched propeller with the vessel bottom and walls was stronger than the
measurements in larger vessel.
(a)
(b)
Figure (4.7): (a) The (r - z) flow field map for the RTP-D propeller without draft tube
configuration in the vessel core, (b) The (r - z) velocities contour map for the RTP-D
propeller with draft tube configuration in the vessel core.
As a result of the presence of the cone at the base of the vessel, a radial flow in the
discharge stream begins to appear relatively and gradually, where the majority of the
discharge flow is radial near the bottom. The other secondary circulation below the
impeller exists as in the propeller and draft tube configuration but with less intensity.
(Mavros et al. 2002) reported a close observation for pitch blade impeller (Mixel TT
- 159 -

type); where a downward flow jet has been released by the impeller blades and then
the flow continues its circulation as a general loop through the vessel and also a small
reversible circulation is created at the area below impeller and near the vessel center.
The observation by (Mavros et al., 2002) was confirmed with the CFD simulation
applied by (Khopkar et al., 2004).
In the Figure 4.7a, the highest axial velocity in the discharge bulk and directly below
the impeller blades is 0.345Vtip; this is greater than that produced in the propeller and
draft tube configuration by 18%. In the upper part above the impeller the highest axial
velocity in the inward flow bulk is about 0.1 Vtip, which is same in the propeller and
draft tube configuration, while at above the RTP-D propeller the highest downward
velocity is 0.21 Vtip, this value being close to that produced in the propeller and draft
tube configuration. For the reversible secondary loop below the RTP-D propeller
beside the shaft, higher upward axial velocities are produced in both propeller and
draft tube configuration and propeller alone configuration and are of the order
0.11Vtip.
The main zone for the radial velocities in the vessel core is the lower part near the
vessel bottom and beside the cone, the value of radial velocity in the discharged flow
in this part is of the order 0.11Vtip, comparing it with the propeller and draft tube
configuration achieved velocity for the same zone, the highest radial velocity that
obtained with propeller alone configuration is of the order 0.217 Vtip and it is greater
than that obtained in the propeller alone propeller and draft tube configuration by 7%.
Above the RTP-D propeller in the inward flow region, the highest radial velocity is of
the order 0.09Vtip,. This dropping in the radial velocity value illustrates the effect of
draft tube presence on the flow pattern in this region clearly.
Comparing the downward axial discharge velocity produced by the RTP-D propeller
with the other observations for the induced velocities by down-pumping axial
impellers, (Bittrof et al. 2000) reported that the highest downward axial velocity in the
discharge flow below a PBD impeller (4-bladed, 45o pitched angle, Cim/ dim = 1.5 and
dim/Tv= 0.19) is of the order 0.1Vtip. While a highest downward axial velocity is of the
order 0.5Vtip in the discharge zone below a three-bladed 45o pitched blade HF
impeller (15o twisted blades, dim/Tv=0.27) has been reported by (Kumaresan & Joshi
2006). (Khopkar et al. 2004) reported that the maximum axial velocities are of the
order 0.4Vtip and 0.2Vtip respectively at below and above a 45o Mixel TT impeller
(dim/Tv= 0.5).
A reversible secondary circulation is created below the RTP-D propeller beside the
lower part of the shaft and the cone. This circulation without the draft tube is smaller
and do not reach the lower edges of impeller blades. The highest upward velocity in
this secondary circulation is of the order 0.1Vtip which is same for propeller with draft
tube configuration. A similar observation has been reported by (Aubin et al., 2001)
- 160 -

for six-bladed PBT impeller (dim/Tv=0.5), but they found the highest velocity in the
secondary circulation was of the order of 0.07Vtip.
From Figure 4.8 the radial profile of the normalized axial velocity above (z/h= 0.62)
and below (z/h=0.55) the RTP-D propeller in the two different measurement planes A
and B is plotted. The divergence for the two planes is small. The convergence of the
normalized axial velocities values below the propeller without the draft tube shows
clearly the effect of the draft tube and of its baffles (Lbf/Lf = 0.6) on the produced
discharge flow below the RTP-D propeller. A light difference still exists between the
two planes, which may be due to the different measurement techniques PIV and LDV.
0.05

0
-0.05
Vz/Vtip, (-)
-0.1
-0.15
-0.2
-0.25
-0.3
-0.35
0.2
0.4
0.6
0.8
r/R, (-)
Figure (4.8): The dimensionless axial velocity profile at different liquid heights for
RTP-D propeller without draft tube (below impeller z/h=0.55, above impeller
z/h=0.62), Cpr /Tv =0.2, N= 5 rps in two different measuring planes; plane A at 00 and
plane B at 900 from the posterior baffle respectively.
It has to be mentionned also that for the propeller alone configuration, sometimes with
the rotational speed N=4.6 rps, a swilling appears from the liquid surface and extends
to the impeller baldes tips (See Fig. 4.9). This phenomineon becomes more
continuous starting from 5.8rps, where the vortex is formed for the implemented
geometry. Despite of using the baffled tank, this vertex occurs because of low ratio of
dpr/Tv = 0.15 is implemented, where the baffles (Lbf/Tv=0.1) are somehow far away
- 161 -

from the impeller to prevent forming this vortex above the impeller as it can be
observed in the Figure 7.9. With the down-pumping direction for the propeller and
draft configuration no swirling was formed despite the elevation of the impellers
rotation speed higher than 6rps.
Figure (4.9): The creation of vortex in the down-pumping mode with draft tube.
4.2.2. Up-Pumping Mode

The flow patterns and mean velocity fields measurements were performed by the PIV
in the same plane that applied for the down-pumping. Figures 4.10 a and b show the rz vector plot and velocity contour plot of the flow pattern generated by the RTP-U
propeller with draft tube presence (Lf = 0.1 m) for up-pumping direction mode in the
vessel core. It is observed that beside the main circulation loop, two secondary
circulations are formed one above the impeller near the shaft which extends to the
water surface, the other one which is a small one formed beside the lower edge of the
draft tube.
Compared with the down-pumping mode the upper secondary circulation loop seems
that it is formed with the discharge flow generated of the by the RTP-U propeller and
the position of this loop is changed upward or downward according to the pumping
direction of the impeller. The flow returns to impeller region and discharged axially
due to draft tube presence.
The recirculation beside the impeller blades tip has disappeared in the up-pumping
mode. In the inward flow region of the impeller inside the draft tube the flow is
entrained predominantly as an axial flow, whilst the flow is clearly propelled upward
with increasing radial component progressively moving toward the liquid surface.
- 162 -

The RTP-U propeller discharges high upward axial velocities with the maximum
velocity of the order 0.315Vtip above the impeller. The upward axial velocities drawn
by the impeller are distinguished with their elevated axial values inside the draft tube
in the inward zone below the impeller as observed in Figure 4.10(a and b). The
highest upward velocity below the RTP-U propeller is 0.28Vtip; (Bro et al., 2004)
have measured the axial velocity field in the vicinity of 4-bladed PBT up-pumping
impeller (pitch angle 45o, dim/Tv=0.2, h/Tv=1.2, Cim/Tv=1.0) with a longer draft tube
Lf = 0.54.The draft tube had four baffles placed in the upper part at the discharge side
of the impeller; they found that upward axial velocity below the impeller was of the
order 0.45Vtip. The highest downward velocity in the secondary loop above the
impeller was of 0.15Vtip. With respect to the secondary loop that localized in the
lower part of the draft tube and close to its wall it is found the highest downward
velocity is of 0.13Vtip.
(a)
(b)
Figure (4.10): (a) The (r - z) flow field map for the RTP-U propeller with draft tube
configuration in the vessel core, (b) The (r - z) velocities contour map for the RTP-U
propeller with draft tube configuration in the vessel core.
- 163 -

The radial velocities are found mainly in the discharge zone above the RTP-U with
highest value of the order 0.22Vtip and noticeably in the intake flow zone in the lower
part of the vessel core below the draft tube with the highest radial velocity of the order
0.19Vtip.
The maximum radial velocity produced in the upper part secondary loop near the
liquid surface of 0.15Vtip. With respect to the radial velocities that achieved in the
down-pumping flow mode with propeller and draft tube configuration these values are
slightly higher by 9% and 15% in the discharge and inward zones by the RTP-U
propeller respectively.
Figure 4.11 illustrates the quantitative analysis of the normalized axial velocities
above and below the up-pumping RTP-U propeller with the draft tube in the two r-z
measurement planes A and B as explained before.
As it is observed the presented radial profiles for the axial velocity above the impeller
are relatively similar, with the highest generated velocities of the order of 0.315Vtip
and 0.3Vtip for both the planes A and B. Whilst below the impeller an obvious
divergence appears between the achieved axial velocities between the two
measurement planes A and B, where the normalized axial velocities in plane A are
much higher than the axial velocities when the normalized propeller radius is less than
0.5.
It seems that the draft tube baffle presence in this part affects the up-flow and reduces
the entering flow to swept volume of the impeller in the plane B. For r/R > 0.5 the
velocity field profile in plane B becomes relatively higher but with clear difference of
the velocity values in plane A.
- 164 -

0.35

0.3
0.25
Vz/Vtip, (-)
0.2
0.15
0.1
0.05
0
-0.05
-0.1
0.2
0.4
0.6
r/R, (-)
0.8
RTP-U propeller with draft tube (below impeller z/h=0.55, above impeller z/h=0.62),
Cpr /Tv =0.2, N= 5 rps in two different measuring planes; plane A at 00 and plane B at
900 from the posterior baffle respectively.
Similar to the propeller and draft tube configuration in down-pumping direction, the
effect of the draft tube baffles appears and alters the velocities in the inward zone
below the impeller and inside the draft tube, but on contrary of down-pumping
condition the alteration occurs in the discharge zone. Beside the fact that the
divergence is due to the difference of the measurement planes A and B cannot be
neglected.

The vector plot and velocity contour map of the r-z plane are shown in Figure 4.12 (a)
and (b) respectively. The flow pattern of the up-pumping direction without the draft
tube in the vessel core generally is not different from the propeller and draft tube
configuration with up-pumping direction, but the propeller alone configuration is
distinguished by the fact that the flow in the inward zone below the RTP-U propeller
is less intensive and has more radial component than that produced with draft tube
presence. The small secondary recirculation below the impeller has vanished in this
case. The upper down-ward secondary recirculation loop is occurring even without
the draft tube, but it doesnt extend to the upper side of the RTP-U propeller as the
propeller and draft tube configuration. The RP-U propeller pushes the flow up-ward
- 165 -

axially, but with moving up this axial flow tends to be relatively radial till reaching
the zone near the liquid surface; here the flow becomes totally radial. In the lower part
of the vessel core a noticeable radial flow is observed also, where the effect of the
cone placed at the vessel bottom appears by preventing any possibility of secondary
loop formation or drawing the flow axial by the impeller, so no axial flow have been
seen in this area.
Comparing with other types of pitched blade impellers, (Gabriele et al., 2011)
reported that up-pumping 6-bladed PBT (pitch angle 45o, dim/Tv=0.54, Cim/Tv=0.33)
generates two distinct circulation loops one in the upper part of the vessel and one in
lower part, while (Mishra et al., 1998) have reported same observation for a flow
created by the APV-B2 hydrofoil 4-bladed impeller (dim/Tv=0.45, Cim/Tv=0.43,
solidity ratio =1), beside a secondary small loop formed below the impeller. For both
the PBT and APV-B2 impellers the flow enters axially into the swept volume from
the inflow zone below the impeller, with a considerable flow that is entrained radially
to the swept volume near the impeller lower tip. The discharge flow by the RTP-U is
mainly axial when it leaves the blades and then it moves gradually in radial direction
till it becomes fully radial near the liquid surface, whilst (Gabriele et al., 2011; Mishra
et al., 1998) reported the outward flow is pushed radially from upper tip of the blades.
This is the main cause of generating two distinct loops in upper and lower part of the
vessel.
Considering the axial velocities in this case the highest upward velocity is observed in
the discharge flow above the RTP-U propeller. It is of the order 0.305Vtip, where the
produced velocities above the impeller are mainly axial. The highest achieved upward
axial velocity below the impeller is of the order 0.22Vtip. The highest downward axial
velocity in the upper secondary is 0.15Vtip. Comparing these velocities with those
reported by (Gabriele et al., 2011; Mishra et al., 1998), (Gabriele et al., 2011) found
the highest axial velocities above and below the PBT are of the order 0.3Vtip and
0.12Vtip respectively, whilst (Mishra et al., 1998) found the highest axial velocities
above and below the APV B-2 are of the order 0.3Vtip and 0.2Vtip respectively. It is
obvious that the axial velocity below both the impellers PBT and APV B-2 have
smaller values than that produced by RTP-U. The flow is mainly entrained into the
swept volume by the RTP-U in an axial way from the inflow zone, when on the other
hand an important ratio of the flow enters the swept volume for both PBT and APV
B-2 radially and from above the impeller.
- 166 -

(a)
(b)
Figure (4.12): (a) The (r - z) flow field map for the RTP-U propeller without draft
tube configuration in the vessel core, (b) The (r - z) velocities contour map for the
RTP-U propeller with draft tube configuration in the vessel core.
The up-pumping RTP-U propeller alone generates high radial velocities in the
discharge flow especially of the upper part of the vessel core near liquid surface. The
maximum radial velocity in this part is 0.3Vtip, whilst in the lower part of the vessel
core the radial velocities are of the order of 0.156Vtip. These values are relatively
close to those reported by and (Mishra et al., 1998) for the APV B-2 of 0.3Vtip in the
impeller discharge and 0.15Vtip in the lower zone below the impeller, while the
reported radial velocities are higher than of those reported by (Gabriele et al., 2011)
for PBT of 0.15 Vtip in the discharge flow and 0.12Vtip in the lower zone below the
impeller.
- 167 -

0.35

0.3

Vz/Vtip, (-)
0.25
0.2
0.15
0.1
0.05
0
0.2
0.4
0.6
0.8
r/R, (-)
RTP-U propeller without draft tube (below impeller z/h=0.55, above impeller
z/h=0.62), Cpr /Tv =0.2, N= 5 rps in two different measuring planes; plane A at 00 and
plane B at 900 from the posterior baffle respectively.
Figure 4.13 illustrates the quantitative analysis of the normalized axial velocities
above and below the up-pumping RTP-U propeller without the draft tube in the two rz measurement planes A and B at the vertical heights of z/h= 0.62 and z/h=0.55
respectively. As it is noticed the axial velocities profile above and below the propeller
in the planes A and B are generally similar, with highest axial velocities above the
impeller of the order 0.31Vtip and 0.3Vtip in the planes A and B respectively. These
values are close to those of propeller and draft tube configuration. Below the RTP-U
propeller the highest axial velocities are of the order of 0.21 Vtip and 0.2 Vtip
respectively for the planes A and B. With these values of the axial velocities it is quite
clear in this case where the draft tube effect no longer exists, that the produced
velocities especially in planes below the impeller are relatively similar. The axial
velocities below the RTP-U propeller are significantly different from those in the
propeller and draft tube configuration case, where maximum axial velocities of
0.27Vtip and 0.18Vtip are achieved for the planes A and B respectively.
4.3. Power Consumption

