600282052
SIMPLIFIED INJECTION OF OXYGEN GAS INTO AN
ACTIVATED SLUDGE PROCESS
by
Lloyd D. Hedenland
Las Virgenes Municipal Water District
Calabasas, California 91302
and
Ralph L. Wagner
VTN Consolidated, Inc.
San Bernardino, California 92402
Grant No. S802356
Project Officer
Richard C. Brenner
Wastewater Research Division
Municipal Environmental Research Laboratory
Cincinnati, Ohio 45268
MUNICIPAL ENVIRONMENTAL RESEARCH LABORATORY
OFFICE OF RESEARCH AND DEVELOPMENT
U.S. ENVIRONMENTAL PROTECTION AGENCY
CINCINNATI, OHIO 45268
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DISCLAIMER
This report has been reviewed by the Municipal Environmental Research
Laboratory, U.S. Environmental Protection Agency, and approved for publication.
Approval does not signify that the contents necessarily reflect the views and
policies of the U.S. Environmental Protection Agency, nor does mention of trade
names or commercial products constitute endorsement or recommendation for use.
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FOREWORD
The Environmental Protection Agency was created because of increasing
public and government concern about the dangers of pollution to the health
and welfare of the American people. Noxious air, foul water, and spoiled
land are tragic testimonies to the deterioration of our natural environment.
The complexity of that environment and the interplay between its components
require a concentrated and integrated attack on the problem.
Research and development is that necessary first step in problem solu-
tion; it involves defining the problem, measuring its impact, and searching
for solutions. The Municipal Environmental Research Laboratory develops new
and improved technology and systems to prevent, treat, and manage wastewater
and solid and hazardous waste pollutant discharges from municipal and com-
munity sources; to preserve and treat public drinking water supplies; and to
minimize the adverse economic, social, health, and aesthetic effects of
pollution. This publication is one of the products of that research and
provides a most vital communications link between the researcher and the
user community.
This report describes a feasibility study of a novel,covered tank,
oxygen activated sludge treatment concept. An inflated polyvinyl dome was
utilized to prevent the enriched-oxygen off gas from escaping to the atmos-
phere. Existing air aeration equipment was employed to recirculate the off
gas through the mixed liquor. The information documented in this report
should be of interest to design engineers and municipal officials who are
considering techniques for converting air activated sludge systems to oxygen
activated sludge systems.
Francis T. Mayo, Director
Municipal Environmental Research
Laboratory
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ABSTRACT
The concept of using a single-stage, covered, activated sludge aeration
basin and conventional aeration equipment for high purity oxygen injection
was devised in 1970 by the Cosmodyne Division of Cordon International. The
concept was termed the Simplex process and is investigated in this report.
The Las Virgenes Municipal Water District conducted this study at their
Tapia Water Reclamation Facility in Calabasas, California. The objectives
of the study were:
1. to determine the practicality and cost of using this
simplified method of high purity oxygen injection as an
upgrading tool.
2. to determine if aeration basin throughput could be at
least doubled,
3. to determine the extent of waste sludge volume reduction, and
4. to compare the results with data collected using conventional
air aeration.
Conversion of the existing facilities involved covering an aeration
basin with a heavy-duty plastic cover. The air blower system was modified
to draw atmosphere from under the plastic cover and recirculate it thorugh
the aeration basin mixed liquor using the existing coarse bubble air diffuser
system. Pure oxygen from a tanker truck was vaporized and fed to the mixed
liquor through a new fine bubble diffuser system (Saran wrapped tubes),
which was installed at the head end of the aeration basin. A portion of the
atmosphere under the plastic cover was bled off to remove carbon dioxide.
Primary effluent was added to the aeration basin using step feed. Mixed
liquor was clarified using conventional secondary clarifiers from which
settled sludge was withdrawn for recycling and wasting.
The following conclusions were reached from this study:
1. The oxygen dispersing equipment required more operating
time than conventional air compressors.
2. The process performed best at higher BOD loadings.
3. Nitrification was easily accomplished.
4. Low detention times, 1.7 hrs, produced nearly 90 percent
BOD removal.
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5. Relatively high SVI values were produced.
6. Solids production was lower than in a conventional activated
sludge process.
7. The use of a polyvinyl basin cover proved to be
unsatisfactory because of leakage.
8. The process did not produce an effluent capable of
meeting the District's strict local discharge requirements of
5 mg/1 BOD and 5 mg/1 suspended solids.
Based primarily on conclusions 1, 5, 7, and 8 above, the use of the
Simplex process as a full-piant conversion was not considered to be economv
cally attractive for the District.
This report was submitted in fulfillment of Grant No. S802356 by the
Las Virgenes Municipal Water District under partial sponsorship of the U.S.
Environmental Protection Agency and covers the period of July 1971 through
February 1974.
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CONTENTS
Foreword iii
Abstract 1v
Figures vii
Tables viii
Acknowledgements x
1. Introduction 1
2. Conclusions 2
3. Background 4
4. Physical Description 8
5. Operation and Evaluation Phase 17
Phase 1 17
Phase 2 21
Phase 3 26
Phase 4 30
Phase 5 35
Phase 6 39
6. Discussion 44
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FIGURES
Number Page
1 Schematic diagram of the Simplex process 5
2 RAS and WAS diversion box 9
3 Aeration tank tent and step feed piping 10
4 Aeration mixing blower, rotometer, and sample parts 11
5 Liquid oxygen tanker 13
6 Oxygen vaporizer and gas feed control system 14
7 SVI versus F/M loading 48
8 Sludge production versus MCRT 53
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TABLES
Number Page
1 Experimental Program 7
2 Phase 1 Primary and Secondary Effluent Characteristics .... 18
3 Phase 1 Loading Parameters 19
4 Phase 1 Sludge Characteristics 20
5 Phase 1 Oxygen Utilization and Nitrogen Conversion 21
6 Phase 2 Primary and Secondary Effluent Characteristics .... 23
7 Phase 2 Loading Parameters 24
8 Phase 2 Sludge Characteristics 25
9 Phase 2 Oxygen Utilization and Nitrogen Conversion 26
10 Phase 3 Primary and Secondary Effluent Characteristics .... 27
11 Phase 3 Loading Parameters 28
12 Phase 3 Sludge Characteristics 29
13 Phase 3 Oxygen Utilization and Nitrogen Conversion 30
14 Phase 4 Primary and Secondary Effluent Characteristics .... 31
15 Phase 4 Loading Parameters 32
16 Phase 4 Sludge Characteristics 33
17 Phase 4 Oxygen Utilization and Nitrogen Conversion 34
18 Phase 5 Primary and Secondary Effluent Characteristics .... 36
19 Phase 5 Loading Parameters 37
20 Phase 5 Sludge Characteristics 38
21 Phase 5 Oxygen Utilization and Nitrogen Conversion 39
viii
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Number Page
22 Phase 6 Primary and Secondary Effluent Characteristics .... 41
23 Phase 6 Loading Parameters 42
24 Phase 6 Sludge Characteristics 43
25 Phase 6 Oxygen Utilization and Nitrogen Conversion 43
26 Nitrogen Changes in the Secondary Clarifiers 45
27 Nitrogen Losses 46
28 Oxygen Data 50
29 Alkalinity Consumption 52
30 Operation Cost Comparison 54
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ACKNOWLEDGEMENTS
This project was funded by the Environmental Protection Agency under
Grant No. S802356. Mr. Richard C. Brenner, Municipal Environmental Research
Laboratory (MERL), Cincinnati, Ohio, was the Project Officer, Mr. Richard
L. Fowle, also of MERL, reviewed this report.
Special thanks is due Mr. Ralph Wagner of VTN Consolidated, Inc. who
served as operations consultant throughout the study and coauthored this
report.
The research consultants, Dr. Andrew Gram and Mr. Milton Spiegel, pro-
vided valuable guidance on all phases of the project.
The personnel at the Tapia Water Reclamation Facility should be
acknowledged for their analytical and operational work. The key persons
were:
Mr. William Ruff, Senior Lab Technician
Mr. Roy Gull, Plant Foreman
Mr. Robert Hensley, Lab Technician
Mr. William McCumber of Cosmodyne Corporation provided instrumentation
and installation services at the start of the project.
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SECTION 1
INTRODUCTION
This report discusses a study which investigated the use of high purity
oxygen gas as a simplified upgrading technique for activated sludge processes.
This simplified technique utilized a single-stage, covered air aeration basin
and minor modifications to existing blower and air injection equipment in an
attempt to minimize the expense of upgrading an activated sludge process with
oxygen. The objectives of the study were:
1. to determine the practicality and cost of using this
simplified method of high purity oxygen injection as an
upgrading tool,
2. to determine if aeration basin throughput rate could be
at least doubled,
3. to determine the extent of reduction of waste activated
sludge production,and
4. to compare results with baseline data collected on a
parallel conventional air aeration system.
The Las Virgenes Municipal Water District (LVMWD) has constructed approxi-
mately one-fifth of the total treatment capacity necessary for ultimate growth.
The District's prime objective in participating in the study was to ascertain
whether savings could be realized in future construction costs (reduced tank
capacity) while still achieving a superb quality effluent.
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SECTION 2
CONCLUSIONS
The following conclusions were reached from this study:
1. Operation procedures for the Simplex system did not require
any new operation techniques that would be unfamiliar to
normal activated sludge operation. However, operation of
the oxygen dispersing equipment did require more time than
the operation of a standard air compressor utilized in an
air activated sludge system.
2. The Simp!ox system performed best at the higher BOD
loadings. This coupled with the relative ease and low cost
for conversion of an existing air activated sludge system
to an oxygen enriched system could make the Simplex system
a viable system for facilities with seasonal high BOD waste
loads.
3. Nitrification was easily accomplished when mixed liquor
dissolved oxygen concentrations were high enough to
maintain dissolved oxygen in the sludge blanket of the
secondary clarifiers.
4. Aeration detention times as low as 1.7 hrs were studied
with BOD removals still approaching 90 percent. Even at
1.7 hrs and the low dissolved oxygen (D.O.) conditions
observed in the mixed liquor, nitrification was 83 percent
complete.
5. The process produced relatively high sludge volume index
(SVI) values; however, high SVI values will not necessarily
reflect poor effluent quality, as long as secondary
clarifier loadings are somewhat conservative.
6. The amount of excess solids production per pound of
substrate removed was less than in the conventional
activated sludge process. This is because the "Simplex"
process provides a greater oxygen transfer rate to the
mixed liquor, allowing a higher mixed liquor volatile
suspended solids (MLVSS) concentration.
7. Use of a polyvinyl tent to cover the aeration basin is
totally unsatisfactory. Leaks in the tent wasted quantities
of oxygen and required constant mending and remending.
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8. Oxygen consumption calculations are not felt to be completely
accurate for any of the phases. Discrepancies in correcting
for leaks in the tent are the cause of the discrepancies.
9. Although the Simplex process produced an excellent effluent
quality based on nationwide standards, the process could not
produce an effluent quality satisfactory to Las Virgenes
Municipal Water District needs, i.e., BOD and suspended solids
concentrations of 5 mg/1 each.
10. Use of the Simplex process is not economically attractive for
the Las Virgenes Municipal Water District from an operation
and maintenance viewpoint.
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SECTION 3
BACKGROUND
The concept of using a single-stage, covered/aeration basin and conven-
tional air aeration equipment for high purity oxygen injection was devised in
1970 by the Cosmodyne Division of Cordon International and given the name
Simplex and is schematically illustrated in Figure 1. In April 1971, the Las
Virgenes Municipal Water District applied to the United States Environmental
Protection Agency (EPA) for a matching grant to investigate the Simplex
concept.
In June 1971, EPA awarded a grant of $186,000 to the LVMWD. In addition
to the EPA grant, LVMWD contributed $62,000 for a total project cost of
$248,000. The study was conducted at the LVMWD Tapia Water Reclamation
Facility, Calabasas, California, which had sufficient capacity for the study
due to a recent expansion.
