master’s dissertation submitted in partial fulfilment of the … · 2017-11-30 · in the...

Master’s dissertation submitted in partial fulfilment of the requirements for the joint degree of

International Master of Science

in Environmental Technology and Engineering

an Erasmus+: Erasmus Mundus Master Course jointly organized by

Ghent University, Belgium

University of Chemistry and Technology, Prague, Czech Republic

IHE-Delft Institute for Water Education, Delft, the Netherlands

Academic year 2015 – 2017

Evaluation of the Speece cone oxygen transfer capacity

in clean water and activated sludge in a pilot-scale

MBR

Host University:


Ivania Margarita Ochoa Barahona

Promotor: Prof. Damir Brdanovic, PhD (IHE-Delft)

Co-promoter: Héctor García Hernández, PhD (IHE-Delft); Tineke Hooijmans, PhD, MSc (IHE-

Delft)

ES.17.24

This thesis was elaborated at IHE-Delft and defended at IHE-Delft within the framework of the European

Erasmus Mundus Programme “Erasmus Mundus International Master of Science in Environmental Technology

and Engineering " (Course N° 2011-0172)

© [2017] [Delft], [Ivania Margarita Ochoa Barahona], Ghent

University, all rights reserved.

Evaluation of the Speece cone oxygen transfer capacity in clean water and

activated sludge in a pilot-scale MBR

Master of Science Thesis

by

Ivania Margarita Ochoa Barahona

Supervisors Prof. Damir Brdanovic, PhD, MSc (IHE-Delf)

Mentors Héctor García Hernández, PhD, MSc (IHE-Delf)

Tineke Hooijmans, PhD, MSc (IHE-Delft)

Examination committee

Prof. Damir Brdanovic, PhD, MSc (IHE-Delft)

Héctor García Hernández, PhD, MSc (IHE-Delft)

Tineke Hoiijmans, PhD, MSc (IHE-Delft)

Eldon Rene Raj, PhD, MSc (IHE-Delft)

Aradaí Herrera, MSc (JCI Industries, Inc. – Ext)

This research is done for the partial fulfilment of requirements for the Master of Science degree at the


Delft

August 2017

Although the author and IHE-Delft Institute for Water Education have made every effort to

ensure that the information in this thesis was correct at press time, the author and IHE-Delft do

not assume and hereby disclaim any liability to any party for any loss, damage, or disruption

caused by errors or omissions, whether such errors or omissions result from negligence,

accident, or any other cause.

© Ivania Margarita Ochoa Barahona 2017.

This work is licensed under a Creative Commons Attribution-NonCommercial 4.0 International License.

http://creativecommons.org/licenses/by-nc/4.0/

http://creativecommons.org/licenses/by-nc/4.0/

i

Abstract

Aeration is a key operational process in the biological wastewater treatment. Enough oxygen

has to be provided for the respiration process of the activated sludge where the organic matter

in the in the wastewater is completely oxidized. The treatment capacity of the biological process

in the wastewater treatment plant, is mainly influenced by the amount of mixed liquor in the

aeration basin, that is, the mixture of activated sludge and wastewater stream. The higher the

mixed liquor the higher the treatment capacity.

Membrane bioreactors (MBR) can handle mixed liquor suspended solids (MLSS) concentration

of approximately 10-15 g/L, three times as high as in conventional activated sludge systems.

However, as the MLSS concentration increases, there is a higher need of oxygenation in order

to supply enough oxygen for the biological demand of the activated sludge. The provision of

enough oxygen represents one of the main process limitations in the wastewater treatment,

especially when using high MLSS concentrations. The most commonly used methods for

aeration, mainly diffused aeration systems, provide oxygen at low efficiencies ranging in 15-

38%. However, these efficiencies can be improved up to five-fold if aeration technologies using

high purity oxygen are used.

Using high purity oxygen in the wastewater treatment can cope with the aeration process

limitations and supply the required oxygen even at high MLSS concentrations. The Speece cone

is an emerging technology for the provision of high purity oxygen for the biological processes.

However, this technology has not been studied regarding its oxygen transfer performance in the

wastewater sector.

A pilot-scale MBR with a Speece cone system as the oxygen provision technology was studied

regarding its oxygen transfer performance in both clean water and process activated sludge

water at low MLSS concentration. The Speece cone system was able to reach a maximum

oxygen transfer rate 3.8±0.2 kgO2/d and 2.2±0.3 kgO2/d, a maximum oxygen transfer efficiency

of 96% and 78%, and a maximum overall oxygen transfer coefficient of 2.1 and 1.5 h-1 in clean

water and activated sludge respectively. To be able to find these maximum values, four

operational parameters were varied, namely pressure, inlet velocity, retention time, and oxygen

flow. Additionally, the maximum alfa factor, that is, the coefficient between the oxygen transfer

efficiency in process water to clean water, was 0.96.

According to these results, it seems that the Speece cone is highly efficient regarding the oxygen

transfer performance, and it can be seen as a potentially vialbe option for the provision of

oxygen in highly concentrated systems, such as MBR.

iii

Table of Contents

Abstract i

List of Figures v

List of Tables ix

Abbreviations x

Introduction 2 1.1. Research objectives 3

1.1.1. General objective 3

1.1.2. Specific objectives 3

Literature Review 4 2.1. Conventional activated sludge process in wastewater treatment 4

2.1.1. CAS key process parameters 5

2.2. Aeration for wastewater treatment 8 2.2.1. Aeration systems 13

2.3. Membrane bioreactors (MBR) in wastewater treatment 20 2.3.1. Membrane separation process 20 2.3.2. Membrane bioreactor (MBR) 20

2.4. Turbo Membrane Bioreactor (tMBR) 22

Methodology 24 3.1. Experimental setup 24

3.1.1. Speece Cone 24 3.1.2. Membrane Bioreactor aerobic tank 25

3.1.3. Other equipment 27 3.1.4. Clean water and activated sludge 27

3.2. Experimental procedure 28 3.2.1. Evaluation of the Speece cone performance in clean water 28 3.2.2. Evaluation of the Speece cone performance in process activated sludge

water 41 3.3. Data analysis 44

3.3.1. Oxygenation capacity determination by numerical integration 44 3.3.2. Oxygenation capacity determination by AQUASIM model 47

3.3.3. Alfa factor (α) 48 3.4. Analytical procedure 48

Results and Discussion 50 4.1. Evaluation of the Speece cone performance in clean water 50

4.1.1. Dissolved oxygen concentration gradient in the MBR aeration tank 50 4.1.2. Intrusion experiments 52

iv

4.1.3. Preliminary evaluation of the operational parameters on the

performance of the Speece cone 53 4.1.4. Evaluation of the operational parameters on the performance of the

Speece cone 61 4.1.5. AQUASIM model for clean water experiments 68

4.2. Evaluation of the Speece cone performance in process activated sludge water 72

4.2.1. Phase 1 72 4.2.2. Phase 3 73 4.2.3. Alfa factor 79

Conclusions 81

References 83

v

List of Figures

Figure 2-1. Conventional Activated Sludge process flow diagram. Adapted from (Roš, 1993) ............................ 4

Figure 2-2. Mixed liquor suspended solids (MLSS) constituents (Henze et al., 2008) .......................................... 5

Figure 2-3. Oxygen requirement of the activated sludge in the aeration basin. * for Nitrogen removal systems.

Adapted from (von Sperling & International Water Association, 2007) ........................................................ 7

Figure 2-4. Influence of mixed liquor suspended solids concentration in KLa20 (Germain et al., 2007) .............. 10

Figure 2-5. Impact on alfa factor according to mixed liquor suspended solids concentration (Judd, 2007) ........ 11

Figure 2-6. Pure oxygen systems for activated sludge aeration and examples. (a) EPA (1973), (b) Metcalf et al.

(2004), (c) Blue in Green , (e) Linde Group , (f) Gray (2004), (g) Mueller et al. (2002) ................................... 15

Figure 2-7. Speece cone original design at bench scale (left) and at pilot scale (right) (Speece et al., 1971) ...... 16

Figure 2-8. Newman Lake’s Speece cone aeration system diagram (Moore et al., 2012) ................................... 16

Figure 2-9. Summary of results of the experimental research on oxygen transfer efficiency parameters of a lab-

scale Speece cone (Ashley et al., 2008) ........................................................................................................ 17

Figure 2-10. Summary of results of the experimental research on oxygen transfer efficiency parameters of a

pilot-scale Speece cone (Ashley et al., 2014) ............................................................................................... 19

Figure 2-11. Classification of membrane separation processes according to their selectivity (ZENA, 2017) ..... 20

Figure 2-12. Membrane cleaning methods (Judd, 2007) ...................................................................................... 22

Figure 3-1. Scheme of complete MBR system and Speece cone ......................................................................... 24

Figure 3-2. Speece cone before (left) and after installation (right) in the pilot membrane bioreactor (C. Barreto,

2015) ............................................................................................................................................................. 25

Figure 3-3. Membrane module before installation in the membrane bioreactor (C. Barreto, 2015) .................... 26

Figure 3-4. Distribution system for air scouring purposes (C. Barreto, 2015) ..................................................... 26

Figure 3-5. Return Activated Sludge collection tank in Harnaschpolder wastewater treatment plant ................. 28

Figure 3-6. Maximum and minimum recirculation flows (m3/h) of sludge pump as a function of Speece cone

pressure at gauge (psi) and drive frequency (%) ........................................................................................... 29

Figure 3-7. Location of the dissolved oxygen concentration probe in the MBR aeration tank for the preliminary

evaluation experiments ................................................................................................................................. 35

Figure 3-8. Locations chosen for Experiment 1 to measure dissolved oxygen gradient concentration in the MBR

aerobic tank................................................................................................................................................... 36

Figure 3-9. Locations chosen for Experiment 2 to measure dissolved oxygen gradient concentration in the MBR

aerobic tank................................................................................................................................................... 37

Figure 3-10. Dissolved Oxygen determination points according to depth in the aerobic tank for the optimal

conditions tests ............................................................................................................................................. 38

Figure 3-11. Phases of experiments in process activated sludge water ................................................................ 41

Figure 3-12. Sub-system selected for mass balance analysis ............................................................................... 44

vi

Figure 3-13. Simplified schematic representation of the MBR and Speece cone system for AQUASIM

modelling ...................................................................................................................................................... 47

Figure 4-1. Mean absolute differences in dissolved oxygen concentrations between a mobile probe with different

depths and locations vs fixed probe at the middle of the tank at 30 cm depth .............................................. 50

Figure 4-2. Dissolved oxygen concentration measurements for Experiment 2 at different locations, depths and

mixing conditions: (a) Low mixing and (b) High mixing ............................................................................. 51

Figure 4-3. Change in dissolved oxygen concentration level due to intrusion in (a) low mixing conditions

(preliminary evaluation) and (b) high mixing conditions (evaluation of optimal parameters) ..................... 52

Figure 4-4. Influence of low, medium, and high pressure levels on standard oxygen transfer rate (SOTR) at two

recirculation flows: (a) 6 m3/h, and (b) 3 m3/h (mean±SD) found during preliminary evaluation experiments

in clean water ................................................................................................................................................ 54

Figure 4-5. Influence of low, medium, and high pressure levels on standard oxygen transfer efficiency (SOTE)

at two recirculation flows: (a) 6 m3/h, and (b) 3 m3/h (mean±SD) found during preliminary evaluation

experiments in clean water ........................................................................................................................... 55

Figure 4-6. Influence of low, medium, and high pressure levels on overall oxygen transfer efficiency corrected

by temperature (KLa_Transference20) at two recirculation flows: (a) 6 m3/h, and (b) 3 m3/h (mean±SD) found

during preliminary evaluation experiments in clean water ........................................................................... 57

Figure 4-7. Influence of low, medium, and high pressure levels on the theoretical maximum standard oxygen

transfer rate (Max_SOTR) at two recirculation flows: (a) 6 m3/h, and (b) 3 m3/h (mean±SD) found during

preliminary evaluation experiments in clean water ...................................................................................... 58

Figure 4-8. Influence of inlet velocity on standard oxygen transfer rate (SOTR) at constant pressure (mean±SD)

found during preliminary evaluation experiments in clean water ................................................................. 60

Figure 4-9. Influence of inlet velocity on standard oxygen transfer efficiency (SOTE) at constant pressure

(mean±SD) found during preliminary evaluation experiments in clean water ............................................. 60

Figure 4-10. Influence of inlet velocity on overall oxygen transfer efficiency corrected by temperature

(KLa_Transference20) at constant pressure (mean±SD) found during preliminary evaluation experiments in clean

water ............................................................................................................................................................. 60

Figure 4-11. Influence of inlet velocity on the theoretical maximum standard oxygen transfer rate (Max_SOTR)

at constant pressure (mean±SD) found during preliminary evaluation experiments in clean water ............. 61

Figure 4-12. Influence of low and high pressure levels on standard oxygen transfer rate (SOTR) at 3 m3/h and

3.43 m/s (mean±SD) found during evaluation of optimal conditions in clean water .................................... 62

Figure 4-13. Influence of low and high pressure levels on standard oxygen transfer efficiency (SOTE) at 3 m3/h

and 3.43 m/s (mean±SD) found during evaluation of optimal conditions in clean water ............................. 62

Figure 4-14. Influence of low and high pressure levels on overall oxygen transfer efficiency corrected by

temperature (KLa_Transference20) at 3 m3/h and 3.43 m/s (mean±SD) found during evaluation of optimal

conditions in clean water .............................................................................................................................. 63

vii

Figure 4-15. Influence of low and high pressure levels on theoretical maximum standard oxygen transfer rate

(Max_SOTR) at 3 m3/h and 3.43 m/s (mean±SD) found during evaluation of optimal conditions in clean

water ............................................................................................................................................................. 63

Figure 4-16. Influence of low and high inlet velocities on standard oxygen transfer rate (SOTR) at 6 m3/h and

3.4 m/s (mean±SD) found during evaluation of optimal conditions in clean water ...................................... 64

Figure 4-17. Influence of low and high inlet velocities on standard oxygen transfer efficiency (SOTE) at 6 m3/h

and 3.4 m/s (mean±SD) found during evaluation of optimal conditions in clean water ............................... 65

Figure 4-18. Influence of low and high inlet velocities on overall oxygen transfer efficiency corrected by

temperature (KLa_Transference20) at 6 m3/h and 10 psi (mean±SD) found during evaluation of optimal

conditions in clean water .............................................................................................................................. 65

Figure 4-19. Influence of low and high inlet velocities on the theoretical maximum standard oxygen transfer rate

(SOTR) at 6 m3/h and 3.4 m/s (mean±SD) found during evaluation of optimal conditions in clean water .. 66

Figure 4-20. Influence of retention time on standard oxygen transfer rate (SOTR) at 10 psi and 3.4 m/s

(mean±SD) found during evaluation of optimal conditions in clean water .................................................. 67

Figure 4-21. Influence of retention time on standard oxygen transfer efficiency (SOTE) at 10 psi and 3.4 m/s

(mean±SD) found during evaluation of optimal conditions in clean water .................................................. 67

Figure 4-22. Influence of retention time on overall oxygen transfer efficiency corrected by temperature

(KLa_Transference20) at 10 psi and 3.4 m/s (mean±SD) found during evaluation of optimal conditions in clean

water ............................................................................................................................................................. 68

Figure 4-23. Influence of retention time on the theoretical maximum standard oxygen transfer rate (Max_SOTR)

at 10 psi and 3.4 m/s (mean±SD) found during evaluation of optimal conditions in clean water ................ 68

Figure 4-24. Influence of low, medium, and high pressure levels on overall oxygen transfer efficiency corrected

by temperature (KLa20) calculated with AQUASIM program at two recirculation flows: (a) 6 m3/h, and (b) 3

m3/h (mean±SD) found during the preliminary evaluation experiments in clean water ............................... 69

Figure 4-25. Influence of inlet velocity on overall oxygen transfer efficiency corrected by temperature (KLa20)

calculated by AQUASIM program at 10 psi (mean±SD) found during the preliminary evaluation

experiments in clean water ........................................................................................................................... 70

Figure 4-26. Influence of low and high pressure levels on overall oxygen transfer efficiency corrected by

temperature (KLa20) calculated by AQUASIM program at 3 m3/h and 3.43 m/s (mean±SD) found during the

preliminary evaluation experiments in clean water ...................................................................................... 70

Figure 4-27. Influence of low and high inlet velocities on overall oxygen transfer efficiency corrected by

temperature (KLa20) calculated by AQUASIM program at 6 m3/h and 10 psi (mean±SD) found during the

evaluation of optimal conditions experiments .............................................................................................. 71

Figure 4-28. Influence of retention time on overall oxygen transfer efficiency corrected by temperature (KLa20)

calculated by AQUASIM program at 10 psi and 3.4 m/s in clean water (mean±SD) found during the

evaluation of optimal condition experiments in clean water ........................................................................ 71

Figure 4-29. Foam formation in the activated sludge after running re-oxygenation experiment ......................... 73

viii

Figure 4-30. Mixed liquor suspended solids (MLSS) concentration in activated sludge for experiments with

process activated sludge water ...................................................................................................................... 73

Figure 4-31. Influence of pressure on standard oxygen transfer rate (SOTR) at 3 m3/h psi and 3.43 m/s in

activated sludge (mean±SD) ......................................................................................................................... 74

Figure 4-32. Influence of pressure on standard oxygen transfer efficiency (SOTE) at 3 m3/h psi and 3.43 m/s in

activated sludge (mean±SD). ........................................................................................................................ 75

Figure 4-33. Influence of pressure on overall oxygen transfer efficiency corrected by temperature

(KLa_Transference20) at 3 m3/h psi and 3.43 m/s in activated sludge (mean±SD) ............................................... 76

Figure 4-32. Influence of pressure on the theoretical maximum standard oxygen transfer rate (Max_SOTR) at 3

m3/h psi and 3.43 m/s in activated sludge (mean±SD) ................................................................................. 76

Figure 4-35. Influence of retention time on standard oxygen transfer rate (SOTR) at 10 psi and 3.4 m/s in


Figure 4-36. Influence of retention time on standard oxygen transfer efficiency (SOTE) at 10 psi and 3.4 m/s in


Figure 4-37. Influence of retention time on overall oxygen transfer efficiency corrected by temperature

(KLa_Transference20) at 10 psi and 3.4 m/s in activated sludge (mean±SD) ....................................................... 78

Figure 4-38. Influence of retention time on the theoretical maximum standard oxygen transfer rate (Max_SOTR)

at 10 psi and 3.4 m/s in activated sludge (mean±SD) ................................................................................... 79

Figure 4-39. Effect of pressure on alfa factors (α) at 3 m3/h and 3.43 m/s in activated sludge (mean±SD) ........ 79

Figure 4-40. Effect of inlet velocity on alfa factors (α) at 3 m3/h psi and 10 psi in activated sludge (mean±SD) 80

Figure 4-44. Influence of retention time on the alfa factor at 10 psi and 3.4 m/s in activated sludge (mean±SD)80

ix

List of Tables

Table 2-1. Alfa factors found for different aeration methods using clean water with detergent. (a) Eckenfelder

(1989) cited by Roš (1993), (b) Mueller et al. (2002). ................................................................................... 11

Table 2-2. Aeration systems for aeration of activated sludge (Gray, 2004; Metcalf et al., 2004; Winkler, 1981) 13

Table 2-3. Efficiency parameter values in clean water and process water of different types of aerators. (a) Henze

et al. (2008), (b) Gray (2004), (c) Mueller et al. (2002), (d) Metcalf et al. (2004) ............................................ 14

Table 3-1. Membrane module characteristics (C. Barreto, 2015; Membranes Modules Systems, 2005) ............. 26

Table 3-2. Selected pressure levels for the preliminary evaluation experiments in clean water .......................... 29

Table 3-3. Experimental design of preliminary evaluation .................................................................................. 32

Table 3-4. Selected pressure levels for the evaluation of the operational parameters on the performance of the

Speece cone in clean water ........................................................................................................................... 38

Table 3-5. Inlet velocities used in the evaluation of the operational parameters on the performance of the Speece

cone in clean water ....................................................................................................................................... 39

Table 3-6. Experimental design of the evaluation of the operational parameters on the performance of the Speece

cone in clean water ....................................................................................................................................... 40

Table 3-7. Experimental design of the evaluation of the operational parameters on the performance of the Speece

cone in process activated sludge water ......................................................................................................... 43

Table 3-8. Variables used in the AQUASIM model for the MBR and Speece cone system ................................ 49

Table 3-9. Compartments used in the AQUASIM model for the MBR and Speece cone system ........................ 49

Table 3-10. Links used in the AQUASIM model for the MBR and Speece cone system .................................... 49

Table 4-1. Absolute mean differences between dissolved oxygen concentration measurements for Experiment 2

at different locations, depths and mixing conditions .................................................................................... 52

Table 4-2. Comparison between mean standard oxygen transfer rates (SOTR) and their theoretical maximum

found during preliminary evaluation experiments in clean water ................................................................. 59

Table 4-3. Comparison between mean standard oxygen transfer rates (SOTR) and their theoretical maximum at

3 m3/h and 3.43 m/s found during evaluation of optimal conditions in clean water ..................................... 64


6 m3/h and 10 psi found during evaluation of optimal conditions in clean water ......................................... 66

Table 4-5. Endogenous respiration, specific endogenous respiration and MLVSS/MLSS ratio results ............... 74


3 m3/h and 3.4 m/s found in process activated sludge water......................................................................... 77

x

Abbreviations

AE Aeration Efficiency

AS Activated Sludge

ATP Adenosine Triphosphate

BNR Biological Nutrient Removal

BOD Biological Oxygen Demand

CAS Conventional Activated Sludge

COD Chemical Oxygen Demand

DO Dissolved Oxygen

F/M Food to Microorganism ratio

FIT Flow Indicator Transmitter

HPO High Purity Oxygen

HRTn Nominal Hydraulic Retention Time

HWWTP Harnaschpolder Wastewater Treatment Plant

iMBR Immersed Membrane Bioreactor

ISS Inorganic Settable Solids

MBR Membrane Bioreactor

MF Microfiltration

MLSS Mixed Liquor Suspended Solids

MLVSS Mixed Liquor Volatile Suspended Solids

NF Nanofiltration

OC Oxygenation Capacity

OHO Ordinary Heterotrophic Organisms

OLR Organic Loading Rate

OTE Oxygen Transfer Efficiency

OTR Oxygen Transfer Rate

OUR Oxygen Uptake Rate

PIT Pressure Indicator Transmitter

PLC Programmable Logic Controller

PSD Particle Size Distribution

RAS Returned Activated Sludge

RO Reverse Osmosis

SAE Standard Aeration Efficiency

sMBR Side-stream Membrane Bioreactor

SOTE Standard Oxygen Transfer Efficiency

SOTR Standard Oxygen Transfer Rate

SRT Sludge Retention Time

tMBR Turbo Membrane Bioreactor

TMP Transmembrane Pressure

TSS Total Suspended Solids

UF Ultrafiltration

VSS Volatile Settable Solids

WAS Waste Activated Sludge

Introduction 1

Introduction 2

CHAPTER 1

Introduction

Currently, in the conventional activated sludge (CAS) wastewater treatment it is common to

use traditional aeration systems to fulfill the aeration requirement of the biological processes.