Table 4.1 shows the power numbers measured for different configurations in this
work or in the literature; the power number seems in the down-pumping mode higher
than up-pumping mode approximately by 25%, and the presence of the draft tube
- 168 -

doesnt have any effect of the power number. The power number in the up-pumping
mode is slightly higher without draft tube configurations. The measured power
number for different types of pitch blades impeller are generally agree with the
reported values of RTP propeller, In other word the power number with downpumping is higher than up-pumping except that found by (Aubin et al., 2001) for
PBT.
Table (4.1): The RTP propeller and several pithed blade impellers power numbers
with different configurations and pumping modes.
Impeller
Type
RTP-U
(Lf = 0.1m)
RTP-D
(Lf = 0.1m)
Cim/Tv
dim/Tv
0.2
Np
Np
(With DT*)
(Without DT)
0.15
0.53
0.54
0.2
0.15
0.61
0.61
HFa -D
0.50
0.33
0.20
0.27
--
PBTb -U
0.3
--
PBTc -D
PBTc -U
0.25
0.33
0.5
0.33
0.5
--
0.25
0.43
0.45
0.45
---
0.33
0.5
--
0.25
0.45
--
1.16
0.18
1.4
--
0.33
0.33
--
0.33
0.33
0.3
0.35
--
APV-B2d -D
APV-B2d -U
MTTc -D
MTTc -U
EEe -D
EEe -U
PBTf U
(Lf =0.54 m)
HF4g D
(Lf = 0.125)
(Lf = 0.167)
Propellerd
*
0.40
0.41
0.46
1.3
1.2
1.1
1.93
2.58
0.95
0.91
0.74
0.67
2.1
1.7
0.89
c
DT : Draft tube ; Kumaresan and Joshi (2006); Chapple et al. (2002) ; Aubin et al. (2001)
d
Mishra et al. (1998); e Zhu et al. (2009); f Broz et al. (2004); g Kumaresan et al. (2005)
Table 4.1 illustrates clearly that the power characteristics for all types of pitched blade
impellers are very dependent on the impeller to tank diameter ratio or and on the offbottom clearance to tank diameter. As observed, the power number changes for the
same pitched blade impeller when the geometry is changed, so it is difficult to
compare the power characteristic between the pitched blade impellers with different
geometrical ratios. This agrees with (Chapple et al., 2002) who found that the power
number for pitched blade impellers is highly related to impeller position in the tank.
(Chapple et al., 2002) explained this changing as the flow at the impeller interacts
strongly with the proximity of the tank walls, so changes in the position of the
- 169 -

impeller in the tank can have a significant impact on the power number due to
changes in the flow patterns. The power number for RTP propeller is generally
smaller than that of other impellers except for HF impeller which has lowest power
number in down-pumping mode. It seems that the draft tube presence has very little
effect on power number.
4.4. Pumping Capacity

The pumping capacity can be determined by the pumping number NQp (Flow number)
and circulation number NQc of the impeller. They are calculated according to
equations 2.11 and 2.14. The pumping numbers and circulation numbers of RTP
propeller and several pitched blade impellers determined in previous works are
presented in Table 4.2 for up-pumping and down-pumping modes and for both with
and without draft tube configurations.
Table (4.2): The RTP propeller and several pithed blade impellers pumping and
circulation numbers with different configurations and pumping modes.
Impeller
Type
RTP-U
(Lf = 0.1m)
RTP-D
(Lf = 0.1m)
HFa -D
PBTc -D
PBTc -U
APV-B2d D
APV-B2d -U
MTTc -D
MTTc -U
EEe -D
EEe -U
PBTf U
(Lf =0.54 m)
HF4g D
(Lf = 0.125)
(Lf = 0.167)
Propellerd
*
NQp
dim/Tv
0.2
0.15
0.2
0.15
0.50
0.33
0.20
0.27
0.33
0.5
0.25
0.43
0.45
0.45
0.33
0.5
0.25
0.45
1.16
0.18
--
0.5
--
0.33
0.33
0.33
0.41
0.65
--
--
0.33
0.67
0.89
1.3
NQc
NQp
Cim/Tv
1.186
(With DT)
1.22
(With DT)
-0.96
-0.91
1.82
1.19
0.975
1.13
1.35
---
(With DT*)
(Without DT)
0.56
0.51
0.55
0.495
-0.59
-0.75
0.68
0.83
0.58
0.67
0.61
0.73
0.72
------
0.73
DT : Draft tube ; Kumaresan and Joshi (2006); Chapple et al. (2002) ; Aubin et al. (2001)
d
Mishra et al. (1998); e Zhu et al. (2009); f Bro et al. (2004); g Kumaresan et al. (2005)
For the RTP propeller the pumping numbers in the up-pumping mode are slightly
higher than those in the down-pumping mode for both configurations with and
- 170 -

without draft tube. Similar observations were reported for the rest except for APVB2d, where the down-pumping has higher pumping and circulation numbers. The
pumping numbers are slightly elevated with draft tube, but this elevation is not always
led to elevation of the discharge flow because part of the discharge flow is redirected
upwards as seen in the down-pumping mode. Same thing may exist in the up-pumping
mode but in a different way, as much flow participates to the secondary loop above
the impeller for propeller and draft tube configuration.
The RTP propeller pumping numbers are relatively close the HF impeller, but
comparing with the others, the RTP propeller has lower pumping number for the uppumping and down-pumping modes. The pumping numbers for the pitched blade
impellers show a noticeable dependence on the flow direction and the impeller
position in the tank, as the observed pumping numbers data are slightly affected with
changing geometry and impeller position.
Despite the low Cim/Tv applied in this work, the down-pumping circulation number
for the RTP-D propeller appears close to the reported range of circulation number for
the MTTc-D and APV-B2d-D circulation numbers, and RTP-D are higher than those
reported for the PBT and HF impellers.
The circulation numbers are strongly related to pumping numbers. As a consequence
the up-pumping mode seems to have higher circulation numbers than in the downpumping mode, the only exception is with MTTc, which shows a reversible behavior
with the pumping number. Generally for the propellers the circulation numbers are
between 1.2 and 1.5 to their pumping numbers. In this work the circulation number
are 2.12 and 2.22 for up and down pumping modes respectively. The slightly elevated
values of circulation numbers are due to the effect of the draft tube.
4.5. Agitation Index and Liquid Volume Quantification for the

Down-Pumping Mode with Draft Tube Configuration
The mixing index for the RTP propeller in the down-pumping mode and with draft
tube configuration was calculated according to the equations (2.15 to 2.17). Table 4.3
shows the agitation indices values (the volume-weighted average velocity) for RTP-D
and for several pitched blade impellers that have been determined in previous works
(Aubin et al., 2001; Mavros and Baudou, 1997). The agitation index for the RTP-D
propeller is noticeably lower than the agitation indices for pitched blade impellers.
This low agitation index is due to low impeller to tank diameters ratio (d pr/Tv=0.15)
rather than the applied grid area applied in this work. Comparing with pumping
number, the agitation index didnt show an important dependence on the pumping
direction mode: it is higher for PBT-U than PBT-D and it is of the same order for
MTT-D and MTT-U (Aubin et al., 2001).
- 171 -

Table (4.3): The agitation Index for RTP-D and various pithed blade impellers, (this
work and literature).
Impeller
Type
RTP-D
(Lf = 0.1m)
PBTa -D
PBTa -U
MTTa -D
MTTa -U
MTTb -D
b
Lightnin A310 -D
*
DT : Draft tube ;
Cim/Tv
dim/Tv
0.2
0.15
0.33
0.5
0.33
0.5
0.33
0.5
6.7
16.0%
22.5%
14.1%
13.7%
12.0%
0.5
6.7
8.0%
0.33
Aubin et al. (2001) ;
Ig
(rps)
3.47%
(With DT*)
Mavros and Baudou (1997)
The liquid volume quantification or distribution for the down-pumping mode with the
draft tube is illustrated in the Figure 4.14, where the liquid volumes distribution in the
vessel is related to determine ranges of the composite mean local 2-D velocities in the
vessel. The composite mean velocities are illustrated as a 3-D surface chart as seen in
Figure 4.14; the liquid volume movement is elucidated in the vessel except in the
swept volume of the impeller, in the volume of the cone at the vessel base, and in the
one of the shaft and baffles edges. From figure 4.14 it is found that in the majority of
the liquid volume, the mean velocity is low, smaller than of the order 0.11Vtip except
in three zones. The most elevated is the discharge and intake zone of the RTP-D
where mean velocity is less than 0.34Vtip excluding a very small part in this with
mean local velocities of the order between 0.34Vtip and 0.45Vtip. The other two zones
near the vessel bottom and walls have lower local velocities less than 0.23Vtip; these
two zones are generated due to the interaction of the flow induced by the RTP-D
propeller with both the vessel bottom and walls.
4.6. Mixing Time

4.6.1. The Effect of RTP Propeller Rotational Speed
The mixing time tm experiments are performed for different RTP propeller rotational
speed levels, N, of (1.67 - 5 rps) to determine the mixing performance for two
configurations with and without draft tube. The system is operated without the turbine
at the top. The flow regime is always kept in the turbulent state. The mixing time
measurements are conducted by the colourisation-decolouration method.
The investigations are applied for two modes of agitation: up-pumping and downpumping, to understand the effect of draft tube in both configurations with and
- 172 -

without draft tube (See Figure 4.15). The injection position of the decolorizing
solution is always the same at the surface of water in order to measure correctly the
mixing time achieved due to the circulation of air bubbles entrapped from the water
surface, as the position of the injection may cause difference in (Kawase and MooYoung, 1989).
Figure 4.15 illustrates the effect of RTP propeller rotation speed N on the mixing time
tm in a single liquid phase (water) of both down-pumping and up-pumping modes for
the two configurations with and without draft tube. From the results it is found that
the mixing time tm is reversely related with the RTP propeller rotation speed for all
tested runs. This agrees with (Bouaifi and Roustan, 2001; Hadjiev et al., 2006;
Houcine et al., 2000).
Figure 4.15 shows also that the mixing time for the up-pumping mode with draft tube
configuration is slightly higher than the down-pumping with and without draft tube
configurations, that agrees with (Patwardhan and Joshi, 1999) for impeller clearance
ratio of Cim/Tv < 0.33. The mixing time tm for down-pumping mode without draft
tube condition is higher than with draft tube; (Aeschbach and Bourne, 1972; Houcine
et al., 2000; Kumaresan et al., 2005; Tatterson, 1982) reported similar behaviour in
their works. (Kumaresan et al., 2005) used a longer draft tube (Lf = 0.167m).
From Figure 4.15 it can be observed for up-pumping mode, the mixing time is lower
without draft tube at moderate propeller speeds, this can be explained by the fact that
the up flow is enhanced due to the presence of a cone at the vessel bottom, which
contributes to redirect the flow upward to propeller intake and also to remove the dead
zone below the propeller, (Nere et al., 2003) have presented similar performance for
PBT impeller. It is noticed also that at higher rotation speed for down-pumping mode
with draft tube lowest mixing time is achieved.
The low values of mixing time in single phase condition are due to low ratio of (dpr/Tv
= 0.15), where this ratio leads to weak flow interaction between the RTP propeller and
the walls and the bottom of the vessel.
- 173 -
- 174 -
0.5-0.6
0.4-0.5
0.3-0.4
0.2-0.3
0.6
0.1-0.2
0.5
0-0.1
- 175 -
vm, (m S-1)
0.4
390
360
330
300
270
240
210
0.3
0.2
0.1
0
180
280
260
150
240
220
120
200
180
160
Axial Direction, z,(mm)
Radial Direction, r, (mm)
90
140
120
60
100
80
60
30
40
20
Figure (4.14): The distribution of the liquid volume in the vessel related with mean local composite velocity ranges for RTP-D propeller with
draft tube configuration.
- 176 -

Comparing the achieved mixing times for the given rotation speed of up and down
pumping modes with and without draft tubes as shown in Figure (4.15) with the
investigated flow patterns for each case shows that, for up-pumping mode the shorter
mixing time is accomplished without draft tube complies the fact that the radial
discharge flow without the draft tube is more intensive. The highest radial velocity in
this zone without the draft tube is greater by 18%. For the down-pumping mode the
shorter mixing time with the draft tube is in compliance to the fact that the radial
discharge flow with the draft tube near the vessel bottom and beside the cone is more
intensive. The highest radial velocity in this zone with the draft tube is greater by
30%.
For the given impeller rotation speed (N = 5rps) and with draft tube configuration, the
higher circulation number for the down-pumping mode produced by RTP-D propeller
(NQc =1.22) agrees with the shorter mixing time (tm=30s) for this mode when
compared with up-pumping mode that produced by RTP-U propeller, with lower
mixing index (NQc =1.186) and longer mixing time (tm=40s) are achieved, as shown in
Figure 4.15.
In the decolourization step, the colour was disappearing in a similar manner for downpumping mode and for up-pumping mode with and without draft tube. The colour was
disappearing gradually first near the bottom and the walls for the down-pumping,
whilst it was disappearing first near the water surface zone for the up-pumping mode.
For the down-pumping mode, the last zone where the colour was disappeared and
delaying the good mixing is the zone that corresponds to the centre of principal
circulation as shown in Figure 4.3, for the other configuration same behaviour was
observed.
140
Down-pumping with DT
Up-pumping with DT
Down-pumping without DT
Up-pumping without DT
120
tm, (s)
100
80
60
40
20
0
1.50
2.50
3.50
N, (rps)
4.50
5.50
Figure (4.15): The effect of RTP propeller rotational speed on the mixing time for
different flow and geometry conditions, h/dpr=1.47, Cpr/Tv = 0.2.
- 177 -