Conversion of existing facilities involved covering a existing aeration
basin with a heavy duty plastic cover (the tent). The existing air blower
system was modified to draw atmosphere from under the tent and recirculate
it through the aeration basin mixed liquor using the existing coarse bubble
air diffuser system. Liquid oxygen from a tanker truck was vaporized and fed
to the mixed liquor thorugh a new fine bubble diffuser system (Saran wrapped
tubes), which was installed at the head end of the aeration basin. A portion
of the atmosphere under the tent was bled off to remove carbon dioxide pro-
duced by metabolic stabilization of substrate. Primary effluent was added to
the aeration basin using the step feed flow reqime. Mixed liquor was clari-
fied using a conventional secondary clarifier(s) from which settled sludge
was withdrawn for recycling and wasting.
THEORY OF PROBLEM SOLUTION
Maintenance of D.O. in the mixed liquor of the activated sludge process
is essential for growth of aerobic organisms. It is generally agreed,
however, that D.O. concentrations above a certain critical concentration have
no effect upon the rate of substrate stabilization. Various researchers have
reported this critical D.O. concentration to be in the range of 0.1 - 0.5 mg/1,
the discrepancies probably occurring since it is actually the D.O. concen-
tration within the microbial floe which is critical, rather than the gross
mixed liquor D.O. concentration.
Transfer of oxygen from the gas phase to the liquid phase has tradi-
tionally been accomplished in the activated sludge process by coarse bubble
aeration, which in addition to transferring oxygen to the liquid phase also
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provides mixing energy to prevent solids deposition and scrubs metabolically
produced carbon dioxide from the mixed liquor. According to Henry's Law, the
D.O. saturation concentration in the mixed liquor is proportional to the
oxygen concentration in the gaseous phase. Therefore, use of a gaseous phase
with a higher oxygen concentration than if air aeration is used. More
importantly, however, the rate of mass oxygen transfer between the gaseous and
liquid phases is proportional to the difference in concentration between
saturation D.O. concentration and operating D.O. concentration. For example,
when the mixed liquor D.O. is 6 mg/1, the driving forces of air and a
50 percent oxygen atmosphere are as follows:
System Saturation Cone. @ 20°C Driving Force
Air 9.3 mg/1 9.3-6.0 = 3.3 mg/1
50% 02 atm 22.3 mg/1 22.3-6.0 = 16.3 mg/1
Therefore, aeration using a 50 percent oxygen atmosphere system will transfer
oxygen to the liquid phase at a rate of 16.3/3.3, or 4.94, times as fast as
aeration with air.
The significance of using a higher oxygen concentration atmosphere for
aeration is, therefore, not that a higher dissolved oxygen could be attained
per se, but rather that a higher oxygen mass transfer rate is possible from
the gas to the liquid phase. A higher transfer rate will allow a higher
microbial concentration to be maintained in the aeration basin, and therefore,
a smaller aeration basin is required to maintain the same food-to-micro-
organism (F/M) loading.
In situations where an existing activated sludge plant is organically
overloaded, use of the Simplex process will provide this higher oxygen mass
transfer rate, thus allowing a higher MLVSS concentration to be maintained and
allow the F/M loading to be lowered to a value which will produce better
sludge settling in the secondary clarifier.
The project was initially planned to be divided into a five-phase study.
Basically, each phase was intended to provide successive decreases in aeration
detention time and successive increases in secondary clarification overflow
rates.
Decreasing detention time was to be achieved through a combination of
increased flow to the aeration basin and partitioning off of a portion of the
aeration basin to decrease its volume. Increasing secondary clarifier over-
flow rates were obtained by increasing flow, but were restricted from going
too high by supplementing the rectangular secondary clarifier with a circular
secondary clarifier.
Due to a number of unforeseen circumstances, deviations were required
from the planned experimental program. Principal among these circumstances
were a lack of available primary effluent step feed ports in the tent,
equipment malfunctions, and bulking in the secondary sedimentation basin(s).
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Table 1 presents a listing of phases as they were actually conducted and as
they are designated in this report. A more detailed discussion of operational
problems will be included with the discussion of results for each individual
phase.
TABLE 1. EXPERIMENTAL PROGRAM
Phase 1 Phase 2 Phase 3 Phase 4 Phase 5 Phased
Influent flow, m3/day 3785 7116 3838 4940 5545 7014
(mgd) (1.000) (1.880) (1.014) (1.305) (1.465) (1.853)
Amount of aeration basin 100 100 46 46 46 46
used, %
Aeration basin detention 7.20 3.53 3.24 2.39 2.14 1.70
time (Q+QR), hr
Number of secondary 112222
clarifiers used
Clarifier overflow rate: 17.0 22.1 14.0 11.0 12.1 16.4
Rectangular, m3/day/m2 (417) (543) (343) (270) (297) (402)
(gpd/ft2)
Circular, m3/day/m2 13.4 16.8 19.2 22.4
(gpd/ft2) (330) (412) (471) (550)
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SECTION 4
PHYSICAL DESCRIPTION
EQUIPMENT FABRICATION
' As a result of a recent expansion of the Tapia Water Reclamation Facility
kto 31,280 m3/day (8 mgd), a portion of the original plant was available for use
in this study. Primary sedimentation basin effluent was obtained from the new
plant addition.
r Aeration Basin
The aeration basin had dimensions of 35.4 m long x 9.14 m wide x 4.57 m
deep (116.25 ft x 30 ft x 15 ft) and an effective volume of 1,481,300 1
(391,361 gal). The entire aeration basin was used for Phases 1 and 2. Prior
to beginning Phase 3, a partitioning baffle was placed in the aeration basin
to decrease its effective volume. Placement of this baffle decreased the
aeration basin length to 16.46 m (54 ft) resulting in an effective volume of
687,980 1 (181, 764 gal). Addition of primary effluent to the aeration basin
was by step feed. Primary effluent was added to the aeration basin at three
locations in Phases 1 and 2 at the head end, 9.14 m (30 ft) down the aeration
basin, and 18.3 m (60 ft) down the aeration basin. In Phases 3, 4, 5,and 6,
primary effluent was only introduced at the first two of these locations.
Return activated sludge was introduced at the head end of the aeration basin
in all phases (Figure 2).
The aeration basins were covered by a tent to allow maintenance of an
oxygen concentration of about 50 percent (Figure 3). The tent was constructed
from polyethylene. The tent was attached to the aeration basin by wrapping
it under sections of lumber, which were in turn bolted to the concrete at the
outer surface of the aeration basin. The tent had a door fabricated in it to
allow the entry of personnel into the tent for operation and maintenance
| purposes. Throughout all phases of the study, leaks were routinely noted in
if the tent, either visually or by smoke detection. These leaks were attributed
to normal wear of the fabric, stress caused by wind, and bullet holes caused
by pranksters. As holes were detected, they were sealed using glue and
polysheets.
Atmosphere from under the tent was recirculated through the existing air
headers in the aeration basin. A new 849.4 I/sec (1,800 cfm) blower was
equipped to perform this function (Figure 4). Suction for the blower was
l taken from under the tent near the effluent end of the aeration basin. The
^ rate of atmospheric recirculation was established to provide both sufficient
mixing and a mixed liquor D.O. concentration above the critical concentration.
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Figure 4. Aeration mixing blower, rotometer, and sample ports,
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During Phase 1, recirculation was at a rate of 307.2 I/sec (651 cfm) with a
mixed liquor D.O. of about 8.5 mg/1. Recirculation rate followed a downward
trend as the study progressed, ending with a rate of 109 I/sec (231 cfm) and a
mixed liquor D.O. of about 1.0 mg/1 in Phase 6. Sufficient mixing was
provided even at the lower flow rate.
In addition to the existing air header system used for atmosphere recircu-
lation, a separate system was installed for addition of pure oxygen to the
system. This system consisted of a submerged header, 12.19 m (40 ft) long,at
the end of the aeration basin and on the opposite side from the aeration
diffusers. On the header, Saran wrapped tubes were installed for fine bubble
diffusion of the pure 02 into the mixed liquor.
Secondary Clarification
Two secondary clarifiers were used in this study. A rectangular secondary
clarifier with dimensions 36.58 m long x 6.10 m wide x 3.05 m average surface
water depth (1.20 ft x 20 ft x 10 ft) was used in all phases of the study. The
surface area of this clarifier was 223 m2 (2,400 ft2), and total weir length
was 40.54 m (133 ft). The rectangular clarifier was equipped with a mechani-
cal surface skimming device. Sludge was withdrawn from the rectangular
clarifier using variable speed, centrifugal pumps with a maximum capacity of
25.1 I/sec (400 gpm) each.
A circular secondary clarifier was also used in parallel operation with
the rectangular secondary clarifier during a portion of Phase 2 and
during all of Phases 3, 4, 5, and 6. The circular clarifier was 13.72 m
(45 ft) in diameter and had an average surface water depth of 3.05 m (10 ft).
The total surface area was 148.2 m2 (1,595 ft2), and total weir length was
43.07 m 041.3 ft). The circular clarifier was equipped with a mechanical
surface skimming device. Sludge was withdrawn using a variable speed,
centrifugal, underflow pump, capable of operating at a pumping rate up to
25.2 I/sec (400 gpm).
Parallel operation of the rectangular and circular secondary clarifier
is referred to as the combined secondary sedimentation system.
Pure Oxygen Storage and Delivery System
Pure oxygen was delivered to the Tapia facility in a 1.5 m^ (4,000 gal)
tank truck by Cosmodyne (Figure 5). Liquid 02 was transferred from the
delivery trucker to the Og storage tanker. Pure oxygen in liquid form was
withdrawn from the tanker and passed through an oxygen vaporizer at ambient
temperatures (Figure 6).
Instrument and Control System
In an effort to achieve optimum results from a minimum expenditure,
spare components were used in the existing Tapia plant instrumentation
system.
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A flow diagram was developed,and points requiring instruments and
controls were established. A comprehensive control system was applied to this
composite using instruments compatible with existing new system. Instrumen-
tation and controls are divided into numbered loops for ease of identification.
Instrumentation loops 1 through 8 are metering only, using propeller type
meters. Only loop 1, influent flow, is recorded and totalized on the main
control panel.
Loop 9 is probably the most complex, metering and controlling oxygen flow
to the oxygen diffuser. Using cascaded feed-forward control, the primary
variable is sensed by the oxygen analyzer as percentage of oxygen in the
atmosphere under the tank cover. A controller with proportional plus reset
modes using primary variable as process input, has its output cascaded into
the set point of the oxygen flow controller which uses the secondary variable,
oxygen flow, as process input. Thus the oxygen flow control is a closed loop,
the set-point of which is continually adjusted by the oxygen analysis
controller to maintain the set percentage of oxygen in the atmosphere under
the cover.
Loop 10 controls the recirculation flow of the under-cover gas through
the blower and the diffusers to provide tank agitation and additional oxygen
uptake. Recirculation gas flow is sensed as AP across an orifice plat and
controlled by a butterfuly valve in the blower suction line. A manual bypass
is provided around the blower to maintain blower air flow above the surge
point if low recirculation flows are desired. Loops 11 through 16 are flow
elements in the recirculation gas line to provide balancing flow indication
to adjust flow to individual diffusers.
Loop 17 senses structure pressure and bleeds off excess undercover gas to
maintain structure support pressure.
Using the new plant spare instrumentation facilities and main instrument
panel all recorded variables were electronically transmitted and recorded on
the main panel.
Electronic controllers are also located on the main panel so that the
process variables may be monitored and set from the main control room.
Control valves were pneumatically operated from electronic input signals
using current to air transducers.
Sewage and sludge flows are locally metered and manually controlled.
During the course of the test, difficulties were encountered which
prevented the instrumentation and control system from operating as planned
and installed.
The major difficulty was encountered in leakage of the gas under the tank
cover to the outside atmosphere. Because of leakage, between the concrete
wall and the fabric structure at this pressure, it was impossible to maintain
adequate support for the structure without adding atmospheric air to the
closed system. This was accomplished by bleeding air from the aeration air
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serving an adjacent tank into the suction of the blower for the test tank.
This air quantity was measured by a large positive displacement gas meter
initially and by a rotometer after the gas meter failed in service.
The leakage also prevented the cascaded feed-forward control of the
oxygen content of the atmosphere under the cover, so the oxygen feed was
manually adjusted using flow controller FIC-9. As surplus gas was never
encountered, bleed to the atmosphere through loop 17 was not needed except as
a safety backup and FCV-17 remained closed thorughout the test.
The propeller meters used for determining influent and RAS flows worked
well on influent but were difficult to keep in operation in sludge service.