These aeration systems can be classified in two groups as mechanical or diffused aeration

systems. The mechanical systems are usually located at the surface of the activated sludge and

use turbulence to improve the oxygen transfer process from the atmosphere to the activated

sludge. On the other hand, the diffused systems are usually submerged where they release a

pressurized stream of air which creates bubbles in the activated sludge where the oxygen can

be transferred. As air only contains 21% of oxygen, high air flows must be injected into the

system to supply with the required oxygen demand (Metcalf, Eddy, & Tchobanoglous, 2004).

Even though diffused aeration systems usually result in higher oxygen transfer efficiencies

compared to mechanical systems, their efficiencies in clean water are still quite low ranging

between 15-38% (Metcalf et al., 2004; Mueller, Boyle, & Popel, 2002). Additionally, the

oxygen transfer efficiency in process water relative to clean water of these systems, as known

as alfa factor (α), range between 0.2 and 0.75 (Mueller et al., 2002; Roš, 1993). The alfa factor

of these systems is highly influenced by the suspended solids concentration. As shown in

several studies, there is an inverse relation between the alfa factor and the suspended solids

concentration, at higher suspended solids concentration, the lower the alfa factor (Germain et

al., 2007; Gunder, 2001; Krampe & Krauth, 2003).

These traditional aeration systems have high energy demand, and their operational cost usually

represents 45-75% of the total costs of a wastewater treatment plant (Henze, van Loosdrecht,

Ekama, & Brdjanovic, 2008). Additionally, the CAS systems require large areas to fit the

necessary basins for the wastewater treatment, such as aeration basins where the biological

process occur, and secondary clarifiers where the liquid-solid separation process occur.

In order to reduce the footprint of the CAS systems in wastewater treatment, the application of

membrane bioreactors (MBR) has become more common. The MBR does not require

secondary clarifiers, as the membrane module separates the solid phase from the liquid phase

and, due to its selectivity characteristics, it usually results in a high quality effluent (Water

Environment Federation, 2012). Additionally, the MBR can usually handle higher

concentration of suspended solids compared to the CAS systems, which reduces the area needed

for the aeration basin (Brepols, 2011).

However, the MBR can have higher aeration needs compared to CAS systems, mainly due to:

(a) higher oxygen demand for biological process as it can handle higher suspended solids

concentration, and (b) aeration need for membrane fouling reduction (Judd, 2007). Moreover,

due to higher aeration demand, the demand for energy can be double as in CAS systems,

influencing also in the related costs (Wagner, Cornel, & Krause, 2002).

Introduction 3

Alternative aeration systems with higher oxygen transfer efficiencies are clearly needed to

overcome the process limitations regarding the high oxygenation demand of wastewater

treatment systems when working at high MLSS concentration, such as MBR. Due to this reason,

the application of high purity oxygen (HPO) systems has increased in the last decade. As these

system use HPO as oxygen source, the oxygen transfer rate immediately increases by five-fold,

resulting in a higher efficiency compared to traditional aeration systems. Even though these

systems can have higher efficiencies, they have not been widely applied mainly due to high

costs for HPO generation (EPA, 1973).

There are several HPO systems used within the wastewater treatment sector, such as UNOX, I-

SO, and SDOX, which have proven to reach efficiencies as high as 90% (Blue in Green; Gray,

2004; Mueller et al., 2002). An emerging technology within the HPO systems is the Speece

cone, which is a pressurized vessel where the activated sludge is super-saturated with HPO and

sent back to the aeration basin. Up to date, there are several full-scale applications of the Speece

cone, but mainly in lakes and reservoirs to avoid the anoxia of hypolimnion (Moore et al., 2012).

However, the Speece cone has high application potential for the wastewater treatment sector,

especially in the systems with high concentration of suspended solids as it could cope with their

high oxygenation needs. This technology has few full-scale applications with this purpose

mainly in the USA by ECO Oxygen Technologies, LLC.

This research aims to contribute to the current knowledge of the Speece cone as a HPO system

for its application in the wastewater treatment sector. It aims to evaluate the Speece cone oxygen

transfer performance in clean water and at low suspended solids concentrations, i.e. 5 g/L.

1.1. Research objectives

1.1.1. General objective To evaluate the performance of a concentrated oxygen delivery system (Speece cone) at

different operational conditions both in clean water and activated sludge in a pilot-scale

membrane bioreactor (MBR).

1.1.2. Specific objectives i. To evaluate the initial Speece cone performance in clean water.

ii. To determine the effect of pressure, inlet velocity, retention time, and oxygen flow on

the oxygen transfer efficiency and aeration coefficients of the Speece cone in clean

water.

iii. To determine the effect of pressure, inlet velocity, retention time, and oxygen flow on

the oxygen transfer efficiency and aeration coefficients of the Speece cone in process

activated sludge with low MLSS.

iv. To find a computational model using AQUASIM that allows the prediction of the

Speece cone behaviour.

Literature Review 4

CHAPTER 2

Literature Review

2.1. Conventional activated sludge process in wastewater treatment

Activated sludge (AS) is a widely applied biotechnology for biological wastewater treatment.

It is called “activated” because the existing bacteria are actively breaking down organic

pollutants into their most oxidized forms, i.e. carbon dioxide (CO2) and water. This is the most

relevant function of the AS process (Roš, 1993).

In the Conventional Activated Sludge systems (CAS, Figure 2-1), the complete degradation of

organic pollutants, or mineralization, occurs in an aeration basin where the wastewater and AS

are in contact with each other; the mixture of these two streams is known as mixed liquor. The

mixed liquor is a suspension of bacteria and other microflora and fauna, which are in charge of

the organic pollutants degradation. After an appropriate retention time in the aeration basin, the

mixed liquor is sent to a secondary clarifier or settling tank, where the AS is separated by gravity

the resulting supernatant which contains a very low concentration of organic pollutants is

commonly known as treated effluent (Gray, 2004; von Sperling & International Water

Association, 2007).

The AS separated in the secondary clarifier is divided into two streams, namely Return

Activated Sludge (RAS) and Waste Activated Sludge (WAS). The RAS is sent back to the

aeration basin as an inoculum in order to keep a highly diverse microbial population as well as

to control the Solids Retention Time (SRT), and the WAS is discarded as excess biomass for

further sludge treatment (Gray, 2004; Roš, 1993).

Figure 2-1. Conventional Activated Sludge process flow diagram. Adapted from (Roš, 1993)

Literature Review 5

CAS is efficient in the removal of organic pollutants from wastewater, in combination with

other treatment processes, such as Biological Nutrient Removal (BNR) and effluent filtration

or tertiary treatment, a high quality effluent can be achieved depending on the intended use for

the treated water.

2.1.1. CAS key process parameters The main process parameters influencing the performance of CAS systems are: amount of

solids in the aeration basin (biomass concentration), Hydraulic Retention Time (HRT), SRT or

sludge age, plant loading or organic loading rate (OLR), sludge settling characteristics, sludge

activity, sludge production, sludge recirculation rate, oxygen requirements, nutrient

requirements, mixing regime, among others. Due to the scope of this research, only some of

these parameters will be further explained in the sections below.

2.1.1.1 Mixed Liquor Suspended Solids (MLSS) and Mixed Liquor Volatile Suspended

Solids (MLVSS)

The mixed liquor suspended solids (MLSS) are comprised of both microbial biomass or active

fraction, as well as inorganic substances or inert fraction (Figure 2-2). The MLSS concentration

in CAS is usually between 1.5-5 g/L (Casey, 1997). As the MLSS concentration increases the

AS process becomes more efficient in assimilating the influent’s substrate which can lead to

smaller aeration tanks. However, von Sperling and International Water Association (2007)

identified two main practical aspects to be taken into account when the MLSS concentration is

increased:

- Higher MLSS concentrations in the aeration basin will require larger secondary

clarifiers, and

- The oxygen transfer process can be particularly affected at higher MLSS concentrations.

Figure 2-2. Mixed liquor suspended solids (MLSS) constituents (Henze et al., 2008)

Mixed Liquor Suspended Solids (MLSS)

Organic Volatile Setteable Solids (VSS)

Ordinary Heterotrophic Organisms (OHOs)

Endogenous residue

Un-biodegradable suspended and setteable

organics from the influent

Inorganic Setteable Solids (ISS)

Inorganic setteable and suspended constituents

Precitable soluble inorganics

Literature Review 6

Usually, the mixed liquor volatile suspended solids (MLVSS) is used to describe the mixed

liquor’s active fraction however this is only an indication of the viable biomass concentration

since MLVSS has three constituents (Figure 2-2):

- Ordinary heterotrophic organisms (OHOs), which are the truly active or viable fraction

of the MLSS; these organisms perform the degradation of organic pollutants using the

available nutrients and dissolved oxygen (DO) to create new biomass.

- Endogenous residue, which is the non-biodegradable material left behind as a result of

endogenous respiration; it is mainly comprised by cell wall material.

- Un-biodegradable suspended and settable organics from the influent, which intrinsically

mix with the OHOs and endogenous residue (Henze et al., 2008; Winkler, 1981).

The ratio between the MLVSS and MLSS is typically around 0.70-0.80 (Casey, 1997; von

Sperling & International Water Association, 2007).

2.1.1.2 Nominal hydraulic retention time (HRTn)

The nominal hydraulic retention time (HRTn) is the relationship between the volume of the

aeration basin and the flowrate; meaning the time the mixed liquor stays in the aeration basin.

In CAS, the HRTn should be enough to allow the MLVSS to perform adsorption and

mineralization of the organic pollutants; it typically ranges between 6-8 hours (Gray, 2004;

Henze et al., 2008).

2.1.1.3 Sludge retention time (SRT)

Similarly to the HRT, the sludge retention time or sludge age gives an indication of the time

that solids (biomass) spend in the system. SRT is defined as the relationship between the total

mass of organisms in the reactor and the total mass of organisms leaving the system per day. It

is considered as one of the most important parameters in the AS process, as it influences the

effluent quality, the amount of required oxygen, biomass production rates, and the reactor

volume. A short sludge age is typically selected for CAS systems, i.e. SRT between 1-10 days,

which mainly removes the chemical and biological oxygen demand (COD and BOD

respectively1). However, the short sludge age produces an unstable WAS which will require

further stabilization, e.g. through sludge digestion (Al-Malack, 2006; Henze et al., 2008;

Metcalf et al., 2004; von Sperling & International Water Association, 2007).

For CAS and MBR systems, the HRTn and SRT are separated, using secondary clarifiers and

membrane separation process respectively. In this systems the SRT is usually > HRTn. Note

that there is not a proportional relationship between these two parameters, as they depend on

the organic loading rate (OLR), the MLSS concentration and sludge recirculation and waste

rates.

1 The COD is only an approximation of the organic content of the influent, as it is uses a dichromate solution which

oxidizes both organic and inorganic compounds of the sample. It is considered to be rapid but dirty method as the

results are obtained within the same day of analysis, but it generates residual wastes of mercury, hexavalent

chromium, sulfuric acid, silver and other acids. The BOD5 analysis, although its laborious process, is the standard

BOD analysis; a sample is incubated during 5 days where the available oxygen is used by microorganisms for

organic pollutants oxidation (carbonaceous oxygen demand) (Henze et al., 2008; Rice, American Public Health,

& American Water Works, 2012).

Literature Review 7

2.1.1.4 Plant loading: Organic loading rate (OLR)

The organic loading rate (OLR) refers to the amount of organic matter that the aeration basin

receives per day; it is usually expressed as the mass of COD or BOD5 per volume of the aeration

basin per day (kg COD or BOD5/(m3*d)).

2.1.1.5 Food to organism ratio (F/M)

The food to microorganism ratio (F/M) represents the relationship between the substrate load

(usually in kg of BOD5) and the existing amount of microorganisms in the mixed liquor (kg

MLVSS*d). In theory, this parameter has a close relation with the SRT, as at higher F/M ratio,

higher nutrients available thus higher biomass activity. If the biomass is more active, more

biomass will be produced, which will subsequently require to increase the SRT in order to

maintain a constant sludge concentration (Gray, 2004; von Sperling & International Water

Association, 2007).

The OLR and the F/M are not basic design parameters, but can be useful for comparison to

historical data and observed operating conditions CAS systems (Metcalf et al., 2004).

2.1.1.6 Oxygen requirement

The AS requires oxygen for its respiration process. In this process, electrons from the

carbonaceous organic pollutants or nitrogenous matter (electron donors) are transferred to

oxygen (electron acceptor) resulting in the production of adenosine triphosphate (ATP). ATP

is an energy source for the OHOs or nitrifying bacteria, which is used for growth and energy

production (Figure 2-3) (Spanjers, Vanrolleghem, Olsson, & Dold, 1998). The total oxygen

requirement can be also expressed as Oxygen Uptake Rate (OUR) in units of mass per time

(e.g. kg/d).

Figure 2-3. Oxygen requirement of the activated sludge in the aeration basin. * for Nitrogen removal systems. Adapted from

(von Sperling & International Water Association, 2007)

The respiration process when carbonaceous organic pollutants are present, is known as

exogenous respiration. But, if the AS is left without substrate addition, the respiration rate will

decrease until reaching a decaying phase; this respiration is known as endogenous respiration.

When there is no external substrate addition, the biomass is respiring its own reserve materials

or the lysis products of dead biomass; there is no growth in the absence of substrate (Loosdrecth,

Nielsen, Lopez-Vasquez, & Brdjanovic, 2016; Spanjers et al., 1998). The sum of the exogenous

and endogenous respiration is the total sludge respiration (Equation 1).

Oxygen requirement

Oxidation of carbonaceous organic pollutants

Bacterial synthesis

Energy production

Oxidation of the nitrogenous matter*

Literature Review 8

𝑟𝑂2 = 𝑟𝑂2,𝑒𝑛𝑑𝑜 + 𝑟𝑂2,𝑒𝑥𝑜 (Equation 1)

Where:

rO2 = Total respiration (mgO2/L)

rO2,endo = Endogenous respiration (mgO2/L)

rO2,exo = Exogenous respiration (mgO2/L)

If the MLVSS concentration is known, specific exogenous or endogenous respiration can be

calculated by dividing the endogenous respiration value with the MLVSS concentration value

(unit of mgO2/mgMLVSS).

2.2. Aeration for wastewater treatment

Aeration is a key operational process in biological wastewater treatment, however it is high

energy-consuming, and it usually represents 45-75% of the total costs of a wastewater treatment

plant (Henze et al., 2008). The main purpose of aeration is to provide enough oxygen for the

total respiration process of organisms in the mixed liquor. In CAS, the aeration has also a role

in the mixing of wastewater and AS, promoting a continuous state of agitated suspension of the

sludge flocs which ensures maximum contact between flocs and wastewater. The organic

pollutants are adsorbed by suspended AS flocs, and are oxidized by aerobic heterotrophic

bacteria using the available oxygen as electron acceptor (Gray, 2004; Roš, 1993).

Aeration is defined as “a gas-liquid, mass transfer process in which interface diffusion occurs

when a driving force is created by a state of disequilibrium” (Roš, 1993). The mass transfer

process of oxygen into mixed liquor can be affected by several factors, such as temperature of

the gas and liquid, elements present in the liquid (e.g. salinity, surfactants, organic compounds,

suspended solids), gas partial pressure in the atmosphere, gas solubility, bubble surface area,

and depth at which the process is occurring (Metcalf et al., 2004; Roš, 1993).

The mass transfer process of oxygen into the mixed liquor can be described by the two-film

theory model. Metcalf et al. (2004) states that “under steady-state conditions, the mass transfer

rate of gas through the gas film must be equal to the rate transfer through the liquid film”; this

relation can be expressed as in Equation 2.

𝑁 = 𝐾𝐿𝐴(𝐶𝑠 − 𝐶𝐿) = 𝐾𝐺𝐴(𝑃𝑔 − 𝑃) (Equation 2)

Where:

N = mass transfer per unit time

KL = liquid film coefficient defined as DL/YL (Diffusion coefficient / Liquid film

thickness)

A = cross-section area through which diffusion occurs

Cs = saturation or equilibrium concentration of gas in the liquid

CL = concentration of oxygen in the liquid

KG = gas film coefficient defined as Dg/Yg (Coefficient of diffusivity through the

gas film / Gas film thickness)

Pg = partial gas pressure in the gas-phase

P = partial gas pressure at the interface in equilibrium with Cs

Literature Review 9

However, being oxygen a gas with low solubility, the degree of mass transfer is governed by

the liquid film resistance, therefore the second part of Equation 2 can be neglected as it does

not constitute a rate controlling step on the mass transfer process (Roš, 1993). Equation 2 can

be then simplified into Equation 3.

𝑁 = 𝐾𝐿𝐴(𝐶𝑠 − 𝐶𝐿) (Equation 3)

Taking into account the wastewater or mixed liquor volume (V), Equation 3 can be transformed

into Equation 4 to express the mass transfer process in concentration units. From this equation,

the KLa, or overall oxygen transfer coefficient2, can be obtained (where A/V = a, or specific

surface or interfacial area). Equation 4 can be then simplified into Equation 5.

1

𝑉𝑁 =

𝑑𝐶

𝑑𝑡= 𝑲𝑳

𝑨

𝑽(𝐶𝑠 − 𝐶𝐿)

(Equation 4)

𝑑𝐶

𝑑𝑡= 𝑲𝑳𝒂(𝐶𝑠 − 𝐶𝐿)

(Equation 5)

Furthermore, the oxygen-consuming processes from microbial activity should be taken into

account; in Equation 6 the total sludge respiration is included in the two film theory equation.

𝑑𝐶

𝑑𝑡= 𝐾𝐿𝑎(𝐶𝑠 − 𝐶𝐿) − 𝑟𝑂2

(Equation 6)

The KLa is used to characterize aeration devices regarding their aeration efficiency (Winkler,

1981). A numerical value of KLa is unique for each situation, because it can be influenced by

the following factors:

- Wastewater physic-chemical characteristics.

o Suspended solids. Dissolved or other suspended solids in the liquid influence

the environmental oxygen saturation concentration, as at higher concentration

of these particles, the lower the liquid’s capacity to hold a gas.

The saturation concentration as a function of salinity has been widely studied;

there are tables available describing this relationship based on the early study

made by Weiss (1970).

Other commonly found suspended solids are surfactants or surface active agents,

e.g. detergents, soaps, or organic acids. At high concentrations, these

constituents tend to reduce the surface tension of the water which will decrease

the size of the bubbles coming from the aeration system. The reduced bubbles

size, increases the surface area through which diffusion can occur which might

have a positive impact on the saturation concentration (Roš, 1993). However, at

low concentrations, the surfactants act as a barrier for molecular diffusion in the

liquid-gas interface which will reduce the oxygen transfer rate (Wang, Pereira,

Hung, & Shammas, 2009).

2 The terminology of KLa is very diverse; it can be found as volumetric mass-transfer coefficient, overall film

coefficient, oxygen transfer rate, or absorption coefficient (Gray, 2004; Metcalf et al., 2004; Roš, 1993; Winkler,

1981).

Literature Review 10

To take into account the influence of suspended solids in the oxygen transfer

rate (OTR), the correction factor β is used (Equation 7).

𝛽 =𝐶𝑠(𝑤𝑎𝑠𝑡𝑒𝑤𝑎𝑡𝑒𝑟)

𝐶𝑠 (𝑐𝑙𝑒𝑎𝑛 𝑤𝑎𝑡𝑒𝑟)

(Equation 7)

Regarding to mixed liquor, the MLSS concentration seems to be the main

controlling parameter for KLa, resulting an inversely proportional relation as

shown in Figure 2-4 (Germain et al., 2007).

Figure 2-4. Influence of mixed liquor suspended solids concentration in KLa20 (Germain et al., 2007)

o Temperature. Wastewater temperature also influences the KLa. There is an

inversely proportional relationship between temperature and oxygen saturation;

at higher liquid temperature, lower oxygen saturation concentration. This is

caused due to reduction in gas-film resistance (higher molecular diffusivity

resulting in a faster gas release), and decrease of gas solubility (Cs) in the liquid

(Wang et al., 2009). Similar to salinity, there are tables describing saturation

concentration as a function of temperature, e.g. available at Metcalf et al. (2004).

