EPA-670/2-74-080
DECEMBER 1974
Environmental Protection Technology Series
STUDY OF AN INTEGRATED
POWER, WATER AND WASTEWATER
UTILITY COMPLEX
National Environmental Research Center
Office of Research and Development
U.S. Environmental Protection Agency
Cincinnati, Ohio 45268
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EPA-670/2-74-080
December 1974
STUDY OF AN INTEGRATED POWER, WATER AND
WASTEWATER UTILITY COMPLEX
By
New York State Atomic and
Space Development Authority
New York, New York 10017
Project No. 17080 HHV
Program Element 1BB043
Project Officers
John R. Trax and Patrick Tobin
U.S. Environmental Protection Agency
Washington, D. C. 20460
NATIONAL ENVIRONMENTAL RESEARCH CENTER
, OFFICE OF RESEARCH AND DEVELOPMENT
U.S. ENVIRONMENTAL PROTECTION AGENCY
CINCINNATI, OHIO 45268
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REVIEW NOTICE
The National Environmental Research Center -- Cincinnati
has reviewed this report and approved its 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 consti-
tute endorsement or recommendation for use.
11
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FOREWORD
Man and his environment must be protected from the adverse
effects of pesticides, radiation, noise and other forms of
pollution, and the unwise management of solid waste. Efforts to
protect the environment require a focus that recognizes the
interplay between the components of our physical environment—
air, water, and land. The National Environmental Research
Centers provide this multidisciplinary focus through programs
engaged in
• studies on the effects of environmental
contaminants on man and the biosphere, and
(
• a search for ways to prevent contamination
and to recycle valuable resources.
Distillation has long been known as a method for producing
very pure water. This report discusses how distillation of
wastewater with heat from an electric power plant might be
utilized for producing reuseable water.
A. W. Breidenbach, Ph.D.
Director
National Environmental
Research Center, Cincinnati
ill
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ABSTRACT
This study evaluates, technically and economically, a new
approach to siting power generation, wastewater treatment
and water supply facilities. It is concluded that the
integrated facility results in more efficient utiliza-
tion of land and water resources, produces a net reduc-
tion in undesirable process effluents, and achieves at
a reduced cost many of the environmental quality goals
sought today. In particular, the use of waste heat for
the beneficiation of wastewater treatment was determined
to be sufficiently promising to merit further investiga-
tory research.
The integrated facility studied will supply 1000 MW of
electric power at 9.1 mills/kw-hr, will provide secon-
dary treatment for 50 MGD of wastewater for 15C/1000 gal.,
and will produce 47.5 MGD of high quality potable water
for approximately 62C/1000 gal. utilizing low quality
steam and waste heat.
A three phase follow-on research and demonstration pro-
gram is defined and is directed toward the development
of the further design and performance information neces-
sary to permit the undertaking of full scale integrated
facilities.
This report was submitted in fulfillment of Project
Number 17080 HHV by the New York State Atomic and Space
Development Authority under the sponsorship of the
Environmental Protection Agency.
IV
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TABLE OF CONTENTS
SECTION
I CONCLUSIONS 1
II RECOMMENDATIONS 4
III INTRODUCTION 6
IV REGION AND SITE DESCRIPTION 9
Regional Geography 9
Population Projections 9
Water Resources, Supply and Demand 12
Demand For Electrical Energy 22
Wastewater 23
Site Description 27
V DESCRIPTION OF INTEGRATED FACILITY 32
VI NUCLEAR STEAM SUPPLY. 37
VII POWER GENERATION AND COOLING .
SYSTEM DESCRIPTION 41
VIII WASTEWATER TREATMENT 48
IX DISTILLATION PLANT 53
X OVERALL PLANT INTEGRATION AND EVALUATION 62
Alternative Methods of Wastewater Treatment 64
Utilization of Heat in Wastewater Treatment 66
Wastewater Treatment Heat Balance Analysis. 67
Approaches to Heat Addition 69
Alternative Approaches of Integration 72
Benefits of Integration 73
Comparison of Integrated and Conventional
Separate Facilities 75
v
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SECTION
XI PILOT SCALE DEMONSTRATION PROJECTS 89
Requirements for Design and Operating Data..... 90
Phase I Component Tests and Laboratory
Screening Studies 94
Phase II System Performance Demonstrations..... 96
Phase ill integrated Pilot Scale
Demonstrations •• 98
XII ACKNOWLEDGEMENTS 99
XIII REFERENCES 100
XIV LIST OF PATENTS AND PUBLICATIONS 103
XV GLOSSARY OF TERMS , ABBREVIATIONS 104
APPENDIX A
TEMPERATURE EFFECTS ON WASTEWATER TREATMENT PROCESSES
Contents 107
Figures 1 108
Summary of Findings 112
Introduction 113
Physical-Chemical Processes 115
Sludge Handling Processes 130
Biological Processes 141
Disinfection 167
Advanced Treatment Processes 170
References 183
Glossary 193
APPENDIX B
ENGINEERING AND ECONOMIC FACTORS AFFECTING THERMAL
HEATING OF WASTEWATER 196
Energy Requirements 196
Available Heat Sources 198
Overall Cost of Heat Addition. 209
References 223
Abbreviations and Symbols 225
vi
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FIGURES
NO
1 Reference Site Location 10
2 Nassau County Water Budget 16
3 Present Suffolk county Generalized
Water Budget 17
4 Suffolk County Generalized Water
Budget for the Year 2020 18
5 Reference Site Topography 29
6 Schematic Diagram-Integrated Facility 35
7 Site Plan-Integrated Facility 36
8 Energy Balance-Integrated Facility. 43
9 Wastewater Treatment plant Process
Flow Sheet 49
10 Integrated Facility Overall Heat &
Material Balance Diagram 54
via.
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TABLES
NO
1 Long Island Population Projections 11
2 Suffolk County Town Population Projections 13
3 Projected Water Usage for the Years 1990
and 2020 for Disposal Districts 11 and 13 21
4 Electrical Power Demand for Nassau and
Suffolk Counties 24
5 Population and Wastewater Flow Projections 26
6 District Areas by Township. 27
1 Principal Land Use, Percentages. 36
8 Estimate for 1972 Operations 1100 Mwe Size 39
9 50 MGD Wastewater Treatment Plant Cost
Estimate-Integrated System, Plant Heated
by Barometric Condensers. 52
10 Water Quality Criteria 59
11 capital Cost Distillation Plant 60
12 Distillation Plant Water Production Costs 61
13 Comparison of Unit sizes for Conventional and.
Integrated, Heated 50 MGD Treatment Plants 68
14 Heat Balance Analysis 70
15 50 MGD Wastewater Treatment Plant Cost
Estimate-Conventional System (Noninte-
grated, unheated) 76
16 50 MGD Wastewater Treatment Plant Operating
Costs conventional System (Nonintegrated,
Unheated)- • 77
viii
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TABLES
NO
17 50 MGD Wastewater Treatment Plant cost
Estimate Integrated System, Plant Heated
By Barometric condenser 78
18 50 MGD Wastewater Treatment Plant Annual
Operating Costs Integrated System, Plant
Heated By Barometric Condenser • 79
19 50 MGD Wastewater Treatment plant Cost
Estimate Integrated System, Plant Not
Heated 81
20 50 MGD Wastewater Treatment Plant Cost
Estimate Integrated System, Plant Heated
by Shell & Tube Heat Exchanger. 82
21 Capital Cost - 50 MGD, Integrated Facility
Heat Added by Heat Exchanger 83
22 Comparison of Costs of Various Waste
Treatment Plant Cases 84
23 Conventional Nonintegrated 50 MGD VTE
Distillation Plant 86
24 Integrated Facility Cost Summary. 87
25 Integrated Facility Products and Revenues 88
IX
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SECTION I
CONCLUSIONS
Conceptual designs were developed to determine the technical
and economic feasibility of integrating a 1000 Mwe nuclear
power plant, a 50 MGD wastewater plant utilizing waste heat
to enhance the treatment process, and a 47.5 MGD distilla-
tion plant producing potable water from wastewater, based on
a reference site located on the north shore of Long island
near Riverhead in Suffolk County, New York.
Detailed evaluations of installing and operating the multi-
ple utility complex at that location, and of utilizing waste
heat to elevate the temperature for processing wastewater
resulted in the principal conclusions that:
(a) The integration of facilities for nuclear power
production, wastewater treatment and distillation
of secondary treated wastewater is technically
feasible and economically attractive.
(b) The use of power plant waste heat to improve the
efficiency of the unit operations in primary and
secondary treatment of wastewater is indicated to
be sufficiently practicable to warrant further in-
vestigation and demonstration.
It was determined that the reference design integrated
facility would cost $386 million and would provide electri-
cal power for 9.1 mills/Kw-hr, wastewater treatment for
15.1^/1000 gal. (with no effluent discharge) and potable
water for 62.4/^/1000 gal. Separate facilities having the
same capacities would cost $284 million, $25.5 million, and
$83.8 million for the power, wastewater treatment, and de-
salination facilities, respectively.
The savings of $7,300,000 in capital cost and $179,000 in
annual operating cost (which does not give effect to the ad-
ditional fuel cost savings available) are offered by the
integrated facility as compared to the individual facilities
it replaces. In addition the integrated facility accom-
plishes environmental benefit which the conventional,
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individual facility alternative does not — its wastewater
treatment operations do not release liquid effluents to the
environment.
The cost of distilling secondary treated wastewater should
be less than the cost of desalting seawater in a combination
electric power/seawater distillation plant in view of the
lower concentrations of dissolved material in the treated
wastewater. Moreover, the distillation process is an effec-
tive method of ammonia control, and the use of this process
should preclude the build up of nitrates in recycle or re-
charge systems. Because of the addition of heat, ammonia
can be removed on a year-round basis and is not subject to
the limitations of conventional ammonia strippers. It should
be noted that the distillation of raw and primary treated
wastes is not considered to be technically or economically
feasible based on the present information.
By using waste heat to increase wastewater temperature from
65 F to 93 P, the performance of the grit chamber, the pri-
mary and secondary clarifiers, the aeration tanks and the
sludge thickeners would increase in efficiency by 18 to 30
percent. The improved performance of these processes would
result either from improved settling rates due to the de-
crease in density and viscosity of the wastewater, or from
increase in the biological activity in the organic waste
assimilation processes, or from a combination thereof.
Heating of the untreated wastewater to 93 F by use of steam
from the low pressure stages of the distillation plant, in-
jected by means of a barometric leg condenser — the conser-
vative approach selected for this study — would result in
some additional capital and operating costs since the dis-
tillation and wastewater plants must accommodate the 1.4 MGD
of condensate added to the wastewater. The estimated added
cost of $1,990,000 for this arrangement is of the same mag-
nitude as the $1.5 million saving which could be realized
through the reduction in size of a heated wastewater plant.
The potential economic advantage for heating wastewater lies
in demonstrating (either by verifying the capability of
existing equipment, or by advancing the technology) that
surface type heat exchangers are capable of realizing heat
transfer coefficients in the range of 150 to 350 Btu/hr-ft2-
F, and that such exchangers can be installed in a cost
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range of $12 to $27 per square foot. Under these conditions,
surface exchangers utilizing wastewater could be used for
cooling the product water, thereby eliminating the 1.4 MGD
of additional processing capacity in the wastewater treat-
ment and distillation plants and accordingly reducing the
cost of these facilities by the $1,990,000 indicated above.
In addition, the product water/seawater exchanger which is
estimated to cost $1,869,000 would be eliminated. The mag-
nitude of the saving realized by substituting product water/
wastewater exchangers will depend on the heat transfer co-
efficients and heat exchanger cost reductions achieved in
the demonstration project.
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SECTION II
RECOMMENDATIONS
The economic and technical results of this study indicate
that a research and demonstration program should be under-
taken toward the development of the technology, design and
performance information that will be required to effect
future full scale integration of the facilities for power
generation, wastewater treatment, and water supply. A pro-
gram addressed to these objectives would be conducted in
three phases:
Phase I Laboratory Screening Studies and
Component Tests
Phase II System Performance Tests and Demon-
strations
Phase III Integrated Pilot Scale Demonstration
Tests
The work to be performed under Phase I would consist of two
separate but related studies — a wastewater laboratory
screening study and a heat exchanger component testing study.
The laboratory screening studies to be conducted in Phase I
would be designed to,
(a) Determine actual performance characteristics of
activated sludge treatment operations at the
elevated temperatures considered in this analysis.
(b) Evaluate the qualitative changes in the nature of
the activated sludge treatment process at elevated
temperatures.
(c) Determine the performance characteristics of non-
biological treatment operations at elevated
temperatures, for example, chemical treatment and
chemical sludge disposal.
Laboratory screening results would be analyzed in terms of
4
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performance and economic implications to arrive at a prior-
ity ordering of unit operations for pilot plant verification.
The heat exchanger component testing program would be de-
signed to contribute to the development of surface condensers
and heat exchangers capable of transferring low grade energy
to wastewater from power plant exhaust steam condenser cool-
ing water or distillation plant product water reliably and
economically. To accomplish this objective, tests would be
designed to establish,
(a) fouling factors and overall heat transfer
coefficients;
(b) the corrosion resistance of candidate waste-
water heat exchanger materials;
(c) methods and procedures for the reduction and
removal of scale.
The system performance tests and demonstrations of Phase II
planned on the basis of Phase I laboratory screening results
and component tests, would be undertaken to verify the per-
formance characteristics at elevated temperatures of
selected wastewater treatment processes. The system perform-
ance tests would provide an opportunity for verification of
the suitability of heat transfer equipment identified in the
previous phase. Performance data would be developed on the
ability of distillation plants to remove ammonia and produce
water of high quality reliably and continuously from treated
wastewater. These tests would also establish requirements
for post-distillation treatment and plant product water
quality control. Finally, the tests and demonstrations would
be analyzed to identify performance and economic aspects
relevant to the development of specific process flow sheets
and selection of heat transfer and other equipment for the
Pilot Scale Demonstration.
In Phase III heated wastewater treatment followed by distil-
lation would be demonstrated at the site of an existing
power plant and wastewater treatment plant having a capacity
of up to 4 MGD.
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SECTION III
INTRODUCTION
In many areas of the United States, communities are experi-
encing difficulty in providing adequate supplies of water to
meet the needs of a rapidly growing population. This growth
and an increased per capita water usage are reflected in an
increasing need for sewage treatment facilities as well as
for storm water and domestic sewage conveyance systems all
of which accelerate the depletion of potable water supplies
in regions which are dependent upon ground water.
In the urban areas of this country, the rate of water ex-
traction from our rivers, lakes, and aquifers has exceeded
the rate of replenishment through the natural hydrologic
cycle. Suburban areas in Long Island are experiencing a
growth rate substantially above the national average at a
time when groundwater resources, the only source of supply,
is deteriorating in quality and diminishing in quantity.
Moreover, the rapid rate of growth has created a concomitant
demand for additional power generation.
Recently, considerable public attention has been focused on
the environmental impact associated with the siting of major
facilities such as power plants, reservoirs, and wastewater
treatment plants. Not only is the selection of sites becom-
ing more expensive due to the environmental studies required
and the inclusion of more stringent environmental safeguards,
but in many areas of the country, the prospects for finding facility
locations are extremely limited. Suburban development, in-
dustrial activities, recreation facilities and other compet-
ing uses for land have already claimed many prime facility
sites.
Historically, organizations responsible for siting major
public and private facilities have proceeded in an independ-
ent manner. For example, sanitary districts seek and ac-
quire sites for wastewater treatment plants, water utilities
locate reservoirs and treatment plants, and electric
•utilities acquire sites for future generating stations. In
each case, the specific environmental impact of waste prod-
ucts must be considered. Federal, state and local standards
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now regulate the quality of air and water discharges from
the separate facilities. The cost of adequate effluent
treatment is passed on to the eventual consumer of the prod-
uct or services.
The hypothesis that a new approach to utility siting, that
of jointly siting power generating, wastewater treatment and
water supply facilities, results in a more efficient utili-
zation of land and water resources, a net reduction in un-
desirable process effluents, and the accomplishment of many
of the new environmental quality goals at a reduced cost
when compared to separate facilities, is the subject of this
study.
Thus, the overall objective of the project was to determine
the feasibility of combining electric power production,
wastewater treatment, and potable water supply utilities,
and to evaluate the economic and environmental advantage of
such a combined facility for treating wastewater for reuse
as an alternate water supply. The specific objectives were:
1. To define the technical considerations involved in the
integrated siting of power generation, wastewater treat-
ment, water supply and distillation facilities.
2. To select an illustrative site and prepare a design of
an integrated facility.
3. To determine the effectiveness of the integrated facil-
ity in meeting future water demand.
4. To establish the principal design and performance aspects
associated with integrated utility complexes.
5. To establish the economic benefits and costs of a full
scale integrated facility.
6. To identify the beneficial effects of heat addition in
wastewater processing.
7. To define the necessary development and demonstration
projects.
In order to determine the benefits associated with an in-
tegrated facility, an engineering and economic analysis was
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performed. The power production, wastewater treatment and
water supply systems were analyzed to identify the most
effective mode of overall integration. The integrated
facility was compared to the conventional method of using
individual facilities for power, wastewater and water ser-
vices. Since the design and economic evaluation of both
the integrated facility and the individual facilities was
strongly site dependent, a specific location was identi-
fied.
The New York State Atomic and Space Development Authority
undertook this study pursuant to its responsibilities for
conduct and fostering the use of atomic energy for produc-
tive purposes. The Authority was joined in this work by
environmental science and engineering firm of Quirk,
Lawler and Matusky, which was responsible for the analyt-
ical and design work and economic analysis of the waste-
water treatment systems of the integrated plant discussed
in Sections V, VIII and X, as well as for the determination
of the effect of elevated temperatures upon wastewater
treatment processes work presented in Appendix A and by
Hittman Associates, Inc. The latter firm reviewed and
assessed the information presented in Section IV, per-
formed the analytical design and economic calculations
for the nuclear power facilities presented in Sections V
through VII, the distillation plant design presented in
IX, those portions of the overall plant integration and
evaluation pertinent to their work, as presented in Sec-
tion X, and the evaluation of engineering and economic
factors affecting addition to wastewater treatment pro-
cesses, as presented in Appendix B.
Section XI, which sets forth the development and recommended
further demonstration projects, was prepared jointly by
Quirk, Lawler and Matusky, Engineers and Hittman Associates.
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SECTION IV
THE REGION AND ITS RESOURCES
REGIONAL GEOGRAPHY
Suffolk County, the easternmost county on Long Island,
covers a land area of approximately 920 square miles. The
county is bounded on the north by Long Island Sound, on the
east and south by the North Atlantic Ocean, and on the west
by Nassau County. Its Nassau County boundary is within 15
miles of the city limits of New York City and within 30
miles of central Manhattan.
The county is approximately 86 miles long and 21 miles wide
at its widest point which is along its western boundary.