Because of continuing clogging and loss of suction in syphoning sample
lines to the mixed liquor 0? analyzers, redesign and installation of purge
units was necessary. Once this was accomplished, continuous monitoring
records were fairly dependable. Oo analyzers A-18, A-19,and A-20 showed D.O.
levels to be equal throughout the length of the tank.
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SECTION 5
OPERATION AND EVALUATION PHASE
The study was initiated on April 6, 1972 with a brief shakedown period,
during which instruments were checked out, calibrated, and adjusted. Mixed
liquor solids were built up by the addition of waste activated sludge (WAS)
from the main plant. Within a few days, mixed liquor solids reached a
sufficiently high concentration, and no further WAS from the main plant was
introduced. By April 17, 1972 effluent quality had stabilized at a BOD of
about 15 mg/1 and SS of about 10 mg/1 at a primary effluent flow of
0.0307 m3/sec (0.7 mgd). Pure oxygen was first added to the aeration basin
on April 17, 1972 and after a week long acclimation period, Phase 1 was begun.
As previously discussed, initial plans were to conduct the study in five
separate phases, but deviations in the planned study were necessary due to a
number of operational and equipment problems. In all, six separate phases
were conducted as shown in Table 1. Each of the six phases is discussed
separately in the following sections.
PHASE 1
Background
Phase 1 began on April 25, 1972 and was completed on July 31, 1972, a
period of 98 days. Phase 1 was planned to be an ultra-conservative phase
which would allow comparison with the latter stages of the study. Flow to
the aeration basin during Phase 1 was 0.0438 m3/sec (1 mgd), with an average
return activated sludge rate of 30 percent. The flow was continuous, with
no diurnal variation, and was taken directly from the primary effluent stream
going to the main air-activated sludge process. Aeration basin detention time
was 7.2 hrs and mean cell residence time (MCRT) was 108 days. No solids were
wasted from the system except secondary clarifier skimmings and suspended
solids in the secondary effluent.
Operational Problems
The major operational problems encountered in Phase 1 were bearing
problems on the recirculation blower, leaks in the tent, and failures of the
aeration basin D.O. analyzer sampling pots.
The blower problems resulted from a motor shaft which was initially
defective and caused rapid wear of the shaft bearings. The blower was shut
down for maintenance several times during Phase 1, usually for only an hour
or two, but once for a 2-day period (May 30 and 31). Pure oxygen was added
continuously to the aeration basin, even during periods when recirculation
blower was down.
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As previously mentioned, leaks in the tent were a persistent problem
throughout the entire study. In order to maintain a positive pressure in the
tent, air was bled into the atmosphere under the tent at an average rate of
3.54 I/sec (7.5 cfm). Numerous smoke tests were conducted throughout Phase 1
to locate leaks in the tent.
Aeration Basin-Secondary Clarifier Efficiency
Overall system efficiency can be best judged by removals through the
aeration basin secondary clarifier complex. As shown in Table 2, removals
were exceptionally good for organic substrates, suspended solids, and ammonia
nitrogen and total Kjeldahl nitrogen (TKN). Almost complete nitrification
occurred, as shown by the 97 percent reduction of ammonia nitrogen. A pH
decrease also occurred, due to removal of alkalinity.
As a result of the low substrate concentrations in the primary effluent
and the large volume of the aeration basin, F/M loadings were very low, as
may be seen in Table 3.
For example, the F/M loading was only 0.073 mass BOD applied/day/mass
MLVSS. These low loadings resulted in a slightly high average SVI, 115 ml/gm ,
but did not adversely affect substrate removal efficiency, as seen in Table 2.
Probably a key factor in the excellent substrate removal was the low secondary
clarifier loading rate and rapid removal of settled sludge for recycle, as
well as the low basin and weir overflow rates in the secondary clarifier, as
shown in Table 3.
TABLE 2. PHASE 1 PRIMARY AND SECONDARY EFFLUENT CHARACTERISTICS
Primary Secondary Percent
Effluent Effluent Change
BOD, mg/1
T-COD, mg/1
S-COD, mg/1
Suspended solids, mg/1
Volatile suspended solids, mg/1
NH3-N, mg/1
NOa-N, mg/1
Total TKN, mg/1
Alkalinity, mg/1
Turbidity, JTU
pH, units
82
153
58
73
72
13
242
7.5
2
35
16
9
7
0.4
16.2
1.2
145
1.4
6.6
97.6
77.1
72.4
87.7
90.3
96.9
40.1
18
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TABLE 3. PHASE 1 LOADING PARAMETERS
F/M Ratios
mass BOD applied/day/mass MLVSS 0.073
mass T-COD applied/day/mass MLVSS 0.133
mass S-COD applied/day/mass MLVSS 0.053
Volumetric Loadings
kg BOD applied/day/m3 0.215
(Ib BOD applied/day/loo ft3) (13.43)
kg T-COD applied/day/m3 0.392
(Ib T-COD applied/day/1000 ft3) (24.48)
hg S-COD applied/day/m3 0.156
(Ib S-COD applied/day/loo ft3) (9.75)
Substrate Removals
mass BOD removed/day/mass MLVSS 0.070
mass T-COD removed/day/mass MLVSS 0.102
mass S-COD removed/day/mass MLVSS 0.036
Rectangular Clarifier Loadings
Surface overflow rate, nr/day/m2 25.14
(gpd/ft2) (617)
Weir overflow rate, m3/day/m 93.39
(gpd/ft) (7519)
Solids loading rate, kg MLSS/day/m2 62.15
(Ib MLSS/day/ft2) (12.73)
Circular Clarifier Loadings
Surface overflow rate, m3/day/m2
(gpd/ft2)
Weir overflow rate, m3/day/m
(gpd/ft)
Solids loading rate, kg MLSS/day/m2
(Ib MLSS/day/ft2)
19
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Sludge Characteristics
The low loading resulted in an extremely long MCRT of 108 days, in spite
of a MLVSS of 2,950 mg/1. As a result of the long MCRT, the sludge growth
was well into the endogenous growth phase. This is also evidenced by the low
sludge production rates, as shown in Table 4, and the absence of any direct
wasting of activated sludge. The only solids wasted from the system were
secondary clarifier skimmings, which averaged 32.65 kg (72 Ib) VSS/day, and
effluent suspended solids, which averaged 26.48 kg (58.4 Ib) VSS/day. The
long MCRT did not significantly affect the return activated sludge (RAS)
concentration, which averaged 14,324 mg/1 and 82.7 percent volatile content.
Oxygen Utilization and Nitrogen Conversion
Oxygen utilization efficiency (Table 5) was relatively high, averaging
85.2 percent, as was the mixed liquor D.O. which averaged 7.5 mg/1. The
amount of oxygen consumed per quantity of substrate stabilized was extremely
high, even after correcting for oxygen required for biological oxidation of
ammonia by nitrifying bacteria. The explanation for the high oxygen consump-
tion values is most likely the large amount of oxygen required for endogenous
respiration of cellular material. This theory is enhanced by considering
that only 59.13 kg (130.4 Ib) of volatile solids were wasted per day, while
302.58 kg (667.2 Ib) of 6005 were oxidized per day.
TABLE 4. PHASE 1 SLUDGE CHARACTERISTICS
MLSS, mg/1 3,692
MLVSS, mg/1 . 2,950
Return Activated Sludge TSS, mg/1 14,324
Return Activated Sludge Rate, % Q
Rectangular clarifier 30
Circular clarifier
SVI, ml/gm 115
Mean Cell Residence Time, days 108
Sludge Production
mass VSS produced/mass BOD removed 0.195
mass VSS produced/mass T-COD removed 0.132
mass VSS produced/mass S-COD removed 0.372
20
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TABLE 5. PHASE 1 OXYGEN UTILIZATION AND NITROGEN CONVERSION
Oxygen supplied,
kg 02 supplied/nr 0.327
(Ib 02 supplied/105 gal) (2733)
mass 02 supplied/mass BOD removed 4.10
mass 02 supplied/mass T-COD removed 3.16
Oxygen consumption
kg 02 consumed/m3 0.279
(Ib 02 consumed/106 gal) (2329)
mass 02 consumed/mass BOD removed 3.50
mass 02 consumed/mass T-COD removed 2.70
Oxygen utilization efficiency, % 85.2
Average aeration basin DO, mg/1 7.5
Ammonia oxidation, % 97
Rate of nitrification,
mass NH3-N oxidized/day/mass MLVSS 0.0203
Alkalinity consumption,
mass alkalinity consumed/mass NH3-N oxidized 7.7
Almost complete nitrification of ammonia, 97 percent,occurred in Phase 1
This could be anticipated,however, because the MCRT of 108 days exceeds the
approximate 7-day MCRT required for nitrification, and rapid sludge removal
in the secondary clarifier prevented anaerobisis of the sludge.
Conclusions
High purity oxygen aeration with an aeration time of 7.2 hrs and a mean
cell residence time of 108 days provides both exceptionally good effluent
quality and exceptionally low solids wasting requirements. These results
are expected, however, because MLVSS are obviously maintained well into the
endogenous growth phase at such a high mean cell residence time.
PHASE 2
Background
Phase 2 began on September 11, 1972 and was terminated on November 13,
1972,a period of 64 days. The period between Phases 1 and 2 was necessary
21
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for factory repair of the recirculation blower motor and buildup and accli-
nation of MLVSS following return of the blower motor. Because primary
effluent available for the study was insufficient to allow operation at the
plannad flow rate, a portion of effluent from the combined secondary sedi-
mentation system was returned to the head works section of the plant to mix
with raw sewage. In order to prevent a substantial dilution of substrate
concentration by return of this secondary effluent, all but one of the weirs
in the existing primary clarifier were blocked off in an attempt to decrease
efficiency of the primary sedimentation basin. This turned out to be an
extremely difficult process to regulate because the sludge blanket depth in
the primary clarifier was not far below the weirs. At times, this method
of operation allowed unxepectedly large concentration of substrate and solids
to pass over the one remaining weir, and thus overload the aeration basin.
Average flow to the aeration basin in Phase 2 was 7225.8 m^/day
(1.88 mgd), and the return activated sludge flow averaged 42.5 percent. Average
aeration basin detention time was 3.53 hrs, and average mean cell residence
time was 48.6 days.
Operational Problems
It was observed that in addition to loss of atmosphere through leaks in
the tent, atmosphere was also being lost through the discharge line from the
aeration basin to the secondary clarifiers. Several attempts were made to
correct this condition, but they resulted in flow restriction during periods
of peak flow and overflow of the aeration basin.
Problems were also encountered in maintaining a sufficient RAS pumping
rate from the rectangular secondary clarifier. This resulted in periodic
solids bulking in the rectangular secondary clarifier. The problem was
finally solved by decreasing the flow to the rectangular clarifier, and thus,
the solids loading rate. The concentration of the RAS then decreased, and the
rate of pumping went up.
Another problem deserving mention is an unexplainable pH increase in the
primary effluent for a week in early October. Aeration basin pH increased
gradually from 6.3 to 7.3 over a 5-day period, then abruptly dropped back
to 6.3 in a 1-day period.
Aeration Basin-Secondary Clarifier Efficiency
Removal efficiencies in Phase 2 (Table 6) were affected by period high
substrate and solids carryover from the primary clarifier aeration basin, as
a result of secondary effluent recirculation and the intentional blockage of
primary clarifier weirs. This method of primary clarifier operation, while
serving to increase substrate concentration, resulted in a higher than normal
percentage of substrate being in the particulate form. Although BOD, SS, and
VSS were essentially the same as Phase 1, COD and TKN removals were much lower.
The low COD removals are felt to be due to a higher percentage of non-
biodegradable COD in the primary effluent compared to effluent from a normally
operated primary sedimentation basin. As in Phase 1, nearly complete nitrifi-
cation occurred as evidenced by a 98.8 percent decrease in ammonia. A pH
22
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decrease again occurred due to destruction of alkalinity during nitrification.
The decreased TKN removal efficiency is also felt to be related to primary
clarifier operation, as well as cellular organic nitrogen contained in micro-
bial floe during periods when sludge in the rectangular clarifier bulked.