The KLa can be corrected for temperature using Equation 8.

𝐾𝐿𝑎20 = 𝐾𝐿𝑎𝑇𝜃(20−𝑇) (Equation 8)

Where:

KLa 20 = KLa at standard temperature (20°C)

KLa T = KLa at experiment temperature (1/h)

θ = Temperature correction factor (1.024) (-)

T = Temperature of the experiment (°C)

The typical value for θ is 1.024, which is the most used value for both

mechanical and diffused aeration systems, however the values of this coefficient

changes according to the conditions of the study; e.g. Bewtra, Nicholas, and

Polkowski (1970) reported values between 1.01768 and 1.02082, Metcalf et al.

(2004) reported values between 1.015 and 1.040, and Wang et al. (2009)

reported values between 1.016 and 1.047.


- Mixing intensity and type of aeration system. Higher mixing intensity can increase

turbulence which will reduce the liquid and gas films thickness described by the two-

film theory model; the thinner the films, the faster the mass transfer (Wang et al., 2009).

The effect of the type of aeration system on the oxygen transfer efficiency can be

described with the alfa factor (α) (Equation 9). The alfa factor values can vary greatly

depending on the conditions, however there seems to be a relationship between this

factor and the type of aeration system used; the alfa factor is lower for diffused aeration

systems than for mechanical systems (Table 2-1). If the alfa factor is closer to 1, the

oxygen transfer in wastewater approaches the oxygen transfer value in clean water,

suggesting higher OTR.

𝛼 =𝐾𝐿𝑎 𝑤𝑎𝑠𝑡𝑒𝑤𝑎𝑡𝑒𝑟

𝐾𝐿𝑎 𝑐𝑙𝑒𝑎𝑛 𝑤𝑎𝑡𝑒𝑟

(Equation 9)

Table 2-1. Alfa factors found for different aeration methods using clean water with detergent. (a) Eckenfelder (1989) cited by

Roš (1993), (b) Mueller et al. (2002).

Type of aerator Alfa factor

Mechanical surface aerator 0.60-1.20(a)

Fine bubble diffuser 0.40-0.60(a)

Coarse bubble diffuser 0.65-0.75(a)

0.31-0.53(b)

Porous diffuser 0.20-0.52(b)

The alfa factor also accounts for the effect of suspended solids in wastewater on oxygen

transfer, and it is usually lower at higher MLSS concentrations (Figure 2-5) (Brepols,

2011). According to Judd (2007), “the principal impact of [MLSS] concentration is on

the interfacial area a, which decreases with increasing solids level whilst leaving the

mass transfer coefficient KL largely unaffected. This has been attributed to the

promotion of bubble coalescence by suspended solids, and the effect is also aeration

rate-dependent.”

Figure 2-5. Impact on alfa factor according to mixed liquor suspended solids concentration (Judd, 2007)


Besides KLa, other efficiency parameters to compare aeration systems are:

- Oxygen Transfer Rate (OTR) and Standard Oxygen Transfer Rate (SOTR). The

OTR parameter expresses the amount of oxygen transferred to the liquid by the aerator

per unit time, no matter how efficient it is. It is defined using Equation 5, and it is usually

expressed in units of kgO2/d. The SOTR is used to describe OTR in standard conditions,

i.e. zero DO, zero salinity, 20°C, and 1 atm (Henze et al., 2008). SOTR is also known

as Oxygenation Capacity (OC) (Casey, 1997).

- Aeration Efficiency (AE) and Standard Aeration Efficiency (SAE). The AE

expresses the aerator’s efficiency regarding the use of energy. It is usually defined by

Equation 10. SAE is the corrected value of AE for standard conditions.

𝐴𝐸 =𝑂𝑇𝑅

𝑃

(Equation 10)

Where:

AE = Aeration efficiency (kgO2/kWh)

OTR = Oxygen Transfer Rate (kgO2/h)

P = Power drawn by the aerator (kWh)

- Oxygen Transfer Efficiency (OTE) and Standard Oxygen Transfer Efficiency

(SOTE). This parameter is mainly used for submerged aerators. It is usually expressed

in % and it is defined by a mass balance equation regarding the amount of oxygen going

in and out of the liquid (Equation 11). Due to the nature of OTE calculation, this

parameter cannot be defined for aeration systems that do not inject air or high purity

oxygen into the liquid, i.e. surface aerators. SOTE is the corrected value of OTE for

standard conditions.

𝑂𝑇𝐸 =(𝑂2,𝑖𝑛 − 𝑂2,𝑜𝑢𝑡)

𝑂2,𝑜𝑢𝑡

(Equation 11)

Where:

OTE = Oxygen Transfer Efficiency (%)

O2,in = Oxygen mass flux going into the test water (mgO2)

O2,out = Oxygen mass flux going out the test water (mgO2)

The values of SOTE for clean water and wastewater can be also used to calculate the

previously mentioned alfa factor, using Equation 12.

𝛼 =𝛼𝑆𝑂𝑇𝐸

𝑆𝑂𝑇𝐸

(Equation 12)

Where:

αSOTE = Standard Oxygen Transfer Efficiency of wastewater (%)

SOTE = Standard Oxygen Transfer Efficiency of clean water (%)


2.2.1. Aeration systems The most commonly used aeration methods in CAS can be divided into 2 systems (Table 2-2);

the choice between them depends on the aeration purpose, size of the reactor, and installation

and operation cost (Metcalf et al., 2004). These aeration systems are also differentiated by the

air flow requirement (only for diffused aeration systems) which is related with aeration costs,

and efficiency parameters, e.g. SAE and SOTE (Table 2-3).

Table 2-2. Aeration systems for aeration of activated sludge (Gray, 2004; Metcalf et al., 2004; Winkler, 1981)

Aeration

system

Aeration

method Description Types of aerators

Mechanical

aeration

Surface

aerators

The aeration equipment is located at or

near the surface of the liquid in the

aeration basin; the equipment has rotary

parts that create turbulence which

enhances oxygen transfer from the

atmosphere to the liquid.

Turbine aerators

(vertical shaft)

Rotary-brush

aerators (horizontal

shaft)

Submerged

aerators

The aeration equipment is submerged in

the aeration basin. For the vertical shaft

types, it is common to find both

mechanical and diffused aeration

systems; the mechanical system helps to

break down the bubbles coming from the

diffused system and also provides more

mixing.

Turbine aerators

(vertical shaft) +

fine/course bubble

diffuser

Disc aerators

(horizontal shaft)

Diffused

aeration

Submerged

aerators

The aeration equipment releases a stream

of pressurized air generating bubbles in

the mixed liquor.

Fine bubble or

porous diffusers

(e.g. ceramic

diffusers, membrane

tubes)

Coarse bubble or

Non-porous diffuser

(e.g. perforated

tubes)

Liquid jets


Table 2-3. Efficiency parameter values in clean water and process water of different types of aerators. (a) Henze et al. (2008), (b) Gray (2004), (c) Mueller et al. (2002), (d) Metcalf et al. (2004)

Type of

Aerator Description

SAE

(kgO2/kWh)

SOTE

(%)

αSOTE

(%)

Turbine

aerators

High-speed surface aerator 0.9-1.3 (a)

1.1-1.4(d)

N/A N/A

Low-speed surface aerator 1.5-2.1(a)

1.5-2.3(b)

1.5-2.1(d)

N/A N/A

Rotary brush

aerators

Kessner brush Up to 2(b)

1.5-2.1(d)

N/A N/A

Fine bubble

diffuser

(-) 3.6-4.8(a)

2.0-2.5(b)

Porous diffuser Ceramic disc, at 4.6 m

submergence

4.1-6.1(c) 25-38(c)

25-35(d)

9-12%(c)

Ceramic dome, at 4.6 m

submergence

3.4-6.0(c) 23-35(c)

27-37(d)

6-17%(c)

Coarse bubble

diffusers

Perforated tubes, mid-width, at

4.6 m submergence

1.5-1.6(c) 11-13(c)

10-13(d)

5%(c)

(-) 0.6-1.5(a)

0.8-1.2(b)

Liquid jets (-) 1.2-1.8(a)

1.7-2.0(c)

15-24(c)

Directional, at 3.8 m

submergence

7-11%(c)

2.2.1.1 High purity oxygen (HPO) systems

Since air is mainly composed by Nitrogen (by around 78%), and Oxygen constitutes only 21%

of it, aeration methods using high purity oxygen (HPO) have been designed expecting an OTE

increase by a factor of ~5; higher partial pressure of Oxygen in HPO gas results in a higher

saturation which increases the gas-liquid interface and consequently the gas mass transfer

process to the liquid phase (Barber, Ashley, Mavinic, & Christison, 2015). With higher OTE,

smaller aeration basins are required, as higher MLSS concentration can be used to treat the

same amount of wastewater (Mueller et al., 2002). Additionally, the HPO systems are

characterized for using mixing equipment mainly for keeping sludge flocs in suspension and

not to create high turbulence in the mixed liquor. This is due to a significant difference between

Cs – C (see Equations 2-6)3 that result in high levels of residual dissolved oxygen without

increasing KLa by for example, increasing liquid’s turbulence (Gray, 2004). The HPO methods

can be divided in 2 systems (Figure 2-6).

3 The saturation concentration value (Cs) at standard conditions using HPO would be increased from 9.09 to 43.4

mgO2/L.


Figure 2-6. Pure oxygen systems for activated sludge aeration and examples. (a) EPA (1973), (b) Metcalf et al. (2004), (c) Blue

in Green , (e) Linde Group , (f) Gray (2004), (g) Mueller et al. (2002)

The closed system, as known as the UNOX process, consists in a closed aeration basin where

the HPO is injected into the head space. There are submerged or surface mechanical aerators

for mixing purposes which increase liquid’s turbulence and oxygen transfer process. The

UNOX process usually has several compartments; the HPO is injected in the first compartment,

and the mixed liquor with the dissolved HPO flows from one compartment to another. The

advantages of the UNOX process is that the MLSS concentration can be increased to 5-6 g/L,

the F/M can be doubled, shorter SRT can be used, and there is less production of WAS

compared to CAS. However, the existing wastewater treatment plant infrastructure has to be

upgraded to apply this system, and there is potential danger of explosion due to enclosed oxygen

(Gray, 2004; Mueller et al., 2002).

The open systems in the other hand do not need to upgrade the existing infrastructure, and there

is better access for maintenance of the system. Nevertheless, the system must be highly effective

to ensure a minimum wastage of HPO (Gray, 2004); according to Mueller et al. (2002) the

SOTE found for this system is around 92%. In the side-stream systems, a stream of AS is sent

to the aeration method located outside the open aeration basin, usually a pressurized vessel,

where it will be saturated with HPO and recirculated to the basin (Mueller et al., 2002).

Although the HPO systems have been studied since 1940’s, they have not been widely applied

to conventional wastewater treatment, identifying the high cost of HPO gas production as the

main constraint of their application (EPA, 1973). However, as there is less WAS produced, the

sludge treatment costs are considerably less compared to CAS (Mueller et al., 2002).

Due to the scope of this research, only the Speece cone will be further explained in the next

section.

High purity oxygen systems

Closed-systems UNOX(a),(f)

Open-systems

I-SOTM (g)

Venturi injectors

Diffuser hoses (SOLVOX-B) (e)

Side-stream systems

Speece cone(b)

U-Tube contact aerator(b)

SDOX(c)

SOLVOX-R(e)

Vitox(f)


2.2.1.2 Speece Cone

The Speece cone is the common name for the downfall bubble contact aeration system proposed

by Professor Richard Speece in 1971 (Speece, Madrid, & Needham, 1971). The original design

of the Speece cone was an inverted Imhoff cone, where water enters at high velocity at the top,

flowing downwards, and leaves at the bottom of the cone with low velocity (Figure 2-7). HPO

is continuously injected at the top of the cone, close to the water inlet, which forms a gas

compartment. The inlet water velocity is designed to be greater than buoyant velocity, and the

outlet water velocity to be smaller, thus retaining the oxygen within the cone which forces its

dissolution in water (Speece et al., 1971).

Figure 2-7. Speece cone original design at bench scale (left) and at pilot scale (right) (Speece et al., 1971)

Since its creation, the Speece cone has been used for the oxygenation of the hypolimnion of

lakes and reservoirs in order to avoid the development of anoxia and further release of

Phosphorous and Nitrogen (Moore et al., 2012). The first Speece cone installed with this

purpose was in 1990 at the Logan Martin Dam, Alabama, USA. It is located outside the dam,

and it treats 741.6 m3/s of water with a 90% oxygen adsorption efficiency if the flow is

maintained constant. >2,000 kgO2 is delivered by this Speece cone per day (US Army Corps of

Engineers, 2009). Later, the first submerged Speece cone was installed into the Newman Lake,

Washington, USA (Figure 2-8). It has a height of 5.5 m and 2.8 maximum diameter, and it was

designed to deliver 1,360 kgO2/d (Moore et al., 2012).

Figure 2-8. Newman Lake’s Speece cone aeration system diagram (Moore et al., 2012)


Besides the original study of Speece regarding the design parameters and absorption efficiency

of the downfall bubble contact aerator (Speece et al., 1971), there is little scientific literature

published regarding the oxygen transfer efficiency parameters of the Speece cone. McGinnis

and Little (1998) designed a model to study the bubble dynamics and oxygen transfer in a

Speece cone, however the model was tested using experimental data from a full-lift

hypolimnetic aerator and not a Speece cone, therefore results should be taken with careful

consideration.

Recently, a research group from the University of British Columbia, Canada, has been

developing experimental research regarding the Speece cone efficiency in water for

hypolimnetic aeration purposes. Their first study used a laboratory-scale Speece cone to

evaluate the oxygen transfer efficiency parameters (KLa20, SOTR, SAE, and SOTE) according

to different values of flow rate, outlet port discharge velocity, and oxygen flows (Ashley,

Mavinic, & Hall, 2008). A total of 76 experiments were developed combining the different

variables. Figure 2-9 shows a summary of the results of this research. According to these results,

there seems to be a linear relationship between the discharge velocities and KLa20, as well as an

increase in the KLa20 value at greater oxygen flow rates; similar trend is observed with SOTE.

A significant decrease in SOTR and SAE were observed at higher discharge velocities and

lower oxygen flow rates.

Figure 2-9. Summary of results of the experimental research on oxygen transfer efficiency parameters of a lab-scale Speece

cone (Ashley et al., 2008)

Later, the same research group performed a similar research in order to evaluate the oxygen

transfer efficiency parameters in water of a pilot-scale Speece cone (Ashley, Fattah, Mavinic,

& Kosari, 2014). They studied the influence of the inlet water velocity and oxygen gas


flow/water flow rate ratio (%) on the oxygen transfer efficiency parameters, using a greater

number of experiments, i.e. 560, combining the different variables. Figure 2-10 shows the

results of these experiments. KLa20, SAE and SOTR seem to have a linear relationship with

increasing oxygen gas flow/water flow rate ratio and higher values at lower inlet velocities. The

SOTE had an inverse relationship with the oxygen gas flow/water flow rate ratio, but

efficiencies close to 100% were achieved at the lowest oxygen gas flow/water flow rate ratio,

i.e. 0.5%.

Even though Speece et al. (1971) also recommended the Speece cone for AS aeration, there is

only one published scientific study regarding this application up to date. The study was

performed by a research group from UNESCO-IHE at Delft, The Netherlands, using a pilot-

scale MBR at the Harnaschpolder wastewater treatment plant (C. M. Barreto, Garcia,

Hooijmans, Herrera, & Brdjanovic, 2017). The aim of the study was to evaluate the

performance of the MBR regarding effluent quality, membrane permeability, sludge

filterability, DO concentration and OUR, at high MLSS concentrations. According to this study,

the Speece cone was used to maintain a DO concentration level at 2 mgO2/L in every MLSS

concentration level. The Speece cone resulted effective in maintaining this DO concentration

level at MLSS from around 8-23 g/L (C. M. Barreto et al., 2017). Nevertheless, the Speece cone

has not been evaluated yet regarding its oxygen transfer efficiency parameters in sludge, which

can be used to obtain α values.


Figure 2-10. Summary of results of the experimental research on oxygen transfer efficiency parameters of a pilot-scale

Speece cone (Ashley et al., 2014)


2.3. Membrane bioreactors (MBR) in wastewater treatment

2.3.1. Membrane separation process The membrane separation process uses a selective material, which according to its selectivity

characteristics, separates particles and substances from wastewater in a concentrated stream as

known as retentate. The substances passing by the membrane, usually water, form another

stream as known as permeate. The flow of water passing through a specific membrane are is

known as flux.

The membrane separation process is merely physical, therefore the separated substances keep

their original chemical and biological features (Gunder, 2001). It is a passive process, as it

requires a driving force to transport substances through the membrane; in wastewater treatment

applications the main driven force is pressure (Judd, 2007).

The main membrane selectivity characteristic is pore size, usually in units of µm, which

determines which type of substance or particles are rejected by the membrane. There are 4 types

of pressure-driven membrane separation processes according to their selectivity characteristics:

Reverse Osmosis (RO), Nanofiltration (NF), Ultrafiltration (UF), and Microfiltration (MF)

(Figure 2-11).

Figure 2-11. Classification of membrane separation processes according to their selectivity (ZENA, 2017)

2.3.2. Membrane bioreactor (MBR) The membrane separation process has been coupled with the suspended-growth AS biological

treatment to separate the MLSS compounds (e.g. organic compounds, bacteria, and viruses

depending on membrane selectivity) and treated water, role that has been traditionally fulfilled

by secondary clarifiers in CAS systems. The coupling of the aeration basin with a membrane

module has been defined as membrane bioreactor (MBR). There are 2 main MBR

configurations: immersed (iMBR) and pumped side-stream (sMBR), with greater application

of iMBR in medium-large wastewater treatment plants. The use of sMBR may result in higher

operational costs due to pumping of mixed liquor to the external membrane filtration tank

(Gunder, 2001; Judd, 2007).


Although the MBR technology has been applied in the wastewater treatment sector for several

decades, it has not yet taken over the market of traditional systems, such as CAS, as it is

considered to have higher investment and operational costs. However, the cost of membrane

technologies is decreasing with time due to the economies of scale and improvement of

technology, mainly due to longer lifespans (Judd, 2007).

2.3.2.1 Advantages of MBR over CAS

The use of MBR substitutes the role of secondary clarifiers in CAS, immediately reducing the

area required for wastewater plants, which results in a smaller footprint with the same treatment

capacity (Water Environment Federation, 2012). Additionally, when a MBR is used, all mixed

liquor substances and particles greater than the pore size of the membrane will be rejected,

therefore there is no dependence in the settling characteristics of the mixed liquor which is an

important parameter for the separation process in the CAS system (Judd, 2007).

MBR can usually work with higher MLSS concentrations than in CAS, 10-15 g/L (Brepols,

2011), increasing F/M ratios, and cope with higher OLR. Additionally, the MBR reduces the

production of sludge, which can reduce the sludge treatment costs (Gunder, 2001).

The effluent resulting from MBR has a high quality, depending on membrane selectivity, the

effluent can have drinking water quality therefore MBR can be used in areas with water scarcity

issues. Moreover, MBRs can be applied in areas with strict effluent discharge limits regarding

organic and inorganic compounds such as phosphates, nitrogenous compounds, or COD (Water

Environment Federation, 2012).

2.3.2.2 Disadvantages of MBR

During the membrane separation process, the retentate particles tend to accumulate in the

membrane surface, which ultimately increases resistance against the membrane separation

process. This process is known as membrane fouling4, and it can considerably reduce the flux

or increase the transmembrane pressure (TMP) (Gunder, 2001; Judd, 2007).

Membrane fouling can be classified as physically reversible, chemically reversible, and

irreversible. Frequent maintenance activities have to be applied to reduce reversible fouling

such as:

- Membrane aeration or air scouring. The purpose of aeration of the membrane module

is to increase turbulence in the membrane surface thus reducing the fouling rate due to

materials deposited on top of the membrane; the air scouring is usually performed by

coarse bubbles diffusers. However, there has to be careful selection of aeration flow as

intensive aeration can damage the floc structure and reduce their size. Additionally,

aeration is an energy intensive process, therefore it is recommended to calculate the

specific aeration needs of the membrane module (Brepols, 2011; Judd, 2007; Water

Environment Federation, 2012).

- Influent pre-treatment. Higher levels of the influent pre-treatment is essential to

protect the membrane module, such as use of fine screen equipment. Without fine

4 Membrane fouling can be also caused by substances precipitation (scaling) or aging of membrane material

(polimerization) (Gunder, 2001).


screens, higher membrane fouling rates or damage to the membrane are more frequent

due to the presence of hair and other inert fibrous materials (Brepols, 2011).

- Membrane module cleaning. To ensure MBR’s performance with time, membrane

cleaning activities should be applied regularly. These activities are divided into 2 main

methods (Figure 2-12).

o The physical methods include back-flushing, which consists in changing the

direction of the flux, and relaxation, which is simply stopping the membrane

separation process. These 2 techniques can be applied simultaneously with air

scouring to enhance the removal of membrane fouling (Judd, 2007).

o The chemical methods include the use of bases, acidic or oxidant substances to

remove the membrane fouling. This techniques can be done insitu or exsitu, i.e.

applying the chemical substances within the MBR or using a separated tank to

submerge the membrane module respectively (Judd, 2007).

Figure 2-12. Membrane cleaning methods (Judd, 2007)

The application of membrane cleaning activities involve investment in financial and human

resources which increases operational costs (Water Environment Federation, 2012).