The major land mass extends east-northeast from the Nassau
County line for 42 miles to Riverhead; east of Riverhead
the land mass is bifurcated into peninsulae extending east-
ward and separated by a series of bays, all shown on Figure
1, which also identifies the reference site location. The
north fork, extends approximately 28 miles east of River-
head; the larger southern fork is approximately 44 miles
long and terminates at Montauk Point, the easternmost point
of New York State.
POPULATION PROJECTIONS
Suffolk County is composed of 10 towns ranging in size from
11 square miles (Shelter Island) to 252 square miles (Brook-
haven) . Until recent years, the county was almost entirely
agricultural in nature, but now agricultural uses predomin-
ate only in the eastern areas. In recent years the increas-
ing population encroachment from Nassau County and extension
of the continuing growth surrounding the New York metropoli-
tan area have resulted in rapid increases in population and
population density in the western part of Suffolk County. A
70 percent increase in the total population of the county,
from 667,000 to 1,127,000, was observed between 1960 and
1970 (1). Approximately 97 percent of this increase was ob-
served in the five western towns which, in 1970, were found
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\SMITHTO WN\
P-A
' I5LIP
BROOKHAVEN
SUFFOLK
Long IsSond
Figure 1. Reference Site Location
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to have a total population of approximately 1,043,000, or
approximately 93 percent of the total in the county situated
on approximately 37 percent of the total county land area
(2).
Although the population growth rate in the United States
has shown a decrease over the last decade and this trend is
expected to continue, the population of Suffolk County is
forecasted to grow at rates substantially higher than the
national average. The dominant factor in this growth will
be the increasing scarcity of suitable land for new home
sites closer to the New York metropolitan area, with the
most rapid growth occurring in the westernmost towns.
TABLE 1
LONG ISLAND POPULATION PROJECTIONS
(In Thousands)
Nassau County Suffolk County
Year (Ref.4) (Ref.3) (Ref.4)
1970 1429 1127 1127
1975 1565 1276 1328
1980 1651 1515 1570
1985 1723 1753 1825
1990 1743 1978 2190
1995 1791 2198 2360
2000 1836 2379 2500
2005 1879 2645
2010 1928 2790
2015 1977 2920
2020 2021 3050
Table 1 presents population projections for both Suffolk
County and adjacent Nassau County. These data are based
primarily on a report commissioned by the Nassau-Suffolk
County Regional Planning Board (4). New York State data
(3) for the years 1970-2000 for Suffolk County are also
11
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included for comparison purposes. Close agreement may be
noted, with slightly lower population growth estimated by
the state. Comparable data were also obtained from a re-
port prepared in evaluation of sewerage needs for eastern
Suffolk County (2). This report indicated a permanent pop-
ulation of approximately 2,046,000 in Suffolk County in the
year 1990 and 2,960,000 in the year 2020. As there is no
significant conflict between these projections, the data
from the County Regional Planning Board were utilized
throughout this report in the development of further analy-
ses of power and water requirements based on population
growth.
In addition to the overall county population information,
estimates of the population in each of the ten individual
Suffolk County towns for the years 1970 through 2020 were
extracted from Reference 4 and are presented in Table 2.
Current New York State projections for the years 1970 to
2020 are presented parenthetically for comparison purposes
(3). Each town covers an extensive area and is comprised
of a number of small villages and hamlets.
WATER RESOURCES, SUPPLY AND DEMAND
Groundwater constitutes virtually all the water supply
available to Suffolk County. Of the 44 inch average annual
precipitation, approximately two percent is lost through
direct runoff and 48 percent returns to the atmosphere via
evapotranspiration, leaving approximately 50 percent of
the average annual rainfall available for groundwater re-
charge. This natural recharge corresponds to about one
million gallons per day per square mile.
The surface water supplies are quite meager since there are
no major lakes within its boundaries. The largest body of
water is Lake Ronkonkoma which covers an area of only 245
acres, and most of the streams are estuarine; that is, salty
in their lower reaches where they flow into bays or directly
into the ocean or sound. It is estimated that 95 percent of
the fresh water streamflow is due to discharge from the
groundwater reservoir, with the rest being direct runoff
(5).
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TABLE 2
SUFFOLK COUNTY TOWN POPULATION PROJECTIONS
(In Thousands)
Town*
Babylon
Hun ting ton
Islip
Smith town
Brookhaven
Southampton
Riverhead
East Hampton
Southold
Shelter Island
Total Suffolk
1970
(10/5)
(18/1)
(16/3)
( 9/0)
(25/1)
(12/0)
( 1/0)
( 3/0)
( 3/0)
( 1/0)
County
200
200
276
115
235
40
19
17
18
2
1127
( 204)
( 201)
( 279)
( 114)
( 244)
( 36)
( 19)
( ID
( 17)
( 2)
(1127)
1980
262
280
358
160
375
65
30
39
27
3
1600
( 243)
( 252)
( 353)
( 167)
( 395)
( 44)
( 26)
( 14)
( 20)
( 2)
(1515)
1990
283
315
378
170
700
116
63
70
42
3
2190
( 281)
( 311)
( 400)
( 208)
( 626)
( 59)
( 39)
( 20)
( 30)
( 4)
(1978)
2000
292
325
385
175
840
170
130
110
73
4
2500
( 327)
( 354)
( 456)
( 239)
( 800)
( 76)
( 53)
( 26)
( 41)
( 7)
(2379)
2010
298
330
400
179
890
224
198
142
120
8
2800
2020
300
333
405
181
920
267
260
165
214
12
3050
*-The first number in parenthesis after the town indicates the number of villages or
hamlets within the town and the second number indicates the number having a popu-
lation of over 15,000.
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The earliest public water supply systems on Long Island de-
pended on surface water for their source of supply. When
the original Brooklyn system was completed in 1862, the sup-
ply consisted entirely of surface water with gravity flow
distribution. In 1872 the first Brooklyn pump stations were
installed. This second water supply system on Long Island
augmented the original system by pumping water from ponds.
As the meager surface water supplies diminished, the growing
population turned to groundwater for its increased needs. In
1874 the first groundwater for public supply was used on
Long Island by Long Island City. By 1880 Brooklyn also began
exploiting groundwater. Thus, by 1902 there were 120 MGD
being used for public water supply on Long Island, 65 MGD
from surface sources and 55 MGD from the ground. Brooklyn
alone required 85 MGD, and obtained 60 MGD from surface sup-
plies and 25 MGD from wells.
As the western part of Long Island developed rapidly, the
increased impermeability of the urbanized areas reduced the
infiltration capacity, while establishment of sanitary and
storm sewers resulted in removal of water which would normal-
ly recharge the aquifers. As a result, groundwater mining
became commonplace. In the 1930's, excessive pumpage in
Brooklyn lowered local groundwater levels to as much as 35
feet below sea level. This condition caused salt water to
contaminate a large portion of the groundwater reservoir,
and forced a suspension of further withdrawals. Recovery
from this condition of depletion is extremely slow due to
the high degree of urbanization and the construction of sewer
outfalls to the ocean.
With the eastward migration of population on Long Island, the
attendant water supply problems also move eastward. Recent
estimates (4) of consumptive use and of the groundwater re-
sources of Nassau County show that there is an average annual
decrease in the groundwater storage in the county of 5 MGD
due to groundwater mining. Consumptive use is defined as the
rendering of water unavailable for reuse until such water has
passed through the precipitation part of the hydrologic cycle.
Consumptive use includes water discharged through sewer sys-
tems to the sea, infiltration of groundwater into the sewer
system, evapotranspiration following irrigation and lawn
sprinkling, and the relatively small amount consumed and
14
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evaporated during general use. Figure 2 shows the water bud-
get for Nassau County (5).
A comprehensive study of the public water supply and require-
ments of Suffolk and Nassau Counties was recently completed
and published (4) which indicated a present per capita water
usage in Suffolk County of approximately 100 gallons per day
(gpd), and a future per capita usage of approximately 150
gpd by the year 2020. On the basis of the present and pro-
jected population, the requirement for water exclusive of
agricultural uses is estimated to be approximately 120 MGD
at the present time and approximately 450 MGD by the year
2020
In general, Suffolk County has not felt the acute water sup-
ply problems caused by development and increased population.
Its water supply is sensitive to precipitation levels and
during periods of extended drought of the type which occurred
in the mid-19601s. Such problems could be quite serious for
many of its communities. Figure 3 shows the present gener-
alized water budget for Suffolk County. As the population
of the county grows, however, large increases in the amount
of human consumptive use will occur, due, in part, to the
general increase in population, in part to the increase in
the amount of sewered area, and in part to the general trend
toward increased per capita use which accompanies increased
personal income. Even if no additional sewers are con-
structed, consumptive use is expected to approach 100 MGD by
the year 2020 (3, 4) .
Although the Suffolk County water budget for the year 2020
is somewhat a matter of conjecture at this time, some gen-
eralized conclusions can be drawn with the aid of Figure 4.
Population estimates place the population of Suffolk County
in the year 2020 at about 2 1/2 times the present population.
Assuming a conservative increase in the amount of storm
sewer systems, direct runoff would probably increase to at
least 100 MGD. With evapotranspiration remaining constant,
the amount of precipitation available for groundwater re-
charge would, necessarily, decrease. Human consumptive use
is shown to increase significantly by the year 2020, as a
result of a more than fourfold increase in the area serviced
by sanitary sewers.
15
-------
TOTAL INFLOW-
PRECIPITATION
404 MGD =43.5
DIRECT '
RUNOFF
4 MGD
GROUNDWATER
RECHARGE
210 MGD
EVAPO-
TRANSPIRATION
190 MGD
Human
Consumptive
Use of Ground
Water
45 MGD
Subsurface
Outflow
83 MGD
Ground Water
Flow to
Streams
76 MGD
Spring Flow
5 MGD
Change in
Storage
5 MGD
(Decrease)
Ground
Water
Evapotran-
spiration
6 MGD
TOTAL OUTFLOW
409 MGD
Figure 2.
Nassau County Water Budget
(Based on Ref. 5)
16
-------
TOTAL INFLOW-
PRECIPITATION
1460 MGD
J I
DIRECT
RUNOFF
70 MGD
GROUNDWATER
RECHARGE
690 MGD
EVAPO-
TRANSPIRATION
700 MGD
HUMAN CONSUMPTIVE
USE OF GROUNDWATER
35 MGD
GROUNDWATER
OUTFLOW TO STREAMS,
SPRINGS, ETC.
280 MGD
SUBSURFACE
OUTFLOW
375 MGD
TOTAL OUTFLOW
1460 MGD
Figure 3. Present Suffolk County Generalized Water Budget
17
-------
TOTAL INFLOW-
PRECIPITATION
1460 MGD
I
DIRECT
RUNOFF
100 MGD
GROUNDWATER
RECHARGE
660 MGD
HUMAN CONSUMPTIVE
USE OF GROUNDWATER
150 MGD
GROUNDWATER OUTFLOW TO
STREAMS, SPRINGS, ETC.
210 MGD
TOTAL OUTFLOW
1460 MGD
EVAPO-
TRANSPIRATION
700 MGD
SUBSURFACE
OUTFLOW
300 MGD
Figure 4. Suffolk County Generalized Water Budget for the
Year 2020.
18
-------
A comparison of Figures 3 and 4 shows that by 2020 the pro-
jected increase in consumptive use of water will reduce sub-
surface outflow and groundwater outflow to streams and
springs. This increased consumptive use could also result
in depletion of the groundwater reserve as is occurring in
Nassau County. Thus, the 20 percent decrease in subsurface
outflow and the 25 percent decrease in the groundwater out-
flow to streams and springs would have a profound effect on
the groundwater balance. Accordingly, streamflow will show
a marked decrease, especially during periods of drought, as
was evidenced in Nassau County during the drought period of
mid-1960's by the zero flow condition recorded in some
streams in areas that were predominantly sewered (4). As
subsurface outflow and strearaflow decrease or as ground-
water reserves are directly depleted, a marked increase in
the amount of saltwater encroachment will also occur, further
deteriorating the water supply situation.
The safe water yield in Suffolk County in 1957 was estimated
to be approximately 501 MGD (7), which would, if such yield
could be sustained in the face of the currently projected
development, meet the average per capita demand of 150 gpd
for all purposes for a population of 3,340,000. Considering
the uncertainty in such yield and population projections,
the 13-17 percent increase in summer population over that of
winter, and the recognized variation in peak daily and sea-
sonal demands from the annual average, the sustained safe
groundwater yield may well be exceeded by the demand during
the period under consideration (1970-2020). Even if the
county-wide yield is not exceeded by the demand, the antici-
pated large imbalance in the population distribution will
necessitate an extensive water distribution system providing
either groundwater from the central portion of the county or
water from sources outside the county for the western towns.
It must be noted, however, that Suffolk County is the only
county on Long Island where demand does not already exceed
supply. Consequently, if an additional source of water for
the county is to be found it must be imported from the
mainland (i.e., from upper New York State through the New
York City metropolitan water system), through desalting of
seawater, or through recovery and reuse of wastewater which
would otherwise be discharged to the Atlantic Ocean or to
Long Island Sound.
19
-------
An additional consideration in the water supply picture of
Suffolk County is the continued pumping and groundwater min-
ing in adjacent Nassau County. If the water levels in Nassau
continue to decline, Suffolk County will eventually feel the
impact of reduced groundwater storage since hydrologic con-
ditions do not respect political boundaries.
If water now lost through human consumptive use can be re-
cycled through wastewater treatment and distillation, a
significant step will have been taken toward meeting future
water supply needs. Probably the most critical area within
Suffolk County is the north fork, containing the town of
Southold and about half the town of Riverhead. Here, the
fresh groundwater reservoir is restricted to relatively thin
lenses in the glacial or Upper Pleistocene deposits. With-
drawal of water from these lenses is thus limited to shallow
wells of low capacity and, therefore, well spacing and dis-
tance from the sea are of great importance.
The towns of Riverhead and Southold, closest to the proposed
site, have present water requirements of approximately 3 MGD
each. These requirements are anticipated to increase by a
factor of three to four by the year 2000 and to further in-
crease by an additional factor of approximately three in the
following two decades. Thus, depending on the population
growth, the water requirements in each of these towns by the
year 2000 and 2020 are expected to be on the order of 30 MGD
and may exceed 50 MGD during peak periods.
The projected population increases and groundwater with-
drawals, the greater use of sewers, the additional runoff
and higher evapotranspiration due to development may be ex-
pected to lower the area's groundwater level and quality.
Seasonal demands currently result in heavy withdrawals of
groundwater, such as in summer months, when both agricul-
tural and summer resort residential demands increase sig-
nificantly. Agricultural pumpage has already created local
problems of saltwater intrusion.
Moreover, during periods of drought such as that experienced
in the 1960's, local lowering of the water tables occurs.
The low groundwater levels and reduced surface flow under
such conditions have resulted in some water quality problems
locally. In addressing this projected need, the water
20
-------
produced from the proposed integrated facility can be con-
sidered as a new source of fresh water. The wastewater from
which it is derived would normally be discharged to the sea.
By recycling this water, the supply of fresh water in the
region is augmented and the other sources of fresh water are
conserved and protected. The product water from the plant
could be used to recharge the groundwater aquifers or for
industrial purposes. Alternatively, distilled water product
could be made available for distribution. In this case
essentially 100 percent of the recovered water would be put
to use, whereas if used for groundwater recharge approxi-
mately 50 percent would be lost. In any case, the quality
of the water produced by distillation would be suitable for
either recharge or direct reuse.
TABLE 3
PROJECTED WATER USAGE
FOR THE YEARS 1990 AND 2020
FOR DISPOSAL DISTRICTS 11 AND 13*
i don
Popu_lation
229,000
84,000
313,000
Total
Water
Usage
(MGD)
59.6
19.0
78.6
9D9O
£,\J £t\J
Population
357,000
218,000
575,000
Total
Water
Usage
(MGD)
61.2
29.5
90.7 '
District
11
13
Total
*-Based on Reference 4.
Table 3 shows the projected total water usage for Disposal
Districts 11 and 13, which would supply wastewater to the
integrated facility for the years 1990 and 2020. If an in-
tegrated facility should be operational by 1990, about 61
percent of the average annual water requirements for these
districts could be recovered by the integrated plant. By
2020, when the average annual requirement has increased to
90.7 MGD, the plant recovery would represent 53 percent.
21
-------
A portion of the product water would be used in Suffolk
County to reduce withdrawals in critical areas and to pro-
vide adequate reserves for drought periods through reduction
in demand on groundwater.
In the near-term one potential use for the product water
would be in neighboring Nassau County, where there exists a
groundwater mining situation such that new sources of fresh
water in the amount of 94 MGD in 1990 (6) will be required
to prevent saltwater infiltration and/or encroachment. Ex-
porting water to Nassau County will also benefit Suffolk
County since the lowering of the groundwater levels in Nas-
sau County would cause a lowering of the groundwater in
Suffolk County. Secondary treated wastewater could also be
used to supply make-up water for a power plant, which at the
level of 1000 Mwe requires approximately 18 MGD of make-up
water.
DEMAND FOR ELECTRICAL ENERGY
While the population of the United States has been growing
at a relatively stable rate of about 25 million persons per
decade over the past 20 to 30 years, the demand for electri-
cal power has grown at a rate of about eight percent per
year, doubling each decade. This phenomenon of dispropor-
tionate growth of population and electrical energy consump-
tion is partially attributable to greater use of air condi-
tioners, major appliances, television, commercial and street
lighting, and a greater power consumption in the industrial
sector. On a per capita basis, electrical power use is ex-
pected to increase over the next 40 years (8, 9). For Long
Island, assuming the per capita electrical power demand will
approximate the previous historical trend, and further as-
suming a relatively constant plant load factor (actual pro-
duction of power divided by maximum possible production) of
about 53 percent, it is possible to project the area power
demand and the installed capacity needed to meet that demand.
The Long Island Lighting Company's relatively low plant load
factor (53 percent as compared to the national average of 64
percent) is attributed to the population variability caused
by an influx of transients during the summer peak load
period, and to the non-industrial character of the service
area.
22
-------
Electrical power requirements for Nassau and Suffolk Counties
(10, 11) are shown in Table 4. The per capita use of elec-
trical power in Nassau and Suffolk Counties was, in 1970,
only about half the national average. Projected per capita
power use, based on Edison Electric Institute and other pro-
jections for national averages, was accordingly reduced by
50 percent to correspond to the existing situation on Long
Island. The plant factor was allowed to increase gradually
from the 1970 value of 53 percent to 56 percent in 2020 in
order to account for expected improvement in system efficien-
cies .
The 1970 generating capacity for Long Island was approximate-
ly 2330 megawatts. Subsequently, gas turbine capacity of 117
megawatts and 386 megawatts of fossil capacity were added in
1972 . The Long Island Lighting Company has scheduled the
Shoreham nuclear plant of 820 megawatts to enter service in
1976. These additions of 1323 megawatts will bring the sys-
tem capacity to 3653 megawatts, substantially under the pro-
jected requirement for 1980 of 4740 megawatts. The differ-
ence will be accounted for by the construction of plants not
yet announced or by the importation of power from interties
with the New York Power Pool and the New England Power Ex-
change .