TABLE 6. PHASE 2 PRIMARY AND SECONDARY EFFLUENT CHARACTERISTICS
Primary Secondary Percent
Effluent Effluent Change
BOD, mg/1
T-COD, mg/1
S-COD, mg/1
Suspended solids, mg/1
Volatile suspended solids, mg/1
NH3-N, mg/1
NOa-N, mg/1
Total TKN, mg/1
Alkalinity, mg/1
Turbidity, JTU
pH, units
143.5
145.8
43.0
214
200
6.7
12.6
183
7.2
3.7
94.9
20.1
13.4
9.5
0.08
13.7
2.7
126
2.8
6.6
97.4
34.9
53.3
93.7
95.3
98.8
78.6
31.1
F/M loading (Table 7) increased in response to the increase in primary
effluent substrate concentration. In Phase 2, they were nearly at an
"optimum level" (relative to air aeration), yet the SVI of 179 (Table 8) was
noticeably higher than in Phase 1. This higher SVI, however, only reduced the
RAS concentration by about 9 percent compared to Phase 1. RAS concentration
was still very acceptable, averaging 13,193 mg/1.
During all of Phase 2, both the rectangular and circular clarifiers were
used. As a result, secondary clarifier surface and weir overflow rates were
essentially unchanged from Phase 1. Solids loading rate did increase slightly
due to a higher MLSS concentration.
Sludge Characteristics
Mixed liquor volatile suspended solids were maintained at a relatively
high level, 3,046 mg/1. The principal factor allowing this high a MLVSS was
the relatively high volatile solids concentration of the return activated
sludge, VSS = 9,400 mg/1. The ability to produce such a concentrated sludge
resulted from low surface overflow rates and relatively low weir overflow
rates in the combined secondary clarification system. In addition, the solids
loading rate was relatively high, which could have partly offset the high
SVI of 179.
MCRT decreased to 48.6 days, partially due to increased effluent VSS, but
mainly due to an increase in surface skimmings from the secondary clarifier.
There was no waste activated sludge from the system until the last week of
Phase 2. Solids production on a BOD basis was only slightly higher than in
23
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TABLE 7. PHASE 2 LOADING PARAMETERS
Ratios:
mass BOD applied/day/mass MLVSS 0.236
mass T-COD applied/day/mass MLVSS 0.224
mass S-COD applied/day/mass MLVSS 0.065
Volumetric loadings:
kg BOD applied/day/m3 0.700
(Ib BOD applied/day/1000 ft3) (43.75)
kg T-COD applied/day/m3 0.692
(Ib T-COD applied/day/1000 ft3) (43.20)
kg S-COD applied/day/m3 0.203
(Ib S-COD applied/day/1000 ft3) (12.70)
Substrate Removals:
mass BOD removed/day/mass MLVSS 0.230
mass T-COD removed/day/mass MVLSS 0.075
mass S-COD removed/day/mass MLVSS 0.035
Rectangular Clarifier Loadings:
Surface overflow rate, m3/day/m2 17.76
(gpd/ft2) (436)
Weir overflow rate, m3/day/m 97.72
(gpd/ft) (7868)
Solids loading rate, kg MLSS/day/m2 68.59
(Ib MLSS/day/ft2) (14.05)
Circular Clarifier Loadings:
Surface overflow rate, m3/day/m2 22.12
(gpd/ft2) (543)
Weir overflow rate, m3/day/m 76.11
(gpd/ft) (6128)
Solids loading rate, kg MLSS/day/m2 86.26
(Ib MLSS/day/ft2) (17.67)
24
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TABLE 8. PHASE 2 SLUDGE CHARACTERISTICS
MLSS, mg/1 3,896
MLVSS, mg/1 3,046
Return Activated Sludge TSS, mg/1 13,193
Return Activated Sludge Rate, % Q
Rectangular clarifier 55.0
Circular clarifier 26.8
SVI, ml/gm 179
Mean Cell Residence Time, days 48.6
Sludge Production
mass VSS produced/mass BOD removed 0.248
mass VSS produced/mass T-COD removed 0.493
mass VSS produced/mass S-COD removed 0.896
Phase 1, which as in Phase 1, was due to the long MCRT, and,therefore,main-
tenance of the cells in an endogenous growth phase for substantial periods of
time. Solids production on a COD basis was substantially greater than in
Phase 1 due principally to a decrease in COD removal efficiency.
Oxygen Utilization and Nitrogen Conversion
As shown in Table 9, oxygen utilization efficiency was much lower than in
Phase 1. This was due to an increased leakage rate through the tent and an
increased amount bearing in the aeration basin effluent, a result of higher
mixed liquor D.O. and higher plant flow. Oxygen consumption per pound of BOD
removed dropped sharply from Phase 1. More than likely, this was due to a
decrease in MCRT to 48.6, days and, therefore, a decrease in the time that
the microbial cells were undergoing endogenous respiration.
Essentially, complete nitrification was achieved, as measured by a
98.8 percent reduction of ammonia nitrogen. The degree of nitrification
dropped from Phase 1 because the NH3-N concentration was only about half as
much due to recycling secondary effluent to the head end of the plant.
As in Phase 1, a greater amount of alkalinity was lost from the system
than stochiometrically required for ammonia oxidation. It is probable that
some carbon dioxide was lost from the system due to the relatively low pH
levels in the mixed liquor and secondary clarifier.
25
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TABLE 9. PHASE 2 OXYGEN UTILIZATION AND NITROGEN CONVERSION
Oxygen Supplied
kg 0? supplied/m3 0.166
(Ib 02 supplied/106 gal) (1385)
mass 02 supplied/mass BOD removal 1.86
mass 02 supplied/mass T-COD removed 5.30
Oxygen Consumption
kg 02 consumed/m3 0.099
(Ib 02 consumed/106 gal) (824)
mass 02 consumed/mass BOD removed 1.08
mass 02 consumed/mass T-COD removed 3.08
Oxygen Utilization Efficiency, % 59.4
Average Aeration Basin DO, mg/1 9.04
Ammonia Oxidation, % 98.8
Rate of Nitrification
mass NHs-N oxidized/day/mass MLVSS 0.014
Alkalinity Consumption
mass alkalinity consumed/mass NH3-H oxidized 9.14
PHASE 3
Background
Phase 3 began on January 22, 1973 and was terminated on April 3, 1973,a
period of 72 days. Phase 3 was intended to have a shorter aeration time than
Phase 2 as well as lower surface overflow rates in the secondary clarifiers.
To accomplish this, a partitioning baffle was placed at approximately the
midpoint of the aeration basis, thereby decreasing the volume by about half.
Flow at one-half the rate in Phases 1 and 2 thus produces the same aeration
detention time, for a constant recycle rate.
Average flow in Phase 3 was 3875.8 m3/day (1.024 mgd) with an average
return activated sludge rate of 33 percent. Aeration basis detention time
was 3.24 hrs,and the mean cell residence time was 84 days. There was no
skimming during Phase 3; solids were only removed in the waste activated sludge
and in the effluent.
26
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Operational Problems
The major operational problems encountered in Phase 3 were failure of the
recirculation blower, leaks in the tent, and bulking of the final clarifiers.
As in prior phases, recurrent leaks occurred in the tent. Usually they
were caused by wind, but they still indicate the drawbacks of using a plastic
cover.
Bulking in the final clarifiers was attributable to several factors;
density currents caused part of the problem, these could not be alleviated,
however, because of the nature of clarifier design. A higher SVI than in the
previous phases also influenced the frequency of bulking. The higher SVI
resulted in maintenance of al.68 to 1.98 m (5.5 to 6.5 ft) deep sludge blanket,
despite a 33 percent return activated sludge recycle rate and low secondary
clarifier overflow rates.
Aeration Basin-Secondary Clarifier Efficiency
Efficiency of the aeration basin-secondary clarifier complex is presented
in Table 10. Substrate removals were still very high, although somewhat less
than in Phase 1. The BOD concentration applied to the aeration basin was less
than in Phase 2, because the primary clarifier was operated in a normal
fashion, but both total and soluable COD concentration increased. Most
notable is the decrease in the reduction of ammonia, averaging only 86.5 per-
cent. A pH decrease was alkalinity by nitrifying bacteria during nitrifi-
cation.
TABLE 70. PHASE 3 PRIMARY AND SECONDARY EFFLUENT CHARACTERISTICS
Primary Secondary Percent
Effluent Effluent Change
BOD, mg/1
T-COD, mg/1
S-COD, mg/1
Suspended solids, mg/1
Volatile suspended solids, mg/1
NH3-N, mg/1
N03-N, mg/1
Total TKU, mg/1
Alkalinity, mg/1
Turbidity, JTU
pH, units
88.8
189
82.4
44.0
38.4
11.83
15.6
225.8
7.43
2.6
30.9
23.9
4.8
3.6
1.6
8.0
3.0
160.2
1.7
6.74
47.1
83.7
71.0
89.1
90.6
86.5
80.8
29.1
In Phase 3 (Table 11), the BOD F/M loading was lower than Phase 2 because
primary effluent BOD decreased considerably. Conversely, total and soluble
COD F/M loadings were higher than Phase 2. It is important to note that the
ratio of the BOD to T-COD F/M loading in Phase 3 was 0.47, while in Phases 1
27
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and 2 it was 0.55 and 1.05, respectively. Primary effluent substrate charac-
teristics were,therefore,closer in Phases 1 and 2.
Surface and weir overflow rates and solids loading rate were lower than
Phase 2, because total flow was less in Phase 2. These low loadings are
considered very conservative.
TABLE 11. PHASE 3 LOADING PARAMETERS
F/M Ratios:
mass BOD applied/day/mass MLVSS 0.174
mass T-COD applied/day/mass MLVSS 0.372
mass S-COD applied/day/mass MLVSS 0.161
Volumetric Loadings:
kg BOD applied/day/m3 0.497
(Ib BOD applied/day/1000 ft3) (31.03)
kg T-COD applied/day/m3 , 1.057
(Ib T-COD applied/day/1000 ft3) (65.97)
kg S-COD applied/day/m3 , 0.461
(Ib S-COD applied/day/1000 ft3) (28.77)
Substrate Removals:
mass BOD removed/day/mass MLVSS 0.163
mass T-COD removed/day/mass MLVSS 0.312
mass S-COD removed/day/mass MLVSS 0.101
'ť*)Ť,..
Rectangular Clarifier Loadings:
Surface overflow rate, m3/day/nr 13.97
(gpd/ft2) (343)
Weir overflow rate, m3/day/m 76.94
(gpd/ft) (61.95)
Solids loading rate, kg MLSS/day/m2 53.34
(Ib MLSS/day/ft2) (10.72)
Circular Clarifier Loadings:
Surface overflow rate, m3/day/nr 13.44
(gpd/ft2) (330)
Weir overflow rate, m3/day/m 46.38
(gpd/ft) (3734)
Solids loading rate, kg MLSS/day/m2 50.19
(Ib MLSS/day/ft2) (10.28)
28
-------
Sludge Characteristics
Although the secondary clarifiers were operated at lower loading rates
than Phase 2, the concentration of the recycled activated sludge was less than
Phase 2 (Table 12). The primary reason is felt to be the higher SVI, 285, for
Phase 2. The mixed liquor solids were lower than Phase 2 because the RAS
concentration was less and because activated sludge was wasted from the system.
As a result of a higher MCRT for Phase 3, sludge production per unit of sub-
strate removed decreased from Phase 2.
TABLE 12. PHASE 3 SLUDGE CHARACTERISTICS
MLSS, mg/1 3,726
MLVSS, mg/1 2,861
Return Activated Sludge TSS, mg/1 11,498
Return Activated Sludge Rate, % Q
Rectangular clarifier 32.7
Circular clarifier 33.0
SVI, ml/gm 185
Mean Cell Residence Time, days 84
Sludge Production
mass VSS produced/mass BOD removed 0.220
mass VSS produced/mass T-COD removed 0.111
mass VSS produced/mass S-COD removed 0.293
Oxygen Utilization and Nitrogren Conversion
Oxygen utilization efficiency was very low in Phase 3, 23.9 percent as
shown in Table 13. The explanation of the low efficiency is the large quan-
tity of oxygen lost through leaks in the tent, meaning that a large amount of
oxygen added to the tent was simply to maintain pressure under the tent.
Oxygen consumption per pound of substrate removed was less in Phase 3
than in Phases 1 and 2, in spite of a longer MCRT than Phase 2. The expla-
nation is felt to be the smaller mass of microbes under aeration, and hence
there would be less oxygen used for endogenous respiration on a daily basis.