Figure 4.16 presents the dimensionless mixing time number relation with propeller
Reynolds number. The mixing time appears to be not very effected by the elevation of
Re. As a consequence this indicates that the dimensionless mixing time number Ntm
in a single phase (non-aerated) condition can be fairly determined apart from the
occurred flow pattern and the bubble presence in the liquid phase for adapted rotation
speed (1.67 -3.33 rps). From Figure 4.16 it is observed that the dimensionless number
mixing tmN stays relatively close for each flow modes and vessel configurations, this
indicates that these configurations and flow modes can be characterized by their
mixing number.
For the entire tested configuration, it is important to mention here that the low mixing
time numbers means that efficient mixing for the given geometrical configuration is
achieved.
260
Down-pumping with DT
Up-pumping with DT
240
Down-pumping without DT
Up-pumping without DT
Ntm, (-)
220
200
180
160
140
120
20
30
40
50
60
-3
Re*10 , (-)
70
80
Figure (4.16): The relation of Reynolds number with the dimensionless mixing time
for different pumping modes and geometry configurations , h/dpr = 1.47, Cpr /Tv = 0.2.
4.7. Conclusions
In this chapter the flow patterns and r-z velocity vector fields and contour maps in a
single phase for the RTP propeller agitated vessel were measured by applying both
PIV and LDV for up-pumping and down-pumping operation modes with and without
draft tube configuration. The LDV measurements were restricted inside the draft tube
and in the measurement plane at 90o from the posterior tank baffle.
For the down-pumping mode, mainly one recirculation loop is generated with RTP-D
propeller with two notable secondary loops; one below the propeller and the second
- 178 -

beside the propeller blades. When the draft tube is removed the flow field did not
show an important change in the vessel core, except that one of the two secondary
loops has disappeared. For the up-pumping mode in the vessel core an overall
circulation loop is generated beside that, one important secondary loop is formed
above the propeller. Similar velocity vector map is observed when the draft tube is
removed. The quantitative analyses were made for the axial velocities for both planes
at 0o and 90o from the posterior baffle of the tank. The results showed that the velocity
profiles are different for these two planes with draft tube presence especially in the
region near draft tubes (i.e.) below the propeller. This difference is lessened when the
draft tube is removed. Starting with propeller rotation speed 4.6 rps, a surface vortex
appears and extends to the propeller blade tip in the down-pumping mode without
draft tube configuration, this phenomena is created.
Generally the pumping number is found very dependable on the flow direction. The
pumping capacity of the RTP propeller doesnt exhibit a great change when the draft
tube is placed. The up-pumping mode shows a slight elevation for the pumping
number compared with the down-pumping mode for both configurations with and
without draft tube. The power consumption for the RTP propeller has a relative
dependence on the flow configuration, as in down-pumping mode a limited increase
in power consumption is observed compare to up-pumping mode.
The down-pumping mode agitation index showed low mean local velocities with
respect to those in the previous works for pitched blade impellers. These low mean
velocities are due to the influence of the low propeller diameter to vessel diameter
ratio dpr /Tv = 0.15.
Mixing time shows an obvious dependence upon the variation of rotation speed in the
single phase conditions with different influence intensity. The mixing time is
increased by 20% without the presence the draft tube in the up-pumping mode. For
the down-pumping mode the mixing time is enhanced by 15% when the draft tube is
placed around the RTP-D propeller. From these observations it is very difficult to
clarify that one of the RTP configurations and pumping modes has an apparent
influence to improve the RTP propeller performance. In summary with respect to RTP
propeller pumping capacity, power consumption and mixing time for the various
tested configurations and operation modes, it is found that no general approach can be
recommended to improve the performance of the RTP propeller that is more efficient
than the already approved down-pumping mode with draft tube.
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- 180 -
- 181 -
- 182 -
Chapter Five: Hydrodynamics and Flow Pattern in Aerated Agitated Tank
Chapter Five
Hydrodynamics and Flow Pattern in Aerated Agitated Tank
5.1. Introduction
In this chapter the flow patterns and velocity fields for the aerated vessel have been
investigated. Two phases air-water are occurring since the water droplets are sprayed
in the air and then with their impacting at the water surface a sufficient amount of air
bubbles are entrapped and entrained inside the water bulk. To evaluate the aeration
performance of the system and to estimate the contribution of each element (Turbine,
RTP propeller and draft tube) on the process different experimental sets were
performed.
The PIV measurements for the flow pattern and mean velocity were carried out in the
r-z plane (Plane A) same as described in chapter four, where r-z measurement plane
was positioned directly on posterior baffle (0o) and at (60o) from the two lateral
baffles from front face (See Figure 4.1). The number of image pairs was (300) and
exposure time delay (t) between each two sequent images was (2500 s) depending
on operation condition such as upper turbine tip velocity and image scale factor
adapted during the acquisition. A pixel interrogation of (32%x32%) with (50%
overlap) was applied for the velocity measurements.
The conical shape turbine was placed at the water surface level (C/Tv=0.313, D/Tv=
0.238, h=0.28m). The used vessel is the 3-baffled cylindrical vessel as described
earlier (Tv =0.8 m). The four 45o pitched and twisted blade propeller RTP effect was
studied at clearance (Cpr/Tv = 0.2 m) with and without the draft tube (df =0.15 m, Lf =
0.1 m) that was positioned in the vessel with clearance (Cf = 0.1m). The RTP
propeller and the draft tube were removed in a part of the measurements. The
measurements were performed with turbine rotation speed of (N= 2.08 rps) to ensure
staying always in turbulent regime for created flow field (Re = 75000) and prevent
excessive air bubbles presence with higher rotation speed, which leads to
miscalculation of r-z mean velocities vector plot by PIV due to the high reflection of
laser bean occurred by these bubbles. The PIV measurements were performed in the
entire tank to identify the effect for the both the draft tube and/or RTP propeller
configurations on the produced mean velocity field and liquid flow patterns.
5.2. Flow Pattern and Mean Velocity Field in the Aerated Tank
The flow pattern and mean velocity fields in the aerated tank for the whole system is
illustrated in the Figures 5.1 and 5.2. From these Figures the time average velocities
vector plot and contour inside the aerated tank at the r-z plane at 0o of the posterior
- 183 -

baffle shows that a main circulation loop occurs in the tank. This circulation is mainly
generated by the turbine, where the falling water droplets at the water surface creates
a strong radial stream of the water and air bubbles near the water surface toward the
walls, which then this flow is redirected by these walls down-ward to the tank bottom.
The circulation node is approximately positioned in the mid-distance from the tank
bottom to the water level and closer to the walls than the tank center (r = 270-290mm,
z = 110-130mm). A part of the returning upward flow in the main recirculation loop is
induced radially without passing through the turbine intake region, as this flow is
induced partially due to the impingement and the plunging of the water droplets in the
subsurface water area so a significant amount of the circulation flow returns to the
plunging position without being induced by the turbine. The part of the flow that
drawn by the turbine blades is propelled as water droplets through the atmospheric air
mostly by the vertical side of the turbine blades and especially from the blades upper
tip; these droplets then falls down at the water surface and plunges in the water
subsurface zone and creates the main recirculation loop. The air is entrained with the
droplets impingement at the water surface as air bubbles and then is dispersed in the
tank by the liquid recirculation loop.
One secondary recirculation loop is generated in the central region of the tank above
and below the RTP-U propeller. The discharged flow above the propeller and below
the turbine is mainly drawn by the turbine blades except that some of the liquid is
moved toward the shaft and then goes down to the intake zone of the propeller and
forming the secondary loop. The flow in the secondary recirculation loop extends
down till the upper part cone, where the cone assists to redirect the liquid flow upward
to the propeller.
As shown in Figures 5.1 and 5.2 there are three high velocity zones in the tank; the
first one is at the tip of the turbine blade. The second high velocity zone is the water
impingement and plunging position at the water surface and subsurface. The third
zone is near the tank wall, where the generated descending flow from the water
droplets plunging zone has relatively kept its high velocities due to the interaction
with the tank wall. The liquid flows in the third zones losses its high velocity
gradually with going down toward the bottom of the tank. Very few studies
concerning the mean velocities profile and flow pattern measurements have been
made of the aerated water inside the vessel for similar surface aeration configuration,
but generally the studies were performed for the stirred aerated tanks with air supply
sources. (Lee et al., 2001) have performed CFD simulation for the flow profile that
generated by the spray impingement at the water surface for a tank that aerated by
surface aerator turbine, whilst (Kang et al., 2001) have carried out CFD simulation for
similar tank that used by (Lee et al., 2001) but for two impellers, where an impeller
was added below the turbine at the same shaft. The flow pattern with that determined
by (Kang et al., 2001) for the dual impeller configuration ( D/T= 1/3, T/h=1/1.84)of
the conical type turbine at the water surface and a PBTD impeller in the liquid has
two distinct circulation loops for each impeller, the upper one is similar to that occurs
in our system, the down pumping PBTD impeller apparently produces the lower loop,
- 184 -

where the flow comes from the upper loop is pushed down-word and returns to the
impeller intake zone above the impeller.
The produced time average velocity in the aerated tank as shown in Figure 5.2 has the
highest axial velocity in the droplets plunging water subsurface zone and is of the
order 0.415Vtip. The highest generated radial velocity exists in the same zone is of the
order 0.64Vtip, which is greater than the highest radial velocity that produced in the in
the tested case of the agitated vessel with the up-pumping RTP-U propeller and draft
tube only (See chapter four)by 66%.
In the high flow zone near the vessel wall the highest axial velocity in the downward
flow bulk is about of the order 0.4Vtip, which is close to the highest axial velocity that
produced in the high flow zone near the vertical side the turbine blades that of the
order 0.42Vtip. In this zone a high radial velocity is achieved of the order 0.47Vtip. The
upward axial velocity in the discharge flow zone of the RTP-U propeller which is in
the same time the intake zone of the upper turbine is of the order 0.17Vtip, while this
velocity is raised to the order 0.32Vtip just below the turbine blades, the radial velocity
just below the turbine blades reaches 0.08Vtip.
5.3. The Effect of the Propeller and Draft Tube

The vector plot and velocity contour map of the r-z plane when both the RTP
propeller and the draft tube were removed are illustrated in the Figures 5.3 and 5.4
respectively. The flow pattern in the aerated vessel without the RTP propeller and the
draft tube shows a slight difference with that obtained in the previous section 5.2. The
only difference is in the region near the shaft, where the secondary recirculation loop
is not distinguished as it was for the whole system configuration, where a weak flow
zone is presented in this region and it extends downward to the cone. From the Figure
5.3, it seems that the propeller and the draft tube presence results in a slightly more
intensive flow than that produced by the turbine alone.
Figure 5.4 confirms that the generated flow by both the turbine and propeller with
draft tube is slightly more intensive. Figure 5.4 shows also that the high flow zone
near the vertical side of the turbine blades is disappeared. The other two high flow
zones near the vessel wall and at water subsurface regions are relatively smaller.
Considering the axial velocities when the propeller and draft tube are removed, the
highest downward velocity is observed in the high flow zone near the vessel wall is of
the order 0.377Vtip. The highest generated upward axial velocity below the turbine
blades is of the order 0.288Vtip. The highest upward axial velocity in the zone near the
vertical side of the turbine blades is of the order 0.17Vtip. The upward velocity in the
intake zone for the turbine is around the order 0.16Vtip; this velocity is slightly smaller
- 185 -

than that produced with the presence of the propeller and the draft tube at the same
region by 5%.
The radial velocities in the case of the turbine alone configuration shows same
behavior as the axial velocities, where their values are slightly lower than the whole
system configuration. The highest radial velocity generated in the water subsurface
zone is of the order 0.6Vtip.
Whilst below the turbine blades the highest radial velocity is of the order 0.29Vtip and
the generated radial velocity near the vertical side of the turbine blades is of the order
0.294Vtip, which is less than that obtained in the whole system configuration by 35%.
5.4. The Effect of the Draft Tube

The generated flow by the RTP-U propeller and the turbine without the presence of
the draft tube as presented in the Figure 5.5 is generally similar to the produced flow
that presented in the whole system and the turbine alone configurations. A weak flow
region near the shaft exists same as the turbine alone configuration. Figure 5.6 shows
the flow pattern in the aerated vessel has two distinguished high flow zones at the
water subsurface region and near the vessel walls, whilst the high flow region near the
vertical side of the turbine blades is mostly disappeared. From Figure 5.6 it can be
found that the produced flow when the draft tube was removed is slightly less
intensive than the induced flow with the draft tube as seen in Figure 5.2.
The maximum axial velocity that produced by both the RTP-U propeller and turbine
without draft tube presence is of the order 0.367Vtip, this velocity exists in the
downward high flow zone near the vessel wall and it is less than that generated at the
same region in the whole system configuration by 7%. The highest upward velocity
just below the turbine blades is 0.32Vtip, whilst this velocity is of the order 0.17Vtip in
the turbine intake zone or in other word this zone is the RTP-U propeller discharge
zone, which is similar to the produced velocity in the whole system configuration.
The highest axial velocity in the zone near the vertical side of the turbine blade is of
the order 0.32Vtip.
With respect to the radial velocities that achieved here, these velocities are found
slightly lower that induced in the whole system configuration, in the zone localized
near the vertical side of the turbine blades a highest radial velocity of 0.147Vtip is
achieved, while in the water subsurface zone that created due to the falling droplets
plunging into the liquid bulk the highest generated radial velocity is of the order
0.54Vtip, these radial velocities are lower by 16.5% and 31% than the radial velocities
that achieved with draft tube presence respectively. The maximum radial velocity that
produced at the lower edge of the turbine blades is of the order 0.05Vtip.
- 186 -
- 187 -
Figure (5.1): The (r - z) flow field map for the turbine and the RTP-U propeller with draft tube (Whole system) configuration in the entire vessel.
- 188 -
- 189 -
Figure (5.2): The (r - z) velocities contour map for the turbine and the RTP-U propeller with draft tube (Whole system) configuration in the
entire vessel.
- 190 -
- 191 -
Figure (5.3): The (r - z) flow field map for the turbine alone configuration in the entire vessel.
- 192 -
- 193 -
Figure (5.4): The (r - z) velocities contour map for the turbine alone configuration in the entire vessel.
- 194 -
- 195 -
Figure (5.5): The (r - z) flow field map for the turbine and the RTP-U propeller configuration in the entire vessel.
- 196 -
- 197 -
Figure (5.6): The (r - z) velocities contour map for the turbine and the RTP-U propeller configuration in the entire vessel.
- 198 -

5.5. Turbine Pumping Number and System Circulation Number
The turbine pumping number NQp (Flow number) and the system circulation number
NQc for the three aerated configurations that investigated in this chapter are
determined by equations 2.11 and 2.14 respectively, the used diameter in the
calculation is the turbine diameter and illustrated in Table 5.1.
Table (5.1): Turbine pumping number and system circulation number.