WASTEWATER
Latest data on wastewater flows in the Suffolk County area
indicate that present requirements for wastewater treatment
are on the order of 75 to 100 gallons per capita day (gpcd)
(2) . The lower figure represents current estimates of domes-
tic and minor commercial consumption; the higher value in-
cludes an allowance for infiltration of groundwater into the
sewer system. On this basis, if the County were to be
entirely sewered, domestic wastewater flow would approximate
250 MGD by the year 2000, 300 MOD by the year 2020. In
order to provide for treatment of wastewater resulting from
increased industrial development such as that observed at
present in western areas of Long Island (i.e., light "dry"
industries such as electronics, metal fabrication, plastics,
etc.), and for typical peak-to-average domestic consumption
ratios on the order of 1.5, the wastewater treatment
23
-------
TABLE 4
ELECTRICAL POWER DEMAND
FOR NASSAU AND SUFFOLK COUNTIES**
Year
1970
1975
1980
1985
1990
1995
2000
2005
2010
2015
2020
Population
2,556,
2,893,
3,221,
3,548,
3,933,
4,151,
4,336,
4,524,
4,718,
4,897,
5,071,
000*
000
000
000
000
000
000
000
000
000
000
Per Capita
Use
Kw*
0.43
0.60
0.78
0.97
1.20
1.40
1.60
1.80
2.00
2.25
2.50
1
1
2
3
4
5
6
8
9
11
12
Total
Use
(Mw)
.10 x
.74 x
.51 x
.44 x
.72 x
.81 x
.94 x
.14 x
.44 x
.02 x
.68 x
103*
103
103
103
103
103
103
103
103
103
103
Plant Load
Factor
0
0
0
0
0
0
0
0
0
0
0
.53*
.53
.53
.54
.54
.54
.55
.55
.55
.55
.56
Required Capacity
Mwe*
2
3
4
6
8
10
12
14
17
20
22
.33 x
,28 x
.74 x
.37 x
.74 x
.76 x
.62 x
.80 x
.16 x
.04 x
.64 x
103
103
103
103
103
103
103
103
103
103
103
*-Actual
**-Based on Long Island Lighting Company data, (10, 11)
-------
capability must be at least twice the average domestic flow,
or approximately 500 MOD by the year 2000, and 600 MOD by
the year 2020.
The wastewater collection areas under consideration for sup-
ply correspond to sewerage Districts 11 and 13 described in
Reference 2. These districts are approximately coincident
with the surface drainage patterns.
District 11 encompasses an area of approximately 114 square
miles and has a projected 1990 sewered population of 229,000
or an average density of about 3.1 persons per acre. This
district is forecast to have a 1990 wastewater yield of ap-
proximately 40 MGD, or about 175 gpcd, including industrial,
commercial, residential, and extraneous sources.
District 13 encompasses an area of approximately 27 square
miles and has a projected 1990 sewered population of 85,000
or an average density of about 5 persons per acre. This
district is expected to have a 1990 wastewater yield of ap-
proximately 9 MGD or about 107 gpcd including industrial,
commercial, and residential sources. Tables" 5 and 6 sum-
marize the available data for the districts under consider-
ation.
An integrated utility complex could therefore expect approx-
imately 50 MGD from Districts 11 and 13 by 1990. Additional
wastewater could be obtained from District 12, thus provid-
ing another 12 MGD, if the transportation cost were justi-
fied. Approximately 7 of a total of 17 miles of required
trunk sewer could constitute a common system with the llth
District; such an arrangement would permit reaching 50 MGD
prior to 1990.
Wastewater collection within the districts is planned to be
primarily by gravity flow to one or two collection loca-
tions. The District 11 gravity collection location is
planned for the Village of Riverhead, approximately at the
junction of the Peconic River and Flanders Bay, and is about
7 miles from the proposed integrated facility. The static
head between the collection location and the proposed plant
site is estimated to be about 60 feet. Two gravity collec-
tion locations are planned for District 13; one at Mattituck
and one at Greenport. The piping distance from Greenport to
25
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TABLE 5
POPULATION AND WASTEWATER FLOW PROJECTIONS
N)
District
Number1 Year
•j
Population^
Average
Daily
Flow
(MGD) 2
11
13
12
1990
2020
1990
2020
1990
2020
229,
356,
84,
218,
70,
91,
000
990
000
300
000
400
39
61
8
18
12
16
.59
.72
.85
.98
.30
.55
Peak
Daily
Flow
(MGD)3
79
123
17
37
24
33
.18
.44
.70
.96
.60
.10
Minimum
Daily
Flow
(MGD) 3
24
37
5
11
7
10
.8
.0
-3
.4
.4
.0
Estimated
Distance
(miles) ^
7
7
15
15
17
17
Static
Head
(feet)
60
60
60
60
60
60
1 - District numbers correspond to the districts established by Bowe, Albertson, and
Walsh in Reference 2.
2 - Based on Reference 2.
3 - Assumed peak-to-average flow of 2:1 and minimum-to-average flow of 0.6:1.
4 - Estimated piping distance from the point of district wastewater collection to
the proposed treatment plant site.
-------
Brook-
haven
42
River-
head
50
South-
old
__
South-
ampton
22
TABLE 6
DISTRICT AREAS BY TOWNSHIP1 (Square Miles)
District
Number haven head old ampton Total
11 42 50 -- 22 114
13 — ~ 27 — 7
12 — — — 25 25
1 - The sewage district boundaries are approxi-
mately the same as the surface water drain-
age divides, thus the surface drainage areas
are roughly equivalent to the sewage district
areas. The area figures are taken from
"Comprehensive Public Water Supply Study,
Suffolk County, New York, Vol. II, 1970" by
Holzmacher, McLendon, and Murrell Consulting
Engineers. (4).
Mattituck is approximately 12 miles and from Mattituck to the
site is 3 miles, with a total increase in elevation of ap-
proximately 60 feet. Pumping and force main facilities would
be required to transport wastewater to the treatment plant.
Such facilities should have the capacity to deliver peak
flows reliably as well as to maintain minimum velocities at
low flow conditions to prevent the settling of solids.
SITE DESCRIPTION
The reference site for the integrated facility is a tract of
land measuring approximately 4400 by 5200 feet (about 500
acres), located adjacent to the Southold Town Line at the
northeastern corner of Riverhead Town, Suffolk County, New
York. The site is at latitude 40° 59'30" North and at
longitude 72° 36' West. The tract is bounded by Sound Ave-
nue to the south, the Southold-Riverhead Town Line to the
east, the Camp Carey access road to the west, and Long Is-
land Sound to the north. The site encompasses almost all
of the abandoned Camp Carey, and approximately 10 private
27
-------
residences. Major natural features of the site include Lily
Pond, Hallocks Pond, and Jacobs Hill. With the exception of
the range of sand hills directly adjacent to the shoreline,
the site is relatively flat, averaging about 65 feet in
height above sea level. A topographic map of the site is
presented in Figure 5.
Since a nuclear plant is the choice for power generation,
the site must be evaluated in terms of the Atomic Energy
Commission criteria. These criteria require, among other
things, a thorough demographic analysis of the region sur-
rounding a potential site.
In the near vicinity of the site,, population densities are
quite low, ranging from about 0.5 to 3.0 persons per acre of
land. Further west, toward the Nassau County line in the
Town of Babylon, population densities reach 11.0 or more
persons per acre, reflecting the spread of suburban communi-
ties eastward from New York City.
Within a 10-mile radius of the site, the largest community
is Riverhead, with a 1970 population of 7585 permanent
residents, and an anticipated growth to approximately 10,000
by 1980. This community is approximately six miles south-
southwest of the site. The area within a 10-mile radius of
the site is approximately 58 percent water.
Within a 20-mile radius of the site, the largest community
is Yaphank in the Town of Brookhaven with a 1970 population
of 8793 residents. The boundary of this community is approx-
imately 20 miles west-southwest of the site. Neither River-
head nor Yaphank are expected to have a 1980 population ap-
proaching 25,000 residents, the figure stipulated for a
"Population Center" under the AEC criteria. Therefore, on
Long Island, the distance to a Population Center is in ex-
cess of 20 miles, which represents a very conservative
value. The approximate distance between the site and any
point on the coast of Connecticut is also 20 miles. The
distance to the nearest community of 25,000 or more residents
is approximately 26 miles to East Haven, Connecticut with a
1970 population of 25,120 or New Haven, Connecticut with a
population of 137,707 in 1970.
28
-------
SOUTHOIX) P?C°NIC-/ BAYV.EW
NOMTH BRIDeEmMPTON
AREA
Figure 5. Reference Site Topography
29
-------
Land use in Suffolk and Nassau Counties has been treated in
detail by the Nassau-Suffolk Regional Planning Board (18) .
The information reported therein supports the conclusion
that the site is surrounded by a sparsely populated area, as
inferred from the population figures.
Land within a 10-mile radius of the site is within the Towns
of Riverhead, Southold, and Southampton. Principal land use
percentages for these towns and Suffolk County as a whole
are shown in Table 7.
TABLE 7
PRINCIPAL LAND USE, PERCENTAGES
Categories
of Use
Residential
Commercial/
Industrial
Institutional
Agricultural
Recreational
Vacant
(including waters)
Other
Suffolk
County
14
2
4
9
7
54
10
100
River-
head
South-
old
South-
ampton
1
1
45
8
23
18
100
1
3
35
7
42
8
100
2
4
13
6
62
6
100
Prom the physical and demographic data discussed previously,
from a review of information previously compiled for prepar-
ation of a Preliminary Safety Analysis Report for a nuclear
facility at the site (11), and from information presented in
the LILCO Shoreham plant licensing application, the prospec-
tive multipurpose facility site may be characterized as
follows:
1. The population density surrounding the site is low and
is expected to remain so and the "Population Center"
distance will be in excess of 20 miles; therefore, the
30
-------
population distribution will not preclude site use.
2. Geological features of the area are relatively well
documented. No faults are known to exist in the
vicinity of the site and the nearest seismically
active area is along the St. Lawrence Valley some
500 miles to the north.
3. Surface soil deposits at the site consist of un-
consolidated sands; therefore, substantial attention
to foundation design for site structures will be
required.
4. Fresh groundwater lies at an elevation slightly
above sea level at the site, or about 60 feet below
the planned final grade. Drainage is directly
northward toward Long Island Sound. Hydrologic
conditions at the site are satisfactory.
5. Eastern Long Island is well-ventilated, with rela-
tively high wind speeds. The metorological condi-
tions are generally favorable for a nuclear site;
however, consideration must be given in design for
protection against hurricane force winds.
31
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SECTION V
DESCRIPTION OF THE INTEGRATED FACILITY
The integrated plant design calls for the utilization of a
3400 Mwt light water reactor producing approximately 1000
Mwe energy for off-site utilization, providing process steam
for a distillation plant capable of recovering approximately
47.5 MOD of high quality water, and providing all necessary
electrical energy for the distillation and 50 MGD wastewater
treatment plants. Steam from the final stage of the distil-
lation unit will be used to elevate the temperature of the
incoming raw wastewater and enhance the primary and secon-
dary treatment processes. The secondary treated water will
serve as feed for the distillation plant, which will recover
95 percent of the water that would otherwise be discharged
to the sea.
The reference site selected for the design of the integrated
complex encompasses over 500 acres and is located on the
north shore of Long Island at the eastern er.d of the Town of
Riverhead in Suffolk County.
Although the population density in the immediate area of the
site and in most of eastern Suffolk County is relatively
low, the rapid growth of population in Nassau County and in
the western towns of Suffolk County will provide a demand
for electrical power that appears to justify full utiliza-
tion of economies of scale in sizing of the power generating
facility. It is also clear that the future water require-
ments of the area can be augmented by the high quality water
produced by the distillation plant. The factor which limits
the water recovery capacity of the integrated facility is
the supply of wastewater which can be economically conveyed
to the facility for treatment. Waste Disposal Districts 11
and 13, comprising the Towns of Southold and Riverhead in
the immediate vicinity of the reference site, the eastern
sector of the Town of Brookhaven and the northern edge of
the Town of Southampton, are projected to have a combined
wastewater flow of approximately 50 MGD by the year 1990.
Importation of wastewater from other districts in eastern
Suffolk County could allow sizing of the wastewater treatment
32
-------
facility for a feed rate of 70 MGD or greater, achieving
some further economies of scale. However, insufficient in-
formation relative to schedules for installation of sewers
and the probable cost of conveyance systems for such impor-
tation is available at this time. The conceptual design of
the integrated plant was therefore based on a feed rate of
50 MGD and is intended to allow expansion of the facility
if required.
\
The nuclear power plant could utilize either a boiling water
reactor (BWR) or a pressurized water reactor (PWR) as will
be discussed in Section VI. A reboiler will be utilized to
isolate the turbine feed steam, which potentially contains
minor amounts of radioactivity, from the steam used in the
distillation facility. Extraction of steam for this pur-
pose will reduce the electrical output of the turbogenerator
from the nominal 1100 Mwe associated with present generation
nuclear plants to approximately 1000 Mwe. In all other re-
spects, the nuclear power plant will be similar to those
used for single purpose power generation. Radioactive waste
disposal systems will be current state-of-the-art systems
meeting all applicable regulations for environmental protec-
tion. Turbine condenser cooling will be provided by a once-
through cooling system utilizing water from Long Island
Sound and returning the water through an outfall structure
designed to be totally compatible with environmental protec-
tion requirements. The nuclear steam supply system and the
power generation and cooling system are described in Sec-
tions VI and VII, respectively, of this report.
Wastewater from Disposal Districts 11 and 13 (2) will be
pumped to the site and will be treated in a modified acti-
vated sludge plant operating at an elevated temperature.
Unit operations will consist of bar screening and pumping
performed off-site, and thermal enhancement, grit removal,
primary sedimentation, and activated sludge treatment (aera-
tion and solids separation) performed on-site. No disinfec-
tion facilities, outfall sewer, or separate administration
facilities are required. Primary and activated sludge will
be combined, gravity thickened, digested, and dewatered on
vacuum filters, then trucked to a landfill area.
The distillation plant, a 19-stage vertical tube evaporator
(VTB) unit, will be operated on 285°F steam generated in a
33
-------
reboiler. The distillation plant will receive approximately
50 MGD of secondary treated wastewater, from which 47.5 MGD
of product water will be produced. The difference, 2.5 MGD,
will appear as evaporator bottoms with a total dissolved
solids content of approximately half that of seawater, and
will be mixed with approximately 1070 MGD of plant cooling
water for discharge to Long Island Sound. The 50 MGD inflow
of wastewater will be diluted with 1.44 MGD of distillation
plant product steam, through injection by a barometric leg
condenser, to provide thermal enhancement of the wastewater
treatment process. Product water cooling required by the
distillation plant will be provided by sharing the seawater
intake and outfall system of the power generating facilities.
Postdistillation carbon filtration, mineralization and
chlorination will be included when the product water is to
be provided for distribution. This post treatment will elim-
inate any volatile organics which may have carried through
the distillation process and assure compliance of the prod-
uct water with all water standards. The distillation plant
system description is presented in Section IX of this report.
An analysis of the economics of wastewater heating is pre-
sented in Appendix B.
A schematic flow diagram of the integrated plant is shown in
Figure 6, and a preliminary site layout showing a possible
arrangement of the nuclear power generating, wastewater
treatment, and distillation facility on the proposed site is
presented in Figure 7.
34
-------
Electricity
1000 Mwe
t
Nuclear
Steam
Supply
Steam } —
„ j , generator
Condensate &
> ,
i f
• T -i- -«-
L. tSeawater Intake and Return
r f 1 740,000 GPM
1 '
' «
. J I 1
I uvap. Bottoms ^ f
W T"^ 1 * T
1 t
T-\ • J. • 1 1 J. • "™
> Plant
49.8 MGD 1.44 MGD Re
Treated cycled stea
•tTT— -_ 4- _ .*. . **
W a s t e t VT "v — —
1.44 MGD Re- Primary
cycled steam and Secondary _^
Waste water-
Treatment
2.4 MOD ''
,L _j
— — — — — -*•*• -• aH
f 1
Product Water Potable
and Cooling Reuse
m 47 R MHn
• ' Raw Waste
50 MG1)
Sludge
Treatment.
L ta. T.snrl-F-i 1 1
Figure 6. Schematic Diagram - Integrated Facility
-------
N
\
\
1-Reactor Containment
2-Turbine Building
3-Distillation Plant
4-Wastewater Plant
Figure 7. Site Plan - Integrated Facility
36
-------
SECTION VI
NUCLEAR STEAM SUPPLY
Present and anticipated power demands in the region of the
proposed integrated utility justify economy of scale in the
selection of the nuclear power generating facility. On this
basis, the nuclear steam supply system has been specified to
be a 3400 Mwt reactor consistent with the capacity now being
offered by reactor manufacturers. Further, for convenience
in preparing this analysis, a light water reactor (LWR) was
selected. This choice, however, does not constitute a re-
quirement for a LWR in preference to a high temperature gas
cooled reactor or other nuclear steam system. Data from
two LWR vendors, General Electric and Westinghouse, were
used to determine specific nuclear steam supply and turbo-
generator characteristics and to develop estimates of capit-
al and operational costs.
A 3400 Mwt nuclear system, when used for the single purpose
of generating electricity, is capable of producing 1100 Mwe.
The present application, however, calls for an integrated
system, with steam being extracted for use in a distillation
plant located on the same site as the reactor. This re-
quirement reduces the net generating capacity of the system
to approximately 1000 Mwe.
The choice of a LWR may be narrowed further to selecting
either a pressurized water reactor (PWR) or boiling water
reactor (BWR) for the proposed integrated facility. Initial
consideration of both types of reactor indicates an advant-
age for the PWR for dual purpose use, in that steam genera-
tors are utilized to isolate the primary reactor coolant
loop from the turbine steam system. Nevertheless, the
nominal isolation in a PWR steam generator may be violated
by small leaks, resulting in transfer of small quantities
of radioactivity from the primary to the secondary system.
With the conventional direct cycle BWR, steam generated in
the nuclear reactor is fed directly to the turbine and
therefore contains significant quantities of short-lived
radioactivity.
37
-------
Total isolation of wastewater reclamation process from the
nuclear steam supply is considered to be essential, in order
to achieve public acceptance. The approach to isolation se-
lected for this study incorporates two high-reliability iso-
lating steam generators, hereafter referred to as reboilers.
The steam, extracted from the turbogenerator cycle of either
the BWR or PWR, enters the reboilers where it generates the
steam to be used in the distillation plant primary stage and
in jet ejector service. The installed cost of these reboil-
ers, including all associated instrumentation, is estimated
to $2,500,000.