In contrast to Phases 1 and 2, incomplete ammonia oxidation occurred.
Since MCRT was more than sufficient for growth of nitrifying organisms, it is
probable that these organisms were killed in the final clarifier due to low
D.O. conditions in the sludge blanket. Since no D.O. measurements were made
on the sludge blanket contents, no proof of this hypothesis exists.
29
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TABLE 13. PHASE 3 OXYGEN UTILIZATION AND NITROGEN CONVERSION
Oxygen supplied
kg 02 supplied/or* 0.368
(Ib 02 supplied/106 gal) (3069)
mass 02 supplied/mass BOD removed 4.35
mass Op supplied/mass T-COD removed 2.76
Oxygen Consumption
kg 0? consumed/m3 0.088
Ob 02 consumed/106 gal) (732)
mass 02 consumed/mass BOD removed 0.90
mass 02 consumed/mass T-COD removed 0.57
Oxygen Utilization Efficiency, % 23.9
Average Aeration Basin DO, mg/1 7.1
Ammonia Oxidation, % 84.6
Rate of Nitrification
mass NH3-N oxidized/day/mass MLVSS 0.020
Alkalinity Consumption
mass alkalinity consumed/mass NH3-N oxidized 6.44
PHASE 4
Background
Phase 4 began on April 4, 1973 and was terminated on April 24, 1973,a
period of 21 days. The principal change from Phase 3 was an increase in the
flow from approximately 3785 to 4920 m3/day (1 to 1.3 mgd). The reduced
volume aeration basin first used in Phase 3 was again used as were the cir-
cular and rectangular secondary clarifiers. Aeration detention time in
Phase 4 decreased to 2.39 hrs. The overall objective of Phase 4 was to study
system performance at higher hydraulic and substrate loadings than Phase 3.
Operational Problems
The principal operational problem of significance was the recurring
problem of holes in the tent. The windy spring weather not only helped create
new leaks, but caused some previously sealed leaks to be reopened. The large
number of leaks resulted in a decreased oxygen concentration in the atmosphere
and a resultant low mixed liquor dissolved oxygen concentration.
30
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Plugging on the dissolved oxygen sampling apparatus at sampling point
one, near the head of the aeration basin was another consistent problem. As
a result of this plugging, no usable dissolved oxygen data was obtained for
sample point one.
Aeration Basin-Secondary Clarifier Efficiency
Substrate removals in Phase 4 (Table 14) dropped slightly from the
efficiencies experienced in the previous three phases. BOD removal exceeded
92 percent, but T-COD and S-COD removal efficiencies averaged only 84.5 and
71.4 percent,respectively. It is felt that the character of the sludge was
responsible for these decreases in substrate and solids removal efficiencies.
That the sludge character had changed is quite evident from the SVI, which
averaged 240. The most likely explanation for this change in sludge character
is the low aeration basin D.O. levels which.,were measured. This hypothesis is
clouded somewhat because no D.O. concentrations were recorded at the head end
of the aeration basin, due to the previously discussed operational problems
with sampling apparatus.
TABLE 14. PHASE 4 PRIMARY AND SECONDARY EFFLUENT CHARACTERISTICS
Primary Secondary Percent
Effluent Effluent Change
BOD, mg/1
T-COD, mg/1
S-COD, mg/1
Suspended solids, mg/1
Volatile suspended solids, mg/1
NH3-N, mg/1
N03-N, mg/1
Total TKN, mg/1
Alkalinity, mg/1
Turbidity, JTU
pH, units
114.1
238.1
99.3
52.2
47.3
14.9
15.7
237.1
7.43
9.0
37.0
28.4
5.2
4.1
4.7
5.9
7.0
180.9
2.1
7.23
92.1
84.5
71.4
90.0
91.3
68.5
55.4
23.7
Ammonia reduction averaged a low 68.5 percent, the lowest level observed
in any of the six phases of study. It is very probable that nitrification was
inhibited by the low dissolved oxygen concentrations measured in the aeration
basin and possibly also in the secondary clarifier. Although dissolved oxygen
measurements were not made in the secondary clarifiers, it it reasonable to
presume that there was less than in the aeration basin effluent, which aver-
aged 1.58 mg/1.
F/M loadings increased (Table 15) as expected, due mainly to an increase
in flow, but also because primary effluent BOD was somewhat greater and because
MLVSS concentration was intentionally decreased. The substrate removal rate
continued to increase in response to the higher substrate loading.
31
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TABLE 15. PHASE 4 LOADING PARAMETERS
F/M Ratios
mass BOD applied/day/mass MLVSS 0.332
mass T-COD applied/day/mass MLVSS 0.689
mass S-COD applied/day/mass MLVSS 0.286
Volumetric Loadings
kg BOD applied/day/m3 0.819
(Ib BOD applied/day/1000 ft3) (51.11)
kg T-COD applied/day/m3 1.709
(Ib T-COD applied/day/1000 ft3) (106.7)
kg S-COD applied/day/m3 0.712
(Ib S-COD applied/day/1000 ft3) (44.46)
Substrate Removals
mass BOD removed/day/mass MLVSS 0.304
mass T-COD removed/day/mass MLVSS 0.583
mass S-COD removed/day/mass MLVSS 0.206
Rectangular Clarifier Loadings
Surface overflow rate, m3/day/m2 11.00
(gpd/ft2) (270)
Weir overflow rate, m3/day/m 60.42
(gpd/ft) (4865)
Solids loading rate, kg MLSS/day/m2 33.64
(Ib MLSS/day/ft2) (6.89)
Circular Clarifier Loadings
Surface overflow rate, nr/day/m2 16.78
(gpd/ft2) (412)
Weir overflow rate, m3/day/m 57.84
(gpd/ft) (4657)
Solids loading rate, kg MLSS/day/m2 51.46
(Ib MLSS/day/ft2) (10.54)
Flow to the circular secondary clarifier was greater than in Phase 3, due
to both a greater system flow and a greater percentage of flow going to the
circular clarifier. The circular clarifier surface overflow rate was still
relatively low, 2.38 m3/day/nr (412 gpd/ft2). -Less flow was diverted to the
rectangular secondary clarifier than in Phase 3. Thus, for the rectangular
clarifier, the surface and weir overflow rates and solids loading rates
decreased from Phase 3.
32
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Sludge Characteristics
As discussed, MLSS and MLVSS were intentionally decreased to increase the
F/M loadings in the aeration basin. This decrease was accomplished by
increasing the rate of sludge wasting. The RAS concentration also decreased
as a result of the increased SVI. The RAS concentration averaged 7,541 mg/1
and was 80.6 percent volatile solids (Table 16).
TABLE 16. PHASE 4 SLUDGE CHARACTERISTICS
MLSS, mg/1 3066
MLVSS, mg/1 " 2480
Return Activated Sludge TSS, mg/1 7541
Return Activated Sludge Rate, % Q
Rectangular clarifier 40.3
Circular clarifier 39.3
SVI, ml/gin 240
Mean Cell Residence Time, days 42.1
Sludge Production
mass VSS produced/mass BOD removed 0.230
mass VSS produced/mass T-COD removed 0.121
mass VSS produced/mass S-COD removed 0.341
The MCRT decreased as a result of the increased rate of wasting activated
sludge. MCRT still remained relatively high, averaging 42 days.
As a result of the increased wasting rate, excess sludge production
increased, but only slightly from Phase 3. Excess sludge production was still
relatively low, however, averaging 0.231 pounds of VSS produced per pound of
BOD removed. This low rate is attributed to the long MCRT and the endogenous
nature of the sludge at such a high MCRT.
Oxygen Utilization and Nitrogen Conversion
Oxygen utilization efficiency continued to be low, averaging only
24.2 percent (Table 17). This low efficiency was due principally to the large
quantity of oxygen lost through leaks in the tent. The aeration basin dissolved
oxygen concentration averaged only 1.52 mg/1. Again, this is principally a
result of leaks in the tent and a concommitant decrease in the oxygen concen-
tration in the atmosphere under the tent.
33
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TABLE 17. PHASE 4 OXYGEN UTILIZATION AND NITROGEN CONVERSION
Oxygen Supplied
kg 02 supplied/rip5 0.246
(Ib 02 supplied/106 gal) (2052)
mass 02 supplied/mass BOD removed 2.85
mass 02 supplied/mass T-COD removed 1.77
Oxygen Consumption
kg Oo consumed/m3 0.072
(Ib 02 consumed/106 gal) (605)
mass 02 consumed/mass BOD removed 0.58
mass 02 consumed/mass T-COD removed 0.36
Oxygen Utilization Efficiency, % 24.2
Average Aeration Basin DO, mg/1 1.52
Ammonia Oxidation, % 68.5
Rate of Nitrification
mass NH3-N oxidized/day/mass MLVSS 0.0297
Alkalinity Consumption
mass alkalinity consumed/mass N^-N oxidized 5.48
It is important to note that a D.O. concentration of zero was often
measured near the middle of the aeration basin. Although D.O. measurements
were not made at the head of the aeration basin, it logically follows from the
above that zero D.O. concentrations also occurred at the head of the aeration
basin.
Oxygen consumed per mass of BOD removed, after accounting for oxygen
required for ammonia oxidation, averaged only 0.25, the lowest of any phase
studied. This low figure is not felt to be valid, and probably results from
discrepancies in measuring the quantity of atmosphere lost through leaks in
the tent.
The percentage of ammonia converted to nitrate averaged only 68.5 percent.
This low percentage, as discussed previously, is believed to be a result of
low aeration basin D.O., and a probable low D.O. concentration in the secondary
clarifier. Low D.O. concentrations such as this probably resulted in the death
of many nitrifying organisms, especially of the genus Nitrosotnonas.
34
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Conclusions
The purpose of Phase 4 was to illustrate the effect of increased F/M
loadings upon the system performance at a decreased mean cell residence time.
These increased loadings were accomplished by increasing the flow and
decreasing the MLVSS, while MCRT was decreased by increasing the rate of
wasting activated sludge.
The results obtained indicated a decrease in overall percentage removed
of most constituents measured. After an analysis of all parameters, however,
it is believed the decreased system performance was due to a low mixed liquor
D.O., averaging only 1.52 mg/1 for sample points two and three. It is important
to note that zero D.O. was often observed at sample point two, and although
D.O. could not be measured at sample point one due to inoperative apparatus,
it can be presumed that zero D.O. conditions also occurred at the head end of
the aeration basin. The low D.O. conditions are attributed to a low atmosphere
oxygen concentration resulting from the bleed-in of large quantities of air to
compensate for leaks in the tent. The low mixed liquor D.O. is felt to also be
the cause of the high SVI and the low percentage removal of ammonia, both the
result of a probable change in the character of the sludge.
PHASE 5
Background
Phase 5 began on April 25, 1973 and continued until May 14, 1973,a
period of 20 days. The major changes from Phase 4 were an increase in flow to
approximately 4564 nvVday (1.47 mgd) and a decrease in MLVSS. As a result of
the increased flow, the aeration basin detention time decreased to 2.14 hrs.
The reduced volume aeration basin and both the circular and rectangular
clarifiers continued in use from Phase 4.
The objective of Phase 5 was to continue study of system performance at
higher hydraulic and substrate loadings than used in prior phases.
Operational Problems
No new operational problems were encountered in Phase 5. However, leaks
in the tent continued to appear and plugging of the D.O. sampling devices
continued to occur.
Aeration Basin-Secondary Clarifier Efficiency
Substrate removal efficiencies in Phase 5 (Table 18) were equal to or
greater than those observed in Phase 4. BOD and T-COD removal efficiencies
were essentially the same as in Phase 4, being 91.8 percent and 85.1 percent,
respectively. S-COD removal averaged 81 percent substantially better than in
Phase 4. Suspended and volatile suspended solids removals were good,
averaging 93.5 percent and 92.2 percent,respectively.
35
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TABLE 18. PHASE 5 PRIMARY AND SECONDARY EFFLUENT CHARACTERISTICS
Primary Secondary Percent
Effluent Effluent Change
BOD, mg/1
T-COD, mg/1
S-COD, mg/1
Suspended solids, mg/1
Volatile suspended solids, mg/1
NH3-N, mg/1
N03-N, mg/1
Total TKN, mg/1
Alkalinity, mg/1
Turbidity, JTU
pH, units
109.7
276.0
174.6
69.4
62.8
16.4
15.9
237.8
7.42
9.0
41.0
33.2
4.5
4.9
3.6
6.7
5.1
170.7
2.0
7.06
91.8
85.1
81.0
93.5
92.2
78.0
67.9
28.2
Oxidation of ammonia was greater than in Phase 4, averaging 78 percent.