2.4. Turbo Membrane Bioreactor (tMBR)

The Turbo Membrane Bioreactor (tMBR) is a special type of an iMBR which focuses on

improving the membrane bioreactor process performance with support of other two

technologies (C. Barreto, 2015):

- Hydrocyclone for physical pre-treatment. The hydrocyclone separates the mixed

liquor’s heavier particles from the supernatant before sending it to the MBR aerobic

tank. This step could possible reduce the membrane fouling rates and the energy needs

for scouring aeration.

- Speece cone as concentrated oxygen delivery system. The Speece cone dissolves

HPO into the mixed liquor resulting in a supersaturated stream (Section 2.2.1.2) which

allows to overcome the current process limitations in terms of oxygen transfer when

working on high MLSS concentration.


The tMBR is currently under study and its performance compared to conventional CAS and

other MBR configurations has not been determined yet. This research will contribute to the

tMBR state of knowledge regarding the oxygen transfer efficiency parameters of the Speece

cone to water and AS at low MLSS concentrations, i.e. 5 g/L.

Methodology 24

CHAPTER 3

Methodology

3.1. Experimental setup

The experiments of this research were performed in a pilot scale MBR system equipped with a

Speece cone as concentrated oxygen delivery system. The pilot MBR system was comprised of

a series tanks in which different process take place, namely anoxic tank, MBR aerobic tank,

permeate tank, and waste sludge tank (Figure 3-1).

Figure 3-1. Scheme of complete MBR system and Speece cone

3.1.1. Speece Cone The Speece cone used in this system (Figure 3-2), was designed by the company Eco Oxygen

Technologies, LLC, based in Indiana, USA. The Speece cone used had a height of 1.17 m

including the base of the cone (0.92 m of cone plus 0.25 cm of cone base), a maximum diameter

of 0.32 m, and a total volume of 60 L (measured experimentally). However, the total volume

of the Speece cone subsystem is 80 L, as it takes into account the volume of the pipes that

connect the MBR aerobic tank, the sludge pump and the Speece cone.

Screened influent

tank

Air blower

Permeate/backwash

pump

Permeate

Speece

cone

FI

Mass flow

controller

PI

VSD

DO pH T

Data logging

Sludge pump

FIT

PITLS

I-15 Anoxic High MLSS MBRAerobic

Oxygen

gas

B.W.

Oxygen Supersaturated stream

To permeate tank

PIT

RAS

Methodology 25

Figure 3-2. Speece cone before (left) and after installation (right) in the pilot membrane bioreactor (C. Barreto, 2015)

The Speece cone was connected to a high purity oxygen (HPO) source, in this case an industrial

HPO gas cylinder containing 50 L at 200 bar. A series of pressure valves were installed in the

HPO gas cylinder to be able to regulate the oxygen pressure. In addition, a digital mass flow

controller was added to provide a constant flow of HPO to the Speece cone. Two hose were

installed at the inlet and at the outlet of the Speece cone in order to measure the oxygen transfer

capacity immediately before and after the oxygenation process inside the Speece cone.

The Speece cone was operated continuously taking water or activated sludge (AS) from the

MBR aerobic tank in a closed recirculation circuit; using a progressive cavity pump, the low

oxygen concentration water or AS from the MBR aerobic tank was transported to the Speece

cone’s inlet where HPO was applied to produce an oxygen supersaturated stream which was

returned back to the MBR aerobic tank (Figure 3-1).

3.1.2. Membrane Bioreactor aerobic tank The MBR aerobic tank had an installed hydraulic capacity of 1 m3/d, and an immersed

membrane module with a tubular configuration fitted within a round metallic compartment

(Figure 3-3); Table 3-1 shows additional information on the membrane module characteristics.

The tubular membranes are sealed at one end and are connected to a permeate-backwash pump

at the open end through a built-in manifold. The permeate pump creates a suction force that

makes the AS to flow from the outside to the inside of the membrane tubes resulting in a

permeate stream or effluent collected in the permeate tank, while the retentate stream remains

in the aerobic tank as concentrated mixed liquor. The permeate pump can also operate in the

opposite direction for membrane backwashing purposes. The membrane module also has

distribution pipes coupled directly underneath the membrane for air scouring purposes (Figure

3-4).

Methodology 26

Figure 3-3. Membrane module before installation in the membrane bioreactor (C. Barreto, 2015)

Figure 3-4. Distribution system for air scouring purposes (C. Barreto, 2015)

Table 3-1. Membrane module characteristics (C. Barreto, 2015; Membranes Modules Systems, 2005)

Characteristic Description Unit

Type of membrane process UF (-)

Configuration Tubular (-)

Material Polivinylide fluoride (-)

Molecular weight cut off 250 KDa5

Installed membrane area 20 m2

pH resistance range 0 to 12 (-)

Operational fluxes range 10 to 40 L/(m*h)

Maximum TMP range -50 to 40 kPa

Maximum temperature 40 ºC

5 1 Dalton (Da) is equivalent to 1 g/mol.

Methodology 27

3.1.3. Other equipment The following equipment was also part of the MBR system:

- Sludge pump. A progressive cavity pump that sends water or AS from the MBR aerobic

tank to the Speece cone sub-system where it gets super-saturated with HPO gas. It has

a maximum capacity of 12 m3/h at 7 bar in clean water.

- Recirculation pump. It sent water or AS from the MBR aerobic tank to the anoxic

tank. Also used to extract WAS.

- Low pressure blower. Installed to run the conventional aeration system (fine bubble

diffuser).

- Control box. With a Programmable Logic Controller (PLC) that allowed the membrane

filtration process to be automated.

- Electromagnetic flow indicator transmitter (FIT). Measured the flow discharge from

the sludge recirculation pump to the Speece cone sub-system.

- Pressure indicator transmitter (PIT). Installed in the permeate line. It measured the

applied pressure to the membrane module during the filtration or backwashing cycles.

It is connected to the PLC to ensure the TMP remains within the set points (40 kPa

while producing permeate, and -50 kPa while backwashing).

- Level monitoring system. It was an ultrasonic level measurement system installed

above the MBR aerobic tank to measure the level changes in water or AS during the

experiments.

The MBR was set up at the Harnaschpolder Wastewater Treatment Plant (HWWTP) research

hall and a previous IHE-Delft MSc thesis was already conducted using that pilot system (C.

Barreto, 2015). The HWWTP treats wastewater coming from southern region of The Hague

and Delft area. It has a treatment capacity of 1.3 million people equivalent, and it is considered

one of the biggest municipal WWT plants in Europe (Delfluent Services bv).

3.1.4. Clean water and activated sludge The clean water used for the experiments was directly taken from the water supply at HWWTP

location, which is characterized for a neutral pH (7-8) and low turbidity (less than 0.1 NTU)

(Smeets, Medema, & van Dijk, 2009).

The AS for the MBR system was pumped from a sludge pumping well which collects RAS

from the 4 secondary clarifiers located around the well as shown in Figure 3-5 (Delfluent

Services bv). The MLSS concentration of the sludge at HWWTP ranges between 3-5 g/L

depending on the operational conditions.

Methodology 28

Figure 3-5. Return Activated Sludge collection tank in Harnaschpolder wastewater treatment plant

3.2. Experimental procedure

3.2.1. Evaluation of the Speece cone performance in clean water

3.2.1.1 Preliminary evaluation

The preliminary evaluation experiments in clean water were developed to have an initial insight

about the Speece cone oxygen transfer efficiency to water (specific objective i, Section 1.1.2).

The preliminary evaluation experiments had hydraulic mixing performed only by the

recirculation pump, which sent clean water from the bottom of the MBR aerobic tank to the

top. Four different operational parameters were used to evaluate the Speece cone oxygen

transfer performance as follows:

- Recirculation flow rate. The recirculation flow rate was determined by the sludge

pump capacity. The theoretical maximum capacity of the sludge pump was 12 m3/h at

a maximum operational pressure of 7 bar (6.91 atm). However, due to the head loss

caused by the Speece cone, the maximum flow rate of this pump was 7 m3/h at 2 psi at

gauge (0.136 atm). The pump’s performance under the specific hydraulic arrangement

and head losses was characterized at different pump velocities (drive frequency) and

different discharge pressures. According to the drive frequency and cone pressures, the

maximum and minimum pump recirculation flows were determined6 (Figure 3-6).

Two recirculation flows were chosen for the experiments in clean water: (a) 6 m3/h as

high recirculation flow, and (b) 3 m3/h as low recirculation flow. Higher and lower

recirculation flows were possible, however, if used, the sludge pump performance was

not stable and could damage the system.

6 Note that by changing both the drive frequency and the discharge pressure it is possible to obtain any pressure-

flow point within the range illustrated in Figure 3-6.

Return Activated Sludge tank

Methodology 29

Figure 3-6. Maximum and minimum recirculation flows (m3/h) of sludge pump as a function of Speece cone pressure at

gauge (psi) and drive frequency (%)

The recirculation flow rate also influences on the retention time of the liquid in the

Speece cone. The retention time was calculated using Equation 13, with the volume of

the Speece cone equal to 60 L. For the recirculation flow of 6 m3/h, the retention time

is 36 seconds, and for 3 m3/h is 72 seconds.

𝐻𝑅𝑇 =𝑉𝑐𝑜𝑛𝑒

𝑄𝑟𝑒𝑐∗ 𝐶𝐹1

(Equation 13)

Where:

HRT = Retention time (s)

Vcone = Volume of the Speece cone (m3)

Qrec = Recirculation flow (m3/h)

CF1 = Conversion factor 1 = 2.78x10-4

- Cone pressure. According to the manufacturers, the maximum pressure that the Speece

cone can handle is 100 psi (6.8 atm). However, due to sludge pump capacity, it was

only possible to increase the cone pressure up to 50 psi (3.4 atm) at a recirculation flow

rate of 2 m3/h (the pressure of the cone was dependant on the recirculation flow rate).

According to these limitations, the selected pressures for the preliminary evaluation

experiments are detailed in Table 3-2.

Table 3-2. Selected pressure levels for the preliminary evaluation experiments in clean water

Recirculation flow (m3/h) Selected pressure level (psig)

High Medium Low

6 15 10 2

3 40 25 10

- HPO flow relative to Speece cone’s maximum oxygen delivery capacity rate. The

HPO flow for the experiments in the preliminary evaluation experiments were chosen

as percentages of the Speece cone’s maximum oxygen delivery capacity rate (O2Del)

0

1

2

3

4

5

6

7

8

0 20 40 60

Flo

w (

m3/h

)

Pressure (psi)

100% frequency

90% frequency

80% frequency

70% frequency

60% frequency

Methodology 30

in specific pressure and recirculation flow conditions. The Equation 14 was used to

calculate the maximum oxygen delivery capacity of the Speece cone. The Equation 15

was used to determine the absolute pressure (PA) required for Equation 14. As HPO was

used in the experiments, the O2Del had to be multiplied by a HPO factor (O2F)

equivalent to 4.7 (-), as HPO contains 21% of Oxygen.

𝑂2𝐷𝑒𝑙 = 𝑃𝐴 ∗ 𝐶𝑠20 ∗ 𝑄𝑟𝑒𝑐 ∗ 𝑂2𝐹 ∗ 𝐶𝐹2 (Equation 14)

Where:

O2Del = Maximum oxygen delivery capacity rate (kgO2/d)

PA = Absolute pressure (atm)

Cs20 = Environmental saturation concentration at standard water temperature

20°C (mg/L), equal to 9.08 mgO2/L

Qrec = Recirculation flow (m3/h)

O2F = High purity oxygen factor (-)


𝑃𝐴 = (𝑃𝑔 ∗ 𝐶𝐹3) + 𝑃𝑠 (Equation 15)

Where:

Pg = Gauge pressure (psi)


Ps = Atmospheric pressure at standard conditions = 1 atm

The value obtained from Equation 15 was considered to be the 100% theoretical oxygen

delivery capacity rate of the experiment. It was decided to evaluate the oxygen transfer

efficiency parameters at different values ranging from low to high flows as follows: 5%,

10%, 20%, 30%, 40%, ~50%, and ~100%.

The ideal gas law (Equation 16) was reorganized into Equation 17 to calculate the

oxygen volumetric flow required for each experiment. The values used to solve

Equation 17 were standard values as the digital mass flow controller requires an input

of standard litre per minute (slpm).

𝑃𝑉 = 𝑛 ∗ 𝑅 ∗ 𝑇 (Equation 16)

Where:

P = Pressure (atm)

V = Gas volume (L)

n = Number of moles in gas (g/mol)

R = Gas constant (atm*L/mol* K)

T = Temperature (K)

Methodology 31

𝑉𝑠 =𝑛 ∗ 𝑅 ∗ 𝑇𝑠

𝑃𝑠∗ 𝑂2𝐷𝑒𝑙

(Equation 17)

Where:

Vs = Oxygen volumetric flow at standard conditions (slpm)

Ps = Pressure at standard conditions = 1 atm

n = Number of moles in Oxygen = 32 g/molO2

R = Gas constant = 0.08205746 (atm*L/mol*K)

Ts = Temperature at standard conditions = 298.15 K

- Cone inlet velocity. According to the Speece cone design, the inlet velocity has “to be

greater than the buoyant velocity of the bubbles” (Speece et al., 1971) so they are able

to stay within the cone and dissolve in the liquid. At higher inlet velocities, there is

higher turbulence in the Speece cone hood, which reduces the gas-liquid films’

thickness increasing in turn the mass transfer process.

According to previous studies, the recommended design inlet velocity is 3.05 m/s

(Ashley et al., 2008). The inlet velocities evaluated during the preliminary evaluation

experiments represented were determined according to the recirculation flow and the

nominal inlet diameter of the Speece cone (42 mm). The evaluated inlet velocities were

1.2 m/s for 6 m3/h, and 0.6 m/s for 3 m3/h which represented 40% and 20% of the

recommended inlet velocity respectively.

According to these operational parameters, a total of 6 treatments and 50 experiments were

performed during the preliminary evaluation experiments; a summary of the experimental

design is shown in the Table 3-3.

As observed in the Table 3-3, not all the experiments were evaluated at the same HPO flows

because when high HPO flows were evaluated during the first set of experiments, i.e. ~50%,

and ~100% in treatments A.2 and A.6, it was observed that the oxygen was escaping as

effervescence. Therefore for the following treatments, it was decided to perform the

experiments until 40% of HPO flows in order to reduce the HPO loss. Additionally, in almost

all treatments (except A.5, which was performed at last) some experiments were performed in

triplicates. A low variation among experiments was observed, therefore, it was decided to

perform only repetition for the rest of the experiments within the treatment.

Reaeration method

The method used to determine the oxygen transfer efficiency during the preliminary evaluation

experiments was the reaeration method defined by the Standards of the American Society of

Civil Engineers (ASCE). This reaeration method requires deoxygenation or the removal of near

all the dissolved oxygen (DO) in the test water, for further reaeration with the desired

equipment, while DO concentrations changes are measured in several points in the tank

(American Society of Civil Engineers, 1992).

Methodology 32

Table 3-3. Experimental design of preliminary evaluation

Treatment

ID

Flow rate

(m3/h)

Pressure

(psi)

Inlet

velocity

(m/s)

HPO flow

relative to

O2Del

Theoretical oxygen

delivery (kg/d)

Oxygen

input (slpm)

Gas flow rate/

Water flow rate

ratio

A.1 6 15 1.20 40% (x 2) 5.28 2.80 2.6%

30% (x 1) 3.96 2.10 1.9%

20% (x 1) 2.64 1.40 1.3%

10% (x 1) 1.32 0.70 0.6%

5% (x 2) 0.66 0.35 0.3%

A.2 6 10 1.20 107% (x 1) 11.70 6.21 5.7%

55% (x 1) 6.07 3.22 3.0%

36% (x 1) 3.94 2.09 1.9%

28% (x 1) 3.03 1.61 1.5%

22% (x 1) 2.43 1.29 1.2%

17% (x 3) 1.88 1.00 0.9%

7% (x 1) 0.75 0.40 0.4%

5% (x 1) 0.56 0.30 0.3%

A.3 6 2 1.20 40% (x 1) 2.94 1.56 1.4%

30% (x 1) 2.26 1.20 1.1%

20% (x 3) 1.51 0.80 0.7%

10% (x 2) 0.75 0.40 0.4%

5% (x 1) 0.38 0.20 0.2%

A.4 3 40 1.20 40% (x 1) 4.90 3.40 4.8%

30% (x 2) 3.68 2.55 3.6%

20% (x 2) 2.45 1.70 2.4%

10% (x 3) 1.23 0.85 1.2%

5% (x 1) 0.60 0.42 0.6%

Methodology 33

Treatment

ID

Flow rate

(m3/h)

Pressure

(psi)

Inlet

velocity

(m/s)

HPO flow

relative to

O2Del

Theoretical oxygen

delivery (kg/d)

Oxygen

input (slpm)

Gas flow rate/

Water flow rate

ratio

A.5 3 25 1.20 40% (x 1) 3.53 1.88 3.4%

30% (x 1) 2.64 1.40 2.6%

22% (x 1) 1.98 1.05 1.9%

15% (x 1) 1.32 0.70 1.3%

7% (x 1) 0.66 0.35 0.6%

4% (x 1) 0.33 0.18 0.3%

A.6 3 10 1.20 53% (x 1) 2.92 1.55 2.8%

35% (x 3) 1.90 1.01 1.9%

27% (x 1) 1.47 0.78 1.4%

21% (x 1) 1.17 0.62 1.1%

16% (x 1) 0.89 0.47 0.9%

5% (x 2) 0.28 0.15 0.3%

Methodology 34

- Deoxygenation. The MBR aerobic tank and the Speece cone were filled with a known

volume of clean water, i.e. 1000 L, and the DO concentration was reduced to a residual

of less than 0.5 mgO2/L by adding sodium metabisulfate (a mixture of NaHSO3 and

Na2S2O5) to act as an oxygen scavenger. According to the relationship between the

molecular weight of this chemical compound (190.10 g/mol) and oxygen (32 g/mol),

5.941 g of sodium metabisulfate were required to remove 1 g of DO in water. The

required sodium metabisulfate doses for the experimental conditions of the day of the

test were calculated according to the Equation 18.

𝑆𝑀 = (𝐶𝑠 ∗ 𝑉 ∗ 𝐶𝐹4) ∗ 5.941 (Equation 18)

Where:

SM = Required sodium metabisulfate for deoxygenation (g)

Cs = Environmental saturation concentration in test water (mg/L)

V = Volume of test water (L)

CF4 = Conversion factor 1 = 1x10-3

The calculated amount of sodium bisulfate for deoxygenation was diluted in a 5 L

bucket and then added at the MBR aerobic tank. The duration of the deoxygenation

process of the test water varied according to test conditions and water temperature;

slower deoxygenation was experienced during low mixing conditions and lower water

temperatures. After the deoxygenation process, HPO was added to the system using the

Speece cone and DO concentration changes were measured.

Note that before deoxygenation, the environmental saturation concentration in test

water (Cs) was measured in a well-mixed beaker during 10 minutes to compensate for

instrument variation.

- DO concentration measurement. For all the DO concentration measurements a

galvanic dissolved oxygen sensor (universal) with integrated temperature compensation

was used. The measurement range of this DO probe is 0 – 50 mg/L ± 0.5% of

concentration value, or 0 – 600% ± 0.5% of concentration saturation value (WTW). The

probes were cleaned and calibrated frequently according to manufacturer instructions,

at least once per week, to ensure that the values were as accurate as possible.

The DO concentration changes were measured in one second interval during 83 minutes

(that is 4,981 data points in total). For the preliminary evaluation experiments, two

determination points were used to monitor the DO concentration changes:

o One probe was located inside the MBR aeration tank placed 30 cm below the

water level and right at the middle of the tank (Figure 3-7). Considering that the

MBR aeration tank was well-mixed, the DO concentration at this point was

representative of the overall DO concentration in the entire MBR aeration tank.

o A second probe was located at the outlet of the Speece cone.

Methodology 35

Figure 3-7. Location of the dissolved oxygen concentration probe in the MBR aeration tank for the preliminary evaluation

experiments

Intrusion

An additional experiment to estimate the overall oxygen transfer coefficient (KLaIntrusion) due to

intrusion into the clean water was performed. During the intrusion experiment, the clean water

was recirculated in the system but no oxygen was introduced using the Speece cone. The mixing

condition were kept the same. The DO probes were located in the middle of the tank using a

metallic stand with fixed probes at 25, 50 and 100 cm. One additional probe was placed at the

Speece cone outlet. The change in DO concentration was monitored and recorded every five

seconds interval for a total time of 6.9 hours (that is 24,959 data points in total).

DO concentration gradient in the MBR aeration tank

During the analysis of the preliminary evaluation experiments, it was noticed that the oxygen

mass balance of the system was not closing, suggesting a concentration gradient within the

MBR aerobic tank. Therefore, two additional experiments were carried out to confirm well-

mixed conditions:

- Experiment 1. A mobile DO probe was placed at different locations and at different

depths at each location in the MBR aeration tank and the DO concentration changes

were compared with the measurements of a fixed probe placed at the same location of

the DO probe used in the previous preliminary evaluation experiments (i.e. Location 1.3

in Figure 3-8). The experiment 1 was performed under the same mixing conditions of

the preliminary evaluation experiments; i.e. hydraulic mixing performed only by the

recirculation pump.

Once a location was set, the mobile DO probe was submerged at that location at 100 cm

depth. The DO concentration was monitored at each depth and recorded for

approximately two minutes interval. Then the probe was moved/lifted to the next depth

(i.e. 80 cm), and again the DO concentrations were monitored and recorded for two

minutes interval; this procedure was repeated for the following depths: 60 cm, 40 cm,

Location of the DO probe

Methodology 36

and 20 cm. The probe recorded the DO concentration change for a total of

approximately 10 minutes per location, then it was moved to the next location. The DO

concentration changes were recorded every second.