The main reboiler would operate on a nominal 10°F tempera-
ture differential, receiving 1.4 x 106 pounds per hour of
295°F process steam extracted from the turbine system and
generating a like quantity of 285°F steam for use in the
distillation plant. A second reboiler operating on 385°F
steam would be used to generate approximately 10,000 pounds
per hour of 100 psi steam for use in the distillation plant
ejectors.
These reboilers and the provisions for interstage extraction
of steam for their operation constitute the only major non-
standard features of the nuclear steam supply and power gen-
erating systems. Radioactive waste monitoring, treatment,
and disposal features of the plant will be specified and de-
signed to meet all requirements of the USAEC and other regu-
latory agencies, as applicable to the site. Similarly,
design features for environmental protection will be incor-
porated in accordance with standard nuclear design practice
and the special requirements of the specific site location.
Table 8 summarizes a current estimate of nuclear plant costs
(15) . The tabulated costs are separated into components as-
sociated with the nuclear steam source and the turbogener-
ator. A 15 percent fixed charge rate was assumed to be ap-
propriate for a New York State private utility and all
capital and.construction costs were based on 1972 dollars
and using an Engineering News Record Construction Cost In-
dex (ENR) 1690.
The levelized fuel cost for the nuclear power plant has been
selected at 20£/105 Btu on the basis of the values reported
by Westinghouse Electric Corporation (16). Other sources
38
-------
TABLE 8
NUCLEAR PLANT INITIAL INVESTMENT COST ESTIMATE
FOR 1972 OPERATIONS
1100 Mwe SIZE*
Unit Cost
Steam
Generator
$/Kwe
• st 40
iction Materials and
>ment 27
iction Labor 33
>nal costs for 1972** 22
iional Service 14
indirect Costs 12
148
:ion During Construction 0
it Durinq Construction 27
Total 175
Turbo-
Generator
$/Kwe
31
18
23
--
11
9
92
0
17
109
Nuclear Plant Cost
*-Based on Reference 15.
$284/Kwe
**-Includes regulation and safety, near-zero radiation
release, additional quality control requirements, and
aesthetics.
(15, 17) generally support that cost, citing a range from
17.9C/106 Btu to 22.1C/106 Btu, with an average of 20.6C/106
Btu.
The cost of prime steam includes the cost of the nuclear
steam generator, fuel costs, and the portion of the opera-
ting and maintenance costs attributed to steam production.
39
-------
Annual Cost of Steam Production
Steam Generator $175/Kwe (1.1 x 106 Kwe)
(15% Fixed Charge Rate) = $28.9 x I06/yr
Operation & Maintenance d> 0.695 mills/
Kw-hr = 5.3 x 106/yr
Fuel (§> 20C/106 Btu = 16.2 x 106/yr
$50.4 x I06/yr
Prime Steam Cost (519°F)
Total Annual Steam Production = 8.1jx
Btu
8 x
The cost of process steam for the distillation plant may be
estimated on the basis of relative energy utilization. The
base cost of steam is 62.3C/106 Btu. This is the cost of
steam assuming that no energy is utilized for the production
of electricity. However, the available energy of prime
steam is much higher than that required for the distillation
plant. Thus, useful work can be performed by the steam be-
fore it is extracted for process purposes. Using the Car-
not efficiency approach the amount of energy extracted from
the steam can be approximated to vary linearly with the
extraction end point. Thus, the energy of 519°F steam will
cost 62.3C/106 Btu and 101°F exhaust steam (the turbine
terminal temperature) will cost nothing. The cost of steam
of intermediate temperatures may be approximated by a
straight line function between these two extremes. The
equation of this line is:
c = 62. 3£ Ts - 101
steam 106 Btu X 519- 101
Thus, the cost of 295°F steam to the main reboiler is 29C/
loo Btu, and the cost of the small additional quantity of
358QF steam required for the high pressure reboiler is ap-
proximately 38C/106 Btu.
40
-------
SECTION VII
POWER GENERATION AND COOLING SYSTEM DESCRIPTION
The 3400 Mwt nuclear reactor/ chosen as the energy source
for the integrated facility, is typical of the light water
reactors currently being purchased for commercial operation
in this country, takes advantage of the economies of scale
and reduces the operating cost per unit of electrical out-
put. The relative ease of transmitting electricity, as com-
pared to water, permits the electrical plant to serve a sig-
nificantly larger service area and population than will the
wastewater and distillation plants.
The nuclear steam supply would permit production of 1100 Mwe
if it were used in a single purpose facility. The multi-
purpose facility described in this study will have a net
electrical output of approximately 1000 Mwe with the distil-
lation plant and wastewater treatment plant in full opera-
tion. The turbogenerator to be used with the plant will
be a conventional 1000 Mwe type, modified for inter-
stage extraction of process steam at temperatures of 295°F
and 385°F. This steam is utilized in isolating reboilers
to provide both process steam to the distillation plant
primary stage and steam for operation of the ejector system
throughout the distillation plant. With the exception of
the provision for extraction of steam, the turbine is fully
condensing, utilizing once-through seawater as a condensing
fluid, as discussed below, with terminal steam conditions of
101°F and 2" Hg pressure. An operating efficiency for the
turbogenerator of approximately 35 percent may be antici-
pated.
A turbogenerator of the required characteristics can be sup-
plied by any of a number of manufacturers. The specific de-
sign characteristics would depend on the manufacturer as
well as the nuclear steam supply system selected.
Recent estimates (15) indicate a total cost of approximately
$109,000,000 for the turbogenerator; i.e., approximately
$109 per installed kilowatt of electrical power. This cost
includes approximately $30 per kilowatt as the basic unit
cost; approximately $40 per kilowatt for construction labor,
41
-------
materials, and equipment; approximately $20 per kilowatt
for engineering professional services, and other indirect
costs; and up to $20 per kilowatt for interest during con-
struction (all cost estimates are expressed in 1972 dollars
using an ENR index of 1690).
Costs of the nuclear portion of the power generating sta-
tion, described in Section VI are estimated to be $175 per
equivalent electrical kilowatt; i.e., since the nuclear
station is sized for the equivalent of an 1100 Mwe generat-
ing capacity, the total estimated cost is approximately
$192,500,000. Combining the steam supply cost with that of
the turbogenerator and auxiliary equipment at an annual
fixed charge rate of 15 percent, fuel cost at 20C/106 Btu,
and operational and maintenance cost at 0.93 mills/Kw-hr,
leads to a total bus bar electric cost for the plant of
approximately 9.1 mills/Kw-hr, as follows:
Annual Fixed Costs (7000 hr/yr
Operation) 6.1 mills/kw-hr
Fuel (10,390 Btu/kw-hr) 2.1 "
Operation and Maintenance 0.9 " "
9.1 mills/kw-hr
The general energy balance for the integrated facility is
shown in Figure 8. As indicated, the total energy produc-
tion is approximately 11.4 x 109 Btu/hr. Of this, 3.4 x 109
Btu/hr is transmitted off^site in the form of electrical
energy. The remaining energy, 8.0 x 109 Btu/hr, although
utilized in the wastewater treatment and distillation pro-
cesses, (e.g., distillation of water and operation of pumps
and other electrical equipment) is not consumed and must,
therefore, eventually leave the site in the form of heat.
An insignificant quantity of heat, approximately 3.9 x 104
Btu/hr leaves the wastewater process with the sludge and
is eventually lost to the atmosphere. It should be noted,
however, that approximately one-third of this heat,
1.3 x 104 Btu/hr. is produced by burning of methane gas
given off in the sludge treatment process and used to in-
crease the reaction rate in the sludge digesters. Approxi-
mately 2 x 108 Btu/hr. of heat is lost to the atmosphere
from the distillation and wastewater treatment process by
radiation and convection from the process equipment.
42
-------
3400 Mwt
(11.4 x 109
Bcu/hr)
2000 Mwe
(3.4 x 10° Btu/hr)
Power Production
Turbine Condenser
6.7 x 109 Stu/hr
E
RS
10 a?
(no
o *-<
Radiative and
Convection Losses
0.06 x JO9 Btu/hr
Product and
Evaporator Bottoms
Distillation
0. 37 x 109 Btu/hr
-
CJ *" I
3 "TM
CD
v
-------
Evaporator bottoms and product water leave the site at tem-
peratures of 35°F and 20°F, respectively, above the temper-
ature of the incoming raw wastewater, carrying with them a
total of approximately 3.7 x 108 BtuA*. The remaining
quantity of heat, approximately 7.4 x 109 Btu/hr, must be
removed and discharged to the environment through use of a
cooling system. The primary source of this waste heat is
the power station turbine condenser discharging 6.7 x 109
Btu/hr. The remainder of the heat, 7.3 x 108 Btu/hr, must
be removed from the distillation plant product condenser and
product water cooling system.
The simplest and most direct approach to condenser cooling
is to utilize water from Long Island Sound on a once-through
basis. A flow of approximately 740,000 gpm will be required
if the temperature rise across the condenser is to be
limited to 20°F, a value which will be near optimum from
engineering and economic considerations. If Long Island
Sound water is to be used for cooling, a large and enviro-
mentally compatible intake structure must be provided and
provision must be made for an outfall system which will dis-
sipate the heat at an approximate depth and over an adequate
area to minimize environmental impact and meet applicable
regulations.
From the Shoreham Nuclear Station now under construction,
the Long Island Lighting Company has designed an outfall
system to conform to the regulations governing thermal dis-
charges. A single pipe will extend from the plant for a
distance of 1600 feet along the floor of the Sound. From
that point, the outfall line will be raised three feet from
the floor of the Sound. The pipe will run an additional
2200 feet in a northerly direction into the Sound in deep
water. Outlet ports will be spaced every 60 feet from the
elevation point and will be alternated on either side of
the pipe to achieve maximum dispersion of the thermal
effluent.
The evaluation performed for the Long Island Lighting Com-
pany Shoreham Plant suggests that once-through cooling,
with discharge of waste heat to Long Island Sound, would con-
stitute an environmentally acceptable approach for the
integrated facility under evaluation. The environmental,
engineering, and economic feasibilities of alternatives to
this approach have also been evaluated.
44
-------
The alternative to the discharge of heat to Long Island Sound
is discharge to the atmosphere, either by evaporative or non-
evaporative cooling towers. Dry (non-evaporative) cooling
towers cost from $30 to $45 per installed kilowatt and have
never been utilized on a plant of the size proposed here.
Moreover, dry cooling requires an increase in terminal tur-
bine steam condensation temperature, resulting in a loss in
turbine efficiency and a concomitant economic penalty (18).
For these reasons, the use of non-evaporative cooling towers
was not considered further.
Evaporative cooling systems may be grouped into three gen-
eral categories: natural draft towers; mechanical draft
towers, and spray ponds. Recirculating cooling ponds other
than spray are discussed below. Cooling ponds transfer heat
to the atmosphere both by evaporation and by radiation and
convection, with evaporation typically resulting in approxi-
mately one half of the heat transfer. This fraction, and
the successful operation of these systems, is highly depend-
ent on local meteorological conditions.
Evaporative cooling systems are characterized by transfer of
heat from water to air in the form of latent heat of evapor-
ation. The disposal of 7.4 x 109 Btu/hr requires, therefore,
the evaporation of approximately 13,000 gpm of water for a
total evaporative loss of approximately 18 MGD. If fresh
water supplies, groundwater, or other water of potable qual-
ity were to be utilized, this arrangement would represent a
significant economic loss. Utilization of seawater from
Long Island Sound would appear to be more economically at-
tractive; however, the extent and effect of salt drift from
the evaporative cooling systems is a factor which must be
evaluated (20).
One alternative which was given major consideration during
this study was the use of partially treated wastewater for
condenser cooling purposes, transferring heat to the atmos-
phere utilizing a cooling pond with floating spray modules.
This approach also represents an integration of wastewater
treatment, power production and water recovery facilities.
In addition to providing a means of heat dissipation, aera-
tion serves to further the wastewater treatment process.
Partial treatment of the wastewater appears to be necessary,
at the present time, because of the uncertainty in values
for the corrosion and fouling.
45
-------
Another advantage initially credited to this approach was
the elimination of the requirement for a specific ammonia
removal step in the wastewater handling. Subsequent analy-
sis, however, discussed in Section IX, showed that ammonia
could be removed directly in the distillation process and
that treatment of the wastewater for this specific purpose
was not required.
A disadvantage of the spray pond system is the space require-
ment of approximately 30 acres of active spray pond and addi-
tional land for approach and return channels. Considering
that the study site contains over 500 acres, this space util-
ization appears to be acceptable. A second and more signif-
icant disadvantage of this system as an alternative evapora-
tive cooling system is the unavoidable loss of approximately
36 percent of the wastewater collected and treated which,
with further treatment and distillation, could be recycled.
The economics of closed cycle cooling have been studied ex-
tensively by Hittman Associates (18) and others (21). Ap-
plication of these data to the integrated facility would
increase the electrical requirements for a closed cycle
spray module system by the equivalent of $1400 a day. This
cost, however, is secondary to the loss in revenue resulting
from an estimated 1.8 percent reduction in power generation
attributed to the increase in turbine heat rate from a value
of 10,390 Btu/Kw-hr with once-through cooling to an esti-
mated 10,575 Btu/Kw-hr on the closed cycle. Based on the
bus bar cost of 9.1 mills/Kw-hr, this production los§ is
equivalent in value to almost $4000 per day.
Theoretically, the calculated spray module cost of $5,300,000
could be more than offset by elimination of much of the
$12,200,000 intake/outfall structure required for the plant.
Complete elimination of the intake/outfall structure is not
practical, however, since some provision for discharge of
secondary treated waste during periods of distillation plant
outage and of distillation plant evaporator bottoms during
normal operation is required. Alternatively, the 2.4 MGD
of 15,000 ppm concentrated evaporator bottoms could be fur-
ther concentrated, dried, and disposed of by incineration
or landfill, if desired.
The use of a fresh water and cooling water storage reservoir
as the primary condenser cooling system was also considered.
46
-------
This reservoir would be supplied not only by product water
from the distillation plant, but also by local surface and
groundwater supplies. The evaporative loss from such a
reservoir, assuming recirculation through the plant conden-
ser, would be approximately 10 MGD. This loss would be com-
posed entirely of high quality water which could otherwise
be returned directly to the Suffolk County water supplies.
The reservoir would require approximately 1800 acres, or
nearly three square miles. Topographic factors in the vicin-
ity of the site are not conducive to installation of such a
reservoir. On the basis of these factors, this alternative
was not considered further in this study.
In summary, the power generation system will consist of a
multistage condensing turbine approximating conventional
nuclear design and with an electrical generating capacity of
1000 Mwe minimum. Provisions will be made for interstage
extraction of 295°F saturated steam at rates up to 1.4 mil-
lion pounds per hour, and 385°F steam at a much lower rate
of 10,000 pounds per hour for use in the distillation pro-
cess. Based on a nominal operational schedule of 7000 hours
per year, the plant will produce 7 x 10^ kilowatt hours of
electricity per year for off-site use, at a bus bar cost of
9.1 mills/Kw-hr. Waste heat from the turbine condenser will
be discharged to Long Island Sound using a once-through cool-
ing system, with a nominal seawater inlet temperature of
71°F, a 20 degree temperature rise across the condenser, and
a 10 degree approach, the turbine will be operated under
condensing conditions of 101°F and 2" Hg abs, with an anti-
cipated heat rate of less than 10,390 Btu/Kw-hr. Protection
of the aquatic environment from the temperature discharge
will be provided by limiting the temperature rise across the
condenser and providing a suitable outfall structure. Total
cooling water utilization for the power and distillation
plants will be approximately 740,000 gpm.
47
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SECTION VIII
WAS TEWATER TREATMENT
The proposed wastewater treatment plant is designed for an
average flow of 50 MGD, with a peak hydraulic capacity of
75 MGD. Minimum flows are expected to be on the order of
30 MGD. Heat is to be added by means of a barometric con-
denser at the inlet of the plant.
Influent BOD and suspended solids are estimated to be 200
mg/1 and 235 mg/1 respectively, and effluent BOD and sus-
pended solids are projected to be 20 mg/1 each. The plant
is designed to remove 90 percent of the incoming BOD or
75,000 pounds per day and 91.5 percent of the suspended
solids or 89,500 pounds per day. Sludge handling and dis-
posal facilities are designed to accommodate 55 tons of
solids per day.
The process flow sheet for the proposed plant is shown on
Figure 9. Because the plant is located on the north shore
of Long Island, adjacent to the Southold town line at an
elevation of approximately 60 feet above sea level, all
wastewater will be pumped to the plant. Pumping stations
will be located at low points in. Districts #11 and #13 (2),
and will be provided with facilities to screen the raw waste-
water. The first unit operation at the plant site will be
thermal enhancement by means of barometric condensers. Steam
from the 19th stage of the distillation plant will be con-
densed to raise the temperature of the raw waste to 93°F in
order to assure a temperature of 86°F (30°C) in the aeration
basins of the activated sludge process.
The grit chamber is designed to remove grit from the thermal-
ly enriched wastewater. The hydraulic capacity of the grit
chamber was increased by approximately three percent to allow
for the increase in flow due to condensation, but the size of
the unit has been reduced by 30 percent to reflect the im-
proved performance due to the elevated temperature.
Primary sedimentation tanks are designed to remove the set-
tleable solids from the degritted wastewater, including the
condensate. The sedimentation tanks were designed on the
48
-------
PROM DISTRICT
PUMPING STATIONS
I
OXYGEN
(FROM AIR)
BAROMETRIC
CONDENSER
H
GRIT
CHAMBER
I
-*
PRIMARY
SEDIMENTATION
TANKS
-^
I
1
AERATION
TANKS
STEAM FROM
DISTILLATION
PLANT
HOLDING
POND AND
DISTILLATIO
PLANT
FINAL
CLARIFIERS
PRIMARY
SLUDGE
WASTE ACTIVATED SLUDGE
GRAVITY
THICKENERS
SLUDGE
TO FILL
SLUDGE
DIGESTERS
Figure 9. WASTEWATER TREATMENT PLANT PROCESS FLOW SHEET
-------
basis of an elevated temperature of 92°F and are approximate-
ly 75 percent of the size of conventional primary sedimenta-
tion tanks designed for operation at ambient temperature.
The activated sludge units have been designed to operate on
heated waste at approximately 86°F (30°C) and reflect a re-
duction in aeration tankage of approximately 30 percent.
The required aeration facilities are comparable in size to
conventional equipment because the increase in oxygen trans-
fer rate is offset by the decrease in oxygen saturation at
elevated temperatures.
Final clarifiers have been reduced 20 percent in size, as a
result of better liquid-solids separation at the elevated
temperatures of 85°F.