Impellers Configuration
NQc
NQp
Whole system
2.42
0.49
Turbine + RTP-U
2.33
0.44
Turbine alone
2.24
0.40
Table 5.1 shows that for the whole system configuration, the turbine pumping number
and the system circulation number are higher than those for both with and without
draft tube and turbine alone configurations. The turbine pumping number is slightly
increased with draft tube presence than the turbine alone configuration. From these
values it is found that the presence of the draft tube assists the RTP-U propeller
performance by redirecting part of the RTP-U propeller radial discharge flow
discharge flow upward as it was seen in the up-pumping in the Mixing mode.
The RTP propeller function seems to have an obvious effect on the turbine pumping
number and the circulation number of the system, where these parameters are
relatively elevated, when the RTP-U propeller was positioned below the turbine, this
elevation occurs due to the improvement of the intake flow of the turbine that
achieved by the propeller action.
5.6. Agitation Index and Liquid Quantification for the Whole System
Configuration
The agitation index (the volume-weighted average velocity) for the whole system
configuration was calculated according to the equations (2.15 to 2.17). The agitation
index is equal to 5.26%, this value is higher than for the non-aerated vessel with
down-pumping mixing mode agitation index of 3.47% (N=5 rps) despite that for the
aerated mode the rotation speed is higher 2.08 rps and the presence of the air bubbles
may hinder the mixing to be efficient easily. The agitation index in the aerated mode
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is higher due to implemented geometry of the turbine compared with that of the RTP
propeller.
The value of the agitation index in this case is low compared to those of classical
configurations with dual impellers (8.8%) (Garcia-Cortes et al., 2006). This low value
is due to the non-classical configuration for which the ratio agitator/tank diameter is
quite low.
The liquid volume quantification or distribution for the whole system configuration
(aeration mode) is illustrated in the Figure 5.7. The liquid volumes distribution in the
aerated vessel is related with determined ranges of the composite mean local 2-D
velocities in the vessel as demonstrated in the chapter four, where the composite mean
velocities are determined as a 3-D surface chart. The liquid volume distribution in
aerated vessel is presented in the vessel except the swept volume of the turbine and
the RTP-U propeller and the volume taken by the cone at the vessel base or that is
taken by the shaft or baffles edges.
Figure 5.7 shows that an important part of the liquid volume has relatively low mean
velocities less than of the order 0.22Vtip except three elevated mean velocity zones,
this observation agrees with the r-z velocity contour (Figure 5.3), the high composite
mean local velocity zones are same as r-z mean velocity contour.
These zones are localized in the liquid subsurface zone, near the vessel wall and near
the vertical side of the turbine blades. The highest composite mean local velocity level
is presented in the zone below the water surface with the ratio of 0.5-0.6 m/s. Both the
zone near the vessel walls and the zone beside the turbine blades have lower local
velocities ratio of 0.4-0.5 m/s.
5.7. The Mixing Time

The mixing time investigations were carried out in the aerated mode, with impellers
rotational speeds N ranges between (1.67 -3.33 rps). In the aeration mode it is very
crucial to let the water droplets spray propelled by turbine blades and then plunged at
water surface within the tank area, for this reason the upper limit for the impellers
speed cannot exceed 3.33 rps. Two geometrical configuration were tested that the
whole system and the turbine alone configurations wherein the RTP-U propeller and
the draft tube were removed.
- 200 -
0.5-0.6
0.4-0.5
0.3-0.4
0.2-0.3
0.6
0.1-0.2
0.5
- 201 -
vm, (m s-1)
(vm) levels
0-0.1
0.4
390
360
330
300
270
240
210
0.3
0.2
0.1
0
180
280
260
150
240
220
Radial direction, r , (mm)
120
200
180
160
Axial direction, z,(mm)
90
140
120
60
100
80
60
30
40
20
Figure (5.7): Distribution of the liquid volume in the vessel related with mean local composite velocity ranges for whole system configuration.
- 202 -

5.7.1. Impellers Rotational Speed Effect
The repeatability for the measurements is illustrated in Figure 5.8, where the results of
two experimental runs for the mixing time number relation with rotation speed N. The
mixing time number Ntm values appear to be relatively close despite the increasing of
N. The mixing time number Ntm in the aeration mode (two phase system) can be
fairly determined between 40-55 for adapted rotation speed range (1.67 -3.33 rps).
The mixing time number shows for the given operated condition is lower than the
values of the mixing mode non-aerated condition.
100
Ntm, (-)
80
60
40
20
1.60
2.10
2.60
N, (rps)
3.10
3.60
Figure (5.8): The relation between impellers rotation speed and the dimensionless
mixing number Ntm for whole system configuration in the aerated vessel.
The relationship of mixing time tm with the impellers rotational speed N is illustrated
in Figure 5.9. For the whole system configuration, the mixing time appears to be very
dependent on the elevation of impellers rotation speed like the single phase condition
(Mixing mode), where it decreases with increasing of the impellers speed. As a
consequence this indicates that the mixing time in the aeration mode is highly effected
by the flow pattern and the bubbles presence in the liquid phase.
For the turbine alone configuration, both the draft tube and the propeller are removed.
Figure 5.9 shows the mixing time tm is highly affected by the elevation of N. The
mixing time tm is lessened remarkably from 33 sec to 17 sec when the rotation speed
was doubled (1.67-3.33).
- 203 -

35
Whole system
Turbine alone
30
tm, (s)
25
20
15
10
1.50
2.00
2.50
N, (rps)
3.00
3.50
Figure (5.9): The effect of impellers rotation speed on the mixing time for two
different geometrical configurations, h/D=1.47, Cpr/T= 0.2, C/T = 0.313 (Aerated
mode).
5.7.2. Effect of Propeller and Draft tube Presence

The propeller and draft tube effect during the aeration is investigated to interpret their
contribution on the achieved mixing time tm. The experiments are carried out for the
rotational speeds N ratio as mentioned in previous section, for the turbine alone and
whole system configurations.
As it is illustrated in Figure 5.9, the whole system configuration has lower tm than of
the turbine alone configuration for the given range of N. The importance of the RTP
propeller and the draft tube contribution was interpreted by the low values of the
accomplished mixing time tm, where the performance of the RTP propeller and the
draft tube assists to achieve the homogenization condition in the entire aerated vessel
within slightly shorter mixing time. The effect of RTP propeller and the draft tube on
the mixing time is interpreted by the relationship between dimensionless mixing time
with Reynolds number for the two different configurations; the whole system and the
turbine alone for different impellers rotation speed as shown in the Figure 5.10, where
the dimensionless mixing time Ntm is shorter with the presence of both the RTP
propeller and draft tube, the improving of Ntm was considerably increased with the
increasing of the Reynolds number.
- 204 -

100
Whole system
Turbine alone
Ntm, (-)
80
60
40
20
20
40
60
80
Re* 10-3, (-)
Figure (5.10): The relation between the Reynolds number and the dimensionless
mixing time of the aerated vessel for whole system and turbine alone configurations,
h/D= 1.47, Cpr/Tv = 0.2, C/Tv = 0.313.
5.7.3. Effect of the Spacing between Two Agitators

The spacing between Sp the turbine and the RTP-U propeller was tested to find out its
effect on the mixing time. An additional lower position was examined of Sp= 83 mm.
This lower position is placed more inside in the draft tube than the primary position
Sp= 76 mm. The mixing time shows a slightly difference compared with first position
as seen in the Figure 5.11. These results agree with (Hsu and Huang, 1997), the
mixing time increased with increasing the spacing may come from the fact that longer
spacing leads to weaker interaction between the two impellers.
- 205 -

36
Sp =76 mm
Sp = 83 mm
31
tm, (s)
26
21
16
11
1.60
2.10
2.60
3.10
3.60
N, (s)
Figure (5.11): The relation between impellers rotation speed and the spacing between
impellers; h/D= 1.47, C/Tv = 0.313, (Aerated mode).
5.7.4. The Effect of the Turbine Blade Submergence

The regular liquid level in the surface aeration is adapted just at the upper edge of the
turbine blades. In other word a 100% submergence of turbine blades corresponds the
ratio S/W=1 or h/D =1.47, where below this level the turbine blades are partially
submerged and higher this level the turbine blades are over submerged. The
experimental tests were performed to find the optimum turbine blade submergence
that should be considered according to its effect on the mixing time within the range
of S/W = 0.58, 1 and 1.42 (h/D=1.42, 1.47 and 1.53). This means that the turbine
submergence at, upper and lower than the turbine upper edge level was examined
respectively (i.e. the clearance of the turbine remains constant only the submergence
is altered).
Figure 5.12, shows the relationship between dimensionless mixing time and Reynolds
number for the investigated turbine blades submergences. It is found that the increase
of turbine submergence leads to a decrease of the dimensionless mixing number, these
results agree with the results founded by (Rodgers et al., 2011). Figure 5.12 shows
that a high dimensionless mixing time was achieved at the turbine submergence ratio
S/W that equals 0.58 because at this level the turbulence intensity was lowered as less
water droplets were propelled by turbine blades and as a result less droplets are
plunging into water surface and slower velocities were generated for the liquid
circulation in the liquid bulk, as consequence longer mixing time takes place. In the
case of higher water level, S/W that equals 1.42 (i.e. turbine blades are over
- 206 -

submerged, h/D= 1.53), the dimensionless mixing number obtained are not much
higher than that of, S/W equal 1.0(water level is just at upper edge of turbine blades).
By these results it can be considered that for each impellers rotation speed the turbine
blades are able to propel a certain quantity of water droplets into the air depending of
the implemented turbine blade submergence.
80
h/D=1.42
S/W=0.58
S/W=1.00
h/D=
1.47
70
S/W=1.42
h/D=1.53
Ntm, (-)
60
50
40
30
20
20
25
30
35
40
-3
Re *10 , (-)
45
50
mixing time of the aerated vessel for three turbine blade submergence levels, Cpr /T=
0.2, C/T = 0.313.
5.7.5. Mixing Time Modelling

From already developed correlations it can be noticed that the mixing time is found
dependant on various relevant parameters that can be taken in account such as; the
operational parameters which it in turn may include; rotational speed, acceleration of
gravity and flow passing through the lower propeller blades (in multiple impellers
case), the property parameters; water density and viscosity, the geometrical
configuration parameters, such as the turbine diameter and blade submergence. Many
other parameters can be involved in the influence on mixing time, where which is
varied according to each system particularity.
The aerated mode mixing time correlation is usually depends on the performance
criteria for the surface aeration system that obtained in the literature and from the
experimental results. The applied investigated parameters for the mixing time in
aerated mixing mode in previous works dont match the current case in order to
correlate a model that can be impeccable for extrapolation and scale-up purposes.
- 207 -

For the aerated condition, the main objective for the modelling is to characterize the
mixing time for the used agitator system (Whole system configuration). According to
the experimental results and observations, the most important relevant parameters for
the target quantity mixing time, tm, are demonstrated in the following relevance list;
(5.1)
By applying Buckingham II of dimensional analysis theory, the general influencing
factors may affect the mixing time in surface aerated condition can represented as:
(5.2)
Since all experimental runs are performed within turbulent regime, Re is eliminated
(Zlokarnik, 1979) and the equation (5.2) is reduced and rearranged as follows:
(5.3)
Equation 5.3 is solved to determine the values of the constants by applying multiple
non-linear regressions for 62 experimental runs; the model will have the following
form:
; S/W= 0.58-1.42;
Fr= 0.054-0.215;
Np= 1.66 5.11
(5.4)
From equation 5.4 it can be found that the decreasing of the turbine blade
submergence lengthens the mixing time within the range tm (S)-0.16. This means
when the submergence of the turbine blades decreased from 24 to 14 mm, the mixing
time would lengthen by the factor 1.1.
The comparison between the model and the experimental results is shown in Figure
5.13. The agreement is quite satisfying. It is somehow hard to compare the constants
in the equation (5.4) with the previous models because of the difference in the
geometry and mode of operation applied, where the already developed aerated
condition mixing time models didnt applied the water droplets projection principles
to achieve the aimed aeration. But generally for the negative exponent of power
number it seem to be logical and refers to that mixing time number is decreased with
higher power consumption (Nienow, 1997). The (S/W) ratio exponent refers to the
mixing time number is decreased when higher turbine blades submergence (higher
liquid level) which doesnt agree with the model that developed by (Kang et al., 2001)
with considering the difference in geometry with their system.
- 208 -

70
R2 =0.92
Ntm, Experimental
60
50
40
30
20
10
0
10
20
30
40
Ntm, Predicted
50
60
70
Figure (5.13): The comparison between experimental and predicted values of

dimensionless mixing time number (tmN) by equation (5.4).
5.8. The Power Consumption in the Aeration Mode

The performance of RTP propeller and the draft tube in the enhancing of the mixing
time in the up-pumping aeration mode is assessed as function of power consumption,
as illustrated in Figure 5.14. In the case of the presence of RTP propeller and draft
tube the Np is slightly increased. The power number is calculated with respect to the
diameter of the turbine.
The power per liquid volume relationship with N for aeration mode is illustrated in
Figure 5.15. With the propeller and the draft tube the power consumed is slightly
elevated due to their performance. The improvement of mixing time that
accomplished for the whole system configuration didnt consumed much higher
power than the turbine alone configuration. For the given implemented geometry, the
power consumption per liquid volume ratio in this case is increasing with increasing
impeller rotation speed N as it is illustrated in the Figure 5.14.
- 209 -

4.5
Whole system
4.0
Turbine alone
Np, (-)
3.5
3.0
2.5
2.0
1.5
1.0
20
30
40
Re*10-3,
50
(-)
Figure (5.14): The effect of Reynolds number on power number for the two different
configurations (Whole system and turbine alone); h/D= 1.47, Cpr/Tv = 0.2, C/T =
0.313, (Aerated mode).
250
Whole system
Turbine alone
P/V, (Watt/m3)
200
150
100
50
4
N,
(s-1)
Figure (5.15): The relation between impellers rotation speed and the consumed power
per water volume for the aerated mode, h/D= 1.47, Cpr/Tv = 0.2, C/Tv = 0.313.
- 210 -