Five different locations were chosen for the Experiment 1 as illustrated in Figure 3-8 as

follows:

o Location 1.1, close to one of the Speece cone discharge pipes7,

o Location 1.2, corner behind the membrane module,

o Location 1.3, in the middle of the MBR aerobic tank,

o Location 1.4, close to the sludge pump suction,

o Location 1.5, close to the other Speece cone discharge pipes.

Figure 3-8. Locations chosen for Experiment 1 to measure dissolved oxygen gradient concentration in the MBR aerobic tank

- Experiment 2. Three fixed DO probes were placed at different depths and locations and

the change in DO concentration was measured at two different mixing conditions: low

mixing condition (same mixing conditions as in the Experiment 1), and high mixing

condition (using two additional mixing pumps of 6.5 and 9 m3/h capacity). Changes in

the DO concentration in the Experiment 2 were recorded every one second interval for

a total time of 83 minutes (that is 4,981 data points in total).

Three different locations were chosen for conducting the Experiment 2 as shown in

Figure 3-9:

o Location 2.1, close to one of the Speece cone discharge pipes at 100 cm,

o Location 2.2, next to the sludge pump suction at 54 cm,

o Location 2.3, in the middle of the MBR aerobic tank at 30 cm depth.

7 The main discharge pipe coming from the Speece cone to the MBR aerobic tank is subdivided into to two pipes,

each close to one corner of the MBR aeration tank and pointing at different directions, one pointing to the middle

of the tank, and other one to the tank wall, in order to increase mixing condition.

1.1

1.5

1.3

1.4

1.2

Methodology 37

Figure 3-9. Locations chosen for Experiment 2 to measure dissolved oxygen gradient concentration in the MBR aerobic tank

3.2.1.2 Evaluation of the operational parameters on the performance of the Speece cone

Following the preliminary evaluation experiments, it was decided to continue with the

experiments to evaluate the operational parameters on the performance of the Speece cone with

improved mixing conditions in the MBR aeration tank. The improved mixing conditions were

achieved by adding extra recirculation pumps of 6.5 and 9 m3/h capacity located at opposite

corners in the MBR aerobic tank. This configuration was supposed to provide well-mixed tank

conditions; however small variations were observed due to intrinsic measurement errors

between DO probes.

Additionally, more DO probes were used to monitor the DO concentration change in the test

waster; six determination points were used:

- Four probes were located at the MBR aerobic tank at different depths: one at deep

location (100 cm depth), two at mid-depth (75 and 50 cm), and one at shallow depth (25

cm). The DO probes in the tank were mounted in a metallic stand using metallic clamps

to keep the probes static and pointing to different directions (Figure 3-10). The metallic

stand was located in the middle of the MBR aerobic tank.

- Two probes were located at the cone: one for the Speece cone inlet and one for the

outlet. The probes were immersed in a bucket with the outflow coming from the Speece

cone’s inlet/outlet hoses.

Similarly to the preliminary evaluation experiments, four different operational parameters were

used to evaluate the Speece cone oxygen transfer performance, which are: recirculation flow

(influencing on the hydraulic retention time), cone pressure, HPO flow relative to Speece cone’s

maximum oxygen delivery capacity rate, and Speece cone inlet velocity.

2.1

2.3

2.2

Methodology 38

Figure 3-10. Dissolved oxygen determination points according to depth in the aerobic tank for the evaluation of optimal

conditions experiments

- Recirculation flow. The same two recirculation flows as in the preliminary evaluation

experiments were chosen for the experiments in clean water: (a) 6 m3/h as high

recirculation flow, and (b) 3 m3/h as low recirculation flow. For these recirculation

flows, the resulting retention time were 36 seconds for 6 m3/h, and 72 seconds for 3

m3/h.

- Cone pressure. For these experiments, it was decided to evaluate only maximum and

minimum pressure levels, as detailed in Table 3-4.

Table 3-4. Selected pressure levels for the evaluation of the operational parameters on the performance of the Speece cone in

clean water

Recirculation flow (m3/h) Selected pressure level (psig)

High Low

6 15 2

3 40 10

- HPO flow relative to Speece cone’s maximum oxygen delivery capacity rate. It was

decided to evaluate the oxygen transfer efficiency parameters at different values ranging

from low to high flows as follows: 5%, 10%, 20%, 30%, and 40%.

- Cone inlet velocity. It was decided to evaluate inlet velocities closer to the

recommended values, i.e. 3.05 m/s. Therefore, it was necessary to place PVC reducers

to the inlet port in order to reduce the inlet diameter of the Speece cone and increase the

Methodology 39

inlet velocity. The Table 3-5 shows the chosen PVC reducer diameters and the resulting

inlet velocity according to the recirculation flow.

Table 3-5. Inlet velocities used in the evaluation of the operational parameters on the performance of the Speece cone in

clean water

Recirculation flow

(m3/h)

PVC reducer

diameter (mm)

Inlet

velocity

(m/s)

Comparison with

recommended inlet velocity

(%)

6 25.0 3.40 111%

3 17.6 3.43 112%

According to these operational parameters, a total of 6 treatments and 50 experiments were

performed during the preliminary evaluation experiments; a summary of the experimental

design is shown in the Table 3-6.

Intrusion

An additional intrusion experiment was also performed using this higher mixing conditions, as

KLaIntrusion is influenced by mixing conditions. Same duration of the experiment and same

locations of the DO probes were used as in the low mixing condition experiment, plus another

probe at 75 cm.

Methodology 40

Table 3-6. Experimental design of the evaluation of the operational parameters on the performance of the Speece cone in clean water

Treatment

ID

Flow

rate

(m3/h)

Pressure

(psi)

Inlet

velocity

(m/s)

HPO flow

relative to

O2Del

Theoretical oxygen

delivery (kg/d)

Oxygen input

(slpm)

Gas flow rate/

Water flow rate

ratio

B.1 6 10 3.40 40% (x 1) 4.39 2.33 2.1%

30% (x 1) 3.30 1.75 1.6%

20% (x 1) 2.20 1.17 1.1%

10% (x 1) 1.10 0.58 0.5%

5% (x 1) 0.55 0.29 0.3%

A.2 6 10 1.20 55% (x 1) 6.07 3.22 3.0%

36% (x 1) 3.94 2.09 1.9%

28% (x 1) 3.03 1.61 1.5%

22% (x 1) 2.43 1.29 1.2%

17% (x 3) 1.88 1.00 0.9%

7% (x 1) 0.75 0.40 0.4%

5% (x 1) 0.56 0.30 0.3%

B.2 3 10 3.43 40% (x 1) 2.20 1.17 2.1%

30% (x 1) 1.65 0.88 1.6%

20% (x 1) 1.10 0.58 1.1%

10% (x 1) 0.55 0.29 0.5%

5% (x 1) 0.28 0.15 0.3%

B.3 3 40 3.43 40% (x 1) 4.86 2.58 4.7%

30% (x 1) 3.64 1.94 3.5%

20% (x 1) 2.43 1.29 2.4%

10% (x 1) 1.23 0.65 1.2%

5% (x 1) 0.60 0.32 0.6%

Methodology 41

3.2.2. Evaluation of the Speece cone performance in process activated sludge water

The evaluation of the Speece cone performance in process activated sludge (AS) water was

developed to reach specific objective iii, focused on AS water at 5 g/L MLSS concentration. It

was decided to evaluate the Speece cone aeration efficiency at 5 g/L which represents low

MLSS concentration, as a base line for further experiments at high MLSS concentration (10,

20, 30, 40 and 50 g/L); additionally, 5 g/L MLSS concentration is representative of most of

conventional activated sludge (CAS) systems (Casey, 1997). The experiments evaluation of the

Speece cone performance in process activated sludge will also include the determination of the

sludge endogenous respiration. These experiments were divided into 3 phases illustrated in

Figure 3-11.

Figure 3-11. Phases of experiments in process activated sludge water

3.2.1.3 Phase 1

The MBR aerobic tank and Speece cone were filled with returned activated sludge (RAS) from

Harnaschpolder WWTP at a known volume, i.e. 1000 L (equivalent to 920 L for the aerobic

tank and 80 L for the Speece cone). The initial mixed liquor suspended solids (MLSS) were

analysed in the RAS as well as final concentration in the MBR tank.

3.2.1.4 Phase 2

When the desired MLSS concentration was reached, the MBR was aerated during a day using

the air scouring installed for the membrane module. This time was considered enough to remove

all possible remaining substrate in the RAS.

3.2.1.5 Phase 3 (C)

The objective of Phase 3 was to calculate the oxygen transfer efficiency parameters and aeration

coefficients (KLa, OTE, OTR, SOTE, SOTR) of the Speece cone using process AS water at 5

g/L MLSS concentration, as well as measuring the AS endogenous respiration.

- Oxygen transfer efficiency parameters and aeration coefficients. For these

experiments, the optimal conditions regarding mixing were used, i.e. hydraulic mixing

performed by the recirculation pump, adding two extra mixing pumps of 6.5 and 9 m3/h

capacity located at the opposite corners in the MBR aerobic tank (Section Error!

Reference source not found.). Additionally, the changes on DO concentration change

were measured using the six determination points selected for the evaluation of the

operational parameters on the performance of the Speece cone with clean water (Section

3.2.1.2).

Phase 1

•Increasing MLSS concentration

Phase 2

•Aeration of activated sludge

Phase 3

• Endougenous respiration

• Determination of oxygen efficiency parameters and coefficients with re-oxygenation method

Methodology 42

The DO concentration change was monitored and recorded every one second interval

for a total time of 83 minutes (that is 4,981 data points in total).

Similarly to the experiments in clean water, four different operational parameters were

used to evaluate the Speece cone oxygen transfer performance, which are: recirculation

flow (influencing on the hydraulic retention time), cone pressure, HPO flow relative to

Speece cone’s maximum oxygen delivery capacity rate, and Speece cone inlet velocity.

A total of 4 treatments and 20 experiments were planned for this phase (Table 3-7).

- Sludge endogenous respiration. The sludge endogenous respiration was measured in

a well-mixed beaker with a 1 L sample of the oxygenated AS after the re-oxygenation

experiment. The change in the DO concentration was measured using a DO probe every

one second interval until it reached a concentration lower than 0.5 mg/L.

Additionally, to calculate specific endogenous respiration (rO2,endo) rate, a 100 mL single

sample was taken after the re-oxygenation experiment to analyse the MLVSS

concentration. The MLVSS concentration was evaluated in duplicates from the same

sample.

Methodology 43

Table 3-7. Experimental design of the evaluation of the operational parameters on the performance of the Speece cone in process activated sludge water

Treatment

ID

Flow rate

(m3/h)

Pressure

(psi)

Inlet

velocity

(m/s)

HPO flow

relative to

O2Del

Theoretical oxygen

delivery (kg/d)

Oxygen

input (slpm)

Gas flow rate/

AS flow rate

ratio

C.1 6 10 3.40 40% (x 1) 4.39 2.33 2.1%

30% (x 1) 3.30 1.75 1.6%

20% (x 1) 2.20 1.17 1.1%

10% (x 1) 1.10 0.58 0.5%

5% (x 1) 0.55 0.29 0.3%

C.2 6 10 1.20 40% (x 1) 3.94 2.09 1.9%

30% (x 1) 3.03 1.61 1.5%

20% (x 1) 2.43 1.29 1.2%

10% (x 1) 0.75 0.40 0.4%

5% (x 1) 0.56 0.30 0.3%

C.3 3 10 3.43 40% (x 1) 2.20 1.17 2.1%

30% (x 1) 1.65 0.88 1.6%

20% (x 1) 1.10 0.58 1.1%

10% (x 1) 0.55 0.29 0.5%

5% (x 1) 0.28 0.15 0.3%

C.4 3 40 3.43 40% (x 1) 4.86 2.58 4.7%

30% (x 1) 3.64 1.94 3.5%

20% (x 1) 2.43 1.29 2.4%

10% (x 1) 1.23 0.65 1.2%

5% (x 1) 0.60 0.32 0.6%

Methodology 44

3.3. Data analysis

3.3.1. Oxygenation capacity determination by numerical integration The overall oxygenation capacity of the Speece cone was determined based on a mass balance

performed on the MBR aerobic tank sub-system (Figure 3-12). The Equation 19 defines the

mass balance within the MBR aerobic tank; this equation has three components:

- The accumulation of DO concentration in the MBR aerobic tank (left side of the

equation).

- The net oxygen delivery by the cone system into the MBR aeration tank (first term on

the right side of the equation); that is, a super-saturated stream coming from the Speece

cone and a less-saturated stream leaving the tank. It represents the net oxygen delivery

in the MBR aerobic tank. This stream have the same flow, e.g. 3 or 6 m3/h depending

on the experiment.

- The oxygen intrusion or escape from the water or AS in the MBR aerobic tank to the

atmosphere (second term on the right side of the equation); intrusion will happen when

the DO concentration in the MBR aerobic tank is lower than the environmental

saturation concentration (Cs), and escape will happen when DO concentration in the

MBR aerobic tank is higher than Cs.

Figure 3-12. Sub-system selected for mass balance analysis

𝑉𝑑𝐶𝑀𝐵𝑅

𝑑𝑡= 𝑄(𝐶𝑐𝑜𝑛𝑒 − 𝐶𝑀𝐵𝑅) + [𝐾𝐿𝑎𝐼𝑛𝑡𝑟𝑢𝑠𝑖𝑜𝑛(𝐶𝑠 − 𝐶𝑀𝐵𝑅)] ∗ 𝑉 (Equation 19)

Where:

V = Clean water or AS volume (L)

dCMBR/dt = DO concentration change in the MBR aeration tank vs time (kg/d)

Q = Recirculation flow (L)

Screened influent

tank

Air blower

Permeate/backwash

pump

Permeate

Speece

cone

FI

Mass flow

controller

PI

VSD

DO pH T

Data logging

Sludge pump

FIT

PITLS

I-15 Anoxic High MLSS MBRAerobic

Oxygen

gas

B.W.

Oxygen Supersaturated stream

To permeate tank

PIT

RAS

Methodology 45

Ccone = Measured DO concentration in the outlet of the cone (mg/L)

CMBR = Measured DO concentration in the MBR aerobic tank (mg/L)

KLaIntrusion = Overall oxygen transfer coefficient in the MBR aerobic tank due to

intrusion (1/h)

Cs = Theoretical environmental saturation concentration in water (mg/L)

Note that the Cs used for these calculations was the theoretical value according to measured

temperature and electrical conductivity of test water8. The theoretical Cs was determined using

the DOTABLES software of the United States Geological Survey (USGS, 2014).

The first term on the right side of the Error! Reference source not found. was used to calculate

the Oxygen Transfer Rate (OTR, in kg/d) of the system, which according to Ashley et al.

(2014), Boyd (1986), Casey (1997) and Mueller et al. (2002), has to be determined using point

1 as 10% of Cs and point 2 as 75% of Cs. However, if the value of point 1 is still at the lag phase

due to residual sodium metasulfide, it is necessary to choose the appropriate value for point 1

after this lag period. The obtained OTR was then corrected for standard temperature using

Equation 20 to obtain Standard Oxygen Transfer Rate (SOTRi). Using the resulting value of

SOTR, the Equation 21 was used to calculate Standard Oxygen Transfer Efficiency (SOTEi, in

%).

𝑆𝑂𝑇𝑅𝑖 = 𝑂𝑇𝑅 ∗ 𝜃(20−𝑇) (Equation 20)

𝑆𝑂𝑇𝐸𝑖 =𝑆𝑂𝑇𝑅

𝐻𝑃𝑂 𝑓𝑙𝑜𝑤∗ 100

(Equation 21)

Based on the Equation 5, the first term of the right side of the Equation 19 can be also used to

estimate a KLaTransference which represents the overall oxygen transfer coefficient from the

Speece cone the MBR aerobic tank (Equation 22). However, as part of this term represents the

situation within the Speece cone, the saturation concentration cannot be the same as for

environmental conditions, as the Speece cone is a closed sub-system dependent on the pressure

conditions chosen for each experiment and in contact with a certain amount of HPO. Therefore,

the required Cs for this component will be determined by the conditions inside the cone

(Equation 23).


𝑑𝑡= [𝐾𝐿𝑎 𝑇𝑟𝑎𝑛𝑠𝑓𝑒𝑟𝑒𝑛𝑐𝑒(𝐶𝑠 𝐻𝑃𝑂 − 𝐶𝑀𝐵𝑅)] ∗ 𝑉

+ [𝐾𝐿𝑎𝐼𝑛𝑡𝑟𝑢𝑠𝑖𝑜𝑛(𝐶𝑠 − 𝐶𝑀𝐵𝑅)] ∗ 𝑉

(Equation 22)

Where:

KLa_Transference = Overall oxygen transfer coefficient in the MBR aerobic tank (1/h)

Cs HPO = Theoretical saturation concentration of high purity oxygen (mg/L)

8 The EC mean value of test water was 585 µS/cm; EC was measured using HACH TitraLabAT1000 series. The

EC value was determined to be constant during all experiments.

Methodology 46

𝐶𝑠 𝐻𝑃𝑂 = 𝑃𝐴 ∗ 𝐶𝑠 ∗ 𝑂2𝐹 (Equation 23)

Where:

PA = Absolute pressure according to experiment (atm)

O2F = High purity oxygen factor (-)

As the KLa_Intrusion was determined during the intrusion experiments, the only unknown term in

Equation 22 was KLa_Transference. To calculate its value, the DO concentration change in the MBR

aeration tank vs time was manually integrated per data point, which was compared to the

calculated values using an assumed KLa_Transference. These two terms were subtracted and

squared, and using the Excel solver tool, the resulting sum of squared residuals was minimized

by finding the optimal value of KLa_Transference.

The resulting KLa_Transference was corrected for standard temperature to obtain KLa_Transference20

using Equation 8 (Section 2.2). When an initial DO concentration value of zero is assumed, the

maximum theoretical SOTR of the system can be calculated (Max_SOTRi); the Equation 24

was used to calculate Max_SOTRi.

𝑀𝑎𝑥_𝑆𝑂𝑇𝑅𝑖 = 𝐾𝐿𝑎𝑇𝑟𝑎𝑛𝑠𝑓𝑒𝑟𝑒𝑛𝑐𝑒20 ∗ 𝐶𝑠𝐻𝑃𝑂 ∗ 𝑉 ∗ 𝐶𝐹4 (Equation 24)

Where:

Max_SOTRi = Standard Oxygen Transfer Rate (kg/d)

CF4 = Conversion factor 4 = 2.04e10-5

For the experiments that have more than one determination points, the average value of SOTR,

SOTE and Max_SOTR was determined using Equation 25.

𝑆𝑂𝑇𝑅 =∑ 𝑆𝑂𝑇𝑅𝑖

𝑛𝑖=1

𝑛

(Equation 25)

Where:

n = Number of determination points

For the experiments using process AS water, the overall oxygenation capacity of the Speece

cone has to take into account the endogenous respiration (based on Equation 6), therefore,

Equation 22 became Equation 26. Additionally, the β factor was included to correct for the

presence of solids. According to American Society of Civil Engineers (1988), this factor can

vary from around 0.8-1 in CAS, therefore it was decided to choose 0.9 for these calculations.

An intrusion experiment with process AS water was also performed.


𝑑𝑡= [𝐾𝐿𝑎_𝑇𝑟𝑎𝑛𝑠𝑓𝑒𝑟𝑒𝑛𝑐𝑒(𝛽 ∗ 𝐶𝑠 𝐻𝑃𝑂 − 𝐶𝑀𝐵𝑅)] ∗ 𝑉

+ [𝐾𝐿𝑎𝐼𝑛𝑡𝑟𝑢𝑠𝑖𝑜𝑛(𝐶𝑠 − 𝐶𝑀𝐵𝑅)] ∗ 𝑉 − 𝑟𝑂2,𝑒𝑛𝑑𝑜

(Equation 26)

Methodology 47

3.3.2. Oxygenation capacity determination by AQUASIM model AQUASIM is a program used for the simulation of aquatic systems such as mixed reactors,

biofilm reactors, advective-diffusive reactors, saturated soil columns, rivers sections, or lakes

(Reichert, 1998). The AQUASIM program was used for the calculation of the KLa_Transference20

of both experiments of clean water and process AS water.

A simplified schematic representation of the system was built (Figure 3-13) to identify the

variables (Table 3-8), compartments (Table 3-9), and links (Table 3-10) required to create an

AQUASIM model9.

Figure 3-13. Simplified schematic representation of the MBR and Speece cone system for AQUASIM modelling

Qin*C_DOin represents the HPO flow from the oxygen cylinder to the Speece cone. The HPO

flow could not be represented by a single variable as in AQUASIM, the variables can only

represent properties of water. Therefore, the HPO flow was represented as a load of oxygen

(C_DOin) in a small water flow (Qin); this small water volume does not have a great influence

on the system’s model and can be neglected for further calculations. C_DOin was calculated

using Equation 27.

𝐶_𝐷𝑂𝑖𝑛 =𝑂2𝐷𝑒𝑙 ∗ 𝐶𝐹6

𝑄𝑖𝑛

(Equation 27)

Where:

C_DOin = Initial DO concentration (mg/L)

CF6 = Conversion factor 6 = 1e6

Qin = Inflow (L/h)

Cone, MBR and Help Reactor are mixed reactor compartments. It is acknowledged that the

mixed reactor compartment might not be a fair representation of the Speece cone as there are a

number of elements influencing on the dissolution of oxygen, such as inlet velocity and

pressure. Additionally, the real system does not have a help reactor, however a compartment

9 For a detailed description of variable, compartment, or links please refer to the AQUASIM 2.0 Manual (Reichert,

1998).