The discharge from the final clarifiers passes directly with-
out disinfection to a buffer reservoir ahead of the distilla-
tion plant. No separate outfall is required because the
entire effluent is used in the distillation plant. In the
event that the distillation plant is shut down, the effluent
is chlorinated with the disinfection facilities of the dis-
tillation plant, using the buffer reservoir as a chlorine
contact tank. The wastewater is then discharged with the
cooling water from the power plant condenser.
Primary sludge and waste activated sludge are combined and
thickened in gravity thickeners which have been reduced in
size, as a result of the 85°F elevated temperatures of the
wastewater, to approximately 80 percent of the capacity of
conventional gravity thickeners operating on unheated sludge.
Thickened sludge passes to anaerobic sludge digesters where
approximately 70 percent of the volatile solids are de-
stroyed, yielding sludge gas (methane and carbon dioxide)
which is used to heat the sludge digesters to 95°F and can
be used to heat the buildings and the air supply for the
ammonia stripping process.
Digested sludge is dewatered on vacuum filters. The sludge
cake, i.e., the product from the vacuum filters, is trucked
to landfill. Selection of landfill as the final disposal
method for the dried sludge is optional and does not influ-
ence the objectives of this study. Alternative means of
final sludge disposal include incineration, disposal at sea,
and wet air oxidation.
50
-------
The barometric leg condenser was selected for incorporation
into the integrated project because under the constraint of
current technology it is the most economical method of heat-
ing wastewater. The proposed treatment process will utilize
heat from the distillation plant and will effectively remove
contaminants from the wastewater. The wastewater treatment
plant effluent is projected to contain approximately 20 mg/1
of NC>3 and 5 to 10 mg/1 of NH3- The effluent turbidity is
estimated to be 5 to 10 Jackson Turbidity Units (JTU) . The
estimated capital costs for the wastewater treatment plant
in the integrated complex are compiled in Table 9, which
summarizes the integrated system with treatment plant heat
supplied by barometric condensers.
Capital costs reflect the construction of a plant to treat
waste at an elevated temperature and do not include disin-
fection facilities, outfall sewer, and administration
facilities which would normally be required for a convention-
al activated sludge plant. Disinfection is not required inso-
far as the subsequent process is one of distillation. The
outfall has been eliminated because the entire output of the
wastewater treatment plant will pass through the distillation
plant, or through the outfall of the power plant cooling
water system in the event of a shut down of the distillation
plant. The administration facilities of the integrated
facility are considered to be adequate to service the needs
of the wastewater treatment plant.
51
-------
TABLE 9
50 MGD WASTEWATER TREATMENT PLANT COST ESTIMATE
INTEGRATED SYSTEM, PLANT HEATED BY BAROMETRIC CONDENSERS*
ENR 1690
Raw Waste Pumping Stations
Barometric Condenser
Grit Chambers
Primary Settling Tanks
Primary Sludge Pump Station
Aeration Basins
Aeration Equipment
Final Clarifiers
Return Sludge System
Gravity Thickeners
Anaerobic Sludge Digesters
Vacuum Filter Facilities
Maintenance Facilities
Yard Piping
Subtotal
Engineering and Contingencies @ 30%
Total
$ 2,415,000
Included in distillation
plant cost-see Chapter IX
580,000
810,000
130,000
2,240,000
1,280,000
1,055,000
395,000
340,000
2,415,000
1,790,000
160,000
2,000,000
$15,610,000
4,685,000
520,295,000
*-Inlet to grit chamber, 93°F; primary settling tanks,
92QF; aeration basins, 86QF (30°C), plant outlet,
850F.
52
-------
SECTION IX
DISTILLATION PLANT
SYSTEM DESCRIPTION
The distillation process employed in the integrated facility
is a falling-film, multiple-effect, vertical tube evaporative
(VTE) process with 19 heat recovery effects. Accumulating
condensate and wastewater feed are flashed separately in a
single flashing stage for each effect. The plant has a per-
formance ratio of 14.5 pounds of product water for each 1000
Btu of input steam. The maximum distillation temperature of
the first stage concentrate is approximately 275°F. Other
temperatures and flow rates are shown on the Process Flow
Diagram, Figure 10.
Input to the distillation plant is approximately 50 MGD of
treated wastewater and 1.4 MGD of distillation plant exhaust
steam previously condensed in the wastewater. Similarly, the
design output of 47.5 MGD does not include the 1.4 MGD of
steam withdrawn from the final stage to heat the wastewater.
The difference between the input and output, approximately
2.5 MGD, consists of sterile evaporator concentrates and
system losses. The evaporator concentrates, which contain
dissolved solids at approximately 15,000 ppm concentration
are discharged through the turbogenerator cooling water
outfall to the waters of Long Island Sound, which typically
contain 33,000 ppm of dissolved solids. Steam condensation
and product water cooling required at the distillation plant
will be provided by once-through cooling with seawater util-
izing the same intake and outfall facilities as the power
generating system. After distillation, the product water
is subjected to treatment such as carbon filtration and
chlorination, depending upon the intended use.
The secondary treated wastewater is pumped from the final
clarifier, located in the wastewater treatment facility, to
a decarbonation basin. Before entering this basin, sulfuric
acid is added to neutralize its alkalinity. In this neutra-
lization reaction, the bicarbonates are decomposed and free
53
-------
Ul
BAROMETRIC CONDENSER
sJ
-HP STEAM
FROM EFFECT
""•"A
BAR SCREEN v
GRIT CHAMBER
RAw
WASTE
PUMPS
i-JL-flV- 1 $
OS At.
V
1
I
^ n PRIMARY SETTLING TANKS
rw= — ==— i
I t-} PRIMARY SLUDGE PUMP
y-^.
'ST -4E.RATION TANKS
_j|------ 1 <5>
ccj
gfCTTM
/T^^E^lI
,CUUM F,LTERS ««£«<> '^^^^-^
PUMPS ,ninGF
r
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THICKENED SLUDGE SL
PUMPS THICK
-BUFFER
RESERVOIR
DECARBONATION
VENTS TO NH, VENT CONDENSER
t t
^1J
ra
CONDENSAie
/
2~o EFFECT 3RD EFFECT 4m EFFECT 5m EFFECT 6™ EFFECT 17m EFFECT 18m EFFECT 19™ EFFECT BRINE BLOW-
FEED PUMP FEED PUMP FEED PUMP FEED PUMP FEED PUMP FEED PUMP FEED PUMP FEED PUMP
6I7TMEFFECT MBfHEFFECT
RECYCLE RECYCLE RECYCLE
Figure 10. INTEGRATED FACILITY OVERALL HEAT & MATERIAL BALANCE DIAGRAM
LINE NUMBER
?LOW. MGD
FLOWi LftyHR.
FLOW, GPW
PPE55URE. INCHES OF He. t
\'/
SQ-o
41WS
w
-SS_
OS
l°£_
sao.
J33_
10,000
3SO
-------
carbon dioxide is evolved. The decarbonation basin consists
of a set of stair-like steps leading down to a retention
pond. The water is introduced at the top step and cascades
down the steps into the retention reservoir, from whose oppo-
site end the water is withdrawn. The fall-splash process at
each step continuously renews the liquid surface to effect
good gas release. The design for the proposed plant consists
of a weir, five steps (each 3 feet tread, 3 feet rise, and 40
feet wide) , and a retention tank (200 feet long, 40 feet wide,
and 4 feet deep) . Total retention time is approximately seven
minutes. Head loss is expected to be approximately 8.5 psi.
The decarbonation basin is to be constructed as part of the
entrance to the buffer reservoir. This buffer reservoir is
expected to hold approximately five hours of normal flow, or
about 10 million gallons. The approximate size is 200 feet
wide, 670 feet long, and 10 feet deep. From the decarbona-
tion basin and buffer reservoir the decarbonated feed is in-
troduced to the last effect preheater section, where its
temperature is raised to approximately 94°F, prior to enter-
ing the deaerator tower. The reasons for deaerating to low
levels (10 ppb of oxygen and 1 ppm of carbon dioxide) are
(1) to minimize corrosion of the evaporator internals; (2)
to minimize condenser gas fouling; and (3) to prevent forma-
tion of carbonate scale.
The deaerator proposed for this plant is a packed tower util-
izing stripping steam for gas release. The equipment consists
of a vertical cylindrical column containing a bed of packing
which breaks up the water and maximizes the exposed surface.
In the deaeration process , the water is spread and sprayed
over the top of the packing bed. As the water trickles down
through the packing, stripping steam is supplied at the bot-
tom and travels upward through the packing. The gases are
separated from the water and are carried away by the stripping
steam in accordance with Henry's Law for concentration of a
dissolved gas. If the steam supply is pure, the concentra-
tion of the dissolved gases in the water can be made to ap-
proach zero. An average liquid loading rate of 20,000 Ibs/
hr-ft2 is considered practical. The flow rate for the strip-
ping steam is a function of the degree of deaeration required,
the amount of steam required to minimize channeling, and the
characteristics of the particular type of packing. The
55
-------
deaeration capacity required for this plant is provided by
two 25-foot diameter by 40-foot high towers in parallel.
(22, 23) .
To the deaerated feed is added an antifoaming agent to im-
prove heat transfer. The feed is then pumped through suc-
cessive preheaters positioned in each effect arriving at
Effect No. 1 at a temperature of 275°F where it enters the
first bundle of vertical falling-film tubes, and proceeds
through the 19 stages of the VTE plant. The VTE plant re-
ceives its heating steam from an intermediate steam reboil-
er, and all of the condensate from the heating steam is re-
turned to this reboiler. This steam (approximately
1.4 x 106 Ibs/hr of reboiled steam) at 285°F and 53.2 psia
(saturated) is admitted to the shell-side of the Effect No.l
falling-film bundle. The steam condenses on the tubes, re-
leasing its heat of condensation to the preheated, decarbon-
ated and deaerated feedwater which is falling as a film on
the interior surface of the tubes. This causes the feed-
water to boil violently, so that approximately five percent
is vaporized as it falls through the vertical tubes. The
mixture of wastewater and newly formed steam flows from the
bottom of the heater tube bundle into the lower section of
the Effect No. 1 evaporator where the wastewater disengages
from the steam and collects in the sump. The slightly con-
centrated feed is then pumped to the top of Effect No. 2
while the steam is directed to the shell-side of the Effect
No. 2 vertical tube bundle and feed preheater tube bundle.
Knitted wire mesh entrainment separators are provided to re-
move entrained droplets of liquid from the steam flow. In
Effect No. 2, the vapor condensing on the outside of the
vertical tubes performs the same function as the heating
steam did in Effect No. 1, boiling more feed on the inside
of the tubes and producing additional steam to be passed on
to Effect No. 3. These essential processes are repeated in
each of nineteen effects through the plant, with each effect
operating at a progressively lower temperature and pressure.
In all effects" (except the first) , the falling-film tubes
are supplied by a pump which withdraws the concentrated feed
from the sump and lifts it to the top of the vertical bundle.
Special distributors (slotted circular weirs) distribute the
concentrated liquid to the tubes through which it returns, by
gravity, to the sump.
56
-------
Ammonia removal is accomplished primarily through venting the
initial three or four stages. Continuous venting of all ef-
fects is necessary to remove noncondensable gases. Effects
operating above atmospheric pressure are vented to a condenser
where the vapors are condensed (except for noncondensables
which are released to the atmosphere) and fed to an ammonia
stripper where any remaining ammonia is removed. The clean
liquid that leaves the stripper is combined with the other
product water streams. The condensed vent vapor flow is ap-
proximately 50 to 100 gpm and is at a temperature of 150°F
when combined with the other product water upstream of the
product water cooler.
Approximately 58 percent of the 100°F steam that is formed
in the last effect (No. 19) is condensed in a final condens-
er. Seawater, at 71 F, from the power plant cooling water
system is used to transfer the latent heat from the condens-
er to Long Island Sound, along with the waste heat from the
power generating facility. Seawater is also used in the
product water cooler to lower the temperature of the warm
product condensate, in the air ejector system to condense
steam, and in the vent system to condense gases.
The remaining 42 percent of the steam (approximately 500,000
Ibs/hr) from the last effect is condensed in a barometric leg
condenser by raw, ungritted wastewater, raising the temper-
ature of the incoming wastewater from 65°F to 93°F before
entering the grit chamber.
The condensate from the final condenser is combined with the
condensate from the air ejector system, condenser vents, and
that which has accumulated from Effects 2 through 19 to form
the product water stream. The combined product stream is at
a temperature of 110°F and must have its sensible heat re-
moved in the product water cooler to reduce the product water
temperature to 85°F. As discussed previously, seawater is
used as the cooling medium.
The cooled product is pumped to an activated carbon absorp-
tion system where any residual COD and odor-forming compounds
will be removed. Because of the sterilizing environment and
phase change in the distillation plant, the final product
would not contain any harmful bacteria or viruses.
The quality of the water produced by the integrated facility
57
-------
is anticipated to consistently meet or exceed the water qual-
ity criteria established by the U. S. Public Health Service
(24) and the World Health Organization (25). The recommend-
ed standard of the American Water Works Association (26) will
also be met, at all times, during normal plant operation.
These various standards are compared in Table 10.
A comparison of the estimated, capital costs for the distil-
lation facility with, and without the provision for heat ad-
dition by use of barometric condenser is presented in Table
11.
The distillation plant as described above, would be capable
of producing 47.5 MGD of high quality water at an estimated
cost of 64C/1000 gals., which includes the cost of the baro-
metric condenser system for heating the wastewater, and
62.4C/1000 gals, where no provision is made for wastewater
heating.
Of the total distillation plant capital cost, $1,992,000 is
attributable to the additional equipment and the increased
throughput capacity required in the distillation plant in
conjunction with the thermal wastewater treatment process.
This incremental capital cost corresponds to an annual fixed
charge of approximately $155,000. Of the total annual O&M
charges listed in Table 12, $131,000 is attributable to the
requirement to transfer heat to the wastewater treatment
process. The total annual cost specifically attributable
to the thermal enhancement of the wastewater is, therefore,
$286,000 which is equivalent to approximately 1.6C/1000 gals.
at a 47.5 MGD production rate. A detailed evaluation of the
economics of utilizing steam from the distillation plant to
heat the incoming wastewater is presented in Appendix B.
58
-------
TABLE 10. WATER QUALITY CRITERIA
DRINKING WATER STANDARDS (mg/1 except as indicated)
en
SUBSTANCE
Turbidity (Units)
Color (Units)
Odor (Threshold Odor No.)
Taste
Alkyl 'Benzene Sulfonate (ABS)
Aluminum (Al)
Arsenic (As)
Barium (Ba)
Chloride (Cl)
Calmium (Cd)
Chromium ( CrT° )
Copper (Cu)
Carbon Chloroform Extract (CCE)
Cyanide (Cn)
Iron (Fe)
Lead (Pb)
Manganese (Mn)
Nitrate (N03)
Phenols (CgH^OH)
Selenium (Se)
Silver (Ag)
Sulfate (S0^~)
Total Dissolved Solids (TDS)
Zinc (Zn)
Fluoride (F)
Hardness (As CaCOo)
Suspended Solids ( S . S . )
Phosphate (POj =)
Calcium (Ca)
Magnesium (Mg)
pH (Value)
Methelene Blue Active Substances (MBAS)
Carbon Alcohol Extract (CAE)
USPHS
Cone. -should not
be exceeded
5
15
3
0.5
0.01
250
1.0
0.2
0.01
0.3
;
0.05
1*5
0.001
250
500
5
1.3
0.5
AWWA
Recommended
Goals
< 0.1
< 3
no odor
Nothing object.
< 0.2
< 0.05
0.01
1.0
0.01
0.05
< 0.2
< 0.0k
0.01
< 0.05
0.05
< 0.01
^5
0.001
0.01
0.05
200
< 1.0
1.3
80-100
1.0
none
0.2
0.15
WHO
Drinking Water
Max. Accept.
5
5
Unob ject .
Unobject .
0.5
200
1.0
0.2
0.3
0.1
0.001
200
5.0
•***^^ ^••••••-~^-^— ^^— -^-^— — ^^— — ^^~— ^
75
50
7-8.5
-------
TABLE 11. CAPITAL COST DISTILLATION PLANT
(ENR-1690)
Case I* Case n**
1. Decarbonation basin and buffer reservoir $ 99,000 $ 96,700
2. Acid and antifoam injection systems 95,400 93,100
3. Deaeration tower 71,200 69,650
4. Multiple effect evaporator 56,442,800 53,176,200
5. Product condenser 1,447,000 3,000,000
6. Product cooler 1,869,000 1,335,000
7. Activated carbon adsorption system 1,276,700 1,276,700
8. Product chlorination system*** 117,000 117,000
9. Intermediate steam reboiler 2,564,000 2,500,000
10. Intake/outfall structure (amount charged
to distillation facility - 70,000 gallons
of 740,000 gallons total) 985,000 1,500,000
11. Ammonia stripper and vent condenser 10,900 10,650
12. Barometric condenser system 179,000 —
$65,167,000 $63,175,000
* case I Distillation plant providing heat to wastewater treatment plant using
barometric leg condenser
** Case II No heat added to wastewater treatment plant
***Including provision for chlorinating 50 MGD of secondary treated wastewater
in the event of distillation plant outage
-------
TABLE 12. DISTILLATION PLANT WATER PRODUCTION COSTS
Case I* Case II**
Distillation Plant Capital Cost $65,167,000 $63,175,000
(See Table 11)
A. Annual Fixed Charges @ 7.823% 5,098,000 4,943,000
B. Annual Operating and Maintenance Charges
1. Distillation Plant O&M
Labor $ 420,000 $ 408,000
Electric Power 710,000 789,000
Chemicals 715,000 694,000
Spare Parts, etc. 918,000 891,000
Steam 3,082,000 2,991,000
2. Reboiler O&M 26,000 25,000
3. Post-treatment O&M 20,000 20,000
4. Barometric condensers
O&M 2,000
Electric Power 56,000
Total Annual O&M 5,949,000 5,818,000
C. Total Annual Charges $11,047,000 $10,761,000
D. Water Cost - 47.5 MGD Production $0.64/1000 gal $0.624/1000 gal
* Case I Distillation plant providing heat to wastewater treatment plant
using barometric leg condenser
**Case II No heat added to wastewater treatment plant
-------
SECTION X
OVERALL PLANT INTEGRATION AND EVALUATION
In accordance with the objectives of this study conceptual
designs were developed for an integrated facility incorpor-
ating 1000 Mwe nuclear power plant, a 50 MGD wastewater
treatment plant, and a 47.5 MGD distillation plant, at a
reference site located on the north shore of Long Island in
the Town of Riverhead, Suffolk County, New York. As dis-
cussed in the preceding sections, the anticipated growth
in population, the increased need for electricity and the
projected inadequacy of ground water supplies make the
concept of integrating electrical, water supply and waste-
water treatment facilities attractive in this area. Sys-
tems to transport the wastewater from Disposal Districts
11 and 13 (and possibly a portion of the waste from Dis-
posal District 12) to the integrated facility could be
planned and installed, well within the time frame required
for its activation. The integrated facility concept, in
accomplishing the objectives of water reuse and resource
conservation satisfies the requirement for tertiary waste-
water treatment and meets all present and proposed federal,
state, and local standards for protection of the environ-
ment. The distillation process, incorporating steriliz-
ing temperatures and phase change, assures the high quality
of the product water. Postdistillation treatment of the
product water, including activated carbon absorption, min-
eralization, and chlorination, can be provided as required.