5.8. Conclusions
The flow patterns and r-z velocity vector fields and contour maps in aerated and
agitated vessel were measured by employing PIV technique, the flow was generated
by the turbine and RTP-U propeller for up-pumping and down-pumping operation
modes with and without RTP-U propeller or the draft tube configuration. The
measurement plane is same that used for PIV measurement in chapter 4 (Plane B).
The flow field didnt show an evident difference when the RTP-U and draft tube were
removed, where the general circulation loop has kept its merits in all three
configurations.
The circulation number and pumping capacity of the turbine showed a relative
improvement when the RTP-U was place beneath. For the turbine and lower RTP-U
propeller configuration: the turbine pumping capacity and the circulation number are
mildly improved with draft tube presence.
Mixing time shows an obvious dependence upon the variation of the impellers
rotation speed in different configurations for the aerated mode with different influence
intensity.
From the results an optimum liquid level can be established. The mixing time is
decreased effectively by lowering the liquid level as well as decreasing the turbine
blade submergence lower than 100%. Moreover the elevation of liquid level as well as
an increase in turbine blades higher than 100% submerged increases the mixing time
but in a lighter manner than what occurs in decreasing the liquid level.
A model have been developed to interpret the mixing time number (Ntm) behaviour
related to more significant and affecting dynamic and geometrical dimensionless
parameters in the surface aerated agitated tank, the model is applicable for within the
limits; (Fr = 0.054-0.215), (S/W =0.58-1.42) and (P/V = 30-200 watt/m3).
From the results and the observations performed in this chapter it is identified that
none of the RTP-U and the draft tube configurations has improved separately the
turbine performance regarding the turbine pumping number and circulation number.
The liquid distribution that ranged according to the mean local velocity in the vessel
agrees with pumping capacity behavior of the turbine. In summary with respect to
turbine pumping capacity, circulation number and mixing time for the various tested
configurations, it is figured out that both the PRT-U and draft tube can together
enhance the performance of the surface aerator turbine.
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The general global purpose of this thesis is to evaluate and interpret the capacity for
the novel technology applied with the investigated surface aerator related to the
possibility to perform the aeration or the mixing by the same equipment.
The first main objective of this study consists in the characterization of the aeration
and the mixing modes capacities related to the operational and system configuration
parameters. The second major objective involved the investigation of the flow pattern
and velocity field takes place in the tank using advanced measuring techniques for
both up-pumping aeration mode and down-pumping mixing mode.
The implemented surface aerator has a distinctive feature of easy putting in place
between the aeration and the mixing modes by reversing the sense of the rotation
where the oversaturation condition is frequently occur in the liquid bulk with
existence of poor aerated zones. This interchanging is performed by applying the
decoupling (clutching) system that enables the lower propeller to carry out the mixing
with setting aside the turbine from the rotation.
A brief description was presented for the nowadays aeration systems for the water
and wastewater treatment with exposing the types and efficiencies for each system,
with emphasizing on the surface aeration.
From the literature review it is found that many modifications and developments are
made on the surface aerators to improve the oxygen mass transfer rate and to reduce
the power consumed. Many parameters have been studied like hydrodynamic, power
consumption, geometric configurations, oxygen mass transfer rate in the water,
modeling and scale up process. It can be presumed that in general the limiting factors
for successful surface aeration can be resumed as follows:
- The turbulent regime should be ensured in the entire water treatment tank that
contains sufficient dissolved oxygen entrained from atmospheric air.
- The generated water flow by surface aerators must be sufficient to reach all parts of
the treatment tank.
- The aeration process accomplished with the aerator turbine must be accompanied by
mixing and agitation performance to ensure an efficient distribution of the dissolved
oxygen.
- These constrains can be overcame by using an appropriate surface aerator with
suitable pumping capacity to handle efficiently and effectively the large quantities of
water.
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The measurements of the flow patterns and mean velocity fields were achieved by
Laser Doppler Velocimetry LDV and Particle Image Velocity PIV. Accurate results
have been obtained by these techniques.
For the water bulk zone it was found that the power consumption is related with
turbine rotation speed and blades submergence more than other parameters in uppumping aeration mode. It has been pointed out that a sufficient rotational speed,
called critical speed, is necessary to accomplish aeration process. As it was observed
the desirable flow patterns in the bulk zone can be obtained when the rotational speed
is over the critical speed. The air bubbles creation appears in the vessel due to
plunging water droplets through the water surface.
The water bulk oxygen mass transfer coefficient kla is highly dependent on the turbine
rotation speed in view of the increased kla in the water bulk zone with increasing of
impellers rotational speed. The rotational speed has to be lower than a roof value to
ensure that the falling water droplets would not hit the vessel wall. An optimum water
level was identified at the blade submergence ratio S/W equal to 1 regarding to the
transferred oxygen mass transfer. The influence of the spacing between the impellers
(turbine and propeller) stays related with the obligatory of remaining the lower
propeller inside the draft tube. The compliance of the kla to the spacing is exhibited
by the highest kla was obtained with the turbine presence with both the draft tube and
propeller. It was found the SAEb is more sensitive to rotational speed on contrary to
water level variation.
Two correlation models are developed for water bulk zone; one for the oxygen
transfer coefficient and another for power consumption. These models have
acceptable confidence factor (R2) values. For surface aeration spray zone, a model is
developed for standard oxygen transfer efficiency correlated with power consumption,
spray Reynolds and Froude numbers and water level ratio. As usual, the model is
applicable for particular limits.
For spray mass transfer zone, in the whole system configuration it is found that both
the oxygen transfer rate OTRsp and spray zone efficiency Esp are dependent on the
rotational speed and the turbine blade submergence. It is worth to mention that the
spray oxygen mass transfer coefficients klad is higher than liquid bulk oxygen mass
transfer coefficients kla. As a result it is figured out that the majority of oxygen mass
transfer is achieved in the spray zone due to high interfacial area and turbulent mass
transfer occurred in this zone.
The discharge flowrate of the water droplets spray Q strongly depends on impeller
rotation speed and water level in the tank: Q is increased when impeller rotation speed
or water level is increased for the given turbine blades submergence. The OTRsp is
calculated with and without the RTP propeller and the draft tube presence. The results
showed that OTRsp in whole system configuration is slightly higher than OTRsp for
the turbine alone configuration.
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The PIV and LDV measurements for the flow pattern were applied in a non-aerated
single phase condition (mixing mode) and in aerated gas-liquid condition (aeration
mode).
For the non-aerated condition, mainly one recirculation loop is generated with RTP-D
propeller with two secondary loops with down-pumping operation mode. When the
draft tube was removed the flow field didnt show any drastic change in the vessel
core. For high rotational speeds, a surface vortex appears and extends to the propeller
blade tip in the down-pumping mode when the draft tube is removed. The quantitative
analyses were made for the axial velocities for both planes at 0o and 90o from the
posterior baffle of the tank. The results showed that the velocity profiles are different
for these two planes with draft tube presence especially in the region near draft tubes
and below the propeller. This difference is lessened when the draft tube is removed.
For non-aerated condition with up-pumping operation mode in the vessel core an
overall circulation loop is generated and one important secondary loop is formed
above the propeller. Similar velocity vector map is observed when the draft tube was
removed. The pumping capacity of the RTP propeller doesnt exhibit a great change
when the draft tube was placed. The up-pumping mode shows a slight elevation for
the RTP-U pumping number compared with the down-pumping mode both with and
without draft tube configurations. The power consumption for the RTP propeller is
relatively dependent on the flow configuration. In the down-pumping mode slightly
increasing in power consumption is noticed compare to up-pumping mode.
Mixing time results in the mixing mode (single phase conditions) show a clear
dependence upon the RTP propeller rotation speed with different influence intensity.
The mixing time is relatively increased when the draft tube was removed in the uppumping mode. Contrary to the down-pumping mode the mixing time is enhanced
when the draft tube was placed around the RTP-D propeller.
It is very difficult to clarify that one of the RTP configurations and pumping modes
has an apparent influence to improve the RTP propeller performance in the single
phase vessel, where with respect to RTP propeller pumping capacity, power
consumption and mixing time for the various tested configurations and operation
modes, it is found that no general approach can be recommended to improve the
performance of the RTP propeller that is more efficient than the already approved of
down-pumping mode with draft tube.
The down-pumping mode showed low mean composite local velocities with respect
to those in the previous works for pitched blade impellers. These low mean velocities
are due to the influence of the low propeller diameter to vessel diameter ratio. The uppumping aeration mode mean composite local velocities were slightly elevated
compared with those in down-pumping mixing mode.
For the aerated condition, the up-pumping flow is generated mainly by the turbine.
The flow pattern didnt show an evident modification when the RTP-U and/or the
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draft tube were removed. The general circulation loop has kept its merits in all three
configurations, whilst the flow intensity was relatively higher when both the draft
tube and RTP-U propeller (Whole system configuration) are positioned in the vessel.
The flow is slightly more intensive when the the RTP-U was placed below the
turbine, but less than the whole system configuration.
The circulation number and pumping capacity of the turbine showed a relative
improvement when the RTP-U was place beneath. For the turbine and lower RTP-U
propeller configuration: the turbine pumping capacity and the circulation number are
mildly improved with draft tube presence.
The tested mixing time in the aerated gas-liquid condition for different configurations
shows high relation with the impellers rotation speed with different influence
intensity. With respect to mixing time it has been figured out that an optimum the
turbine blade submergence (liquid level) can be established when the submergence
was at the water surface level. This optimum parameter enables to achieve the
desirable mixing performance in the bulk zone with shorter time. The mixing time
was decreased submerged effectively by lowering the turbine blade submergence.
Moreover the elevation of turbine blades submergence higher than the water surface
increases the mixing time but in a lighter manner than what occurs in decreasing the
blade submergence. A model is built to interpret the mixing time number (tmN)
behaviour related to more significant and affecting dynamic and geometrical
dimensionless parameters in the surface aerated agitated tank.
From the results and the observations with respect to turbine pumping capacity,
circulation number and mixing time for the various tested configurations, it is found
that both the RTP-U and the draft tube has improved the turbine performance. The
liquid distribution that ranged according to the mean local velocity in the vessel
agrees with pumping capacity behavior of the turbine.
Generally the wide-ranged studies performed in this thesis show that the proposed
surface aerator system is found out efficient with respect to the aeration and mixing
capacities compared to those actually implemented in situ.
The particular studies on the hydrodynamic in the aeration mode carried out in this
thesis gave the necessary information on the used configurations. The whole system
configuration appears to have the most desirable. As the flow pattern has been
relatively improved when the draft tube and the propeller were positioned below the
turbine.
The experimental investigations in this thesis reveal that the system set up is
constructed in the way to be simple and easy to handle to facilitate the operation as a
result the maintenance cost is reduced.
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The experimental studies on the operation condition in the thesis demonstrate that the
critical rotational speed starts with the value 2.08 rps (effective bubbles distribution
and water spray started with this point). The optimum turbine submergence is founded
at the ration S/W=1.0 (higher aeration efficiency and effective flow pattern are
accomplished at this point). The upper reliable limit of the rotational speed is 2.5 rps
after this limit water droplets start to be splashed out of the tank walls. It is founded
also operating the system at optimum conditions (rotational speed, turbine
submergence etc.), this prevents give a rise of the power consumption.
The presented trends of investigations in this thesis are expected to promote further
surface aerator related studies. More knowledge about the water spray zone and the
water surface condition are needed to determine of the oxygen mass transfer as the
most of the oxygen is transferred within this zone and it is mainly controlled by
surface condition. Also it will contribute to the development of the surface aerator
efficiency. More specific flow pattern investigation could be performed on this zone
to characterize water droplets projection and impingement, water spray and the
condition for the water surface, as these parameters can affect drastically the
transferred oxygen in the system. The presented study of the spray zone oxygen mass
transfer in this thesis requires more studies to try to find out more accurate and
competitive methods to measure the dissolved oxygen concentration in the water
spray.
The experimental results obtained in this thesis have given limited information on the
bubbles condition in the liquid phase. More studies are needed to characterize the
bubbles condition inside water bulk. Such as the gas hold-up and the bubble size
distribution could be determined inside the vessel by performing various accurate
methods like photographic method, tomographic scanner, or other gas hold-up
sensors. For the bubble size distribution many methods could be carried out like the
optical sensors or image analysis methods.
Flow pattern in the aerated condition of the liquid phase generated by the up-pumping
turbine involved radial and axial velocities field acquisition. In order to finalize this
study a supplementary study is required on the produced tangential velocities in the
tank. For the tangential velocity component that expected to be substantial, the laser
sheet is positioned horizontally with a mirror placed beneath the stirred tank placed in
inclined plane to elucidate the tangential flow occurred inside the vessel. The other
technique could be implemented to have better flow visualization is the three
dimensional PIV, where in this case two digital cameras are needed (Chung et al.,
2007).
The changing between the aeration mode and the mixing mode is applied using a
coupling (clutching) system. Other ways of exchanging between the mixing and
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aeration could be explored in order to find out competitive methods accomplishing the
same objective.
The mixing process investigation presented in this thesis has been carried out using
one propeller type the RTP propeller. To enlarge this specific study other types of
axial impellers could be investigated such as PBT. It is required to keep the
symmetrical design of the tested propeller geometry in future works in order to
maintain the capability to exchange between aeration and mixing modes.
The geometrical configurations of surface aeration system performed in this thesis
have provided adequate information of mixing and aeration capabilities, however the
results are limited for fixed geometries for the tank and the internals. Same
investigations for the mixing and the aeration could be carried out with different
system geometrical configurations to seek for new system capabilities promoted with
geometrical alteration.
The CFD simulation of single phase mixing mode or gas-liquid phase aeration flow
mode patterns could be performed depending on the experimental results obtained in
this thesis. The CFD simulation could be employed to explore different geometrical
configurations differ from that applied this thesis, which not only enables the
investigation of various geometrical configurations for the vessel, impellers and
vessel internals. But also with CFD implementation, the time, efforts and expenses
could be reduced.
Scale-up for surface aerators can lean on the obtained experimental results to figure
out the demonstration of the general elements for the extrapolation and to decide the
types of the similarities that could be accomplished between the model type and the
proto-type are described. The obtained information from the experimental runs in the
lab scale could be verified for scaling-up purposes by applying this information in a
larger industrial scale surface aerator on the basis of geometrical and dynamic
similarity with that in lab. The SAEb may be regarded as a main objective that should
be maintained through extrapolation process.
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Appendix I: A Review for Biological Wastewater Treatment with Activated Sludge
Appendix I
A Review for Biological Wastewater Treatment with
Activated Sludge
The clean water availability question becomes more and more as a developing
important issue especially during the last decades. The necessity for wastewater
treatment regarded as a very important compromise among the available solutions to
preserve the water resources and to reduce the water pollution. The aerobic biological
purification that implied in the wastewater plants consists of many steps; generally it
can be divided into two main treatments: urban or domestic wastewater treatment and
industrial wastewater treatment.
I.1. Industrial and Domestic Wastewater Treatment

Wastewater from the industry contains diverse constituents of various chemicals and
undesirable deposits such as mineral scales, colloidal matters, corrosion products and
biological growth (Amjad, 2010). These components are separated before the
biological treatment. The removal and treatment of industrial waste water process
depends on the characteristics of the industry that produces the wastewater like
(cooling, desalination, oil production, etc.). Aeration with activated sludge is similar
generally with that applied in domestic wastewater purification, in which the
biological and physical-chemical wastewater treatments are combined (see Fig.I.1),
the primary and secondary processes handle most of the non-toxic waste water. The
primary treatment prepares the waste water for the next step of biological treatment.
All un- dissolved solids, dirt, floatable solids and else are removed by settling or
screening. In the secondary treatment the degradation process for the soluble organic
compounds carried out in the aerated treating vessels and lagoons. In the tertiary
treatment, different processes are used such as filtration (Eckenfelder et al., 2009).
Sometimes the industrial waste water is accompanied with other types of pollutions
such as the thermal pollution, where the waste water effluent is rejected with high
temperature more than 350C, which causes of deterioration of the aquatic systems.
The other type of the industrial
wastewater comes from the nuclear sites or
radioactive waste treatment plants or from the explosion of uranium mines (Roubaty
and Boeglin, 2007) and many of them contain radioactive pollution, similar
treatments steps are exist for the domestic wastewater treatment, and they will be
described in more details.
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Figure (I.1): Process selection diagram for industrial wastewater treatment

(Eckenfelder et al., 2009).
Domestic and industrial wastewater biological treatment is generally consisting in

four main steps; (i) preliminary treatment, (ii) primary treatment, (iii) secondary
treatment and (iv) tertiary treatment (See Figure I.2). There are two additional
(auxiliary) treatments of sludge and odor treatment. These steps of the treatment
combine the three types of treatment; chemical, physical and biological, where they
used to eliminate the existence of pollutants (Sardeing, 2002).
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Figure (I.2): Domestic wastewater treatments steps (Try and Price, 1995).
I.1.1. Preliminary Treatment Process