Methodology 48

for the outflow of Qin is required. The help reactor has a small volume and can be also neglected

for further calculations.

As observed in Table 3-9, the C_DO is active variable for the Cone and MBR compartments,

which resulted in the calculated curves for the DO concentration change in both compartments.

Additionally, C_DOin was also active for both compartments, as they are influenced by the HPO

gas that is being inserted to the system. The initial condition for the Cone compartment is the

measured DO concentration, and for the MBR compartment is the theoretical initial DO

concentration fixed as 0. Additionally, Qrec is a bifurcation of the Cone compartment’s outflow;

it has a constant value depending on the treatment. And, C_DO_XX represents the measured

DO concentration in time of each experiment.

Oxygen transfer process represents the overall oxygen transfer process happening in the MBR

aerobic tank, including the intrusion or escape of oxygen to the atmosphere and the transference

from the Speece cone. It is represented as a dynamic process which had a rate as the two-film

theory equation (Equation 5). It also has a stoichiometric coefficient of 1 for the variable of

C_DO. This process was only active for the MBR compartment. The initial value of the KLa for

this process was set as the experimental value of KLaIntrusion.

The model was run with a step size of 0.01 and 150 steps, resulting in a total modelling time of

1.5 hours. Based on the data provided per experiment, the model simulated the ideal DO

concentration change in time in the Speece cone and the MBR compartments. Later, a new KLa

value was calculated using the parameter estimation tool.

For the experiments with process AS water, the endogenous respiration was included in the

model as a dynamic process called Endo_resp with a stoichiometric coefficient of -1, activated

for the MBR compartment. A Formula variable was created to insert the endogenous respiration

as a constant value using the value obtained during the experiments; this variable was used as

the rate of the Endo_resp process.

3.3.3. Alfa factor (α) For the numerical integration method, the α was calculated using Equation 9 (Section 2.1.1)

according to their respective KLa_Transference20 results. For the AQUASIM results, the α was

calculated using the calculated KLa value.

3.4. Analytical procedure

The analysis used to measure MLSS concentration in Phase 1 of the experiments in process AS

water was the “Standard Method 2540D for Total Suspended Solids dried at 103-105ºC” (Rice

et al., 2012). And, to measure MLVSS concentration in Phase 3, the “Standard Method 2540E

Fixed and volatile solids ignited at 550ºC”(Rice et al., 2012) was used. These analysis were

both performed at the Harnaschpolder WWTP laboratory.

Methodology 49

Table 3-8. Variables used in the AQUASIM model for the MBR and Speece cone system

Name Description Type of Variable Value Unit

Q Flow (Discharge) Program - L/h

Qin Inflow Formula 0.1 (constant) L/h

C_DO Calculated DO concentration Dynamic Volume State - mg/L

C_DOin Initial DO concentration Formula Dependent on oxygen saturation of the experiment mg/L

Qrec Recirculation flow Formula 3000 or 6000 depending on treatment L/h

t Time Program - h

C_DO_XX Measured concentration of DO

of experiment in time

Real list Measured values of DO in MBR per unit of time mg/L*h

KLa Overall oxygen transfer

coefficient

Constant Obtained experimental value according to mixing

conditions

1/h

C_DO_Ini Theoretical DO concentration Constant 0 mg/L

Table 3-9. Compartments used in the AQUASIM model for the MBR and Speece cone system

Name Description Type of

compartment Volume (L) Variables Processes

Initial

condition Input

Cone Speece cone Mixed reactor 80 (constant) C_DO

C_DOin

Oxygen

transfer

C_DO_XX Water inflow: Qin,

Loading: Qin*C_DOin

MBR Membrane

Bioreactor

Mixed reactor 920

(constant)

C_DO

C_DOin

C_DO_Ini -

Help

reactor

Help reactor Mixed reactor 100

(constant)

- - -

Table 3-10. Links used in the AQUASIM model for the MBR and Speece cone system

Name Description Type of link Bifurcations Water flow

C_HR Cone to help reactor Advective Rec Qrec

MBR_C MBR to cone Advective - -

Results and Discussion 50

CHAPTER 4

Results and Discussion

4.1. Evaluation of the Speece cone performance in clean water

4.1.1. Dissolved oxygen concentration gradient in the MBR aeration tank Two experiments were performed to evaluate the mixing conditions in the MBR aeration tank.

The results of Experiment 1 using a mobile DO probe placed at different locations and at

different depths at each location are illustrated in Figure 4-1. This figure shows the absolute

differences between the measurements of the fixed DO probe and the mobile DO probe at

different locations and depths.

Even though some differences between locations were expected due to the movement of the

DO probes and depth of the probes, there were absolute differences as high as 7.12 mgO2/L (in

Location 1.2 at 100 cm). In general, the results of Figure 4-1 suggest that the mixing conditions

and DO probe location used for the preliminary evaluation experiments was not optimal.

Figure 4-1. Mean absolute differences in dissolved oxygen concentrations between a mobile probe with different depths and

locations vs fixed probe at the middle of the tank at 30 cm depth

The results of Experiment 2 using three fixed DO probes placed at different depths and locations

and using low and high mixing conditions are illustrated in Figure 4-2. These results confirmed

that at the high mixing conditions the absolute difference in DO concentration between

locations and depths were reduced; the DO concentration had no noticeable change at the three

0.0

1.0

2.0

3.0

4.0

5.0

6.0

7.0

8.0

100 80 60 40 20

DO

co

nce

ntr

atio

n (

mgO

2/L

)

Depth (cm)

Location 1.1

Location 1.2

Location 1.3

Location 1.4

Location 1.5


locations for the first 35 minutes of the experiment and between the ranges of 0-10 mgO2/L. As

explained in Section 3.3.1, the Speece cone oxygen transfer performance was measured until

75% of the Cs, which was usually below 10 mgO2/L. Therefore, through these experiments it

was confirmed that the addition of the recirculation pumps improved the mixing conditions

during the range of interest. Additionally, the Table 4-1 demonstrates that there was a mean

reduction in the absolute difference of the DO concentration in 51% between the low and high

mixing conditions.

Figure 4-2. Dissolved oxygen concentration measurements for Experiment 2 at different locations, depths and mixing

conditions: (a) Low mixing and (b) High mixing

0.0

5.0

10.0

15.0

20.0

25.0

30.0

0 0.5 1 1.5

DO

co

nce

ntr

atio

n (

mg/L

)

Time (h)

Location 2.1

Location 2.2

Location 2.3

0.0

5.0

10.0

15.0

20.0

25.0

30.0

0 0.5 1 1.5

DO

co

nce

ntr

atio

n (

mg/L

)

Time (h)

Location 2.1

Location 2.2

Location 2.3

(4-2.a)

(4-2.b)


Table 4-1. Absolute mean differences between dissolved oxygen concentration measurements for Experiment 2 at different

locations, depths and mixing conditions

Locations

Absolute mean difference (mg/L)

Difference Low mixing

conditions

High mixing

conditions

Location 2.1 vs Location 2.2 0.59 0.34 58%



Mean 0.92 1.92 51%

4.1.2. Intrusion experiments Intrusion was measured at both mixing conditions, i.e. only with the recirculation flow pump

for preliminary evaluation tests and two extra mixing pumps for the evaluation of the optimal

parameters. The Figure 4-3 shows the results of these intrusion experiments.

Figure 4-3. Change in dissolved oxygen concentration level due to intrusion in (a) low mixing conditions (preliminary

evaluation) and (b) high mixing conditions (evaluation of optimal parameters)

0.0

0.5

1.0

1.5

2.0

2.5

3.0

0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0

DO

co

nce

ntr

atio

n (

mg/L

)

Time (h)

100 cm

50 cm

25 cm

Cone outlet

0.0

1.0

2.0

3.0

4.0

5.0

6.0

7.0

0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0

DO

co

ncn

etra

tio

n (

mg/L

)

Time (h)

100 cm

75 cm

50 cm

25 cm

Cone outlet

(4-3.a)

(4-3.b)


As intrusion is a mass transfer process, it can be described with the two-film theory equation

(Equation 5). From the results of the intrusion experiments, the value of KLaIntrusion was

calculated. The KLaIntrusion value for the low mixing conditions equals to 0.03105 1/h and for

the high mixing conditions equals to 0.05541 1/h. The KLaIntrusion value in the test water

increased by 44% from one mixing condition to the other, demonstrating the influence of

mixing conditions on KLaIntrusion values. Intrusion can be however negligible compared to

resulting oxygen transfer performance, as demonstrated in further results.

4.1.3. Preliminary evaluation of the operational parameters on the performance of the Speece cone

From the experiments in the preliminary evaluation, the effect of pressure and inlet velocity on

the oxygen transfer performance of the Speece cone can be evaluated.

- Effect of pressure. The oxygen transfer efficiency parameters were evaluated at 2, 10

and 15 psi at 6 m3/h of recirculation flow, and 10, 25 and 40 psi at 3 m3/h. Inlet velocities

were constant at each recirculation flow level, 1.2 and 0.6 m/s for 6 and 3 m3/h

respectively. Figure 4-4 shows the results of the pressure effects on SOTR for both

recirculation flows.

As observed in Figure 4-4, there is an increasing tendency of the SOTR as function of

the pressure level and HPO flow; that is, at higher pressure in the cone and higher HPO

flow, higher SOTR. The ability to change the operation pressure is one of the main

features of the Speece cone as a closed system. This together with the use of HPO

instead of air, can enhance the oxygen gas dissolution process.

As shown in Figure 4-4.a, the relation between pressure and SOTR is less evident at 6

m3/h of recirculation flow, where both 15 and 10 psi deliver similar amount of DO in

the system. However, there is a substantial decrease in SOTR of around 55% between

these two pressure levels and minimum pressure, i.e. 2 psi.

Additionally, it can be noticed that the maximum SOTR reached for both recirculation

flows is close to 4 kgO2/d. In the case of the higher recirculation flowrate and higher

HPO flow (i.e. 6 m3/h, 55% HPO flow, at 10 psi) there is a maximum SOTR of 3.8±0.2

kgO2/d. A very similar value of 3.6±0.1 kgO2/d was observed for the lower recirculation

flowrate (3 m3/h), however this was achieved in a different combination of the

operational parameters, i.e. 40% HPO flow and 40 psi. When the system is run at

minimal conditions, i.e. 5% of HPO flow, the SOTR for both recirculation flows and

any pressure level is similar; that is, 0.3±0.2 kgO2/d and 0.4±0.1 kgO2/d for 3 m3/h and

6 m3/h respectively.


Figure 4-4. Influence of low, medium, and high pressure levels on standard oxygen transfer rate (SOTR) at two recirculation

flows: (a) 6 m3/h, and (b) 3 m3/h (mean±SD) found during preliminary evaluation experiments in clean water

From these results, it can be suggested that maximum SOTR is mostly dependent on the

amount of HPO injected into the system, rather than on the pressure or recirculation

flowrate. This was also found by Ashley et al. (2014) (Figure 2-10) which can be

explained by a higher interfacial area of the gas bubbles at higher HPO flows, as well

as higher turbulence caused by the amount of gas injected decreasing the films thickness

and favouring the gas mass transfer process into the liquid.

Additionally, the resulting maximum SOTR at these conditions seem to be significantly

lower than those SOTR obtained in other studies. For instance, the experiment at 3 m3/h

and 40% HPO flow where a SOTR of 3.6±0.1 kgO2/d was observed, can be also

identified as 4.8% oxygen flowrate/water flowrate. The maximum SOTR in a Speece

cone found by Ashley et al. (2014) at 5% of oxygen flowrate/water flowrate was of 13.2

kgO2/d (Figure 2-10), however this important difference can be related to a higher inlet

velocity used (i.e. 5.7 m/s) or the way these authors calculated SOTR, which is based

on the KLa20 value (see explanation later).

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

0% 10% 20% 30% 40% 50% 60%

SO

TR

(kgO

2/d

)

HPO flow relative to maximum oxygen delivery capacity rate (%)

15 psi

10 psi

2 psi

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

0% 10% 20% 30% 40% 50% 60%

SO

TR

(kgO

2/d

)


40 psi

25 psi

10 psi

(4-4.a)

(4-4.b)


Figure 4-5 shows the results of the pressure effects on SOTE for both recirculation

flows. Pressure effects on SOTE are less noticeable than on SOTR. For the experiments

at 6 m3/h (Figure 4-5.a), it can be noticed that maximum and minimum pressure levels

seem less efficient than medium pressure, which provide the highest values for SOTE

for the majority of HPO flows. For the experiments at 3 m3/h (Figure 4-5.b), the higher

pressure level seems more efficient, but the difference between pressure levels is less

evident.

Figure 4-5. Influence of low, medium, and high pressure levels on standard oxygen transfer efficiency (SOTE) at two

recirculation flows: (a) 6 m3/h, and (b) 3 m3/h (mean±SD) found during preliminary evaluation experiments in clean water

In general terms, it can be noticed that the obtained SOTE for 3 m3/h are lower than

those for 6 m3/h, ranging from 41-82% compared to 53-96% respectively. Note the these

experiments also differed in inlet velocities, which can influence on the higher

performance of the experiments at 6 m3/h. Maximum SOTE values are not directly

related to higher HPO flows (as in SOTR) suggesting an inverse relation between these

two parameters; the higher HPO flows, the lower SOTE in both recirculation flows.

Therefore, even though at higher HPO flows there is a higher SOTR, the efficiency is

0%

20%

40%

60%

80%

100%

120%

0% 10% 20% 30% 40% 50% 60%

SO

TE

(%

)


15 psi

10 psi

2 psi

0%

20%

40%

60%

80%

100%

120%

0% 10% 20% 30% 40% 50% 60%

SO

TE

(%

)


40 psi

25 psi

10 psi

(4-5.a)

(4-5.b)


meaningfully affected by oxygen escape to the atmosphere or oxygen accumulation

inside the cone.

These same results were observed by Ashley et al. (2014), where at higher oxygen

flowrate/water flowrates, lower SOTE results were observed (Figure 2-10). These

authors relate this phenomena to the formation of a gas pocket, or accumulation of

bubbles within the Speece cone, and its subsequent release which could represent an

important wastage of HPO gas. Also, McGinnis and Little (1998) predicted this

behaviour using a bubble dynamic model, and stated that large bubbles will dissolve

slower and will probably accumulate within the cone.

A maximum SOTE of 96% was achieved at 17% of HPO flow and 10 psi for 6 m3/h

experiments, and a maximum SOTE of 82% was achieved at 20% of HPO flow and 40

psi for 3 m3/h experiments. Below and above these HPO flows, the SOTE have a

decreasing tendency, which might suggest an efficiency threshold or a breakpoint in

which nearly all the injected HPO will be dissolved into the system minimizing the

wastage. However, according to the results found by Ashley et al. (2014), the SOTE

seems to have a decreasing, almost linear relation as a function of HPO flows.

Therefore, the low SOTE values found in these experiments at HPO flows, could also

entail that they are not sufficient to provide enough oxygen to the system.

Even though the system was not running at optimal conditions, i.e. low mixing and low

inlet velocities, the maximum SOTE found were at 96% and 82% for 6 m3/h and 3 m3/h

respectively. These values of SOTE are much higher than those found in diffused

aeration systems using air as oxygen source. In addition, the SOTE are not affected by

submergence. As observed in Table 2-3, maximum SOTE are in the range of 35-38% at

a 4.6 m submergence using ceramic domes as porous diffusers, which are 54% and 60%

less than the maximum SOTE values of the Speece cone at 3 m3/h and 6 m3/h

respectively.

Figure 4-6 shows the results of the pressure effects on KLa_Transference20 and KLa20 for

both recirculation flows. These results suggests that pressure does not have a great

influence on KLa_Transference20. From these results, it is confirmed again that the HPO

flow is the main factor that influences the oxygen transfer coefficients. There is an

increasing tendency of KLa_Transference20 as a function of HPO flows suggesting that the

overall oxygen mass transfer process is faster at higher presence of oxygen in the

system. However, this also suggests that the oxygen can escape faster to the atmosphere

at higher HPO flows. Ashley et al. (2014) explains this phenomena with 2 reasons:

o At high HPO flows there is a higher interfacial area of gas bubbles, which will

increase the specific surface (a) through were diffusion occurs;

o There is higher turbulence which will promote faster renewal of the liquid

surface, therefore increasing the liquid film coefficient (KL).


Figure 4-6. Influence of low, medium, and high pressure levels on overall oxygen transfer efficiency corrected by

temperature (KLa_Transference20) at two recirculation flows: (a) 6 m3/h, and (b) 3 m3/h (mean±SD) found during

preliminary evaluation experiments in clean water

The obtained KLa_Transference20 values in these experiments are significantly different from

those obtained by Ashley et al. (2014). For instance, at 5% oxygen flowrate/water

flowrate, these authors obtained a KLa20 value of approximately 51 1/h, and with the

preliminary evaluation experiments at 4.8%, the KLa_Transference20 obtained was of 0.98

1/h. These might be explained by a higher inlet velocity used, i.e. 5.6 m/s, compared to

the one used in these experiments, 1.2 and 0.6 m/s. It could be also explained by the

data analysis method employed by these authors. Ashley et al. (2014)used Equation 28

to calculate KLa using Cs as the dissolved oxygen saturation concentration at

environmental conditions. Using Cs at environmental conditions for this calculation,

disregards the fact that when using HPO, the partial pressure of the gas is increased by

a factor of 4.7 (-), which in turn will decrease the KLa value, since the major driving

force for the mass transfer process is the increased partial pressure in the gas phase.

0.0

0.5

1.0

1.5

2.0

2.5

0% 10% 20% 30% 40% 50% 60%

KLa _

Tra

nsf

eren

ce20

(1/h

)


15 psi

10 psi

2 psi

0.0

0.5

1.0

1.5

0% 10% 20% 30% 40% 50% 60%

KLa _

Tra

nsf

eren

ce20

(1/h

)


40 psi

25 psi

10 psi

(4-6.b)

(4-6.a)


𝐾𝐿𝑎 =ln

𝐶𝑠 − 𝐶1

𝐶𝑠 − 𝐶2

𝑡2 − 𝑡1

(Equation 28)

Where:

C1 = Dissolved oxygen concentration at point 1 (mg/L)

C2 = Dissolved oxygen concentration at point 2 (mg/L)

t1 = Time at point 1 (h)

t2 = Time at point 2 (h)

Figure 4-7 shows the results of the pressure effects on the theoretical maximum SOTR

(Max_SOTR) for both recirculation flows. As expected, the Max_SOTR results were

higher than SOTR as it is a function of the value of the CsHPO and the found

KLa_Transference20.

Figure 4-7. Influence of low, medium, and high pressure levels on the theoretical maximum standard oxygen transfer rate

(Max_SOTR) at two recirculation flows: (a) 6 m3/h, and (b) 3 m3/h (mean±SD) found during preliminary evaluation

experiments in clean water

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

4.5

0% 10% 20% 30% 40% 50% 60%

Max

_S

OT

R (

kgO

2/d

)


15 psi

10 psi

2 psi

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

0% 5% 10% 15% 20% 25% 30% 35% 40% 45%

Max

_S

OT

R (

kgO

2/d

)


40 psi

25 psi

10 psi

(4-7.a)

(4-7.b)


Figure 4-7 confirms that even though the KLa_Transference20 values found were not as high

as those ones found by other studies, the Max_SOTR of the Speece is still considerably

high due to the influence of the CsHPO. Additionally, the SOTR values compared to those

obtained when calculating Max_SOTR are fairly similar having only small differences

as shown in Table 4-2. This results suggests that the Speece cone at these experimental

conditions, is almost achieving the maximum capacity of oxygen transfer that it can

theoretically deliver.

Table 4-2. Comparison between mean standard oxygen transfer rates (SOTR) and their theoretical maximum found during


Recirculation

flow (m3/h)

Pressure

(psi)

Mean SOTR

(kgO2/d)

Mean Max_SOTR

(kgO2/d)

SOTR compared to

Max_SOTR (%)

6 15 0.48 0.50 96% ± 0.9%

10 0.41 0.44 95% ± 0.6%

2 0.33 0.36 93% ± 1.0%

3 40 0.50 0.51 98% ± 0.7%

25 0.20 0.21 97% ± 0.4%

10 0.20 0.22 95% ± 0.2%

- Effect of inlet velocity. The oxygen transfer efficiency parameters were evaluated at

low and high inlet velocity levels, i.e. 1.2 m/s and 0.6 m/s (40% and 20% of

recommended value respectively) with a constant pressure of 10 psi. Note that the

retention time is not the same for the experiments, having 36 seconds of retention time

at 6 m3/h and 72 seconds at 3 m3/h.

Figure 4-8 shows the results of the inlet velocity effects on SOTR. The inlet velocity at

this mixing conditions seems to have a greater influence on SOTR than pressure; e.g. at

around 50% of HPO flow, the SOTR increases by 68% and at approximately 5% of

HPO it increases by 50% from low to high inlet velocity. Similarly, notable differences

in KLa_Transference20 values between tested inlet velocities are observed in Figure 4-10.

These results are in fact, contrary to those found by Ashley et al. (2014), who concluded

that “the rate of change of KLa20, SOTR and SAE became slower as the inlet water

velocity increased.”