In addition to sharing a site and administrative facilities,
the three functional components of the complex are inte-
grated through interties at a number of process points. The
nuclear steam supply system, in addition to serving the
power generation facility, provides process steam to the
distillation plant. The nuclear steam supply and power
generating facility will provide all the electrical power
for the pumps and other motor driven equipment as well as
site lighting, heating, ventilating, and air conditioning,
while producing approximately 1000 Mew for off-site trans-
mission.
The distillation plant receives its thermal energy from the
62
-------
nuclear steam supply system, and in turn provides heat for
enhancement of the wastewater treatment process by utilizing
a barometric leg condenser through which approximately 42
percent of the product steam of the final stage is intro-
duced to the wastewater. (This steam would normally be con-
densed by the plant cooling water and the heat disposed of
in the receiving body of water.) This addition of energy is
sufficient to raise the temperature of the wastewater stream,
as received, from a nominal 65°F (18°C) to the initial pro-
cessing temperature of 93°F (34°c).
The distillation plant acts as a tertiary stage of wastewater
treatment which eliminates or replaces treatment steps that
would be required for discharge to the environment. Ammonia
removal is accomplished through venting of the initial stages
of the distillation train. Phosphates, nitrates, and refrac-
tory organics passing through the secondary treatment stage
of the waste treatment plant will also be removed in the dis-
tillation process. Chlorination facilities provided for the
product water would serve the wastewater treatment plant if
the distillation plant were temporarily shut down, eliminat-
ing the need for a separate chlorination station. The sea-
water circulation system, with its elaborate intake and out-
fall structures required for the turbogenerator condenser,
will also provide cooling water for the distillation plant
condenser and product water cooler, afford a convenient means
of disposing of the sterile concentrates from the distilla-
tion plant, and eliminate the need for an outfall system to
serve the waste treatment plant during periods of distilla-
tion plant outages.
The integrated facility substantially reduces the amount of
water removed from and discharged to Long Island Sound as
compared with separate facilities. A 50 MOD treatment plant
would discharge approximately 50 MOD of treated and chlori-
nated wastewater to the sound; whereas the integrated facil-
ity will discharge no wastewater except during times of
distillation plant shutdown. A 50 MGD unintegrated desali-
nation plant would remove approximately 259 MGD of 33,000 ppm
(of chloride) feed and cooling water and discharge approxi-
mately 50 MGD of 60,000-70,000 ppm (of chloride) brine and
159 MGD of cooling water heated to about 20 degrees above
the intake temperatures. The proposed integrated distilla-
tion plant will discharge only 2.4 MGD of 15,000 ppm sterile
evaporator concentrate and approximately 100 MGD of cooling
water.
63
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ALTERNATIVE METHODS OF WASTEWATER TREATMENT
A detailed literature search, conducted to evaluate the ef-
fect of elevated temperature on various wastewater treatment
processes and operations, is described in Appendix A. It
was concluded that existing work is fragmentary in that only
single unit operations have been considered and no system-
atic evaluation has been made of heat input to the overall
process.
Nonetheless, the limited information was sufficient to con-
clude that the operation of conventional treatment units at
elevated temperatures can be expected to improve removal ef-
ficiencies. Thus, under conditions of thermal enhancement,
as contrasted to operation at ambient temperatures, a given
throughput may be handled in a plant of smaller physical
size, while maintaining the same efficiency and effluent
quality.
The processes examined generally fall into the categories of
primary, secondary, or tertiary treatment, all of which in-
clude sludge treatment and disposal. Unit operations for
primary treatment include bar screening, comminution, pump-
ing, grit removal, sedimentation and chlorination.
As an alternative to sedimentation, flotation was considered
but not incorporated into the plant design because flotation
is generally less effective than sedimentation for removal of
solids, although elevated temperatures and steam injection
techniques could improve the flotation process.
For secondary treatment, the alternatives considered in lieu
of the activated sludge process were the use of trickling
filtration and stabilization basins.
For the integrated plant, consideration was given to utiliz-
ing a stabilization basin as a cooling pond for the nuclear
power system, with this concept, the wastewater temperature
would be elevated either by passage through the turbogenera-
tor condenser, or by a separate exchanger through which the
condenser coolant would be passed. This arrangement, in
which the aeration basin could be eliminated, was rejected
because adequate information is unavailable relative to heat
transfer performance under potential fouling conditions
64
-------
associated with wastewater processing, as well as the exten-
sive land area which would be necessary for the cooling pond
stabilization basin.
Trickling filtration was rejected because effluents from that
process have been found to be less desirable than those from
activated sludge effluents for feed to distillation units.
Additionally, use of trickling filters would, by increasing
the rate of heat loss to the atmosphere, require more energy
to be transferred to maintain elevated temperatures in the
secondary clarifiers. While this requirement is consistent
with the objective of dissipating waste heat, the state-of-
the-art of heat transfer technology militated against this
approach.
In view of the foregoing considerations, therefore, the acti-
vated sludge system incorporating primary sedimentation,
anaerobic digestion, and vacuum filtration was selected for
this project because the activated sludge process is the only
treatment system with which there is sufficient experience to
be incorporated into a plant of this size.
In order to produce a quality effluent, a tertiary treatment
process is employed. The choice among various alternatives
for achieving tertiary treatment depends on the specific ef-
fluent standards to be met. In the case at hand, distilla-
tion was selected as the principal tertiary treatment pro-
cess in order to assure a high quality product. A convention-
al activated sludge plant operated with 2,000 to 5,000
milligrams per liter of suspended solids in the aeration
tanks and 6 hours detention time will produce an effluent that
is suitable for distillation. Carbon absorption was added as
a polishing step and a means of insuring the removal of cer-
tain volatile contaminants, such as phenols, alcohol and
aldehydes which might be present in the wastewater and which
could be carried over to produce objectionable tastes and
odors in the product water.
The activated sludge treatment process will produce an ef-
fluent which will be low in BOD but high in nitrogen, which
subsequently will be removed by venting the initial stages
of the distillation plant. The vented vapors are processed
through a condenser and ammonia stripping unit; whereas, in
a separate facility nitrogen removal would be effected by a
65
-------
more costly nitrification-denitrification activated sludge
process. - 1
UTILIZATION OF HEAT IN WASTEWATER TREATMENT
The design criteria for the processing units which are incor-
porated in this plant are developed in Appendix A. The
settling velocities, the rate of biological activity, the
power required to dissolve oxygen in the wastes and the de-
tention time required for effective disinfection with chlo-
rine are all dependent upon temperature. Once the solids
have been separated from the wet stream, the only unit opera-
tion affected by temperature is gravity thickening of com-
bined waste, consisting of activated and primary sludges.
Digestion will take place in the mesophilic range as would
the case in a conventional plant.
Investigators have reported that the optimum temperature for
mesophilic biological growth lies in the range of 30°C to
37°C, while a thermophilic growth lies in the 50OC to 54°C
range. 30°C and 52°C are generally considered to be the
optimum temperatures for mesophilic and thermophilic aerobic
biological treatment, while 36°C and 54°C are considered
optimum for anaerobic digestion processes. Thermophilic pro-
cesses were ruled out because insufficient experience exists
with the process in large scale operation, and the process is
extremely sensitive to changes in temperature.
The choice of mesophilic aerobic process at the temperature
of 30°C for the integrated plant was dictated by the apparent
leveling off of improvement in efficiency of the activated
sludge process beyond that temperature, and by the difficulty
in achieving higher temperatures with the available heat
sources.
Illustrative of the overall system improvement to be gained,
it is projected that increasing the temperature of the waste-
water to 30°C (in the aeration tanks) and keeping the unit
sizes constant results in an increase in overall treatment ef-
ficiency from 90.5 percent to 93.0 percent. Conversely, main-
taining the removal efficiencies and effluent quality equal to
that achieved by an ambient temperature plant permits signi-
ficant reduction in unit sizes. For example, the primary
66
-------
settling tank surface area was reduced from 53,600 ft2 to
42,000 ft2. Similar results were achieved for the secondary
clarifier. The reduction in aeration basin volume was re-
duced from 2,073,600 ft3 to 1,695,000 ft3. A complete com-
parison of the selected unit operations for the conventional
and the integrated, heated 50 MOD treatment plants is shown
in Table 13. It should be noted that sludge handling pro-
cesses such as thermal conditioning and combustion or wet
oxidation, which are performed subsequent to gravity thicken-
ing, would not be affected by the initial heating of the
wastewater. Basically, only the wet stream processes of
primary settling, aeration, final settling, gravity thicken-
ing and chlorination are affected by temperature.
WASTEWATER TREATMENT HEAT BALANCE ANALYSIS
Heat balances were made on the proposed integrated, heat
treatment plant to determine the quantity of the heat lost
by each unit in different seasons, and the quantity of heat
necessary to achieve the design temperature. The efficien-
cies of the primary and secondary clarifiers increase with
temperature in the range of 10°C to 60°C, while the activated
sludge unit reaches constant efficiency at a temperature of
86°F (30°c) ; hence, a heat balance helps to insure that the
plant operates at or near optimal efficiency under a broad
range of external conditions.
The factors affecting heat losses include temperature dif-
ferentials between the wastewater and the atmosphere and
ground, surface wind velocity, solar radiation, relative
humidity, equipment surface area and type of flow considered,
either plug or mixed. These factors were then analyzed for
their applicability to the proposed integrated plant and the
geographical area. In order to determine the average quan-
tity of heat required to maintain design efficiency, the ap-
plicable factors were evaluated under conditions existing
during the four seasons of the year. The most extreme con-
ditions occur in the winter months when the temperature dif-
ferentials between wastewater and the atmosphere and ground
are at maximum values. Wind velocities are at a maximum
value, while solar radiation is at a minimum value. A sta-
tistical approach was used in this analysis, incorporating
mean values of winter temperature, wind velocity, relative
67
-------
TABLE 13
COMPARISON OF UNIT SIZES FOR CONVENTIONAL AND INTEGRATED,
HEATED 50 MGD TREATMENT PLANTS
UNIT
Raw Waste Pumping
Station
Grit Chambers
Primary Settling
Tanks
Primary Sludge Pump
Station
Aeration Basins
UNIT SIZES
CONVENTIONAL
T = 20°C
75 MGD
L = 61.2 ft
INTEGRATED HEATED
T = 30°C
75 MGD
L = 51 ft
53,600 ft2 S.A. 42,000 ft2 S.A.
214 gpm
2,073,600 ft3
Aeration Equipment 1920 HP
Final Clarifier 66,240 ft2 S.A,
Return Sludge System 50 MGD
Gravity Thickener
Anerobic Sludge
Digesters
Vacuum Filter Facil-
ites
11,538 ft2
828,990 ft3
2580 ft2
222 gpm
1,695,000 ft3
1920 HP
51,800 ft2 S.A.
52 MGD
8860 ft2
828,990 ft3
2214 ft2
Chlorine Contact Tank 116,148 ft3 *
Chlorine Feed System 4395 #/Day *
*-Not necessary in integrated plant (heated or unheated)
68
-------
humidity and solar radiation. However, on any particular
week or day, the mean values may be exceeded and heat losses
will rise above the value necessary to maintain maximum
efficiency. A factor of safety may be applied if the quan-
tity of heat added is to be sufficient for a critical month,
week or day. The risk of not supplying enough heat for a
particular period may be weighed against the additional
costs and the reduced quality of effluent water. Table 14
summarizes the results obtained from calculations utilizing
a computer program developed specifically for this purpose.
Average heat input was calculated using values obtained for
the four seasons and thus is not representative of the ex-
treme case. The basic values selected for the climatologic-
al factors were 15 mph for wind velocity, 10°F for air tem-
perature and 70 percent for relative humidity.
APPROACHES FOR HEAT ADDITION
The method used for the addition of heat to the wastewater
treatment process depends in part on the source used to pro-
vide the heat. Seven basic sources of heat were identified
and considered in this study, including; prime or high
temperature steam produced specifically for heating purposes,
low pressure or process steam, turbogenerator exhaust steam,
power plant and distillation plant cooling water streams,
exhaust steam from the final stage of the distillation plant,
extracted steam from intermediate stages of the distillation
plant, and the distillation plant product water. The alter-
natives of high temperature and process steam were rejected
on the basis of economics, insofar as the cost of high tem-
perature prime steam or even process steam is much greater
on a unit heat value basis than low temperature energy, such
as spent steam and distillation plant exhaust steam. Fur-
thermore, based on a review of existing literature, relative
to the effect of heat on wastewater treatment, no signifi-
cant advantage was found in raising the wastewater treatment
temperature above 93°F.
Heating of the incoming wastewater by countercurrent cooling
of warm product water or by condensation of exhaust steam
from the turbogenerator appears to be the most attractive
approach. The energy from these sources is waste heat nor-
mally rejected to the cooling water. However, conflicting
69
-------
TABLE 14
HEAT BALANCE ANALYSIS
Climatological Basis Yearly Averages
Treatment Plant Temperature 86°F{30°C) 104°F(4QOC) 122OF(5QOC)
Temperature (°F)
Raw Wastewater
Grit Chamber
Influent
Effluent
Primary Settler
Influent
Effluent
Aeration Tank
Influent
Effluent
Secondary Settler
Influent
Effluent
65
65
65
93
92
92
91
91
86
86
85
112
111
111
110
110
104
104
103
131
130
130
129
129
122
122
121
Heat Input
(10b BTU/hr)
Heat Exchanger Capacity at
Average Flow
Heat Exchanger Capacity at
Peak Flow
Condensate Flow MGD at
Peak Flow
4.85
7.3
2.0
8.2
11.3
3.4
11.5
17.2
5.1
70
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and insufficient information relative to heat exchange tech-
nology precluded the adoption of these energy sources for the
design of the integrated facility. In the case of surface
heat exchangers, the limited data presently available on
wastewater fouling factors and associated heat transfer coef-
ficients, as well as uncertainties regarding the materials
of construction and the costs of fabrication, suggest that
further information be developed as a prerequisite for con-
sidering this approach.
Another alternative considered was the use of the power plant
and distillation plant cooling water streams. This approach
was rejected because the temperatures of these coolant
streams are constrained, by thermal standards, to values that
will not produce the 93°F necessary in the wastewater.
The exhaust steam from the distillation plant or extracted
steam from the latter stages of distillation constitute the
remaining alternatives. Heat from these stages may be trans-
ferred to the wastewater stream utilizing either surface heat
exchangers of barometric leg condensers. Again, in the case
of surface heat exchangers, performance and cost uncertain-
ties must be resolved. In the case of the barometric leg
condenser, the direct cost of heat transfer is very low, but
high purity product is lost and recycled into the wastewater
system. Maintaining the desired production from the distil-
lation plant under these conditions requires an increase in
the design capacity and operating throughput of all equip-
ment between the point of heat addition in the wastewater
process and the final stage condenser of the distillation
train.
A detailed analysis of the engineering and economic factors
associated with addition of heat to the wastewater treatment
plant was performed and is presented in Apprndix B. This
analysis indicated that, based on available information rel-
ative to heat transfer surface fouling in wastewater service
and expected heat transfer coefficients, utilization of bar-
ometric leg condensers to inject a portion of the steam from
the final distillation stage into the incoming raw waste
would result in the lowest overall cost penalty for the sys-
tem. For this reason this approach was incorporated in the
basic flowsheet for the integrated facility, and resulted in
a reduction in the capital cost of the wastewater treatment
71
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facilities which was, however, less than the increase in
capital cost calculated for the distillation plant.
It should be noted, however, that information provided to
ASDA by various manufacturers (27, 28, 29) indicates that
higher heat transfer coefficients and lower installed equip-
ment costs than those considered in this study may be obtain-
able. For a maximum wastewater treatment temperature of
93°F, surface heat transfer equipment would be competitive
with barometric leg condensers, if the total installed cost
of the former using low grade steam was no more than $2
million. In the case of a product water/wastewater ex-
changer, the total installed cost could be as much as $3.3
million and ill be competitive. The reduction could be
achieved either by a decrease in equipment fabrication or
installation costs or by an improvement in the overall heat
transfer coefficient resulting in a reduction in the total
required heat transfer area.
ALTERNATIVE APPROACHES OF INTEGRATION
A number of alternative approaches of integrating the facil-
ities were considered during the course of this study. One
approach which was considered to be highly promising was based
on the utilization of partially treated wastewater as the
make-up to a closed evaporative cooling system for the power
generation and distillation plant facilities. Since the
blowdown from this system would provide the feed to the dis-
tillation plant, routine discharges to the aquatic environ-
ment would be completely eliminated with the exception of
the evaporator bottoms which could be further concentrated,
dried and disposed of as landfill or by incineration. A
spray pond utilizing floating power spray modules was con-
sidered to be a feasible approach and a logical choice for
evaporative cooling at this site. In addition to discharg-
ing all heat from the turbogenerator and distillation plant
condensers to the atmosphere rather than to an aquatic
heat sink, the spray module approach provides aeration of
the partially treated waste and, thus, serves as an addi-
tional stage of wastewater treatment. The disadvantage of
this approach, and the primary reason it was ultimately re-
jected, was that the cooling requirements for the facility
would result in the evaporative loss of approximately 18 MGD
of water which could otherwise be largely returned to the
water supply of the region.
72
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Furthermore, Long Island Sound provides an adequate cooling
water supply at this site and with proper outfall design
could be utilized as a heat sink without adverse environ-
mental impact under existing heat load conditions and ther-
mal discharge criteria. For sites with more restrictive
environmental constraints on thermal discharge or lacking
in water supply for cooling purposes , the use of partially
treated waste for condenser cooling appears to offer sig-
nificant economic and environmental advantage.
Alternatives to distillation were considered and rejected
in the planning stage of this project. Although other de-
salting techniques, particularly reverse osmosis and elec-
trodialysis, appear to have promise for the recovery of
brackish waters or wastewater, these processes are less
amenable to integration with a large scale power generation
facility and lack the fail-safe characteristics of distil-
lation considered necessary when converting wastewater to a
high quality product.
Multistage flash (MSF) and combined multistage flash and
vertical tube evaporative (MSF-VTE) systems were investi-
gated in addition to the VTE approach which was finally se-
lected on the basis of recent developments and improvements
in the economics of this system (30). The VTE also has
greater flexibility than MSF plants both in mode of opera-
tion and maintainability. Consideration was also given to
operation of the distillation plant solely, or largely, in
off power peak periods in order to obtain greater electric
power outputs during peak demand periods and permit reduc-
tions in cost of steam used for distillation plant purposes.