This treatment is used for gravity separation and accumulation of settleable solids
from the wastewater. where the removal of these settleable and floatable solids such
as (dirt, sand, etc.) by degreasing, sieving, filtering and settling to prevent their
clogging the effluents distribution system (Clesceri, 2009).
I.1.2. Primary Treatment Process

In the primary treatment, the solid particles that didnt settled out by the first step in
the preliminary treatment will be settled out by decantation. These solids are mainly
of organic or heavy metallic basis in the purpose of preparing the wastewater to the
next biological treatment. The colloids and other suspended particles are removed by
coagulation process especially the heavy metals pollutants combined with filtration,
where colloidal particles destabilized by charge neutralization. The primary treatment
contains other types of processes depend on the pollutants kind in the wastewater like
equalization. It is employed to minimize or prevent high concentrations of toxic
materials from entering the biological treatment plant and also to control the
fluctuations in wastewater characteristics. Another process exist in this step is the
neutralization, which is used if the acidic or alkaline materials are existed in
wastewater in order to neutralize the discharge of the primary treatment step before
further chemical or biological treatments (Eckenfelder et al., 2009).
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I.1.3. Secondary Waste Water Treatment
In this stage is the biological treatment is applied for the biodegradable organic
pollutants with aerobic process and anaerobic process. It consists in converting them
to stabilized non-pollutant materials such as nitrates, phosphates and carbonates, and
then they are removed by settling. The aerobic process involves the implementation of
activated sludge in the tanks, where the activated sludge is always in suspension
condition during the aerobic process. The required air for the aerobic process is
supplied by direct injection the air into the tanks or by entraining the atmospheric air
from the wastewater surface through the surface aeration technique. The activated
sludge contains the active bacteria that can be recycled to ensure the continuity the
process, wherein the solid materials are separated in followed settled steps
(Eckenfelder et al., 2009). Filtration is applied as a part of this stage of treatment to
deliver the best feed material to the next step by separating the fine solids and
suspended solids. Diverse filtration types are utilized such as membrane filtration.
The organic or ceramic membranes separate the two phases the waste water and the
micro-organisms to keep them inside the tank, where the membrane provides a
physical barrier to the larger complex organics. The wastewater pumped across the
membrane surface at high flow rate, some pollutants slowly accumulate on the
membrane and that needs periodically cleaning with solutions containing chlorine,
acids and/or bases (Try and Price, 1995).
I.1.4. Tertiary Wastewater Treatment

The tertiary step is to enhance the removal of remaining pollutants such as suspended
solids, organic matters, dissolved solids, toxic substance and nutrients (nitrogen and
phosphorus) in the wastewater after the prior biological treatment steps. To improve
the quality of the effluent removal a disinfection step is used to reduce the bacterial
levels. The residuals that contain the suspended solids, organics or colloidal materials,
are removed by filtration. The adsorption by activated carbon is also applied and used
to remove the remaining organics (Eckenfelder et al., 2009). The disinfection process
is achieved by the addition of chlorine or ozone. The other treatment in this step is
implementation the micro-screening where, they are considered influential to
accomplish the removal. Micro-screening include the use of low-speed continuously
backwashed, rotting-drum screen operating under gravity- flow conditions (Gupta and
Sharma, 1996; Shammas et al., 2006).
I.2. Activated Sludge Process

Activated sludge is applied for both industrial and domestic wastewater treatment.
The concept of this process is to create dispersed bacterial flocs and free-living microorganisms that are suspended within the water (Gary, 2004). The activated sludge is
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applied to eliminate the organics contained in the wastewater for both soluble and
insoluble materials by converting them into a flocculent microbial suspension and
they can be easily separated from wastewater by followed treatment steps like
gravitational solid-liquid process (Eckenfelder et al., 2009).
Generally the activated sludge is employed in the aerated tanks, batch reactors,
oxidation ditches, lagoons or long aeration basins with continuous flow and is mixed
with the wastewater. The produced biological effluents are separated from the
activated sludge and this sludge is returned to the inlet point of the basin or the tank to
ensure the continuous treatment. A rapid rate of microbial growth occurs during the
activated sludge process. The organic materials used to form oxidized end products
like CO2, NO3, SO4 and PO4 or to produce the biosynthesis of new micro-organisms.
The organic materials that are treated with the activated sludge are usually colloidal
and larger sized particulates and they contain also some by contents of light nutrients.
Most of these organic materials will be absorbed and agglomerated onto the microbial
flocs (See Figure I.3).
Figure (I.3): Development and control of biomass in the activated sludge process
(Gary, 2004).
The activated sludge treatment depends on several process factors, especially the
flocculation as described earlier. The successful agglomeration of the various organic
materials is related to their flocculent growth when mixed with activated sludge, the
settled biological materials from this operation are removed later by the separation
from the wastewater. The flocs can be determined as a biomass cluster of several
million heterotrophic bacteria bounded together with some inert organic and inorganic
material (Gary, 2004). The process of activated sludge includes several
implementations. Among of them there are nutrient removal processes:
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I.2.1. Nutrient Removal
Nutrient compounds presence in the wastewater comes from various sources, the
quantities of the nutrients is in continuous raise , which required to be treated in order
to eliminate its disturbance on the ecological balance in the aquatic mediums and
surrounding environment, (the eutrophication), (Nair et al., 2008; Wiesmann et al.,
2007).
A. Nitrogen Removal
Nitrogen presents in different forms in the wastewater due to various oxidation states,
which can be changed from one state to state to another a according to the different
physical and biochemical conditions involving wastewater. The source of organic
nitrogen may come from domestic wastes, farming activities and food industry. Most
of bio-nitrogen in wastewater is present as ammonia of unstable organic compounds,
but in urban wastewater, it contains excess nitrogen of the microbial requirement to
oxidize the amount of carbon present, so only a part of nitrogen is being combined
into the biomass (Crites et al., 2006).
The removal of the ammonia prevents its toxic effects on the aquatic life systems and
the deterioration of water quality (Gupta and Sharma, 1996). In some cases the N2O
gas could be emitted during nitrates removal from waste water as a secondary
product. The emitted gas is controlled by additional techniques like intermittent
aeration (Park et al., 2000).
The nitrogen is used by the biomass effectively, where ammonia is oxidized by
autotrophic nitrifying bacteria to nitrite. The biological treatment of ammonia by
oxidation is occurred in two stages each achieved by different type of nitrifying
bacteria; they utilize the ammonia as an energy source in the first step, or the nitrite
for same purpose in the second step as following:
NH4+ + 1.5 O2
NO2 + 2H+ + H2O
And the second step as:

NO2- + 0.5 O2
0.5 O2 + NO3-
The nitrification implemented with slow growing bacteria therefore, it proceeds at low
rate then, in order to achieve an acceptable treatment it needs maintaining more longer
retention time for the activated sludge in the tank or the basin by keeping maximum
contact with activated sludge to ensure maximum nitrification rate and to avoid washout of the nitrifiers. It is preferred in the nitrification process with activated sludge
maintaining the dissolved oxygen concentration about (2.0 mg/l), where it is remarked
that the nitrification doesnt occur below (0.5 mg/l), where with high dissolved
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oxygen the process is not inhibited just up to (20 mg/l), it is found out the process is
dependent on temperature and several organic matter existence (Gary, 2004).
Another process that occurred with nitrification is the de-nitrification where, nitrate
can be converted through the producing the nitrite to nitrogen gas under low oxygen
dissolved condition, this process occurred in the deprived of oxygen sludge layer of
sedimentation tank after the aerobic biological nitrification (See Figure I.4).
There are many modifications essays to improve the nitrification process like
applying the process involve activate sludge was immobilized onto support materials
like cloth or polyurethane foam (Nair et al., 2008), whereas for high-strength
ammonium wastewater that may contains industrial sources a two-sludge system for
both nitrification and de-nitrification process applied (Carrera et al., 2003),The
activated sludge nitrogen removal process also can be boosted by successful
manipulating control system with using an aerobic/anoxic/oxic scheme (Baeza et al.,
2002).
Figure (1.4): The schematic layout of two-stage nitrification and de-nitrification

(Gary, 2004).
B. Phosphorus Removal
The phosphorus exists in the wastewater in form of orthophosphates (PO4 3-, HPO4 2-,
H2PO4 -2, H3PO4), polyphosphates and organic phosphates, where its source is
generally comes from fertilizing implementation. The total phosphorus concentration
in wastewater is about (5-20 mg P/l), where only (1-2 mg P/l) is removed by
biological treatment (Sincero and Sincero, 2003). The phosphorus in wastewater
stream is removed by both chemical and biological methods (Figures I.5 and I.6);
where the main biological phosphorus removal is anaerobic-oxic system. The bacteria
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are conditioned in the anaerobic step with converting the phosphorus as soluble form,
the biomass of the activated sludge works to absorb this soluble phosphorus in the
followed aeration process step (See Figure I.5). With chemical removal step the final
concentration in the treated wastewater effluent will be (1 mg P/l); in the anaerobic
step the residence time depends on the (BOD: P) ratio in the inlet wastewater (Gary,
2004).
For specific treatment durations, there is deterioration for the biological phosphorus
removal, which is happened when excessive aeration of the activated sludge occurred
in previous treatment steps (Brdjanovic et al., 1998).
Usually coagulants are added after the biological treatment, where for the chemical
part is the phosphorus is removed by the addition of coagulants to wastewater and
then it removed as precipitates the inorganic forms of phosphates, many materials
used as coagulants such as lime, aluminum salts or iron salts. In the case of lime
addition to wastewater, it removes the phosphorus by two steps; in the first step it
reacts with natural alkaline to produce the calcium carbonate:
Ca(HCO3)2 + Ca(OH)2
2CaCO3 + 2H2O
Then the calcium ions combined with orthophosphate to form insoluble and
gelatinous calcium hydroxyapatite.
5Ca 2+ + 4OH - + 3HPO4 2-
Ca5(OH)(PO4)3 + 2H2O
Figure (I.5): The Two stage biological phosphorus treatment (Gary, 2004).
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Figure (I.6): Chemical-biological phosphorus treatment (Phostrip process), (Barth

and Stensel, 1981).
I.3. Aeration Process in the Activated Sludge Treatment

The oxygen here is supplied by diffused or mechanical aeration systems, where the
process is entirely biological treatment as a form of stabilization lagoon or basin used
in order to settling, removal of suspended solids, aeration and equalization beside the
biological treatment. The oxygen supply in aerated lagoons depends on the
mechanical aeration devices properties like the power level used for aeration.
The aerated lagoons has many advantages such as low operational and maintenance
cost, lagoons doesnt pass any harm to the environment in the same time it act very
effectively for wastewater treatment (Clesceri, 2009).
In the last 20 years a similar technique to lagoons is applied that is pond aeration
systems but of course without utilizing activated sludge, which became more and
more in demand to preserve the aquatic life like fish and bio-organisms to sustain this
aim , many air delivering tools employed like paddle wheel aerators (Boyd, 1998).
The pure oxygen is used with activated sludge process instead of using air due to its
effective oxygen dissolution rate during aeration operation. The injection of pure
oxygen to the liquid is applied in two process types that are covered and uncovered
process:
(i) The covered process, in order to transfer about 90% of the oxygen to the liquid
phase the technique of series of covered aeration tanks is used. (See Figure I.7).
(ii) The uncovered process, this method involves the usage of spargers or diffusers of
oxygen in deep position inside the liquid wastewater, so high concentration of
dissolved oxygen that reaches the activated sludge despite that opened cover of
treatment tank (Baeza et al., 2002; Carrera et al., 2003; Mueller et al., 2002).
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Figure (I.7): Covered aerated tank using pure oxygen (Mueller et al., 2002).
Usually the aeration is performed with active sludge in tanks or basins, their depth
vary from (3.5 6.0 m) to about (20-30 m ) equipped with air diffuser below the
liquid and inject the air or oxygen , where these types are preferred to treat industrial
wastewater especially with proper type of aeration tools was chosen. An effective
process that preferred to treat synthetic wastewater can remove more than 95% of
COD from the effluent (Polprasert and Raghunandana, 1985; Sincero and Sincero,
2003).
I.4. Additional Treatments