Nonetheless, Figure 4-9 shows that inlet velocity does not influence strongly on SOTE;

only at the lowest HPO flows (i.e. 5% and 10%) a higher SOTR was obtained for the

higher inlet velocity, but below these HPO flows there was not a clear difference

between inlet velocities. This figure also confirms that the SOTE is mainly determined

by the HPO flows, having an inverse relation, that is, the lower the HPO flows the higher

the SOTE.


Figure 4-8. Influence of inlet velocity on standard oxygen transfer rate (SOTR) at constant pressure (mean±SD) found during


Figure 4-9. Influence of inlet velocity on standard oxygen transfer efficiency (SOTE) at constant pressure (mean±SD) found

during preliminary evaluation experiments in clean water

Figure 4-10. Influence of inlet velocity on overall oxygen transfer efficiency corrected by temperature (KLa_Transference20) at

constant pressure (mean±SD) found during preliminary evaluation experiments in clean water

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

4.5

0% 10% 20% 30% 40% 50% 60%

SO

TR

(kgO

2/d

)


1.2 m/s, 36 s

0.6 m/s, 72 s

0%

20%

40%

60%

80%

100%

0% 10% 20% 30% 40% 50% 60%

SO

TE

(%

)


1.2 m/s, 36 s

0.6 m/s, 72 s

0.0

0.5

1.0

1.5

2.0

2.5

0% 10% 20% 30% 40% 50% 60%

KLa _

Tra

nsf

eren

ce2

0(1

/h)


1.2 m/s, 36 seconds

0.6 m/s, 72 seconds


Figure 4-11 shows the results of the inlet velocity effects on the Max_SOTR for both

recirculation flows. Similar to previous results, the Max_SOTR is slightly higher than

the SOTR values, which is expected.

Figure 4-11. Influence of inlet velocity on the theoretical maximum standard oxygen transfer rate (Max_SOTR) at constant

pressure (mean±SD) found during preliminary evaluation experiments in clean water

4.1.4. Evaluation of the operational parameters on the performance of the Speece cone

From the experiments with optimal conditions, the effect of pressure, inlet velocity and

retention time (based on recirculation flow) on the oxygen transfer efficiency parameters were

evaluated.

- Effect of pressure. The oxygen transfer parameters were evaluated at 10 and 40 psi

keeping all the other variables constant (that is 3 m3/h of recirculation flowrate, at a

constant inlet velocity of 3.43 m/s and a constant retention time of 36 seconds). Figure

4-12 shows the results of the pressure effects on the SOTR. The effect of pressure in

these experiments has the same tendency as in preliminary evaluation experiments,

showing increasing SOTR as a function of pressure and HPO flows. The difference

between pressure levels is higher at higher HPO flows, observing a difference of 2.9±0.9

kgO2/d at 30-40% HPO flows and 0.6±0.2 kgO2/d at 5-20% HPO flows.

As the mixing and inlet velocity conditions were improved during these experiments, a

SOTR higher than preliminary evaluation experiments was expected, however, the

resulting maximum SOTR at 40% of HPO flow and 40 psi is 3.5±0.7 kgO2/d, which is

slightly lower than previous experiment at low mixing conditions. Improved mixing

conditions and inlet velocity seem to have a greater influence at 10 psi experiments, as

all the results obtained at this pressure are slightly higher than in low mixing conditions

and inlet velocity.

Additionally, improved mixing conditions and inlet velocity have a noticeable influence

at lower HPO flows; in the preliminary evaluation experiments the SOTR obtained with

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

4.5

0% 10% 20% 30% 40% 50% 60%

SO

TR

(kgO

2/d

)


1.2 m/s, 36 s

0.6 m/s, 72 s


minimum HPO flows, i.e. 5%, at any pressure level was similar but in these experiments

there was a noticeable increase on SOTR from 10 psi to 40 psi, going from 0.24±0.01

to 0.55±0.01 kgO2/d (56% increase). The results of the evaluation at optimal conditions

also seem to generate a linear tendency of SOTR, which was less evident at preliminary

evaluation experiments.

Figure 4-12. Influence of low and high pressure levels on standard oxygen transfer rate (SOTR) at 3 m3/h and 3.43 m/s

(mean±SD) found during evaluation of optimal conditions in clean water

Figure 4-13 shows the results of the pressure effects on SOTE. Same tendency of results

is observed as in preliminary evaluation experiments, observing an inverse relation, that

is, the higher HPO flows the lower SOTE. Variation between results at 5-20% HPO

flows and at both pressure levels is lower compared to those one at 30-40%.

Additionally, higher SOTE values were observed at these conditions, ranging from 67-

90%, compared to 41-82% at preliminary evaluation experiments.

Figure 4-13. Influence of low and high pressure levels on standard oxygen transfer efficiency (SOTE) at 3 m3/h and 3.43 m/s


Figure 4-14 shows the results of the pressure effects on KLa_Transference20 and KLa20. With

these conditions and at HPO flows higher than 20%, pressure seems to have an inverse

relation with KLa_Transference20; that is, the higher pressure the lower the KLa_Transference20.

0.0

1.0

2.0

3.0

4.0

0% 10% 20% 30% 40%

SO

TR

(kgO

2/d

)


10 psi

40 psi

0%

20%

40%

60%

80%

100%

0% 10% 20% 30% 40%

SO

TE

(%

)


10 psi

40 psi


Figure 4-14. Influence of low and high pressure levels on overall oxygen transfer efficiency corrected by temperature

(KLa_Transference20) at 3 m3/h and 3.43 m/s (mean±SD) found during evaluation of optimal conditions in clean water

When compared with preliminary evaluation experiments, the KLa_Transference20 values

obtained at HPO flows of 5-20% in the evaluation of optimal conditions experiments

have no notable difference with those obtained at preliminary evaluation experiments,

but at higher HPO flow (30-40%) the KLa_Transference20 values increase by 60%. However,

other parameter, such as higher inlet velocity, can be influencing these results.

Figure 4-15 shows the results of the pressure effects on the Max_SOTR. Similar results

were obtained when comparing SOTR and their Max_SOTR, having high similarities

between values as shown in Table 4-3. These results are fairly similar than those found

in the preliminary evaluation experiments, showing that the Speece cone constantly

performs closer to its maximum theoretical oxygen delivery capacity no matter the

operational conditions.

Figure 4-15. Influence of low and high pressure levels on theoretical maximum standard oxygen transfer rate (Max_SOTR)

at 3 m3/h and 3.43 m/s (mean±SD) found during evaluation of optimal conditions in clean water

0.0

0.5

1.0

1.5

0% 10% 20% 30% 40%

KLa _

Tra

nsf

eren

ce20

(1/h

)


10 psi

40 psi

0.0

1.0

2.0

3.0

4.0

5.0

0% 5% 10% 15% 20% 25% 30% 35% 40% 45%

Max

_S

OT

R (

kgO

2/d

)


10 psi

40 psi


Table 4-3. Comparison between mean standard oxygen transfer rates (SOTR) and their theoretical maximum at 3 m3/h and

3.43 m/s found during evaluation of optimal conditions in clean water

Pressure

(psi)

Mean SOTR

(kgO2/d)

Mean Max_SOTR

(kgO2/d)

SOTR compared to

Max_SOTR (%)

40 1.85 1.89 98%±0.2%

10 0.90 0.94 96%±1.9%

- Effect of inlet velocity. The oxygen transfer performance was evaluated at low and high

inlet velocities, i.e. 1.2 and 3.4 m/s (40% and 111% of recommended value), at 6 m3/h

of recirculation flow with a constant pressure of 10 psi. Figure 4-16 shows the results

of the inlet velocity effects on SOTR. Even though an influent velocity higher than the

recommended value was used, the obtained SOTR was not significantly different from

the obtained at lower inlet velocity, suggesting at these conditions, influent velocity is

not an influential factor on the SOTR. These results are opposite from those found in

the preliminary evaluation experiments (Figure 4-8), however the recirculation flow

used in the preliminary evaluation experiments was 3 m3/h. This might suggest that

lower recirculation flows are more sensitive to changes in operational parameters, than

higher recirculation flows.

Figure 4-16. Influence of low and high inlet velocities on standard oxygen transfer rate (SOTR) at 6 m3/h and 3.4 m/s


Figure 4-17 shows the results of the inlet velocity effects on SOTE. According to this

figure, it seems that higher inlet velocities and low HPO flows (5-20%) result in higher

SOTE values. there is a positively influence on obtained SOTE values, but mainly at.

SOTEs close to 100% are reached at these conditions. Similar results were found by

Ashley et al. (2014) which concluded that at higher inlet velocities, higher SOTE values.

0.0

1.0

2.0

3.0

4.0

5.0

6.0

0% 10% 20% 30% 40% 50% 60% 70% 80%

SO

TR

(kgO

2/d

)


3.4 m/s

1.2 m/s


Figure 4-17. Influence of low and high inlet velocities on standard oxygen transfer efficiency (SOTE) at 6 m3/h and 3.4 m/s


Figure 4-18 shows the results of the influence inlet velocity in the KLa_Transference20. An

increase in the KLa_Transference20 value is observed at HPO flows higher than 20%,

suggesting that there is relation between higher inlet velocity and the KLa_Transference20.

These results are similar to those found by Ashley et al. (2014), who concluded that the

KLa values increased with increased values of inlet velocities.

Figure 4-18. Influence of low and high inlet velocities on overall oxygen transfer efficiency corrected by temperature

(KLa_Transference20) at 6 m3/h and 10 psi (mean±SD) found during evaluation of optimal conditions in clean water

Figure 4-19 shows the results of the pressure effects on the Max_SOTR found at 6 m3/h

of recirculation flow and 3.4 m/s of inlet velocity. Similar to previous results, the

obtained SOTR values in these experiments almost achieved their calculated

Max_SOTR (Table 4-4).

0%

20%

40%

60%

80%

100%

120%

140%

0% 20% 40% 60% 80%

SO

TE

(%

)


3.4 m/s

1.2 m/s

0.0

1.0

2.0

3.0

4.0

5.0

0% 20% 40% 60% 80%

KLa _

Tra

nsf

eren

ce20

(1/h

)


3.4 m/s

1.2 m/s


Figure 4-19. Influence of low and high inlet velocities on the theoretical maximum standard oxygen transfer rate (SOTR) at 6

m3/h and 3.4 m/s (mean±SD) found during evaluation of optimal conditions in clean water


10 psi found during evaluation of optimal conditions in clean water

Inlet velocity

(m/s)

Mean SOTR

(kgO2/d)

Mean Max_SOTR

(kgO2/d)

SOTR compared to

Max_SOTR (%)

3.4 2.63 2.78 95%±4.7%

1.2 0.41 0.44 95%±0.6%

- Effect of retention time. The oxygen transfer performance was evaluated at two

different retention times, namely 36 and 72 seconds, with a constant pressure of 10 psi

and 3.4±0.02 m/s of inlet velocity. Figure 4-20 shows the results of the retention time

effects on SOTR. As observed in the figure, there is an inverse effect of retention time

on SOTR values; the higher the retention time, the lower the SOTR. These results are

opposite of expected, as it was assumed that at higher retention times there was a higher

chance for oxygen to dissolve into the liquid within the Speece cone.

As the retention time depends on the recirculation flowrates, these results confirm that

recirculation flowrates have a great influence in the oxygen transfer performance of the

Speece cone, as shown in Figure 4-4.a vs Figure 4-4.b and Figure 4-8.

Figure 4-21 shows the results of the retention time effects on SOTE. The effect of

retention time on SOTE seems to have a similar tendency as SOTR, that is, the lower

the retention time the higher the SOTE; however this relation is only notorious at HPO

flows of 10% and 20%. Additionally, these results show the same tendency as in

preliminary evaluation experiments, but an almost linear negative relation of SOTE as

a function of HPO flow is observed at higher retention time.

0.0

1.0

2.0

3.0

4.0

5.0

6.0

7.0

0% 20% 40% 60% 80% 100% 120%

Max

_S

OT

R (

kgO

2/d

)


3.4 m/s

1.2 m/s


Figure 4-20. Influence of retention time on standard oxygen transfer rate (SOTR) at 10 psi and 3.4 m/s (mean±SD) found

during evaluation of optimal conditions in clean water

Figure 4-21. Influence of retention time on standard oxygen transfer efficiency (SOTE) at 10 psi and 3.4 m/s (mean±SD)

found during evaluation of optimal conditions in clean water

Figure 4-22 shows the influence of retention time on KLa_Transference20. Similar to the

results of SOTR, retention time seems to have an inverse effect on KLa_Transference20

values; the higher the retention time, the lower the KLa_Transference20. As retention time is

influenced by recirculation flow, it might be suggested that the high KLa_Transference20

values are achieved at shorter retention times due to higher turbulence occurring at high

recirculation flows.

Figure 4-23 shows the results of the inlet pressure effects on the Max_SOTR found at10

psi and 3.4 m/s of inlet velocity. Similar to previous results, the obtained SOTR values

are fairly similar to their calculated Max_SOTR.

0.0

1.0

2.0

3.0

4.0

5.0

6.0

0% 10% 20% 30% 40% 50% 60%

SO

TR

(kgO

2/d

)


36 s

72 s

0%

20%

40%

60%

80%

100%

120%

0% 10% 20% 30% 40% 50% 60%

SO

TE

(%

)


36 s

72 s


Figure 4-22. Influence of retention time on overall oxygen transfer efficiency corrected by temperature (KLa_Transference20) at

10 psi and 3.4 m/s (mean±SD) found during evaluation of optimal conditions in clean water

Figure 4-23. Influence of retention time on the theoretical maximum standard oxygen transfer rate (Max_SOTR) at 10 psi

and 3.4 m/s (mean±SD) found during evaluation of optimal conditions in clean water

4.1.5. AQUASIM model for clean water experiments

The AQUASIM model was used to estimate the value of KLa20 which represents the overall

oxygen transfer efficiency in the MBR aerobic tank including the oxygen transfer process from

intrusion and from the Speece cone.

The Figure 4-24, Figure 4-25, and Figure 4-26 show the estimation of the KLa made with the

AQUASIM program for the preliminary evaluation experiments. As observed in all Figures,

the KLa20 values obtained using AQUASIM program are significantly different from the

0.0

1.0

2.0

3.0

0% 10% 20% 30% 40% 50% 60%

KLa _

Tra

nsf

eren

ce20

(1/h

)


36 s

72 s

0.0

1.0

2.0

3.0

4.0

5.0

6.0

0% 10% 20% 30% 40% 50% 60%

Max

_S

OT

R (

kg

O2/d

)


36 s

72 s


KLa_Transference20 obtained with the numerical integration method. For instance, the results shown

in Figure 4-24.a for the experiments at 6 m3/h are between 4 to 11 times higher than those

obtained in the numerical integration method, and the results shown in Figure 4-24.b for the

experiments at 3 m3/h are between 6 to 18 times higher than those obtained in the numerical

integration method.

Additionally, the results of the AQUASIM program reflected a better linear tendency than those

obtained using the numerical integration method. For instance, the results of the experiment at

different inlet velocities (1.2 and 0.6 m/s) and constant pressure (10 psi) shown in Figure 4-25

have a more positive slope, especially for the results of 0.6 m/s, than those obtained with the

numerical integration method. This might suggest that the program might not be taking into

account the potential losses due to oxygen escape as bubbles, and the addition of another

process to reflect these loss was required.

Figure 4-24. Influence of low, medium, and high pressure levels on overall oxygen transfer efficiency corrected by

temperature (KLa20) calculated with AQUASIM program at two recirculation flows: (a) 6 m3/h, and (b) 3 m3/h (mean±SD)

found during the preliminary evaluation experiments in clean water

0.0

2.0

4.0

6.0

8.0

10.0

12.0

14.0

16.0

0% 10% 20% 30% 40% 50% 60%

KLa 2

0(1

/h)


15 psi

10 psi

2 psi

0.0

2.0

4.0

6.0

8.0

10.0

12.0

0% 10% 20% 30% 40% 50% 60%

KLa 2

0(1

/h)


40 psi

25 psi

10 psi

(4-24.b)

(4-24.a)


Figure 4-25. Influence of inlet velocity on overall oxygen transfer efficiency corrected by temperature (KLa20) calculated by

AQUASIM program at 10 psi (mean±SD) found during the preliminary evaluation experiments in clean water

Figure 4-26. Influence of low and high pressure levels on overall oxygen transfer efficiency corrected by temperature (KLa20)

calculated by AQUASIM program at 3 m3/h and 3.43 m/s (mean±SD) found during the preliminary evaluation experiments in

clean water

However, there was a similar tendency of the AQUASIM results of all preliminary evaluation

experiments compared with the numerical integration results, except for the lowest HPO flow

levels (i.e. 5%) were some zero values were generated. This might suggest that this amount of

HPO flow injected was too low to be detected as influential in the oxygen transfer process.

The Figure 4-27 and Figure 4-28 shown the estimation of the KLa made with the AQUASIM

program for the evaluation of optimal conditions experiments. The tendency of the results were

similar to those ones found in the preliminary evaluation experiments, with a significant

increase of the resulting KLa values. Some of these results demonstrate in a better way the

relations between the operational parameter and KLa found in the numerical integration, for

0.0

2.0

4.0

6.0

8.0

10.0

12.0

14.0

16.0

0% 10% 20% 30% 40% 50% 60%

KLa 2

0(1

/h)


1.2 m/s

0.6 m/s

0.0

1.0

2.0

3.0

4.0

5.0

6.0

0% 10% 20% 30% 40%

KLa 2

0(1

/h)


10 psi

40 psi


instance, in Figure 4-27 illustrates better the relation of higher inlet velocity generates higher

the KLa20 than the results found in the numerical integration method.

Figure 4-27. Influence of low and high inlet velocities on overall oxygen transfer efficiency corrected by temperature (KLa20)

calculated by AQUASIM program at 6 m3/h and 10 psi (mean±SD) found during the evaluation of optimal conditions

experiments

Figure 4-28. Influence of retention time on overall oxygen transfer efficiency corrected by temperature (KLa20) calculated by

AQUASIM program at 10 psi and 3.4 m/s in clean water (mean±SD) found during the evaluation of optimal condition

experiments in clean water

According to these results, it is assumed that the AQUASIM model used to analyse the data,

might not be well designed due to these meaningful differences with the results of the numerical

integration model. To improve the AQUASIM model, there might be the need to add more

0.0

2.0

4.0

6.0

8.0

10.0

12.0

14.0

16.0

0% 20% 40% 60% 80%

KLa 2

0(1

/h)


3.4 m/s

1.2 m/s

0.0

2.0

4.0

6.0

8.0

10.0

12.0

14.0

16.0

0% 10% 20% 30% 40% 50% 60%

KLa 2

0(1

/h)


36 s

72 s


processes to the Cone or to the MBR which represent better the specific mass transfer process

within these compartments.

4.2. Evaluation of the Speece cone performance in process activated sludge water

The experiments with process activated sludge (AS) water were performed after filling the

MBR aerobic tank with return activated sludge (RAS) from Harnaschpolder wastewater

treatment plant, and after reaching the desired mixed liquor (MLSS) concentration, i.e. 5 g/L.

The effect of pressure and retention time on the oxygen transfer performance of the Speece

cone were evaluated at different high purity oxygen (HPO) flows.

4.2.1. Phase 1

4.2.1.1 Filling the MBR aerobic tank

The initial measured MLSS concentration of RAS used to fill the tank was 6.2 g/L. Therefore,

the AS in the tank had to be initially diluted with clean water to reach the desired concentration,

i.e. 5 g/L. However, after some experiments were made, there was formation of visible foam

on the surface of the mixed liquor (Figure 4-29). This foam reduced the MLSS concentration

in the AS, additional to some loss of solids when the DO probes were taken out of the tank and

some solids precipitation between the membrane module when the experiment was running.

Consequently, the AS had to be concentrated using the membrane module which extracted

permeate and leaved the retentate in the MBR aerobic tank. During this procedure, air scouring

was applied in order to reduce membrane fouling mitigation. If needed, extra AS was added to

increase the MLSS concentration.

4.2.1.2 MLSS concentration during experiments

Due to loss of solids and sedimentation issues within the membrane module, the MLSS

concentration during the experiments was not maintained at 5 g/L, however mean values are

fairly close to expected. Mean MLSS concentration during the experiments were observed at

4.9±0.8 g/L for the experiments at 3 m3/h and 10 psi, 3.9±0.4 g/L for the experiments at 3 m3/h

and 40 psi, and 4.2±1.1 g/L for the experiments at 6 m3/h and 10 psi. Figure 4-30 shows the

MLSS concentration values obtained at each experiment.


Figure 4-29. Foam formation in the activated sludge after running re-oxygenation experiment

Figure 4-30. Mixed liquor suspended solids (MLSS) concentration in activated sludge for experiments with process activated

sludge water

4.2.2. Phase 3 Due to technical issues with the sludge recirculation pump, not all the expected experiments for

AS were performed as detailed in Table 3-7. Only 14 experiments out of 20 were performed;

all the experiments at 3 m3/h, and 4 experiments at 6 m3/h. However, the maximum flow

achieved with the defective pump using AS was not exactly 6 m3/h; a mean recirculation flow

of 5.7±0.3 m3/h was used for these experiments, which represents 3.2±0.2 m/s of inlet velocity

and 38 seconds of retention time. Additionally, due to incompleteness of the data set, only two

effects on oxygen transfer efficiency parameters could be studied, effect of pressure and effect

of retention time.

4.2.1.3 Endogenous respiration

The mean endogenous respiration (rO2,endo) during the experiments is detailed in Table 4-5. This

table also shows the value of the specific endogenous respiration according to the obtained

MLVSS concentration per experiment, and the ratio of MLVSS compared to MLSS

concentration.