It is interesting to note that 12.8 mg/1 of ammonia was removed, but that the
maximum nitrate concentration increase was only 6.7 mg/1. Nitrate production
may have been slightly less than this, however, but must be expressed as a
maximum since primary effluent nitrate was not measured.
pH decreased due to destruction of alkalinity nitrification but only by
0.36 units. This decrease is substantially less than in other phases, except
for Phase 4, where the pH decrease was only 0.2 units.
F/M loadings increased (Table 19) in response to the combination of
increased flow and the decreased MLVSS concentration. On a BOD basis, the
F/M loading was 0.385. The substrate removal rate also increased, averaging
0.366.
Both secondary clarifiers were more highly loaded than in Phase 4, but
loadings were still relatively conservative. For rectangular secondary^
clarifiers,.surface overflow and weir overflow rates were 12.1 nr/day/m2
(297 gpd/ft^) and 66.57 nrVday/m (5368 gpd/ft),respectively. For the circular
secondary clarifier, these values were 19.2 m3/day/m2 (471 gpd/ft2) and
66.05 m3/day/m (5318 gpd/ft), respectively.
Sludge Characteristics
MLSS and MLVSS were intentionally lower than in any prior phase, averaging
2.712 and 2,277 mg/1, respectively (Table 20). The lower solids concentrations
were achieved by increasing the rate of sludge wasting. The concentration of
the RAS was lower than in Phase 4, although the SVI decreased. This is attri-
buted to an increase in the rate of sludge pumping from the rectangular
secondary clarifier, and a consequent dilution of the RAS. The volatile
fraction of the RAS increased slightly from Phase 4 and averaged 83.9 percent.
36
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TABLE 19. PHASE 5 LOADING PARAMETERS
F/M Ratios
mass BOD applied/day/mass MLVSS 0.385
mass T-COD applied/day/mass MLVSS 0.960
mass S-COD applied/day/mass MLVSS 0.664
Volumetric Loadings
kg BOD applied/day/nr5 0.877
(Ib BOD applied/day/1000 ft3) (54.73)
kg T-COD applied/day/m3 2.192
(Ib T-COD applied/day/1000 ft3) (136.8;
kg S-COD applied/day/m3 1.434
(Ib S-COD applied/day/1000 ft3) (89.50)
Substrate Removals
mass BOD removed/day/mass MLVSS 0.366
mass T-COD removed/day/mass MLVSS 0.816
mass S-COD removed/day/mass MLVSS 0.530
Rectangular Clarifier Loading
Surface overflow rate, m3/day/m2 12.10
(gpd/ft2) (297)
Weir overflow rate, m3/day/m 66.67
(gpd/ft) (5368)
Solids loading rate, kg M.SS/day/m2 32.71
(Ib MLSS/day/ft2) (6.70)
Circular Clarifier Loadings
Surface overflow rate, m3/day/m 19.19
(gpd/ft2) (471)
Weir overflow rate, m3/day/m 66.05
(gpd/ft) (5318)
Solids loading rate, kg MLSS/day/m2 51.75
(Ib MLSS/day/ft2) (10.60)
37
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TABLE 20. PHASE 5 SLUDGE CHARACTERISTICS
MLSS, mg/1
MLVSS, mg/1
Return Activated Sludge TSS, mg/1
Return Activated Sludge Rate, % Q
Rectangular clarifier
Circular clarifier
SVI, rnl/gm
2712
2277
6932
42.2
37.9
220
Mean Cell Residence Time, days 22.4
Sludge Production
mass VSS produced/mass BOD removed 0.396
mass VSS produced/mass T-COD removed 0.162
mass VSS produced/mass S-COD removed 0.371
As a result of increased sludge wasting, the mean cell residence time
decreased to 22.2 days. This MCRT is approximately one-half the MCRT of
Phase 4. Excess sludge production increased to 0.333. This is in response
to an increased rate of sludge wasting, which lowered the MCRT and decreased
the loss of cell mass by endogenous respiration. This excess production is
still quite low compared to conventional activated sludge however, indicating
that substantial cell mass is still being destroyed by endogenous respiration.
Oxygen Utilization and Nitrogen Conversion
As in Phase 4, oxygen utilization efficiency remained very low, averaging
only 26.4 percent (Table 21). Again, this is attributed to the large quantity
of atmosphere and thus oxygen, lost through leaks in the tent, aeration basin
D.O. increased only slightly to 1.69 mg/1. This low value is the result of a
low oxygen concentration in the atmosphere and the rapid rate of utilization
at the F/M loading utilized.
Although D.O. overaged 1.0, 1.9 and 2.1 mg/1 at the head end, middle,
and effluent end of the aeration basin respectively, zero D.O. conditions also
occurred at the head end, but measurements were only made at the head end on
two days, and on these days zero D.O. conditions were not noted at the middle
or effluent end of the basin.
Oxygen consumption still remained at a very low level, 0.37 after
accounting for oxygen required for nitrification. This is not believed to be
38
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a valid figure and is felt to be influenced by unexplainable discrepancies in
correcting for leaks in the tent.
TABLE 21. PHASE 5 OXYGEN UTILIZATION AND NITROGEN CONVERSION
Oxygen supplied _
kg 02 supplied/nr 0.309
(Ib 02 supplied/106 gal) (2583)
mass 02 supplied/mass BOD removed 3.81
mass 02 supplied/mass T-COD removed 1.49
Oxygen consumption
kg 0? consumed/m3 0.082
(Ib 02 consumed/106 gal) (684)
mass 02 consumed/mass BOD removed 0.97
mass 0? consumed/mass T-COD removed 0.38
i c-
Oxygen Utilization Efficiency, % 26.4
Average Aeration Basin D.O., mg/1 1.69
Ammonia Oxidation, % 78.0
Rate of Nitrification
mass NH3-N oxidized/day/mass MLVSS 0.0470
Alkalinity Consumption
mass alkalinity consumed/mass Nh^-N oxidized 5.22
As discussed, only 78 percent of the ammonia was oxidized. Inhibition of
nitrification reactions is felt to have been caused by the occasional periods
of zero D.O. in the aeration basin as well as occasional periods of zero D.O.
in the secondary clarifiers. Again, alkalinity consumption in the nitrification
reaction was less than stochiometric requirements, averaging 5.44 pounds of
alkalinity per pound of ammonia oxidized.
PHASE 6
Background
Phase 6 began the day following termination of Phase 5, May 15, 1973 and
continued for 119 days until September 10, 1973. In addition to increasing
flow rate to 7064.4 m3/day (1.84 mgd), MLVSS was also decreased to 2,121 mg/1.
At this increased flow rate, detention time was a relativey low 1.7 hrs. The
aeration basin and secondary clarifiers were identical to Phase 5.
-------
The objective of Phase 6 was to determine what effluent quality could be
produced at a relatively low aeration detention time. Another objective was
to try to increase mixed liquor D.O. by feeding more pure oxygen to raise the
percentage of oxygen in the atmosphere under the tent.
Operational Problems
The RAS pump on the rectangular clarifier caused problems on several
occasions in Phase 6. On several occasions, the pump was down for service
for up to a day. Near the end of Phase 6, the pump was found to have a cracked
impeller and it was replaced.
Dumping of septic tank pumping trucks caused a degree of operational
difficulty because the feed rate of pure oxygen had to be increased during
periods they were dumping. It is probable that these shock loads of high
strength septic wastes temporarily decreased the efficiency of treatment.
An extensive program of leak detection and patching was conducted
throughout Phase 6. Unfortunately, this did not cut down on the leakage rate,
and the worst leakage rates of the entire study occurred during the August
and September of this phase. This was attributed to the extremely poor
condition of the tent after a year and a half exposure to the elements.
Bulking in the rectangular secondary clarifier was noted on several
occasions. This bulking was believed to have been caused by density currents.
No action could be taken to eliminate these currents.
On several occasions, delivery of oxygen to the site was late and the
system had to be run without pure oxygen feed for periods of several hours.
Aeration Basin-Secondary Clarifier Efficiency
Substrate removal efficiencies decreased to the lowest levels of any
phase except Phase 2 (Table 22). BOD removal efficiency was only 89.4 per-
cent, while T-COD and S-COD removals were 79.0 percent and 68.0 percent,
respectively. Suspended and volatile suspended solids removal efficiencies
were 86.6 percent and 87.2 percent,respectively.
Ammonia removal, 83.1 percent, was better than in Phases 4 and 5. This
is probably a result of maintaining more consistent D.O. levels in the aeration
basin as a result high feed rates of pure oxygen. At times, however, zero
D.O. conditions continued to be observed, but the frequency of occurrence of
these conditions was less than in Phases 4 and 5.
pH decrease averaged 0.77 units, the largest decrease noted except for
Phase 1. The variation in pH decreases between phases cannot be explained,
but probably is a function of buffering capacity not directly related to
alkalinity of the wastewater.
F/M loadings reached their maximum levels during Phase 6 (Table 23).
On a BOD basis, F/M loading averaged 0.454. This represents an increase of
over six times the F/M loading used in Phase 1, with a corresponding decrease
40
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in BOD removal efficiency from 98 percent to 89.4 percent. On a T-COD basis,
however, F/M loading in Phase 6 was over 21 times that in Phase 1, yet the
percentage T-COD removal was greater in Phase 6. Essentially the same
relationships existed for substrate removal rates between Phases 1 and 6.
TABLE 22. PHASE 6 PRIMARY AND SECONDARY EFFLUENT CHARACTERISTICS
Primary Secondary Percent
Effluent Effluent Change
BOD, mg/1
T-COD, mg/1
S-COD, mg/1
Suspended solids, mg/1
Volative suspended solids, mg/1
NH3-N, mg/1
N03-N, mg/1
Total TKN, mg/1
Alkalinity, mg/1
Turbidity, JTU
pH, units
95.3
239.4
100.0
44.1
37.6
18.3
19.7
221.7
7.4
10.2
50.3
32.0
5.9
4.8
3.1
8.1
6.1
150.9
2.5
6.63
89.3
79.0
68.0
86.6
87.2
83.1
69.0
31.9
Secondary clarifier loading conditions remained very conservative even at
the maximum hydraulic loading. Surface overflow rates were 16.38 and 22.41
m3/day/m2 (402 and 550 gpd/ftr) for the rectangular and circular clarifiers,
respectively. Weir overflow rates were 90.2 and 77.2 nr/day/m (7260 and 6214
gpd/ft) for the rectangular and circular clarifiers, respectively.
The sludge characteristics and the oxygen utilization and nitrogen conver-
sion data are shown in Table 24 and Table 25, respectively.