However, the relatively high operating efficiency of the VTE
system and operational control considerations led to the
conclusion that the plant and the power generating facility
should both be operated in a base loaded mode. This ap-
proach also minimizes the need for buffer reservoirs between
segments of the water treatment and recovery processes.
BENEFITS OF INTEGRATION
Integration of power generation, wastewater treatment, and
water supply facilities results in both tangible and intan-
73
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gible benefits. Reuse of water results in conservation of
fresh water supplies that would otherwise be lost if the
wastewater were treated and discharged to Long Island Sound.
The integrated system results in the reduction of the build-
up of phosphates, organics, and nitrates observed in areas
served by septic systems. Buildup of nitrates occurs in
groundwater systems where a substantial portion of the re-
supply is from treated wastewater and nitrates are consid-
ered to cause eutrophication when discharged to certain sur-
face waters. The current advanced water treatment methods
for the removal of nitrogen includes ammonia stripping with
air, nitrification-denitrification, and breakpoint chlorin-
ation. All of these processes have limitations. The
ammonia stripping process is severely limited by low temper-
atures and would be economically prohibitive in a convention-
al 50 MGD treatment plant. Because heat is supplied in the
distillation process, ammonia stripping can be accomplished
on a year-round basis. Furthermore, the ammonia need not be
released to the atmosphere but can be concentrated and used
or converted to nitrogen and harmlessly released to the
atmosphere.
Reuse not only creates a water resource, which could be vit-
al to the maintenance of the water balance of the region in
the future, but also addresses the recently established
criteria for "zero discharge" by eliminating many of the
pollutants that would be discharged from separate facilities.
The high quality of the water which can be produced by this
system would augment water supplies in the area, either
through distribution, recharge to groundwater aquifers, or
supply for industrial purposes.
Utilization of auxiliary and ancillary facilities, jointly,
will result in operational and economic savings. Savings
estimated at $3.9 million will result from elimination of
separate administrative, chlorination, and outfall facili-
ties for the wastewater treatment plant. Other benefits
include reduction in fencing, access roads, number of main-
tenance and security personnel, and other similar items
common to the three plants.
74
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COMPARISON OF INTEGRATED AM) CONVENTIONAL SEPARATE FACILITIES
In the absence of the integrated facility the electrical,
wastewater treatment, and potable water demands of the re-
gion would have to be met by separate conventional power
stations, waste treatment plants, and water supply facilities,
Additional electrical energy production, considering the pro-
jected growth in population and electrical demand in the re-
gion, would likely be provided by large nuclear power gener-
ating stations similar to that proposed for the integrated
facility without the capability for dual purpose operation.
A recent estimate (15) of the cost of such a facility (1972
dollars, ENR index 1690) is approximately $284/Kw or $284
million for 1000 Mwe unit. The calculated bus bar electrical
cost for such a unit if built and operated with once-through
seawater cooling, is approximately 9.1 mills/Kw-hr.
The total projected capital cost for the nuclear steam sup-
ply and power generating portion of the integrated facility
is $301.5 million with the increase in cost attributable to
the larger steam supply system required for the dual-purpose
operation. However, since in the integrated facility steam
will be "sold" at cost to the distillation plant, the bus
bar electrical energy costs remains unchanged at 9.1 mills/
Kw-hr.
The capital and operating costs for a separate 50 MGD con-
ventional activated sludge wastewater treatment plant, in-
cluding outfall structure, chlorination facilities, and
administrative facilities are presented in Tables 15 and 16.
The $25.7 million capital cost of secondary wastewater treat-
ment at such a facility are equivalent to approximately
16.2^/1000 gallons assuming a fixed charge rate 7.842%.
The capital cost allocated to the wastewater treatment por-
tion of the integrated facility totals $20.3 million and is
detailed in Table 17. The operating costs for such a plant
are estimated to be $861,000 annually and are detailed in
Table 18. The wastewater treatment costs, including fixed
charges and operating costs are equivalent to 13.5C/1000
gallons.
The estimates in Tables 15 and 17 for wastewater pumping re-
flect several pumping stations located in the sewage
75
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TABLE 15
50 MOD WASTEWATER TREATMENT PLANT COST ESTIMATE
CONVENTIONAL SYSTEM (NONINTBQRATEDf UNHEATED)
(ENR 1690)
Raw Water Pump Station $ 2,415,000
Grit Chambers 685,000
Primary Clarifiers 1,005,000
Primary Sludge Pumping Station 120,000
Aeration Basins 2,750,000
Aeration Equipment 1,280,000
Final Clarifiers 1,300,000
Return Sludge System 380,000
Gravity Sludge Thickeners 400,000
Anaerobic Sludge Digesters 2,415,000
Vacuum Filter Facilities 1,790,000
Yard Piping 2,060,000
Chlorination Facilities 405,000
Administration Building 160,000
Maintenance Facilities 160,000
Outfall 2,415,000
Subtotal 19,740,000
Engineering and Contingencies d> 30% 5,922,000
TO
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TABLE 16
50 MGD WASTEWATER TREATMENT PLANT OPERATING COSTS
CONVENTIONAL SYSTEM (NONINTEGRATED, UNHEATED)
A. Annual Fixed Charge @ 7.823% of $25,662,000 = $2,008,000
B. Annual Operating and Maintenance Charges
1. Labor $ 320,000
2. Power 172,000
3. Chemicals 62,000
4. SUBTOTAL 5 54,000
5. Miscellaneous @
10% (1 & 2 & 3) 56,000
6. SUBTOTAL 610,000
7. Contingencies d>
10% 60,000
8. SUBTOTAL 670,000
9. Maintenance Budget 270,000
TOTAL Annual 0 & M 940,000
C. TOTAL Annual Charges 2,948,000
D. Water Costs - 50 MGD Treatment 16.2C/1000
77
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TABLE 17
50 MGD WASTEWATER TREATMENT PLANT COST ESTIMATE
INTEGRATED SYSTEM, PLANT HEATED BY BAROMETRIC CONDENSER*
ENR 1690
Raw Waste Pumping Stations $ 2,415,000
Included in distillation
Barometric Condenser , . . /-vu-,.^4-,,,. TV
plant cost-see Chapter IX
Grit Chambers 580,000
Primary Settling Tanks 810,000
Primary Sludge Pump Station 130,000
Aeration Basins 2,240,000
Aeration Equipment 1,280,000
Final Clarifiers 1,055,000
Return Sludge System 395,000
Gravity Thickeners 340,000
Anaerobic Sludge Digesters 2,415,000
Vacuum Filter Facilities 1,790,000
Maintenance Facilities 160,000
Yard Piping 2,000,000
Subtotal $15,610,000
Engineering and Contingencies d> 30% 4,685,000
Total $20,295,000
\
*-Inlet to grit chamber, 93°F; primary settling tanks, 92°F;
aeration basins, 86°F (30°C); plant outlet, 85°F.
78
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TABLE 18
50 MGD WASTEWATER TREATMENT PLANT ANNUAL OPERATING COSTS
INTEGRATED SYSTEM, PLANT HEATED BY BAROMETRIC CONDENSER
A. Annual Fixed Charge @>7.823% of
20,295,000 1,588,000
B. Annual Operating and Maintenance
Charges
1. Labor 300,000
2. Power 172,000
3. Chemical 0,000
4. Sub Total 472,000
5. Misc.d>10% (1&2&3) 47,000
6. Sub Total 519,000
7. Contingencies @10% 52,000
8. Sub Total 571,000
9. Maintenance 290,000
Total Annual O & M 861,000
C. Total Annual Charges 2,449,000
D. Water Costs - 50 MGD Treatment 13.5C/1000 gal.
79
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districts to convey the wastewater via forced main to the
facility site and, thereby, preclude the need for pumping
facilities on the site.
The primary sludge and return sludge pumping stations and
maintenance building sizes and costs were adjusted to re-
flect the 3 percent increase in flow due to the steam con-
densate from the barometric condenser used to supply heat
to the wastewater.
To operate the aeration basins at 86°F (30°C) the tempera-
ture of the water entering the grit chamber is 93°F and in
the primary clarifiers is approximately 92°F. Reductions in
grit chambers and primary clarifier costs reflect operation
at these temperature levels. The required horsepower to
force oxygen into the aeration tanks is reasonably constant
over the temperature range considered. Although the ability
to transfer oxygen increases with temperature, oxygen satur-
ation concentration decreases simultaneously and the two ef-
fects cancel.
Cost data for an integrated, unheated, 50 MGD plant are
given in Table 19. In this case, chlorination facilities,
an outfall structure and an administration building are all
eliminated. Disinfection, by chlorination is not required
since the process following final waste treatment is dis-
tillation. The outfall is not required because the entire
treated effluent flow is passed through the distillation
plant, or in the event of a distillation plant shut down,
through the outfall of the power plant cooling water system.
The administrative facilities of the power plant-distilla-
tion plant complex are considered to be adequate to service
the needs of the wastewater treatment plant.
Comparison of the conventional nonintegrated, unheated plant
and the integrated, unheated plant cost estimates indicates
that savings in treatment plant construction costs on the
order of 15 percent are possible when the wastewater treat-
ment facility is incorporated to the power plant-distilla-
tion plant complex. This saving is due to integration alone
and the effect of heat addition on the cost saving is not
included.
By comparing the capital costs for the integrated unheated
80
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TABLE 19
50 MGD WASTEWATER TREATMENT PLANT COST ESTIMATE
INTEGRATED SYSTEM. PLAJT NOT HEATED
ENR 1690
Total Cost, Conventional System $25,662,000
Cost of items to be eliminated:
• Chlorination Facilities $ 405,000
• Administration building 160,000
• Outfall 2,415,ooo
Subtotal $2,980,000
• Engineering & contingencies 890,000
Total Savings $3,870,000
Total Cost, unheated, integrated system 21,792,000
Net Benefit of integration 3,870,000
Percentage Cost Reduction due to integration . .
(Basis-Total Cost, Conventional System)
and heated plant, it is readily apparent that a reduction of
$1,497,000 is possible. This figure represents an addition-
al 5 percent reduction in the construction cost of the waste-
water treatment facility, over and above the 15 percent
realized by integration. However, if the cost associated
with the addition of heat to the wastewater, which is pri-
marily reflected in increased capital and operating costs
for the distillation plant, is charged to the wastewater
treatment plant, the cost is increased by approximately
1.6C/1000 gallons, for a total cost of 15.1C/1000 gallons.
Table 20 shows the effect upon capital costs by adding heat
to the wastewater treatment facility by means of a shell
and tube heat exchanger. The major cost difference, as com-
pared to the barometric condenser case, is associated with
81
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TABLE 20
50 MGD WASTEWATER TREATMENT PLANT COST ESTIMATE
INTEGRATED SYSTEM,
PLANT HEATED BY SHELL & TUBE HEAT EXCHANGER
ENR 1690
Total Treatement Plant Cost, Barometric Condenser System
From Table 16 $20,295,000
Cost additions, using shell and tube heat exchange:
1. Heat exchange downstream of grit chamber use
conventional grit chamber cost
Cost addition: 685,000-580,000 = $105,000
Cost reductions, reflecting 3% flow reduction due to
removal of barometric condenser:
1. Primary clarifiers: 0.03 x 810,000 = 25,000
2. Aeration basins: 0.03 x 2,240,000 = 67,000
3. Secondary clarifiers: 0.03 x 1,055,000 = 32,000
Total Cost Reductions 124,000
Net cost reduction, including 30% contingency: $ 25,000
Total Treatment Plant Cost, Shell and Tube
the grit chamber. Because of potential erosion of the heat
exchanger, the sewage must be degritted prior to being
heated. Thus, the cost for the grit chamber in the conven-
tional unheated plant is identical to. the value shown in
Table 15, and increases the cost by $105,000. Against this
increase in cost, a 3 percent reduction in cost of the wet
stream plant size is possible with the elimination of the
barometric leg condenser. The saving in cost resulting
from this size reduction is $124,000, which after deducting
the $105,000 increase in the grit chamber cost and applying
82
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TABLE 21
CAPITAL COST - 50 MGD INTEGRATED FACILITY HEAT ADDED BY HEAT EXCHANGER
oo
Temperature
Raw Waste Pump Station
Grit Chamber
Primary clarifier
Primary Sludge Pump Station
Aeration Basin
Mechanical Aeration
Secondary Clarifier
Return Sludge & Pump Station
Thickener
Anaerobic Digestion
Vacuum Filter
Maintenance Building
jYard Piping
Subtotal
Eng. & Contingencies (30%)
20°C
30°C
40°C
50°C
$2,415,000 $2,415,000 $2,415,000 $2,415,000
685,000 685,000 685,000 685,000
710,000
143,000
2,244,000
1,280,000
900,000
408,000
275,000
2,415,000
1,790,000
167,000
2,136,000
1,005,000
120,000
2,750,000
1,280,000
1,300,000
380,000
400,000
2,415,000
1,790,000
160,000
2,060,000
785,000
130,000
2,173,000
1,280,000
1,023,000
395,000
340,000
2,415,000
1,790,000
160,000
2,000,000
620,000
150,000
2,276,000
1,280,000
760,000
418,000
260,000
2,415,000
1,790,000
177,000
2,133,000
16,760,000 15,591,000 15,568,000 15,379,000
5,032,000 4,679,000 4,670,000 4,614,000
TOTAL
$21,792,000 $20,270,000 $20,238,000 $19,983,000
-------
the 30 percent contingency factor used in preparing the plant
estimates yields an additional saving of $25,000, exclusive
of the cost of the heat exchanger.
Table 21 demonstrates the relationship between capital costs
and the temperature at which the plant is operated, again
with heat transferred by means of shell and tube heat ex-
changers. The importance of the size of the aeration basin
in this analysis is evident, particularly at the 30°C level
where the aeration basin cost is at a minimum. Clearly, the
cost saving to be achieved for temperatures in excess of 30°C
are relatively inconsequential compared to the saving asso-
ciated with the 30°C operation.
Table 22 summarizes the foregoing discussions of various al-
ternatives for integrating the wastewater plant and shows
that a 15 percent reduction in the construction cost of the
wastewater treatment facility is possible through integra-
tion, with an additional 6 percent reduction in cost avail-
able through heat addition.
TABLE 22
COMPARISON OF COSTS OF VARIOUS WASTE TREATMENT PLANT CASES
Plant and System
Conventional Separate
Plant, Unheated
Integrated System,
Unheated
Capital Cost
for
50 MGD Plant
(In Million $)
25.66
21.79
Percentage Reduction
Over
Conventional Plant
15
Integrated System,
Heat Via Barometric
Condensers
Integrated System,
Heat via Shell and
Tube Exchanger
20.30
20.27
21
21
84
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Comparison of wastewater treatment costs, however, must take
into consideration the fact that conventional secondary
treatment with discharge to the environment is no longer ac-
ceptable. Tertiary waste treatment is required to approach
the "zero discharge" criterion. The capital cost of waste-
water facilities providing tertiary treatment is estimated
to be more than twice the cost of facilities providing sec-
ondary treatment only and the total annual operating costs
for these facilities are estimated to be approximately three
times the cost of present conventional waste treatment sys-
tems (31). On this basis a separate waste treatment plant
or plants serving the study region and designed to meet the
water quality goals incorporated in the Federal Water Pollu-
tion Control Act Amendments of 1972 (32) would cost in excess
of $50 million and the cost of the wastewater treatment would
be on the order of 55C/1000 gallons. Elimination of the need
for tertiary waste treatment provides an additional benefit
for the integrated facility approach.
An alternate approach to meeting the water demands of the
region is seawater desalting. The projected economics of
this process are indicative of an advantage for large scale
plants similar in size to that proposed for the integrated
facility, i.e., in the range of 40 to 60 MGD. In order to
achieve operating economy, these plants will be coupled to
a dual-purpose power generating facility, most probably nu-
clear fueled. It is, of course, possible for a desalting
plant to have its own thermal energy source. Table 23 pre-
sents the capital cost of such a nonintegrated 50 MGD dis-
tillation plant. The present and projected costs of fossil
fuels, however, indicate that even the most efficient single
purpose fossil plants would incur a cost penalty of at least
25C/1000 gallons over a dual purpose nuclear facility.
Present projections of the cost of large seawater desalting
plants coupled with nuclear power generating facilities such
as the 40 MGD plant at Diablo Canyon, California (capital
cost of $92 million, and product water cost of 92C/1000 gal-
lons) or the pair of 20 MGD flash train evaporators at Encina,
California (capital cost of $72 million and product water cost
of 99C/1000 gallons) (34) illustrates the potential savings
associated with an integrated facility having a distillation
plant capital cost of $63.2 million, and product water cost
of 62£/1000 gallons.
85
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TABLE 23
CONVENTIONAL NONINTEGRATED 50 MGD VTE DISTILIATION PLANT
Unit Operation Cost
Steam & Power Supply-Fossil $ 17,100,000
Distillation Plant 54,200,000
Intake/Outfall Structure 6,100,000
^
Product Steam Condenser 3,000,000
Product Water Cooling 1,700,000
Pretreatment 270,000
Activated Carbon 1,270,000
Post Treatment 200,000
$ 83,840,000
The cost for the integrated distillation plant includes the
cost of product post-treatment and the full cost of chlorin-
ation and administrative facilities shared with the waste-
water treatment plant, but excludes the cost directly appli-
cable to the requirement for thermal enhancement of the "
wastewater treatment process. The unit process cost of
62C/1000 gallons includes 17.7C/1000 gallons ($3.1 million
per year) for steam purchased from the power facility.
In summary, the total capital cost of the integrated facility
is estimated to be $386 million with the total annualized
cost estimated to be $77 million per year.
The facility is designed to be capable of producing 7 billion
Kw/hr per year of electricity at a cost of 9.1 mills/Kw-hr
and 47.5 MGD of high quality product water at 62<:/1000 gal-
lons, while treating 50 MGD wastewater at a cost of approxi-
mately 15.1C/1000 gallons.,
Summaries of the capital and annual cost of the integrated
facility and the revenues required to recover these costs
are presented in Tables 24 and 25.
86
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TABLE 24
INTEGRATED FACILITY COST SUMMARY
I. CAPITAL INVESTMENT (ENR 1690) $106
A. Nuclear Steam Supply and Power Generation
Nuclear Steam Supply @ $175/Kwe $192.5
Turbogenerator @ $109/Kwe 109.0
$301.5
B. Waste Treatment Plant
Basic Plant $ 20.3
Heat Addition 2.0
$ 22.3
C. Distillation Plant
Basic Plant $ 61.9
Product Post-Treatment 1.3
$ 63.2
TOTAL CAPITAL COSTS $386.0
II. ANNUAL COSTS
A. Steam Supply and Power Generation
Fixed Charges @ 15.0% $ 45.2
Fuel 16.2
Operation and Maintenance 5.3
Less: Proceeds from sale of steam to
distillation plant -3.1
$ 63.6
B. Wastewater Treatment
Fixed Charges @> 7.823% $ 1-75
Operation and Maintenance -_99_*
$ 2.74*
C. Distillation Plant
Fixed Charges @ 7.823% $ 4.94
Operation and Maintenance 5.82**
$ 10.76**
TOTAL ANNUAL COSTS $ 77.2
(excluding interutility purchases)
*Includes $130,000 distillation plant costs for wastewater
heating.