I.4.1. Sludge Treatments
Many processes treat the sludge settled down during the wastewater treatment stages.
The sludge produced is different for each one of implemented stages, the treatment of
settled sludge is expensive so in several times it returned to the inlet and settled out
with the primary sludge especially for the sludge that produced in the primary and
secondary steps. Generally the sludge treatment contains warming to (35 oC) for at
least two weeks to ensure that the organic solids were biologically broken down as
what called the digestion treatment, the accompanied produced gas is used to heat the
sludge (Clesceri, 2009; Try and Price, 1995).
The anaerobic and aerobic digestion by biochemical oxidative stabilization of the
sludge used in open or closed tanks separated from liquid process system.
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I.4.2. Odor treatment
As a final step to treat the wastewater stream is malodor gases elimination or what is
called odor treatment, where its implied by either biological or chemical methods, the
biological way is implemented for biological volatile organic compounds are removed
by biological filters or biological scrubbers (Kleinheinz and Wright, 2009), while the
chemical method is applied for odorous gases associated with volatile organic
compounds (VOC) like hydrogen sulfide and ammonia and other gases, the dissolved
gases stripped out of the liquid during the agitation or aeration, this method implies
chemical scrubber system and air/oxygen injection system.
- 222 -
- 222 -
- 235 -
- 236 -
Appendix II: Derivation of Spray Flowrate for the Surface Aeration System
Appendix II
Derivation of Spray Flowrate for the Surface Aeration System
Fluid mechanics analysis of the surface aerator system is performed to find a logical way
to calculate the water spray flowrate Q, where direct experimental measurement of Q
would be extremely difficult and expensive to accomplish this objective (McWhirter et
al., 1995). This calculation depends on the basic physical principles that occurred in the
surface aeration system related to the volumetric flow rate Q of the water spray
discharged by the surface turbine blades into atmospheric air.
The spray flowrate can be de determined by consideration of the overall conservation of
energy applied to a surface aeration system and the measured overall power input to the
aerator (Baylar et al., 2001; McWhirter et al., 1995). (McWhirter et al., 1995) envisioned
the surface aeration system as impeller blades rotating in the free liquid surface of a
relatively large body of water, they considered the surface aeration turbine impeller
blades accelerate a flow rate of liquid Q from a relatively low velocity at the inlet to the
impeller blades up to a relatively high velocity at the discharge from the tip of the aerator
impeller blades.
The general power input to the surface aeration system is a mechanical energy transfer to
the water bulk and then it is converted to kinetic energy in the water spray as explained in
chapter two, this power input can be presented as the following equation (Baylar et al.,
2001).
The fluid friction losses for the water flow across the surface aerator turbine blades is
considered negligible, the efficiency is assumed 100% as a result all the mechanical
energy is converted with 100% to a kinetic energy, so the efficiency, , equals to 1. Ym
represents the height between the upper point of the water spray maximum point.
By applying the general energy balance of Bernoulli equation between two points 1 and
2 as illustrated in Figure (II.1), where the total energy of a fluid in motion consists of the
following components; potential, pressure and kinetic energies (Holland and Bragg,
1995).
- 237 -
Figure (II.1): The energy balance points for the spray discharge flowrate.
Equation II.3 is reduced to equation II.4 as the pressures for the two points are same (the
atmospheric pressure).
Since the spray velocity v2 at the upper point (point 2) of the spray is equal to zero
By replacing the terms in Equation II.4 as following
The water is propelled from the tip of the turbine blades will be subject to the same
kinematic laws of physics as any free falling projectile (McWhirter et al., 1995), so the
velocity at the impingement point equals the velocity at the turbine blades tip.
So Equation II.2 can be written as follows
Equation II.8 is written as follows, for power input in kW and flowrate in liter/h
Where the constant 1.38910-7 includes the unit conversion factor and it depends on
system geometry as the discharge flowrate depends on the turbine geometry.
- 238 -
The flow rate Q is assumed to be constant along the spray, where the flow rate Q into the
atmospheric air at a relatively high total discharge velocity Vsp from the tip of the turbine
blades can be applied at the impingement point at the water surface. With negligible
friction loss all of the aerator shaft power input goes into accelerating the flow rate of
liquid spray Q up to the discharge velocity Vsp (McWhirter et al., 1995), so equation II.8
can be applied to relate the total power input with the discharged spray flowrate and
spray velocity Vsp at impingement point.
- 239 -
- 240 -
List of Symbols
List of Symbols
Latin symbols
a
constant in equations 3.2, 3.3and 3.6 (varying)
a term in kla
interfacial area per unit volume in water bulk, (m2 m-3)
ad term in klad interfacial area per unit volume in water droplets, (m2 m-3)
Ad
interfacial contact area of the water droplets, (m2)
constant in equations 3.2, 3.3and 3.6, (varying)
constant in equations 3.2, 3.3and 3.6, (varying)
Cd
concentration of dissolved oxygen in water droplets, (mg L-1)
Cdt
concentration of dissolved oxygen in water droplets at time t, (mg L-1)
Ce
concentration of dissolved oxygen measured by the probe, (mg L-1)
Cl
dissolved oxygen concentration in liquid, (mg L-1)
CLS
saturated concentration of dissolved oxygen in water, (mg L-1)
CLt
concentration of dissolved oxygen in water at time t, (mg L-1)
Cpr
propeller clearance, (m)
Cs
concentration of dissolved oxygen at saturation, (mg L-1)
Csd
saturated concentration of dissolved oxygen in water droplets, (mg L-1)
Cs10oC
saturation concentration of dissolved oxygen at 10C0, (mg L-1)
Ct
concentration of dissolved oxygen at time t, (mg L-1)
Ctr
turbine clearance, (m)
Co
initial concentration of dissolved oxygen at time t=0, (mg L-1)
constant in equation 3.2, (varying)
turbine diameter, (m)
df
draft tube diameter, (m)
dpr
propeller diameter, (m)
Emd
Murphree contacting efficiency for oxygen mass transfer, (-)

- 241 -
List of Symbols
Esp
spray mass transfer zone efficiency, (-)
(Esp)20
standard spray mass transfer zone efficiency at 20 oC, (-)
temperature correction function to 20 oC for the spray mass transfer

zone, (-)
Fr
Froude number, (ND2/g)
Frsp
modified Froude number for water spray, (gYm3)/qsp
acceleration of gravity (9.80665), (m2 s-1)
water level in the vessel, (m)
height of apparent turbine blade above water surface, (m)
Ig
agitation index, (%)
oxygen mass transfer parameter

(1.12,1.13,1.14 and 1.15), (-)
kla
volumetric oxygen mass transfer coefficient (Liquid bulk mass transfer

zone), (s-1)
klad
volumetric oxygen mass transfer coefficient (Spray mass transfer

zone), (s-1)
kla10oC
volumetric oxygen mass transfer coefficient at 10 oC (Liquid bulk mass

transfer zone), (s-1)
klaT
volumetric oxygen mass transfer coefficient at T oC (Liquid bulk mass

transfer zone),(s-1)
rotation speed, (s-1)
Np
power number, (P/ N3 D5)
NQp
Pumping number, (Qp / ND3)
NQc
circulation number, (Qc / ND3)
OTRb
oxygen transfer rate in water bulk mass transfer zone, (kgO2 h-1)
OTRsp
oxygen transfer rate in water spray mass transfer zone, (kgO2 h-1)
power consumed, (watt)
Pin
the input power to the turbine (that assumed 100% converted to kinetic
energy to propel water spray into the air), (k watt)
volumetric flowrate of the water spray, (m3/s)

- 242 -
(kla20(2/g)1/3),
in
equations
List of Symbols
Qc
circulated flow rate in the vessel, (m3/s)
Qp
flow rate by propeller, (m3/s)
Qpr
radial flow rate by propeller, (m3/s)
Qpz
axial flow rate by propeller, (m3/s)
qsp
water spray volumetric flowrate per spray radius, (m3 m-1)
radial coordinate, (m)
draft tube radius, (m)
rcent
radial coordinate at circulation center, (m)
Re
Reynolds number, (ND2/ )
Resp
modified Reynolds number for water spray, (Q/)
Rm
maximum radius of the water spray, (m)
Rsp
radius of the water spray calculated from the shaft center, (m)
turbine blades submergence, (m)
SOTRb
standard oxygen transfer rate in water bulk mass transfer zone,

(kgO2 h-1)
SAEb
standard aeration efficiency in water bulk mass transfer zone,

(kgO2 kWh-1)
time, (s)
vessel diameter, (m)
tf
flight time of water droplets, (s)
tm
mixing time, (s)
To
torque measured with filled vessel, (N m)
Toe
torque measured with empty vessel, (N m)
Tv
vessel diameter, (oC)
Va
the mean velocity of the water spray radial and tangential velocities
components, (m s-1)
Vd
water droplet volume, (m3)
VL
liquid volume, (m3)

- 243 -
List of Symbols
vm
mean volume weighted velocity, (m s-1)
Vsp
water spray velocity, (m s-1)
Vr
radial component of the water spray velocity, (m s-1)
Vy
axial component of the water spray velocity, (m s-1)
Vtip
propeller blade tip speed (m s-1)
vr
liquid bulk mean radial velocity (m s-1)
vz
liquid bulk mean axial velocity (m s-1)
Vy
axial component of the water spray velocity, (m s-1)
tangential component of the water spray velocity, (m s-1)
Voy
initial axial component of the water spray velocity at time tf =0, (m s-1)
Vr'
radial velocity of the water spray at impingement point, (m s-1)
Vy'
axial velocity of the water spray at impingement point, (m s-1)
water volume, (m3)
turbine blade width, (m)
WT
square tank width in equation (1.10), (m)
sorption number in ref. (Zlokarnik, 1979), (-)
Ym
maximum droplets height, (m)
axial coordinate, (m)
zcent
axial coordinate at circulation center, (m)
Greek Symbols
water density, (kg m-3)
probe time constant, (s)
water dynamic viscosity, (kg m-1 s-1)
water kinematic viscosity, (m2 s-1)
- 244 -
List of Symbols
Impellers types abbreviations
APV-B2 D, U hydrofoil impeller (developed particularly for gas dispersion), down

and up pumping
A310-D
Lightnin A310 impeller, down pumping
EE D, U
elephant ear impeller, down and up pumping
HF-D
hydrofoil impeller, down pumping
HF4-D
hydrofoil impeller (has three blades with a blade angle of 35 and a

twist of 15 in the tip region), down pumping
MTT D, U
Mixel TT impeller, down and up pumping
PBT D, U
pitch blade turbine, down and up pumping
RTP D, U
reversal twisted pitched blade propeller, down and up pumping
- 245 -
- 246 -
List of Figures
List of Figures
Chapter One
Figure (1.1): Plate diffuser aerator,(Permox H ceramic plate diffuser, (Supratec Co.
Ltd.)
10
Figure (1.2): (a)
Tube diffuser aerator,
(b)
Disc diffuser aerator,
(Gemgate GmbH)
11
Figure (1.3): Static tube diffusers (Process Engineering s.r.l)
12
Figure (1.4): Submerged turbine aerator, (ARS-ARS/S Radial submersible aerator,

14
(Caprari S. p. A.)
Figure (1.5): Hydro Jet Aerator, Plaquette Aerodyn, (Biotrade Co.)
14
Figure (1.6): Flow diagram for uncovered pure oxygen aeration (Mueller et al.,
15
2002)
Figure (1.7): Horizontal flow aspirating aerator, Aspirator, (AIRE-O2 Aeration
Industries International)
16
Figure (1.8): Low speed vertical flow aerator Up-ward flow,Praxair Technology.17
Figure (1.9): Low speed vertical flow aerator (Down-ward flow), (a) Turboxal (Aire
Liquide), (b) Praxair (Praxair Technology)
17
Figure (1.10): Low speed horizontal flow aerator (Twin mini rotor aeration,
(Botjheng Water Ltd.)
18
Figure (1.11): High speed surface aerator, Aqua turbo (AER-AS), (AQUATURBO
SYTEMS inc.)
19
Figure (1.12): The surface aeration regimes applying air entrainment from free liquid
surface (A) Direct entraining of the atmospheric air, (B) Spray
formation and entraining the air with droplets impingement at the
surface,(Patwardhan and Joshi, 1998)
21
Figure (1.13): The histogram distribution of measured aeration efficiency for (111)
low speed surface aerators in field, the average is 1.49 kgO2/kWh,
(Heduit and Racault, 1983b)
27
Figure (1.14): The relation between surface aeration impeller blades submergence
and oxygen transfer rate for different blades number, (where H is the
liquid level in the tank) (Backhurst et al., 1988)
30
Figure (1.15): Relation between sorption number (Y) with the Froude number (Fr) for
the conical shape turbine surface aerator (Zlokarnik, 1979)
36
- 247 -
List of Figures
Figure (1.16): Bubble distribution of surface aerator system of rotation speed N=110
rpm, liquid level h=0.66 m. (Lee et al., 2001)
41
Figure (1.17): The relation between the rotation speed and mixing time for surface
aeration dual impeller system, (Kang et al., 2001)
42
Figure (1.18): The relation between the Reynolds number and power number for
three different liquid condition ns, clean water and two types of
44
activated sludge mixed liquid (Takase et al., 1982)
Chapter Two
Figure (2.1): The schematic diagram of the experimental apparatus
52
Figure (2.2): The turbine, propeller and draft tube
52
Figure (2.3): Schematic diagram of experimental apparatus; (a) Up pumping flow

(Aeration mode), (b) Down pumping flow (Mixing mode)
53
Figure (2.4): The Laser Doppler Velocimetry (LDV) testing apparatus, (1) Laser
source. (2) Traverse system, (3) Tested tank, (4) Flow velocity
analyser
55
Figure (2.5): The LDV measuring volume fringe planes
56
Figure (2.6): The measuring volume
57
Figure (2.7): The LDV laser beams positions
58
Figure (2.8): Schematic diagram of the PIV
59
Figure (2.9): The Particle Image Velocimetry (PIV) testing apparatus, (1) Laser
source, (2) Recording camera, (3) Tested tank
60
Figure (2.10): Light scattering by the (10 m) glass particles in the water (Raffel et
al., 2007)
60
Figure (2.11): The pumping number and pumping flowrate measuring volume
65
Figure (2.12): The 3-D liquid volume cell and the vessel volume grids, (GarciaCortes et al., 2006)
67
Figure (2.13): The schematic diagram of the polarographic probe
68
Figure (2.14): The schematic diagram of the optical probe
68
Figure (2.15): The schematic diagram of oxygen mass transfer zones
69
- 248 -
List of Figures
Figure (2.16): (a) Experimental oxygen probes positions in the liquid volume,
(b) Experimental oxygen probes positions in the liquid volume, N=2.5
rps, (h/Tv) = 0.35
72
Figure (2.17): The repeatability of the mass transfer (kla) experimental results for
both oxygen probes, (N= 2.08 rps)
73
Figure (2.18): Optical probe response time verification
74
Figure (2.19): Polarographic probe response time verification
74
and experimental DO values for the optical probe
76
and experimental DO values for polarographic probe
76
Figure (2.22): The droplets zone mass transfer coefficient (klad) measurement
83
Figure (2.23): Schematic diagram illustrates the surface aeration water droplets spray
from turbine blades till the impingement point with some important
relevant dimensions
84
Figure (2.24): Droplets radial and tangential velocities propelled in horizontal plane
by turbine blades
85
Chapter Three
Figure (3.1): Different tested geometric configurations, (a) Whole system, (b)
Turbine alone, (c) Turbine + Propeller, (d) Turbine + Draft tube 92
Figure (3.2): Impellers speed effect on the oxygen transfer coefficient with four
different geometrical configurations; (Whole System, without draft
tube, turbine alone and without propeller), D/Tv=0.24, C/Tv=0.31,
h/D=1.47, Cpr/Tv=0.2, S/W=1
93
Figure (3.3): The DO profile for different rotation speeds, h/Tv= 0.35, C/Tv=0.313,
Cpr/Tv=0.2, temperature =15 oC
94
Figure (3.4): (a) The limits of the tested turbine blades submergence, (b) Mass
transfer Coefficient kla relation with turbine blades submergence and
water height in the liquid bulk for three levels of rotational speed, (b)
Mass transfer Coefficient kla relation with the rotation speed for three
levels of turbine submergence. D/Tv= 0.24, C/Tv = 0.31,
Cpr/Tv=0.2.
96
- 249 -
List of Figures
Figure (3.5): Spacing effect on the oxygen transfer coefficient, D/Tv= 0.24,
C/Tv = 0.31, Sp (Reference) =7.6 mm Figure (3.6): radial flow rate by
propeller, (m3/s)
98
Figure (3.6): (a) The relationship between power number and Reynolds No., for
different turbine blades submergence levels C/Tv = 0.31, (b) The
relation between power consumption and rotation speed,
C/Tv = 0.31
99
Figure (3.7): The spacing
consumption
between
two
impellers
effect
on
the
power
100
Figure (3.8): The effect of various impellers configurations on the power