41
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TABLE 23. PHASE 6 LOADING PARAMETERS
F/M Ratios
mass BOD applied/day/mass MLVSS 0.454
mass T-COD applied/day/mass MLVSS 1.137
mass S-COD applied/day/mass MLVSS 0.478
Volumetric Loadings
kg BOD applied/day/m3 0.996
(Ib BOD applied/day/1000 ft3) (60.31)
kg T-COD applied/day/m3 2.424
(Ib T-COD applied/day/1000 ft3) (151.3)
kg S-COD applied/day/m3 1.013
(Ib S-COD applied/day/1000 ft3) (63.23)
Substrate Removals
mass BOD removed/day/mass MLVSS 0.409
mass T-COD removed/day/mass MLVSS 0.898
mass S-COD removed/day/mass MLVSS 0.325
Rectangular Clarifier Loadings
Surface overflow rate, nr/day/m2 13.97
(gpd/ft2) (343)
Weir overflow rate, m3/day/m 76.94
(gpd/ft) (6195)
Solids loading rate, kg MLSS/day/m2 52.34
(Ib MLSS/day/ft2 (10.72)
Circular Clarifier Loadings
Surface overflow rate, m3/day/m2 13.44
(gpd/ft2) (330)
Weir overflow rate, m3/day/m 46.38
(gpd/ft) (3734)
Solids loading rate, kg MLSS/day/m2 50.19
(Ib MLSS/day/ft2) (10.28)
-------
TABLE 24. PHASE 6 SLUDGE CHARACTERISTICS
MLSS, mg/1 2491
MLVSS, mg/1 2121
Return Activated Sludge TSS, mg/1 8360
Return Activated Sludge Rate, % Q
Rectangular clarifier 37.9
Circular clarifier 41.3
SVI, ml/gm 117
Mean Cell Residence Time, days 32
Sludge Production
mass VSS produced/mass BOD removed 0.269
mass VSS produced/mass T-COD removed 0.122
mass VSS produced/mass S-COD removed 0.331
TABLE 25. PHASE 6 OXYGEN UTILIZATION AND NITROGEN CONVERSION
Oxygen Supplied
kg Op supplied/in3 0.430
(Ib 02 supplied/105 gal) (3586)
mass 02 supplied/mass BOD removed 4.93
mass Og supplied/mass T-COD removed 2.42
Oxygen Consumption
kg 02 consumed/m3 0.208
(Ib 02 consumed/106 gal) (1734)
mass 02 consumed/mass BOD removed 2.38
mass 02 consumed/mass T-COD removed 1.17
Oxygen Utilization Efficiency, % 48.4
Average Aeration Basin D.O., mg/1 1.2
Ammonia Oxidation, % 83.1
Rate of Nitrification
mass NH3-N oxidized/day/mass MLVSS 0.060
Alkalinity Consumption
mass alkalinity consumed/mass NH3-N oxidized 5.73
43
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SECTION 6
DISCUSSION
NITROGEN INTERACTIONS
Conversion of ammonia to nitrates was essentially 100 percent in the
first two phases of the study, but thereafter varied within the range of 68 to
86 percent. This indicates that the growth of nitrifying bacteria was some-
what inhibited in the latter four phases. It is believed that this inhibition
of nitrification resulted from a combination of:
1. low average D.O. conditions in the aeration basin in the
latter three phases,
2. transient zero D.O. conditions in the aeration basin in
the latter three phases,
3. sludge blanket depths between 1.52 and 2.13 m (5 and 7 ft)
in the secondary clarifiers in the latter four phases, and
4. successively shorter detention times with each phase which
may have allowed short circuiting in the aeration basin
during peak flow periods.
The influence of low averaged D.O. conditions upon nitrification in the
study is questionable, since the lowest average D.O. concentration was
1.2 mg/1. It is generally believed that inhibition of nitrifying bacteria
will not occur at D.O. concentrations greater than 1.0 mg/1. However, the
occurrence of transient zero D.O. conditions in the aeration basin during
peak flow periods definitely affected nitrification. Since D.O. measurements
were not made in the secondary clarifier(s) on a continuous basis, it is not
known with certainty whether zero D.O. conditions developed in the sludge
blanket of the secondary clarifiers. Comparison of the nitrate concentration
in the aeration basin with the concentration in the secondary effluent, how-
ever, points strongly to the existence of denitrification in the sludge
blanket, and thus zero D.O. in the sludge blanket during the latter three
phases. Table 26 compares nitrate and ammonia changes through the secondary
clarifier.
This comparison illustrates several important facts. First, the ammonia
concentration decreased in the secondary clarifier in each of the six phases,
indicating some nitrification in the clarifier. This is also demonstrated by
an increase in nitrate in the secondary clarifier in Phases 1, 2, and 3.
Secondly, denitrification occurred in the clarifiers in Phases 4, 5, and 6.
This is demonstrated by a nitrate decrease in the secondary clarifiers. The
44
-------
summation of these two reactions is equal to the net nitrogen change, as shown
in Table 26. In Phases 1 and 2, there was essentially no change within the
limits of normal error. In Phases 3 to 6, it appears that nitrification
occurred in the secondary clarifier, followed by denitrification, resulting in
a net nitrogen loss (as nitrogen gas) from the system. The demarkation
between the nitrification and denitrification reactions most likely corresponds
with depletion of D.O. in the sludge blanket.
TABLE 26. NITROGEN CHANGES IN THE SECONDARY CLARIFIERS
Phase 1 Phase 2 Phase 3 Phase ,4 Phase 5 Phase 6
Nitrate
Aeration Basin 15.70 13.50 7.85 6.61 7.71 9.10
Secondary Clarifier 16.20 13.70 8.00 5.90 6.70 8.10
Change +0.50 +0.20 +0.15 -0.71 -1.01 -1.00
Ammonia
Aeration Basin 0.80 0.16 2.25 5.30 3.79 3.40
Secondary Clarifier 0.40 0.08 1.60 4.70 3.60 3.10
Change -0.40 -0.08 -0.65 -0.60 -0.19 -0.30
Net Nitrogen Change +0.10 +0.12 -0.40 -1.31 -1.20 -1.30
NOTE: All units are in mg/1
A possible correction for this condition in the secondary clarifiers
could have been an increase in D.O. in the aeration basin effluent. However,
when this was tried on several occasions, turbidity of the secondary effluent
would increase and overall effluent quality would deteriorate. In addition
to this correction, return activated sludge pumping rate was also increased.
The increased pumping rates did not alleviate the situation.
Also of interest is the net change in nitrogen content between the
primary effluent and the secondary effluent. In the primary and secondary
effluent, nitrogen is present as total TKN and nitrate. Nitrogen losses from
the system include nitrogen gas produced by denitrification and nitrogen in
waste activated sludge. Table 27 was prepared to compare the estimated amount
of nitrogen wasted from the system as sludge with actual nitrogen lost as
sludge from the system. Phase 1 is not shown in the table because of a lack
of total TKN data and Phase 2 is not included because primary effluent nitrate
concentrations were not measured, and it is felt that they were greater than
zero. For all other Phases, primary effluent nitrate was assumed to be zero.
Although the results of the two calculations don't compare exactly, a
general idea of nitrogen losses in WAS is obtained. The difference between
the two methods is the result of subtracting many numbers of the same general
magnitude, as well as using a cellular composition and assuming that it
remained constant with time.
45
-------
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-------
SLUDGE SETTLEABILITY
The best laboratory measure of mixed liquor settleability is the SVI test.
A variety of activated sludge conditions affect sludge settleability. Princi-
pal among these factors are:
1. volatility of the sludge,
2. nutrient balance in the primary effluent.,
3. formation of nitrogen gas in the sludge due to denitrification,
4. F/M loading rate in the aeration basin.
Sludge volatility is not believed to have been a factor in this study
because it remained within a fairly well defined range throughout. Volatility
of the mixed liquor solids varied from 77 percent to 85 percent. Neither was
nutrient balance in the primary effluent felt to be a factor in the study.
The Tapia facility receives sewage of primarily domestic origin which is
properly balanced with nutrients for the growth of organisms in the activated
sludge process.
Denitrification in the secondary clarifier did have an effect upon sludge
settelability in this study, as previously shown by the denitrification which
occurred in Phases 4, 5, and 6. Whether denitrification occurred within 30
min, the time required for SVI testing, cannot be stated.
F/M loading affects settling by influencing what growth phase the micro-
***** organisms are in at the effluent end of the aeration basin. At low F/M
loadings, microorganisms are in the endogenous or death growth phase and settle
very poorly because of their colloidal nature, low density, and destruction of
the slime layer. At high F/M ratios, microorganisms are in the log growth
phase and contain large quantities of bound water, making them approximately
the same density as the liquid fraction of the mixed liquor. Protozoans and
rotifers are present in small numbers, while free swimming bacteria are com-
prising the majority of the sludge. Poor settling is generally observed
because of the large quantity of free-swimming bacteria. It is generally
recognized that there is an optimum F/M loading between these two growth
phase conditions.
A plot of SVI versus F/M loading (Figure 7) does not indicate the existence
of an optimum F/M loading for this study. A curve is shown in Figure 7, however,
to depict what is felt to define conditions for the study. The curve was drawn
in an effort to place less emphasis on data where SVI is felt to have been in-
fluenced by denitrification. Based upon this curve, optimum F/M loadings appear
to be above 0.5 mass BOD/day/mass MLVSS.
In addition to the effects of the loading characteristics on the highest
SVI's, it is felt that the low rate of sludge wasting had an additional effect
on the SVI. During the periods of time where sludge volume index was higher
than 200,a sample of the sludge was microscopically examined. This examination
revealed the fact that there was a considerable amount of trash dispersed
47
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240
220
200
180
E
05
160
CO
140
120
100
0 0.1 0.2 0.3 0.4 0.5
F/M RATIO, mass BOD applied/day/mass MLVSS
Figure 7. SVI versus F/M loading.
48
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throughout the sludge. The trash consisted of grease particles and inert
materials. This trash buildup seemed to coincide with the instances where
SVI's were considerably high and wasting rates were low. At least three of
these time periods of trash buildup are illustrated on Figure 7 as the highest
SVI points on the graph.
SVI values obtained in this study are higher than those considered to be
"good" in activated sludge processes using air aeration and are considerably
higher than those found in other oxygenation studies. However, it should be
realized that SVI is only an indicator of process performance, and the matter
of ultimate importance is effluent quality. In this study, average secondary
effluent BOD (before chlorination) varied between phases from 2 mg/1 to
10 mg/1, and suspended solids from 4.8 mg/1 to 13 mg/1. This effluent quality
would be considered truly superb for most plants in the United States. How-
ever, the only acceptable effluent quality from the Tapia Water Reclamation
Facility is an effluent with a continuous BOD concentration of less than
5 mg/1 and a suspended solids concentration of 2 mg/1. Because high SVI values
were obtained, this cannot be taken to mean that the process was ineffective,
but only that SVI is a relative parameter and not directly related to plant
effluent quality.
OXYGEN REQUIREMENTS
Provision of oxygen in the activated sludge process is necessary for
cellular synthesis and endogenous respiration. Oxygen consumption in the
process is correlatable with the quantity of substrate removed. Theoretically,
one pound of oxygen is required for the biological stabilization of one pound
of carbonaceous BOD, or slightly more if the microorganisms are maintained in
the endogenous growth phase.
At long mean cell residence times, net oxygen requirements per quality
of substrate removed increased for two reasons. First, the growth of nitrifying
bacteria occurs (assuming other conditions are favorable), requiring oxygen at
a theoretical rate of 4.57 pounds per pound of ammonia removed. Secondly, a
high MCRT allows a longer time for cellular endogenous respiration with a
resultant increase in oxygen consumption.
In this study, oxygen consumption was calculated by an cxygen balance,
according to the formula: Consumption = (pure oxygen added) + (oxygen in
bleed-in air) - (oxygen lost by leaks) - (oxygen in aeration basin effluent).
Results of the oxygen consumption calculations are presented in Table 28.
These figures indicated definite descrepancies in either the measurements or
the assumptions made in the calculations.
This is rather obvious for Phases 2-5 in which, on the average, less than
one pound of 02 was consumed per pound of BODs removed, even though substan-
tial endogenous respiration and nitrification were occurring. Since the cal-
culated estimates for these four phases are in such apparent error, it is
questionable whether the estimates from Phases 1 and 6 can be considered
correct either.
The problem of leaks in the tent compounded the problem of maintaining
D.O. in the aeration basin. A portion of the overall problem was also caused
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by the location of adding bleed-in air to the system. Bleed-in air was added
on the suction side of the recirculation blower, which in turn decreased the
concentration of oxygen in the recirculation stream. Table 28 also presents
average values for air bleed in, recirculation, and percent oxygen in the
atmosphere, and calculated percent oxygen in the recirculation stream.
TABLE 28. OXYGEN DATA
Phase 1 Phase 2 Phase 3 Phase 4 Phase 5 Phase 6
Mass 02 consumed/mass
BOD removed
Total recirculation,
m3/min
(cfm)
Bleed-in air, m3/min
(cfm)
02 in atmosphere, %
Og in recirculation, %
Gas input, m3/m3
(ft3/gal)
3.50
1.08
0.90
0.58
0.97
2.38
18.32
(647.0)
0.21
(7.5)
44.0
43.7
7.48
(i.o)
18.44
(651.0)
0.48
(17.0)
49.5
48.8
3.81
(0.51)
8.51
(300.5)
1.46
(51.6)
39.5
36.3
3.74
(0.50)
6.64
(234.5)
1.64
(57.9)
38.8
34.4
2.39
(0.32)
5.96
(210.3)
1.74
(61.4)
41.1
35.2
2.02
(0.27)
6.54
(231.0)
1.61
(56.7)
54.7
46.2
1.65
(0.22)
The table indicates that bleed-in did not significantly lower the recir-
culation oxygen content in Phases 1 and 2, but in the latter four phases
lowered the recirculation oxygen content by 10 to 15 percent. This decrease
is significant because it decreased the driving force for transfer of oxgyen
to the mixed liquor. This was one factor tending to decrease lower mixed
liquor D.O. concentrations than desired.