0.0

1.0

2.0

3.0

4.0

5.0

6.0

7.0

0% 10% 20% 30% 40% 50%

ML

SS

co

nce

ntr

atio

n (

g/L

)

HPO flow relative to maximum oxygen delivery capacity (%)

3 m3/h, 10 psi

3 m3/h, 40 psi

5.7 m3/h, 10 psi


Table 4-5. Endogenous respiration, specific endogenous respiration and MLVSS/MLSS ratio results

Recirculation

flow (m3/h)

Pressure

(psi)

Mean

endogenous

respiration

(gO2/L*d)

Mean specific

endogenous

respiration

(gO2/gVSS*d)

Mean ratio

MLVSS/MLSS

3 10 0.28±0.05 0.07 0.88±0.03

40 0.18±0.06 0.05 0.88±0.03

5.7 10 0.42±0.31 0.15 0.87±0.03

4.2.1.4 Oxygen transfer efficiency parameters

The effect of pressure and retention time on the oxygen transfer performance of the Speece

cone were evaluated at different HPO flows.

- Effect of pressure. The oxygen transfer efficiency parameters were evaluated at low

and high levels of pressure at 3 m3/h with a constant inlet velocity of 3.43 m/s (112% of

recommended inlet velocity); these conditions were the same as the optimal conditions

evaluated for clean water. Figure 4-31 shows the results of pressure effects on SOTR in

AS. The pressure effects on SOTR of AS follows the same tendency of positive effect

of pressure as observed in clean water; however, the values obtained are approximately

28% lower. The difference between pressure levels is higher at higher HPO flows,

observing a difference of 0.75±0.8 kgO2/d at 20-40% HPO flows and 0.3±0.2 kgO2/d at

5-10% HPO flows; these differences are less evident than in clean water results.

Similarly to results in clean water, there seems to be a linear relation of SOTR as a

function of HPO flows.

Maximum SOTR in AS was also achieved at 40% of HPO flow and 40 psi, reaching

2.0±0.3 kgO2/d which is 43% lower than the results in clean water.

Figure 4-31. Influence of pressure on standard oxygen transfer rate (SOTR) at 3 m3/h psi and 3.43 m/s in activated sludge

(mean±SD)

0.0

0.5

1.0

1.5

2.0

2.5

0% 10% 20% 30% 40%

SO

TR

(kgO

2/d

)


10 psi

40 psi


Figure 4-32 shows the results of pressure effects on SOTE in AS. In general, the

obtained SOTE follow a negative tendency as a function of HPO flow as shown in

previous results. The influence of pressure seems to be inverse, that is, the lower the

pressure, the higher the SOTE, however this is clearer in at 5%, 30% and 40% of HPO

flows. Also, these results seem to have a higher variability than results in clean water.

Maximum SOTE value in these experiments was achieved at 78% for 5% HPO flow

and 10 psi. This result differs from the one obtained at clean water, were maximum

SOTE value was achieved at higher pressure. In general, SOTE ranges from 52-78% at

10 psi, and from 42-64% at 40 psi.

Figure 4-32. Influence of pressure on standard oxygen transfer efficiency (SOTE) at 3 m3/h psi and 3.43 m/s in activated

sludge (mean±SD).

Figure 4-33 shows the results of the pressure effects on KLa_Transference20 in AS. These

results seem to have the same tendency as shown in clean water results, where at lower

pressure there is higher values of KLa_Transference20 especially at higher HPO flows.

It is important to remark, that the obtained KLa_Transference20 values were influenced by

the temperatures of the AS during the experiment which were notably higher compared

to those in clean water. Mean temperature of the experiments were 30.5±1.2°C at 3 m3/h

and 10 psi, 31.3±1.5°C at 3 m3/h and 40 psi, and 31.2±1.0°C at 5.7 m3/h and 10 psi.

Therefore, when KLa_Transference was corrected by temperature, it decreased considerably;

the KLa_Transference20 values are 23%±2% lower than their KLa_Transference values.

0%

10%

20%

30%

40%

50%

60%

70%

80%

90%

100%

0% 10% 20% 30% 40%

SO

TE

(%

)


10 psi

40 psi


Figure 4-33. Influence of pressure on overall oxygen transfer efficiency corrected by temperature (KLa_Transference20) at 3 m3/h

psi and 3.43 m/s in activated sludge (mean±SD)

Figure 4-34 shows the results of the pressure effects on the Max_SOTR found at 3 m3/h

of recirculation flow and 3.43 m/s of inlet velocity. As expected, the obtained SOTR

values in these experiments were higher than obtained in SOTR, but contrary to the

clean water experiments, the SOTR had higher differences with the Max_SOTR as

shown in Table 4-6. This could be explained by the presence of suspended solids which

affect the mass transfer process of the oxygen into the liquid.

Figure 4-34. Influence of pressure on the theoretical maximum standard oxygen transfer rate (Max_SOTR) at 3 m3/h psi and

3.43 m/s in activated sludge (mean±SD)

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

0% 10% 20% 30% 40%

KLa _

Tra

nsf

eren

ce20 (1

/h)


10 psi

40 psi

0.0

0.5

1.0

1.5

2.0

2.5

3.0

0% 5% 10% 15% 20% 25% 30% 35% 40% 45%

Max

_S

OT

R (

kgO

2/d

)


10 psi

40 psi



3.4 m/s found in process activated sludge water

Pressure

(psi)

Mean SOTR

(kgO2/d)

Mean Max_SOTR

(kgO2/d)

SOTR compared to

Max_SOTR (%)

10 0.66 0.77 70%±3.0%

40 1.22 1.39 60%±7.0%

- Effect of retention time. The oxygen transfer efficiency parameters were evaluated at

two different retention times with a constant pressure of 10 psi and 3.3±0.2 m/s of inlet

velocity; these conditions were slightly different as the optimal conditions tests for clean

water due to technical issues with the sludge recirculation pump. Figure 4-35 shows the

results of the retention time effects on SOTR. These results follow the same tendency

as results in clean water, were shorter retention times result in higher SOTR. Similar

results are found for the effect of retention time in KLa_Transference20 (Figure 4-37). Figure

4-36 shows the results of retention time in SOTE, which has the same tendency as

previous experiments, that is, the higher the HPO flow, the lower the SOTE.

Figure 4-35. Influence of retention time on standard oxygen transfer rate (SOTR) at 10 psi and 3.4 m/s in activated sludge

(mean±SD)

0.0

0.5

1.0

1.5

2.0

2.5

3.0

0% 10% 20% 30% 40%

SO

TR

(kgO

2/d

)


38 s

72 s


Figure 4-36. Influence of retention time on standard oxygen transfer efficiency (SOTE) at 10 psi and 3.4 m/s in activated

sludge (mean±SD)

Figure 4-37. Influence of retention time on overall oxygen transfer efficiency corrected by temperature (KLa_Transference20) at

10 psi and 3.4 m/s in activated sludge (mean±SD)

0%

10%

20%

30%

40%

50%

60%

70%

80%

90%

100%

0% 5% 10% 15% 20% 25% 30% 35% 40% 45%

SO

TE

(%

)


38 s

72 s

0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4

1.6

1.8

0% 5% 10% 15% 20% 25% 30% 35% 40% 45%

KLa _

Tra

nsf

eren

ce20

(1/h

)


38 s

72 s


Figure 4-38. Influence of retention time on the theoretical maximum standard oxygen transfer rate (Max_SOTR) at 10 psi

and 3.4 m/s in activated sludge (mean±SD)

4.2.3. Alfa factor

Figure 4-39, Figure 4-40, and Figure 4-41 show the results of the influence of pressure, inlet

velocity and retention time on the alfa factor of the system. As observed in the Figure 4-39, the

maximum α obtained was 0.96 at 3 m3/h, 40 psi and 20% of HPO flow, followed by 0.95 at 3

m3/h, 10 psi and 5% HPO flow. The maximum value found in these experiments is 38% higher

than the maximum value of fine bubble diffuser (Table 2-1); meaning that the overall oxygen

transfer process happening in this system at these conditions is closer to the one of clean water.

In Figure 4-40, it can be noticed that the lowest α values were obtained at the highest

recirculation flow, except at 30% HPO flow; this α seems to be an outlier because there is an

initial negative tendency as a function of HPO flow however more experiments at higher HPO

flows are required.

Figure 4-39. Effect of pressure on alfa factors (α) at 3 m3/h and 3.43 m/s in activated sludge (mean±SD)

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0% 5% 10% 15% 20% 25% 30% 35% 40% 45%

SO

TR

(kgO

2/d

)


38 s

72 s

0.0

0.2

0.4

0.6

0.8

1.0

0% 10% 20% 30% 40% 50%

α(-

)


3 m3/h, 10 psi

3 m3/h, 40 psi


Figure 4-40. Effect of inlet velocity on alfa factors (α) at 3 m3/h psi and 10 psi in activated sludge (mean±SD)

Figure 4-41. Influence of retention time on the alfa factor at 10 psi and 3.4 m/s in activated sludge (mean±SD)

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0% 5% 10% 15% 20% 25% 30% 35% 40% 45%

α(-

)


38 s

72 s

0.0

2.0

4.0

6.0

8.0

10.0

12.0

0% 5% 10% 15% 20% 25% 30% 35% 40% 45%

α(-

)


38 s

72 s

Conclusions 81

CHAPTER 5

Conclusions

This research was developed to evaluate the oxygen transfer performance of the Speece cone

at different operational conditions both in clean water and activated sludge in a pilot-scale

membrane bioreactor (MBR).

According to the results, it can be concluded that the Speece cone can operate at a maximum

oxygen transfer rate (OTR) in clean water of 3.8±0.2 kgO2/d with the following operational

conditions: 6 m3/h of recirculation flowrate, 10 psi of pressure, 1.2 m/s of inlet velocity, and

55% of high purity oxygen (HPO) flow relative to the maximum oxygen delivery capacity rate

of the Speece cone. The maximum OTR in activated sludge was 2.2±0.3 kgO2/d found at the

following conditions: 5.7 m3/h of recirculation flowrate, 10 psi of pressure, 3.2 m/s of inlet

velocity, and 30% of high purity oxygen (HPO) flow.

During this research it was found that the Speece cone can provide oxygen transfer efficiencies

(OTE) in clean water as high as 96%±8% with the following operational conditions: 6 m3/h of

recirculation flowrate, 10 psi of pressure, 1.2 m/s of inlet velocity, and 17% of HPO flow. This

efficiency is comparable with other high purity oxygen delivery systems already applied in the

wastewater treatment sector, such as I-SO or UNOX which achieve efficiencies as high as 90%.

The maximum OTE of the Speece cone found in this research is much higher from those found

at traditional diffused aeration systems, which achieve only up to 35-38% of OTE.

Additionally, the maximum OTE in activated sludge was 78% found at the following

conditions: 3 m3/h of recirculation flowrate, 10 psi of pressure, 3.43 m/s of inlet velocity, and

5% of high purity oxygen (HPO) flow.

Regarding the overall oxygen transfer coefficient (KLa), it was found that in clean water the

value of this coefficient can go up to 2.1 1/h in the following operational conditions: 6 m3/h of

recirculation flowrate, 10 psi of pressure, 1.2 m/s of inlet velocity, and 55% of HPO flow. For

activated sludge, the maximum KLa was 1.5±0.2 1/h found at the following conditions: 5.7 m3/h

of recirculation flowrate, 10 psi of pressure, 3.2 m/s of inlet velocity, and 30% of high purity

oxygen (HPO) flow.

It can be stated that the oxygen transfer performance of the Speece cone in clean water and in

activated sludge is mainly influenced by the amount of high purity oxygen that is injected into

the system. The oxygen transfer rate (OTR) and the overall oxygen transfer coefficient (KLa)

Conclusions 82

have a positive trend at increasing HPO flows, but the contrary was found for the oxygen

transfer efficiency (OTE), where there is a negative trend at increasing HPO flows.

Additionally, it was found that higher pressure levels and higher inlet velocities positively

influence on the oxygen transfer performance of the Speece cone, mainly on the OTR and in a

less extent on the OTE and KLa. However, the retention time seems to have a negative influence

on the on the oxygen transfer performance of the Speece cone, as at the higher inlet velocities,

the lower the OTR, OTE and KLa.

Finally, it was found that the computational model using AQUASIM did not represent the

Speece cone system as expected, as it generated KLa values significantly higher than those one

calculated by the numerical integration method. Therefore, a more rigorous study of the mass

transfer process happening inside the Cone compartment is recommended in order to improve

the AQUASIM model simulation of the Speece cone system.

References 83

References

Al-Malack, M. H. (2006). Determination of biokinetic coefficients of an immersed membrane bioreactor. Journal of

Membrane Science, 271(1–2), 47-58. doi:http://dx.doi.org/10.1016/j.memsci.2005.07.008

American Society of Civil Engineers. (1988). Aeration: A Wastewater Treatment Process: American Society of Civil

Engineers.

American Society of Civil Engineers. (1992). Measuring of Oxygen Transfer in Clean Water (pp. 51). New York, USA:

American Society of Civil Engineers.

Ashley, K., Fattah, K., Mavinic, D., & Kosari, S. (2014). Analysis of Design Factors Influencing the Oxygen Transfer of a

Pilot-Scale Speece Cone Hypolimnetic Aerator. Journal of Environmental Engineering, 140(3), 04013011.

doi:10.1061/(ASCE)EE.1943-7870.0000789

Ashley, K., Mavinic, D., & Hall, K. (2008). Oxygenation performance of a laboratory-scale Speece Cone hypolimnetic

aerator: preliminary assessment. Canadian Journal of Civil Engineering, 35(7), 663-675. doi:10.1139/L08-011

Barber, T. W., Ashley, K. I., Mavinic, D. S., & Christison, K. (2015). Superoxygenation: analysis of oxygen transfer design

parameters using high-purity oxygen and a pressurized column. Canadian Journal of Civil Engineering, 42(10),

737-746. doi:10.1139/cjce-2015-0037

Barreto, C. (2015). Evaluation of a pilot MBR system operated at high MLSS provided with a Speece cone aeration system as

an alternative for sanitation provision in emergencies. (Master of Science), UNESCO-IHE Institute for Water

Education, The Netherlands. (UWS-SE-2015-15)

Barreto, C. M., Garcia, H. A., Hooijmans, C. M., Herrera, A., & Brdjanovic, D. (2017). Assessing the performance of an

MBR operated at high biomass concentrations. International Biodeterioration & Biodegradation, 119, 528-537.

doi:https://doi.org/10.1016/j.ibiod.2016.10.006

Bewtra, J. K., Nicholas, W. R., & Polkowski, L. B. (1970). Effect of temperature on oxygen transfer in water. Water

Research, 4(1), 115-123. doi:http://dx.doi.org/10.1016/0043-1354(70)90023-0

Blue in Green. SDOX Adaptable aeration. Retrieved from https://www.blueingreen.com/sdox-adaptable-aeration

Boyd, C. E. (1986). A Method for Testing Aerators for Fish Tanks. The Progressive Fish-Culturist, 48(1), 68-70.

doi:10.1577/1548-8640(1986)48<68:AMFTAF>2.0.CO;2

Brepols, C. (2011). Operating Large Scale Membrane Bioreactors for Municipal Wastewater Treatment. Water Intelligence

Online, 9.

Casey, T. J. (1997). Unit treatment processes in water and wastewater engineering. Chichester; New York: Wiley.

Delfluent Services bv. Installations: WWTP Harnaschpolder. Retrieved from http://delfluent.nl/en/plant/wwtp-

harnaschpolder/

EPA. (1973). Oxygen Activated-Sludge Wastewater Treatment Systems: Desing Criteria and Operating Experience (pp. 56).

USA: Environmental Protection Agency: Technology Transfer.

Germain, E., Nelles, F., Drews, A., Pearce, P., Kraume, M., Reid, E., . . . Stephenson, T. (2007). Biomass effects on oxygen

transfer in membrane bioreactors. Water Research, 41(5), 1038-1044.

doi:http://dx.doi.org/10.1016/j.watres.2006.10.020

Gray, N. F. (2004). Biology of Wastewater Treatment.

Gunder, B. (2001). The membrane-coupled activated sludge process in municipal wastewater treatment. Stuttgart, Germany:

Technomic Publishing Co., INC.

Henze, M., van Loosdrecht, M. C. M., Ekama, G. A., & Brdjanovic, D. (2008). Biological wastewater treatment : principles,

modelling and design. London: IWA Pub.

Judd, S. (2007). The MBR book: Principles and applications of membrane bioreactors in water and wastewater treatment.

Amsterdam, The Netherlands: Elsevier Science.

Krampe, J., & Krauth, K. (2003). Oxygen transfer into activated sludge with high MLSS concentrations. Water Science and

Technology, 47(11), 297.

Linde Group. SOLVOX® Family. Retrieved from http://www.linde-

gas.com/en/products_and_supply/water_and_wastewater_solutions/solvox_family/index.html

Loosdrecth, M., Nielsen, P. H., Lopez-Vasquez, C. M., & Brdjanovic, D. (2016). Experimental Methods in Wastewater

Treatment. The Netherlands: IWA Publishing.

http://dx.doi.org/10.1016/j.memsci.2005.07.008

https://doi.org/10.1016/j.ibiod.2016.10.006

http://dx.doi.org/10.1016/0043-1354(70)90023-0

https://www.blueingreen.com/sdox-adaptable-aeration

http://delfluent.nl/en/plant/wwtp-harnaschpolder/

http://delfluent.nl/en/plant/wwtp-harnaschpolder/

http://dx.doi.org/10.1016/j.watres.2006.10.020

http://www.linde-gas.com/en/products_and_supply/water_and_wastewater_solutions/solvox_family/index.html

http://www.linde-gas.com/en/products_and_supply/water_and_wastewater_solutions/solvox_family/index.html

References 84

McGinnis, D. F., & Little, J. C. (1998). Bubble dynamics and oxygen transfer in a speece cone. Water Science and

Technology, 37(2), 285-292. doi:http://dx.doi.org/10.1016/S0273-1223(98)00035-3

Membranes Modules Systems. (2005). Submerged Pulse Membrane Modules: OPERATION MANUAL (pp. 9).

Metcalf, L., Eddy, H. P., & Tchobanoglous, G. (2004). Wastewater engineering : treatment, disposal, and reuse. New York

[etc.]: McGraw-Hill.

Moore, B. C., Cross, B. K., Beutel, M., Dent, S., Preece, E., & Swanson, M. (2012). Newman Lake restoration: A case study

Part III. Hypolimnetic oxygenation. Lake and Reservoir Management, 28(4), 311-327.

doi:10.1080/07438141.2012.738463

Mueller, J. A., Boyle, W. C., & Popel, H. J. (2002). Aeration : principles and practice. Boca Raton: CRC-Press.

Reichert, P. (1998). AQUASIM 2.0 - User Manual: Computer Program for the Identification and Simulation of Aquatic

Systems (pp. 219). Switzerland: Swiss Federal Institute for Environmental Science and Technology (EAWAG).

Rice, E. W., American Public Health, A., & American Water Works, A. (2012). Standard methods for the examination of

water and wastewater. Washington, D.C.: American Public Health Association.

Roš, M. (1993). Respirometry of activated sludge. Lancaster, PA: Technomic Pub. Co.

Smeets, P., Medema, G., & van Dijk, J. (2009). The Dutch secret: how to provide safe driking water without chlroine in the

Netherlands. Drinking Water Engineering and Science, 2, 1-14.

Spanjers, H., Vanrolleghem, P. A., Olsson, G., & Dold, P. L. (1998). Respirometry in control of the activated sludge process

: principles. London: International Association on Water Quality.

Speece, R. E., Madrid, M., & Needham, K. (1971). Downflow Bubble Contact Aeration. Journal of the Sanitary Engineering

Division, 97(4), 433-441.

US Army Corps of Engineers. (2009). Savannah Harbor Reoxygenation Demonstration Project: Part H (pp. 35).

USGS. (2014). Dissolved oxygen solubility tables. Retrieved from https://water.usgs.gov/software/DOTABLES/

von Sperling, M., & International Water Association. (2007). Activated sludge and aerobic biofilm reactors. London: IWA

publishing.

Wagner, M., Cornel, P., & Krause, S. (2002). Efficiency of different aeration systems in full scale membrane bioreactors

(Vol. 2002).

Wang, L. K., Pereira, N. C., Hung, Y.-T., & Shammas, N. K. (2009). Biological Treatment Processes: Humana Press.

Water Environment Federation. (2012). Membrane bioreactors. New York: McGraw-Hill.

Weiss, R. F. (1970). The solubility of nitrogen, oxygen and argon in water and seawater. Deep Sea Research and

Oceanographic Abstracts, 17(4), 721-735. doi:http://dx.doi.org/10.1016/0011-7471(70)90037-9

Winkler, M. A. (1981). Biological treatment of waste-water. Chichester, England; New York: Ellis Horwood ; Halsted Press.

WTW. Galvanic dissolved oxygen sensors (universal). Retrieved from https://www.wtw.com/en/products-

categories/sensors-technology/conventional-sensors-lab/galvanic-oxygen-sensors-universal.html

ZENA. (2017). Gallery: Membrane. Retrieved from http://www.zena-membranes.cz/index.php/gallery/membrane-gallery

http://dx.doi.org/10.1016/S0273-1223(98)00035-3

https://water.usgs.gov/software/DOTABLES/

http://dx.doi.org/10.1016/0011-7471(70)90037-9

https://www.wtw.com/en/products-categories/sensors-technology/conventional-sensors-lab/galvanic-oxygen-sensors-universal.html

https://www.wtw.com/en/products-categories/sensors-technology/conventional-sensors-lab/galvanic-oxygen-sensors-universal.html

http://www.zena-membranes.cz/index.php/gallery/membrane-gallery

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