consumption
101
Figure (3.9): The relationship between the standard aeration efficiency SAEb for
three aeration mode configurations (Whole System, turbine + propeller
and turbine alone) with the impeller rotation speed, D/Tv=0.24,
102
C/Tv=0.31, h/D=1.47, S/W=1
Figure (3.10): The relation between the standard oxygen transfer rate SOTRb for three
aeration mode configurations (Whole System, turbine + propeller and
turbine alone) with the impeller rotation speed, D/Tv=0.24, C/Tv=0.31,
h/D=1.47, S/W=1
105
Figure (3.11): (a) The comparison between the kla/N, predicted by the correlation
model (Eq. 3.4) with the experimentally resulted kla/N, (b) The
Comparison between the (Np) values predicted by the Eq. 3.7 with the
experimentally resulted (Np) values
110
water bulk zone (Cdt), and the water spray zone (CLt) at N = 1.67 rps
for two experimental runs
112
112
water bulk zone (Cdt), and the water spray zone (CLt) at N = 2.5 rps for
two experimental runs
113
113
efficiency by plotting Cdt versus CLt for various rotation speeds
114
- 250 -
List of Figures
Figure (3.17): The relation between the oxygen mass transfer coefficient k lad and
impellers rotation speed and in the spray zone, S/W = 1
116
Figure (3.18): The relation between the oxygen transfer rate OTRsp and impellers
rotation speed and, in the spray zone, S/W = 1
119
Figure (3.19): The influence of the impellers rotational speed on the spray zone
oxygen transfer rate OTRsp and in the bulk zone oxygen transfer rate
OTRsp contributions the overall transfer rate OTR, S/W = 1
120
Figure (3.20): Effect of the turbine blade submergence on spray mass transfer zone
dissolved oxygen concentration at rotation speed (N=2.5 rps) and
123
turbine clearance (C/Tv = 0.35)
efficiency by plotting (Cdt) versus (CLt) for various turbine blades
submergence, N = 2.5 rps
124
Figure (3.22): The relation between the spray zone aeration efficiency Esp and turbine
blade submergence ratio, N = 2.5 rps
125
Figure (3.23): The relation between the spray zone oxygen mass transfer coefficient,
klad with the turbine submergence and liquid level impellers rotation
speed, N = 2.5 rps
127
Figure (3.24): The relation between spray zone oxygen transfer rate OTRsp with the
turbine blades submergence and liquid level, N = 2.5 rps
128
Figure (3.25): The influence of the turbine submergence on the spray zone oxygen
transfer rate, OTRsp and in the bulk zone oxygen transfer rate OTRsp
contributions the overall transfer rate OTR, N = 2.5 rps
129
Figure (3.26): The effect of the turbine rotation speed on the spray mass transfer zone
dissolved oxygen concentration at h/D=1.47and C/T=0.35, (Turbine
alone configuration)
131
Figure (3.27): The linear regression of the spray zone aeration efficiency of the plot
(Cdt) versus (CLt) for various rotation speeds, (Turbine alone
configuration)
132
Figure (3.28): The relation between the impellers speed N and the spray zone aeration
efficiency Esp for both turbine alone and whole system configurations,
S/W = 1.0
133
Figure (3.29): The relation between N and the oxygen mass transfer coefficient klad
for both turbine alone and whole system configurations,
S/W = 1.0
133
- 251 -
List of Figures
Figure (3.30): The relation between spray zone oxygen transfer rate with the
rotational speed for the whole system and turbine alone configurations,
S/W = 1.0
135
Figure (3.31): The influence of the tank internals (Propeller and draft tube) on the
spray zone oxygen transfer rate, OTRsp contribution the overall transfer
rate OTR, S/W = 1.0
136
model (Eq. 3.10) with the experimentally resulted (Esp)20
139
model (Eq. 3.11) with the experimentally resulted (Esp)20
140
model (Eq.3.16) with the experimentally resulted (Esp)20
142
Chapter Four
Figure (4.1): LDV implementation for measuring velocity in a (r-z) plane,
(Plane B)
148
Figure (4.2): PIV implementation for measuring velocity in a (r-z) plane,
(Plane A)
149
Figure (4.3): The axial 4-bladed, reversible 45o pitched and twisted blade propeller
RTP propeller (HPM204D, Milton Roy Mixing)
150
Figure (4.4): The (r-z) flow field map for the RTP-D propeller with draft tube
configuration in the entire vessel
153
Figure (4.5): The (r-z) velocities contour map for the RTP-D propeller with draft
tube configuration in the entire vessel
155
Figure (4.6): Dimensionless axial velocity profile at different liquid levels (below
impeller z/h=0.55, above impeller z/h=0.62) for RTP-D with draft
tube, Cpr /Tv =0.2, N= 5 rps in two different measuring planes; plane A
at 00 and plane B at 900 from the posterior baffle respectively
157
Figure (4.7): (a) The (r - z) flow field map for the RTP-D propeller without draft
tube configuration in the vessel core, (b) The (r - z) velocities contour
map for the RTP-D propeller with draft tube configuration in the vessel
core
159
- 252 -
List of Figures
RTP-D propeller without draft tube (below impeller z/h=0.55, above
impeller z/h=0.62), Cpr /Tv =0.2, N= 5 rps in two different measuring
planes; plane A at 00 and plane B at 900 from the posterior baffle
respectively
161
Figure (4.9): The creation of vortex in the down-pumping mode with draft tube 162
Figure (4.10): (a) The (r - z) flow field map for the RTP-U propeller with draft tube
configuration in the vessel core, (b) The (r - z) velocities contour map
for the RTP-U propeller with draft tube configuration in the vessel
core
163
RTP-U propeller with draft tube (below impeller z/h=0.55, above
respectively
165
Figure (4.12): (a) The (r - z) flow field map for the RTP-U propeller without draft
tube configuration in the vessel core, (b) The (r - z) velocities contour
map for the RTP-U propeller with draft tube configuration in the vessel
core
167
RTP-U propeller without draft tube (below impeller z/h=0.55, above
respectively
168
Figure (4.14): The distribution of the liquid volume in the vessel related with mean
local composite velocity ranges for RTP-D propeller with draft tube
configuration
175
Figure (4.15): The effect of RTP propeller rotational speed on the mixing time for
different flow and geometry conditions, h/dpr=1.47, Cpr/Tv = 0.2
177
Figure (4.16): The relation of Reynolds number with the dimensionless mixing time
for different pumping modes and geometry configurations , h/dpr =
1.47, Cpr /Tv = 0.2
178
Chapter Five
Figure (5.1): The (r-z) flow field map for the turbine and the RTP-U propeller with
draft tube (Whole system) configuration in the entire vessel
187
- 253 -
List of Figures
Figure (5.2): The (r-z) velocities contour map for the turbine and the RTP-U
propeller with draft tube (Whole system) configuration in the entire
vessel
189
Figure (5.3): The (r-z) flow field map for the turbine alone configuration in the
entire vessel
191
Figure (5.4): The (r-z) velocities contour map for the turbine alone configuration in
the entire vessel
193
Figure (5.5): The (r-z) flow field map for the turbine and the RTP-U propeller
configuration in the entire vessel
195
Figure (5.6): The (r-z) velocities contour map for the turbine and the RTP-U
propeller configuration in the entire vessel
197
Figure (5.7): Distribution of the liquid volume in the vessel related with mean local
composite velocity ranges for whole system configuration
201
Figure (5.8): The relation between impellers rotation speed and the dimensionless
mixing number Ntm for whole system configuration in the aerated
vessel
203
Figure (5.9): The effect of impellers rotation speed on the mixing time for two
different geometrical configurations, h/D=1.47, Cpr/T= 0.2, C/T =
0.313 (Aerated mode)
204
mixing time of the aerated vessel for whole system and turbine alone
configurations, h/D= 1.47, Cpr/Tv = 0.2, C/Tv = 0.313
205
Figure (5.11): The relation between impellers rotation speed and the spacing between
impellers; h/D= 1.47, C/Tv = 0.313, (Aerated mode)
206
mixing time of the aerated vessel for three turbine blade submergence
levels, Cpr /T= 0.2, C/T = 0.313
207
Figure (5.13): The comparison between experimental and predicted values of
dimensionless mixing time number (tmN) by equation (5.4)
209
Figure (5.14): The effect of Reynolds number on power number for the two different
configurations (Whole system and turbine alone); h/D= 1.47, Cpr/Tv =
0.2, C/T = 0.313, (Aerated mode)
210
Figure (5.15): The relation between impellers rotation speed and the consumed power
per water volume for the aerated mode, h/D= 1.47, Cpr/Tv = 0.2, C/Tv =
0.313
210
- 254 -
List of Figures
Appendix I
Figure (I.1):
Process selection diagram for industrial wastewater treatment

224
(Eckenfelder et al., 2009)
Figure (I.2):
Domestic wastewater treatments steps (Try and Price, 1995)
Figure (I.3):
Development and control of biomass in the activated sludge process

(Gary, 2004)
227
Figure (I.4):
The schematic layout of two-stage nitrification and de-nitrification

(Gary, 2004)
229
Figure (I.5):
The Two stage biological phosphorus treatment (Gary, 2004)
Figure (I.6):
Chemical-biological phosphorus treatment (Phostrip process), (Barth

and Stensel, 1981)
231
Figure (I.7):
Covered aerated tank using pure oxygen (Mueller et al., 2002)
225
230
232
Appendix II
Figure (II.1): The energy balance points for the spray discharge flowrate
- 255 -
238
- 256 -
List of Tables
List of Tables
Chapter One
Table (1.1):
The standard aeration efficiency (SAE) for various aerators types 22
Table (1.2):
The standard aeration efficiency of the fine bubble diffused aeration in

different water treatment basins (Duchene and Cotteux, 2002)
23
Chapter Two
Table (2.1):
Geometrical configuration details
53
Chapter Three
Table (3.1):
The SAEb and SOTRb for different water level ratios (h/D) and turbine
blade width to blades submergence ratio (S/W), D/T=0.24, C/T=0.31,
Cpr/T=0.2, N=2.803 rps, (Whole system Configuration)
103
Table (3.2):
The SAEb and SOTRb for different impellers speed, D/T= 0.24, C/T=
0.31, Cpr/T= 0.2, h/D = 1.47 (Whole system Configuration, Aeration
mode)
104
Table (3.3):
The SAEb and SOTRb for different impellers speed, D/Tv= 0.24, C/Tv=
0.31, Cpr/Tv= 0.2, h/D = 1.47 (Turbine and propeller
configuration)
104
Table (3.4):
The SAEb and SOTRb for different impellers speed, D/Tv= 0.24, C/Tv=
0.31, h/D = 1.47, (Turbine alone configuration)
105
Table (3.5):
The effect the spacing between impellers on SOTRb and SAE b, D/Tv=
0.24, C/Tv= 0.31, h/D = 1.47, Sp=8.3 mm
105
Table (3.6):
The spray droplets mass transfer zone of klad, flight time t f and the
related measured parameters, (of 100% submerged turbine blades, H =
zero), h/D = 1.47, S/W = 1
116
Table (3.7):
Water spray droplets velocity and volumetric flow rate of given

operation and geometric configuration (whole system) for various
rotation speed levels
117
Table (3.8):
Spray zone oxygen transfer rate OTRsp variation with impellers

rotation speed
118
- 257 -
List of Tables
Table (3.9):
The oxygen mass transfer rates for spray and bulk zones OTR b and
OTRsp with the percentage of contribution
120
Table (3.10): The water spray velocity and volumetric flow rate for the whole
system geometrical configuration with various turbine blades
submergence and water levels, N= 2.5 rps
122
Table (3.11): The calculated water spray mass transfer zone (klad), droplets flight
time(tf) and some related measured parameters and turbine blades
submergence and water level, at constant rotation speed (N =
2.5 rps)
126
Table (3.12): Water spray zone oxygen transfer rate (OTRsp) variation with different
turbine blades submergence and water levels, (N = 2.5 rps)
128
Table (3.13): Oxygen mass transfer rate and percentage of contribution for both
water spray and bulk (re-aeration) zones, (N= 2.5 rps)
130
Table (3.14): The spray mass transfer zone (klad), flight time (t f) and some related
measured parameters, S/W=1.00, h/D =1.47, (Turbine alone
configuration)
134
Table (3.15): The spray velocity and volumetric flow rate for various turbine rotation
speeds (Turbine alone configuration)
134
Table (3.16): The Spray zone oxygen transfer rate OTRsp for different rotation speed
levels, (Turbine alone configuration)
136
Table (3.17): The Comparison for the spray zone oxygen transfer rates OTR sp
between the whole system and turbine alone configurations, S/W
=1.00, h/D=1.47
137
Chapter Four
Table (4.1):
The RTP propeller and several pithed blade impellers power numbers
with different configurations and pumping modes
169
Table (4.2):
The RTP propeller and several pithed blade impellers pumping and
circulation numbers with different configurations and pumping
modes
170
Table (4.3):
The agitation Index for RTP-D and various pithed blade impellers,
(this work and literature)
172
Chapter Five
Table (5.1):
Turbine pumping number and system circulation number

- 258 -
199
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5)4&

CHARACTERIZATION AND IMPROVEMENT OF A SURFACE AERATOR

DPMF EPDUPSBMF et discipline ou spcialit 

ED MEGEP : Gnie des procds et de l'Environnement

Mots-cls: Aration de surface, Cuve agite, Transfert de matire, Transfert

d'oxygne, Hydrodynamique, Ecoulement multiphasique, LDV, PIV, Analyse

Introduction and Outlines

Chapter 1: Surface Aeration Process for Water Treatment

1.1. Presentation of Different Aeration Technologies in Water Treatment

1.1.1. Diffused Aeration

I.1. Plate Diffusers

I.2. Panel Diffuser

I.3. Tube Diffuser

I.4. Dome Diffusers

I.5. Disc Diffuser

II. Non Porous Diffusers

II.1. Fixed Orifice Diffusers

II.2. Valved Orifice Diffusers

II.3. Static Tube Diffusers

1.1.2. Submerged Aeration

I. Submerged Turbine Aerators

II. Jets Aerators

1.1.3. Aeration with High-Purity Oxygen

1.1.4. Aspirating Aeration

1.1.5. Surface Aeration

1.2. Surface Aeration for Water Treatment Processes

1.2.1. Types of Surface Aerators

I. Low Speed Surface Aerators

I.1. Low Speed Vertical Flow Aerators

I.2. Low Speed Horizontal Flow Aerators

II. High Speed Surface Aerators

1.2.2. Principals and Characterization

II. Surface Aerator Characterizations

II. Dissolved Oxygen Concentration Gradient Calculation Methodology

IV. Surface Aeration Oxygen Mass Transfer

IV.1. Operational Condition Effects

A. Rotational Speed Effect

B. Number of Impellers Effect

C. Liquid Level Effect

D. Clearance and Submergence Effect

IV.2. Geometry Effect

C. Draft Tube Effect

D. Surface Aerators Geometry

D.1. Surface Aerator Diameter

V. Impeller Position in the Treatment Tank

IV. Temperature Effect

V.1. Flow Patterns

A. Flow Patterns Characterization

V.2. Air Bubble Size Distribution and Hold-up

V.3. Mixing Time

A. Mixing Time Characterization

B. Mixing Time Modeling

V.4. Air Bubbles Entrainment

V.5. Surface Aeration Power Consumption

A. Operational Condition Effect

B. Power Consumption Relation with Oxygen Mass Transfer

VI. Contact Time between Water Droplets and Atmospheric Air

VII. Environmental Effects

Chapter 2: Experimental Setup and Calculation Methods

DPMF EPDUPSBMF et discipline ou spcialit