A second factor promoting low aeration basin D.O. concentration was the
relatively small amount of recirculation required to maintain a "mixed" con-
dition in the aeration basin. Table 28 presents results of calculations for
gas phase inputs (recirculation plus pure oxygen) per gallon of wastewater.
Thus, as wastewater flow increased, the quantity of gas (pure oxygen plus
recirculation) decreased significantly. This decrease, by over a factor of
four, allowed less gaseous-liquid phase contact and,therefore,less transfer
of oxygen to the mixed liquor.
These two factors, dilution of recirculation with bleed-in and decreased
gas phase input per gallon of wastewater, significantly affected the ability
to maintain sufficient mixed liquor D.O. Unfortunately, the second of these
factors was becoming worse at a time when the F/M loading on the system was
being steadily increased.
50
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EXCESS SLUDGE PRODUCTION
Excess sludge produced in the activated sludge process is the summation
of waste activated sludge, skimmings from the secondary clarifiers, and
suspended solids in the secondary effluent. The quantity of excess sludge
produced is primarily a function of the length of time cell material remains
in the system. As mean cell residence time increases, the growth curve of the
microorganisms proceeds further and further into the endogneous phase of
growth. A natural consequence is loss of cell mass by auto-oxidation.
Figure 8 presents a plot of pounds of excess VSS produced per pound of
BOD removed versus MCRT. Each point represents the average of data for a one
month period, except for Phase 1, which is represented by one point. As
expected, excess VSS production decreased as MCRT increased.
More importantly, however, is the small quantity of solids produced per
pound of BOD removed. Values in this study were about 40 percent less than in
a conventional activated sludge process. Decreased sludge production is not
due to "magical" properties of the Simplex process itself, but only that
aeration with high purity oxygen allows more solids to be carried in the
system, and thus allows operation at longer mean cell residence times. As
Figure 8 shows, operation at these long MCRT results in significant decreased
sludge production.
ALKALINITY CONSUMPTION
In the biological conversion of ammonia to nitrite, and thence nitrate,
natural alkalinity is consumed in addition to ammonia alkalinity. The
following equations represent biochemical reactions consuming alkalinity in
the nitrification process:
2NH4HC03 + 402 ---- 2HN03 + 4H20 + 2C02
2HN03 + CA(HC03)2 ---- CA(N03)2 + 2C02
2NH4HC03 + 402 + CA(HC03)2 ---- CA(N03)2 + 4C02 + 6H20
On a stochiometric basis, 7.18 pounds of alkalinity are destroyed for each
pound of ammonia oxidized.
During this study, two phases had alkalinity reduction greater than
stochiometric requirements, while four had less. Table 29 presents calculated
values for alkalinity consumption.
OPERATION CONSIDERATIONS
As was pointed out in an earlier section of this report, one of the
objectives of this study was to ascertain whether or not savings could be
achieved in capital costs due to the supposedly short durations of aeration
time required for proper treatment. This objective is interpreted by the Las
51
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Virgenes Municipal Water District as being a reduction in aeration tank size
requirements for future expansions of the Tapia Water Reclamation Facility.
The reduction of aeration tank capacity has to be of such magnitude to offset
any additional operation and maintenance costs that may be experienced through
the use of an enriched oxygen system. At this point the reader of this report
should fully understand the basis upon which the District makes their interpre-
tation. On a nationwide basis, it has generally been accepted that the design
parameter for the aeration tank capacity has generally been in the area of
6 to 8 hrs of aeration time. To the Las Virgenes Municipal Water District
this 6 to 8 hrs of aeration time is felt to be excessive. Consequently, in
the design and operation of the Tapia Water Reclamation Facility, the air
activated sludge aeration system was designed for 3.6 hrs of aeration time.
It has been proven at the Tapia facility that this 3.6 hrs is more than
adequate for not only carbonaceoas removal but,additionally, for complete
nitrification. Over and above this, the air activated sludge system at the
Tapia facility has performed superbly at aeration times as low as 2 hrs. The
guideline or the limiting factor for the operation at the Tapia facility is
the effluent quality. The only acceptable effluent quality produced at Tapia
is an effluent with a BOD concentration of 5 mg/1 or less and a suspended
solids concentration of 5 mg/1 or less, a turbidity of one Jackson unit or
less, and a total coliform index of less than 2.2/100 ml.
TABLE 29. ALKALINITY CONSUMPTION
Mass Alkalinity Consumed/
Phase Mass NH3 Oxidized
1
2
3
4
5
6
7.70
9.14
6.44
5.48
5.22
5.73
In view of the above information it is evident that in order for the
Simplex oxygenation system to become a viable alternative for treatment at
the Tapia facility, not only must the aeration time be extremely low but at
the same time the oxygen system must produce a superb quality effluent.
Throughout the entire study, effluent quality remained foremost as a
limiting parameter. That is, operational changes were made to continually
produce the best effluent possible.
From an operational control standpoint the oxygenation system did not
require significant changes in normal operating procedures related to an
activated sludge system. However, operator attention to these procedures
appeared to have a more intense requirement than the normal air system.
Table 30 depicts a comparison of costs associated with the Simplex oxygen
system versus the Tapia air system. Operation labor costs for the oxygen
52
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Z w
O 2
U>.
=) ro
Q TJ
21
0_ u
m §
CD o
Q J;
C/3 CO
(A
(O
03
U
X
0)
(0
(O
CO
Ł
0.40
0.35
0.30
0.25
0.20
0.15
0.10
20 40 60 80 100
MCRT, days
120
Figure 8. Sludge production versus MCRT.
53
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system as shown in Table 30 depict this increase. It was found that more
operator time had to be expended on the liquid oxygen control, conversion
and distribution system than would be required on a normal air compressor for
an air activated sludge system. However, it should be pointed out that in
the $31.00/mil gal on the Og system, considerable labor was expended for the
constant repair of tent leakage. Laboratory control of the oxygen system
appears to be about the same as what is required in an air activated sludge
system. Power, as shown in Table 30, is predominately the power requirement
to operate the mixing aeration blower on the oxygen system. As expected, of
course, this is considerably lower than the power requirement for the operation
of air compressors to support the air activated sludge system. Oxygen costs
as depicted on Table 30 include the cost of delivery of liquid oxygen to the
Tapia site.
TABLE 30. OPERATION COST COMPARISON
Simp!ox System Tapia Air System
Operation labor, $/1000 m3
Laboratory, $/1000 m3
Power, $/1000 m3
Oxygen, $/1000 m3
Total, $/1000 m3
8.19
6.34
3.17
12.68
30.38
6.34
6.08
6.34
0
18.76
NOTE: $/1000 m3 x 3.785 = $/106 gal
The four categories that are used for the cost comparison in Table 30
are felt to be the categories that have a direct effect on the total cost in
comparing the two systems. Items such as chemical costs for chlorine used for
disinfection are equal to both systems, therefore were not used. In addition,
the operation labor category for both systems includes the incidental preven-
tative maintenance on equipment such as adjustments of packing on pumps,
bearing inspections, etc.
Although the data illustrated that there was some reduction in the amount
of excess sludge produced in the oxygen system, it was not felt to be large
enough to warrant reduction in aerobic digester sizing at the Tapia Facility.
Therefore no credit was given to the oxygen system for reduction in sludge
handling costs.
From Hie costs shown on Table 30, it is apparent that the increased
oepration costs for the oxygen system over the air system is predominately
the purchased oxygen. Admittedly, amounts of oxygen were lost unnecessarily
through the tent leakage throughout the study. However, if the oxygen costs
were removed entirely, the related costs are still near the same. Further,
since the Tapia air system can produce an acceptable effluent at an aeration
time of only 2.0 hrs as compared to the oxygen system's 1.7 hrs,the difference
of 0.3 hr for aeration time could not justify a large enough savings in capital
costs to offset the total operation and maintenance costs of the oxygen system.
54
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Additionally, at the 1.7 hrs aeration time, the oxygen system cannot continually
produce the effluent quality needed at the Tapia facility.
CALCULATION PROCEDURES
Because of a lack of continuity by various researchers in the definitions
of several important parameters, it is felt that the definition of certain
parameters should be presented at this point to eliminate any confusion in the
interpretation of results. These key parameters were calculated according to
the following definitions:
Mean Cell Residence Time (MCRT) - Mass of volitale suspended
solids in the system (aeration basin, aeration basin effluent
channel, secondary clarifier sludge blanket, and return
activated sludge line) divided by the mass of volatile suspended
solids wasted per day (volatile waste activated sludge, volatile
skimmings, and volatile solids in the final effluent). Units-
days.
Substrate Removal Rate - The mass of substrate (expressed as
BOD, T-COD or S-COD) removed per day per mass of mixed liquor
volatile suspended solids in the aeration basin. Units-
substrate removed/day/mass MLVSS.
Food - Microorganism Loading (F/M Loading) - The mass of
substrate (expressed as BOD, T-COD or S-COD) applied to the
aeration basin per day per mass of mixed liquor volatile
solids in the aeration basin. Units-mass substrate applied/day/
mass MLVSS.
Aeration Detention Time - The average length of time wastewater
remains in the aeration basin. Calculated by dividing aeration
basin capacity by the summation of primary effluent and return
activated sludge. Units-hours.
The definitions of other terms used throughout this report are presented in the
glossary.
55
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TECHNICAL REPORT DATA
(Please read Instructions on the reverse before completing)
1. REPORT NO.
2.
3. RECIPIENT'S ACCESSIOI>*NO.
4. TITLE AND SUBTITLE
5. REPORT DATE
Simplified Injection of Oxygen Gas Into an
Activated Sludge Process
6. PERFORMING ORGANIZATION CODE
7. AUTHOR(S)
Lloyd D. Hedenland
Ralph L. Wagner
8. PERFORMING ORGANIZATION REPORT NO.
9. PERFORMING ORGANIZATION NAME AND ADDRESS
Las 1/irgenes Municipal Water District
4232 Las Virgenes Road
Calabasas, California 91302
10. PROGRAM ELEMENT NO.
AZB1B, D.U.B-113. Task D-1/14
11. CONTRACT/GRANT NO.
Grant No. S802356
12. SPONSORING AGENCY NAME AND ADDRESS
Municipal Environmental Research Laboratory -- Cin., OH
Office of Research and Development
U.S. Environmental Protection Agency
Cincinnati, Ohio 45268
13. TYPE OF REPORT AND PERIOD COVERED
Final. July 1971-Feb. 1974
14. SPONSORING AGENCY CODE
EPA/600/14
15. SUPPLEMENTARY NOTES
Project Officer: Richard C. Brenner (513/684-7657)
16. ABSTRACT
The Las Virgenes Municipal Water District conducted a pilot investigation of the
Simplex process at their Tapia Water Reclamation Facility in Calabasas, California.
The Simplex process, developed by the Cosmodyne Division of Cordon International,
involves covering an activated sludge aeration basin with an inflated dome and
injecting high purity oxygen into the mixed liquor through a conventional coarse
bubble diffuser.
The purpose of the study was to determine the operational and economic advantages,
if any, of this process over a conventional activated sludge system.
The results of the study indicated that although the Simplex process produced
excellent quality effluent throughout the project by most standards, it was considered
to be economically attractive to the District due, primarily, to the inability of
the process to meet strict local effluent discharge requirements of 5 mg/l each for
BOD and suspended solids and the need for constant repairs to the inflated dome.
17.
KEY WORDS AND DOCUMENT ANALYSIS
DESCRIPTORS
b.lDENTIFIERS/OPEN ENDED TERMS C. COSATI Field/Group
13. DISTRIBUTION STATEMENT
19. SECURITY CLASS (ThisReport)
Unclassified
21. NO. OF PAGES
20. SECURIT.Y CLASS (Thispage)
Unclassified
22. PRICE
EPA Form 2220-1 (9-73)
56
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