**Includes $3,082,000 for steam "purchased" from the power
facility.
87
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TABLE 25
INTEGRATED FACILITY
PRODUCTS AND REVENUES
Electricity
$106
7 x 109 Kw/-hr/yr @ 9.1 mills/Kw-hr $ 63.6
Potable Water
47.5 MGD <§> 62C/1000 gallons $ 10.8
Treated Waste
50 MGD @ 15.1C/1000 gallons $ 2.8
TOTAL REVENUES $ 77.2
88
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SECTION XI
PILOT SCALE DEMONSTRATION PROJECTS
As discussed in the preceding sections and the appendices to
this report, limited design and performance data presently
exist on the heating of wastewater, the effects of heat on
wastewater treatment processes, and the use of distillation
processes to produce potable water from treated wastewater.
In addition, public health considerations would preclude the
direct recycle of wastewater, even using the distillation
process, until the ability to produce high quality water con-
tinuously and reliably is demonstrated. For these reasons,
a phased program for the development and demonstration of
the required technology is recommended prior to implementing
plans for a full-scale integrated facility complex. This
program will consist of three phases:
Phase I - Component Tests and Laboratory
Screening Studies
Phase II - System Performance Tests and
Demonstrations
Phase III - Integrated Pilot Scale
Demonstrations
Phase I, Components Tests and Laboratory Screening Studies,
will be directed to developing the fundamental design and
performance data with respect to components and subsystems
needed to validate the assumptions and theoretical predic-
tions made in this study and to provide a basis for the sub-
sequent system tests and pilot scale demonstrations. In
Phase II, System Performance Tests and Demonstrations, a
distillation plant will be combined with heated wastewater
treatment and prototype heat exchangers to simulate an in-
tegrated system at the 50,000 gpd scale. Phase III, Inte-
grated Pilot Scale Demonstrations, will consist of operating
a heated wastewater treatment plant and a distillation plant
at an existing sewage plant site. To be meaningful, the
Phase III program should be based on an integrated plant
having a capacity of at least 500,000 to 5,000,000 gallons
per day.
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REQUIREMENTS FOR DESIGN AND OPERATING DATA
The objectives and goals of the recommended development and
demonstration program can best be defined in terms of the
subsystems that will comprise the integrated facility com-
plex and the currently available and required technology for
the design of these subsystems.
Nuclear Power Plant
As discussed in prior sections, the principal change in the
nuclear power plant will consist of providing the additional
thermal output needed to provide the energy required for
heating the wastewater and for the distillation plant. With
electrical output in the 1000 Mwe range, and with wastewater
and product water flows in the 50 MGD range, the required
additional output is only approximately 10 percent, which
can be provided by selecting currently available reactor de-
signs and matching them with available power generation
equipment. No special development or demonstration projects
will be required.
Radioactive materials will be present in the steam systems
of boiling water reactors and, to a lesser extent, in pres-
surized water systems. For the reasons discussed in Sec-
tion VI, a reboiler is used to provide isolation of this
steam from the wastewater and distillation plants. The use
of reboilers for other purposes is a common practice and
thermal and hydraulic design can be considered to be state-
of-the-art.
The other aspect of power plant integration to be considered
is the possible use of the condensers for wastewater heat-
ing, which is discussed later.
Heated Wastewater Treatment
As discussed in Section VIII and Appendix A, heating of
wastewater improves treatment processes by increasing the
rates of biological activity and improves sedimentation by
decreasing the density and viscosity of the wastewater. The
review of existing literature relative to sewage treatment
at elevated temperatures indicates that fully conclusive
data are not available with respect to the quantitative or
90
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qualitative effects of heat addition. This is exemplified
by the narrow band of elevated temperature experience and
the limited references to such qualitative effects as changes
in the oxidation-synthesis relationships which control bio-
logical treatment.
The limited quantity of available literature has constrained
the work described herein to the use of conservative esti-
mates of the reductions in size of conventional wastewater
treatment operations resulting from heat addition. Addi-
tionally, literature data have not been sufficient to pro-
vide a basis for quantitative process designs for treatment
unit operations other than those employed in conventional
biological treatment flow sheets.
Most significantly, reports appearing in the third quarter
1972 literature of qualitative changes in biological treat-
ment performance due to heat enrichment, and operating ex-
perience with elevated temperature biological treatment of
soluble industrial wastes, strongly indicates that greater
size reductions are possible than were considered for the
reference design, and that continued analytical work be un-
dertaken to define the quantitative and qualitative effects
of thermal enrichment in wastewater treatment operations.
Heating of Wastewater
As discussed in Sections I and XII and Appendix B, there are
three potential sources of energy for wastewater heating:
Power plant condensers
Distillation plant product water cooling
Injection of steam from low pressure stages
of the distillation plant
Of these energy sources, heating of wastewater in the power
plant condensers or the distillation plant product water
coolers appears to be the most attractive, since cooling
water would normally be required to remove the waste heat
from these sources.
Using waste heat from either the power plant condensers or
the distillation plant product water coolers will require
91
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the use of surface type heat exchangers. Data on fouling
factors and attainable heat transfer coefficients, the avail-
able temperature differences, and the materials that must be
used to resist corrosion must be developed in order to eval-
uate the options involving the use of heat exchangers.
The use of injection steam from the low pressure stage using
barometric leg condensers is attractive in that the heat ex-
change equipment costs are low, fouling problems are avoided,
and a system of this type could be built based on present
technology. This approach is economically limited in that
the size of the distillation plant, the flow through the
wastewater treatment plant, and the total energy required
are all increased. The cost of using this approach is esti-
mated to be less than that of surface type wastewater heat
exchangers, conservatively designed on the basis of present
technology and the very limited existing data. Pursuing this
approach further is not recommended because advances in tech-
nology cannot be expected to improve significantly the
economics of this approach.
The principal area in which productive technological oppor-
tunities exist which could make heated wastewater treatment
more economically attractive is the development of reliable
and economic wastewater heat exchangers suitable for use in
transferring waste heat from the power plant condensers or
the distillation plant product water cooler.
The most potentially productive opportunities lie in:
1. Increasing the heat transfer coefficients through
material selection and the use of coatings and
additives to inhibit scaling and fouling.
2. Maintaining desired heat transfer rates through
periodic mechanical or chemical cleaning of the
heat transfer surfaces.
3. Developing heat exchanger designs of high reli-
ability and ease of maintenance. Such designs
would explore the advantages of higher flow
velocities and flow turbulence in enhancing
heat transfer.
92
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Distillation of Treated Wastewater
The distillation plant design presented in Section IX is
based on the vertical tube evaporation process developed for
the desalination of sea and brackish water. Although devel-
opment and demonstration work is continuing to improve the
performance, efficiency, and economics of this process, the
VTE process can be considered state-of-the-art. Further,
multistage flash evaporation and the other existing distil-
lation desalting processes could also be used with waste-
water. Work is also underway on the use of membrane proces-
ses for the treatment of wastewater. Because of fouling,
biological growth, and the possibility of undetectable fail-
ures, it is felt that membrane processes should not be con-
sidered for direct reuse applications.
I
Even though the equipment and processes proposed for waste-
water distillation can be considered to be state-of-the-art,
there are a number of uncertainties with respect to distil-
ling secondary treated wastewater. With the activated sludge
process, a large portion of the nitrogen in the secondary
treated effluent Will be in the form of ammonia. Since
ammonia has a higher vapor pressure than the water in which
it is dissolved, it will be evaporated with the water vapor
in the distillation process. From a theoretical standpoint
and with proper selection of operating parameters, it should
be possible to remove this ammonia in a gaseous form with
the noncondensible gases. Once removed, the ammonia would
be condensed for removal from the process. This will re-
quire a separate system for handling the condensate from the
air ejectors and may require special materials of construc-
tion. For these reasons, the ammonia removal process and
equipment should be demonstrated.
Other materials potentially present are phenols and other
volatile organic materials which have boiling and condensing
temperatures and pressures similar to water or are volatile
and highly water-soluble. Such materials could be carried
over with the product water, but at low concentrations could
be removed inexpensively by carbon absorption as a post-
distillation treatment process.
It will be necessary to have extensive distillation process
operating and performance data to show extremely high reli-
93
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ability before direct product water reuse could be considered,
Even in cases where the product water is for use in ground-
water recharge, industrial processes or as makeup for evapor-
ative power plant cooling systems such as cooling towers, a
high degree of process control and product purity will be
achieved.
PHASE I COMPONENT TESTS AND LABORATORY SCREENING STUDIES
The Phase I Component Tests would consist primarily of bench
scale experiments concentrating on the development of basic
data required to design and build prototype equipment. In
the case of the distillation system, there is minimal need
for component testing and work in this area can concurrently
proceed with the Phase II system demonstrations using waste-
water from a conventional secondary treatment plant. This
approach would allow a longer demonstration period for the
distillation plant and the accumulation of reliability and
maintainability information.
Laboratory Screening Studies of Wastewater Treatment Unit
Operations
Theory indicates that thermal enrichment of wastewater could
reduce the size and cost or improve the efficiency of waste-
water treatment by at least 20 percent. Very recent operat-
ing experience in treating soluble industrial wastes at
elevated temperatures indicates much higher percentages may
be attainable.
The degree to which wastewater treatment processes can be im-
proved must be demonstrated. The Phase I laboratory scale
screening studies of candidate unit operations are designed
to achieve the following purposes:
1. Determine actual performance characteristics of
activated sludge treatment operations at the
elevated temperatures considered in this analysis.
2. Evaluate the qualitative changes in the nature of
the activated sludge treatment process at elevated
temperatures.
94
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3. Determine the performance characteristics of non-
biological treatment operations at elevated temper-
atures, for example, chemical treatment and chemical
sludge disposal.
The unit operations to be screened will include those consti-
tuting the 50 MGD plant flow sheet developed in this study
and additional unit operations that might, based on new in-
formation, be indicated to improve upon this flow sheet. The
unit operations to be tested include grit removal, settling,
activated sludge, chlorination, thickening, digestion and
dewatering.
Two series of screenings tests will be performed. The first
series will be batch analyses to determine individual unit
operation responses. With the information gained through
these tests, a continuous, in line process will be set up to
model the proposed demonstration project and to determine
the interaction of the various processes.
Laboratory scale results will be analyzed in terms of perform-
ance and economic implications to identify a priority ordering
of unit operations for pilot plant verification, and to
project the process benefits and cost savings available
through heat addition to sewage processes. Information
relevant to the optimum process temperatures for the pilot
plant will be developed and potential operational problems
that might hamper treatment at elevated temperature identi-
fied.
Wastewater Heating
The Phase I Component Test Program on wastewater heating
consists of a series of surface heat exchanger heat transfer
experiments using wastes of various concentrations over the
temperature and flow range of interest. For these tests, it
is recommended that electrically heated tubes be used in con-
junction with the portable heat transfer and fouling test
equipment of the type developed and used by the Heat Transfer
Research Institute. This equipment permits the variables
which effect heat transfer and fouling to be investigated.
The tubes used in these heat transfer experiments will also
be used as corrosion test samples and will be examined
metallurgically following the tests. The variables to be
evaluated include:
95
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1. Degree of wastewater treatment (primary and second-
ary) needed prior to heating.
2. Temperatures from 65°F to 110°F and 150°F to 300°F.
3. Flow velocities from 2 to 30 feet per second.
4. Tube materials consisting of carbon steel, stain-
less steel, admiralty brass and other alloys.
5. Inhibitors for corrosion and fouling.
6. Descaling solutions.
In addition to the experimental program, parallel design and
application studies will be conducted on surface heat ex-
change equipment. These studies will review and analyze
various heat exchanger configurations for use in an inte-
grated plant.
PHASE II SYSTEM PERFORMANCE TESTS AND DEMONSTRATIONS
Distillation Plant System Demonstrations
The distillation plant system demonstrations will be conduct-
ed using equipment developed for use with seawater or brack-
ish water. Ideally, equipment already in the possession of the
the Office of Saline Water, U.S. Department of the Interior
would be obtained on loan for this purpose. The distillation
plant will require a source of steam and secondary treated
wastewater. A 50,000 gpd distillation plant for instance
will require about 170 MBtu per day, or approximately 7000
pounds of steam per hour, which is the equivalent of a 200
hp packaged boiler using about 50 gallons of fuel oil per
hour.
The distillation plant system demonstration could alterna-
tively be located at an existing power plant near an exist-
ing wastewater treatment plant. A pumping station could be
used to convey the treated wastewater to the distillation
plant and the plant could use extraction or low pressure
steam from the power plant. Another alternative would be
96
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location of the distillation plant where it can be supplied
with secondary treated wastewater and use of a packaged
boiler.
The distillation plant will require some ancillary subsys-
tems for use with wastewater, including provisions for am-
monia removal and for post-distillation treatment. The am-
monia removal system will consist of separate steam air
ejectors for the removal of noncondensibles and ammonia
vapor from the initial stages, an after-condenser, and a
small ammonia stripping tower for removal of the ammonia
from the condensate. The post-treatment subsystem may con-
sist of parallel carbon adsorption columns with provision
for thermal or chemical recycling if phenols are present and
not otherwise removed.
The test installation would be supported by a water analysis
laboratory, probably the one utilized in the Phase I work.
The test and demonstration program for the distillation plant
would consist of a series of test runs to attain operating
and performance data, periodic inspections to determine
material compatability, corrosion rates, etc., and special
tests to simulate process malfunction.
Specific items to be evaluated would include:
1. Operation at various temperatures.
2. Variable bottoms recycle rates and concentrations.
3. Variable feed/product ratios.
4. Use of acid feed.
5. Use of various methods of pretreatment.
6. Descaling methods.
7. Startup, shutdown, and partial load conditions.
8. Water quality as a function of production.
97
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9. Effect of off-standard conditions on water quality:
a. System flooding;
b. Loss of vacuum;
c. Loss of cooling water; and
d. Loss of feed.
10. Post-distillation removal effectiveness.
11. Carbon column recycle times.
PHASE III INTEGRATED PILOT SCALE DEMONSTRATIONS
The demonstration will be conducted on a scale adequate to
permit extrapolation to full scale. Thus the capacity should
be at least 500,000 gpd and, preferably, on the order of
5,000,000 gpd to approach within a factor of 10, the designs
considered in this study. In this phase, heat addition and
related facilities and equipment will be designed and in-
stalled using the design and operating data developed. The
demonstration will involve addition of heating capability to
a soon to be constructed municipal wastewater treatment
plant. Heat will be provided from a complex of power facil-
ities located on immediately adjacent land.
The Phase III demonstration will provide an actual opera-
tional evaluation of heated wastewater processing. Like a
full scale facility, only a limited number of tests will be
conducted to validate the design and performance predictions.
The major emphasis in the Phase III demonstration will be
the collection of performance, reliability, and maintenance
data and treated water quality information.
98
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SECTION XII
ACKNOWLEDGEMENT
The support of the project by the Water Quality Office of
the Environmental Protection Agency and the guidance and
assistance provided by Mr. Patrick M. Tobin and Mr. John
R. Trax, the Agency's Grant Project Officers, is grate-
fully acknowledged.
99
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SECTION XIII
REFERENCES
1. U.S. Bureau of the Census,"Statistical Abstracts of the
United States: 1970,"(91 Edition) Washington, D.C.
2. Bowe, Walsh & Associates, "Comprehensive Sewerage Studies,
Five Eastern Towns, Suffolk County, New York," WPC-CS-158,
(Preliminary Draft).
3. New York State Office of Planning Services, Unpublished
preliminary population projections, July 1972,
4. Holzmacher, McLendon & Murrell, "Comprehensive Public
Water Supply Study,"Suffolk County, New *ork, CPWS-24,
(3 Volumes) 1968-1970.
5. Division of Water Resources State Conservation Dept. for
State Office of Planning Coordination, "Long Island Water
Resources," January 1970.
6. County of Nassau, State of New York, Greeley and Hansen
Engineers, "Comprehensive Public Water Supply" (CPWS-60)
June 1971.
7. Wiggins, Thomas H., "Report on Comprehensive Plan for
the Development and Distribution of the Available Water
Supply of Suffolk County, Long Island, New York," Suffolk
County Water Authority, January 1957.
8. McGraw-Hill, Electrical World Magazine, September 15, 1970.
9. Battelle Memorial Institute,"A Review and Comparison of
Selected United States Energy Forecasts!' December 1969-
10. Long Island Lighting Company, "Population Survey 1969-
Current Population Estimates for Nassau and Suffolk
Counties"
11. Statistical Department, Long Island Lighting Company,
"Population Survey 1964."
12. United States Code of Federal Regulations, "Criteria for
the Siting of Nuclear Power Facilities," Title 10,
Part 100.
100
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13. Na&sau-Suffolk Regional Planning Board, "Existing Land
Use," 1968,
14. "Preliminary Safety Analysis Report for Surfside
Multipurpose Nuclear Reactor Facility," (Unpublished).
15. Vann, Harold E., "Cost Trends for Nuclear Power Plants,"
Nuclear News/Geneva III, pp. 31-34, October 15, 1971.
16. Hauser, L. G., and R. F. Potter, "The Effect of Es-
calation on Future Electric Utility Fuel Costs," Nuclear
Fuel Division, Westinghouse Electric Corporation, 1971.
17. Leung, Paul, "Cost Separation of Steam and Electricity
for a Dual Purpose Station," Bechtel Corporation,
San Francisco, California, April 1972.
18. Hittman Associates, Inc., Study for EPA Pacific
Northwest Laboratories, "Nomographs for Thermal
Pollution Control Systems," (to be Published).
19. Long Island Lighting Company, "Environmental Impact
Evaluation for the Shoreham Nuclear Power Station,"
December 1971.
20. State of Maryland Dept. of Health & Mental Hygiene
Regulations for the Control of Air Pollution.
21. Dynatech R/D Company, Report for Water Quality Office
EPA, "A Survey of Alternate Methods for Cooling Condenser
Discharge Water-Large Scale Heat Rejection Equipment,"
July 1969.
22. Strobel, J. J., "Overall Economic Considerations of
Desalination Operations," OSW, St. Croix Conference,
December 1971.
23. "Diablo Canyon Prototype Desalting Plant Design Study,"
Kaiser Engineers Report No. 70-60-RE, Dept. of Water
Resources and OSW, California, January 1972.
24. U.S. Dept. of Health, Education, and Welfare, Public
Service Commission, "Public Health Service Drinking
Water Standards," 1962.
101
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25. World Health Organization, Geneva, Switzerland,
"International Standard for Drinking Water," 1963.
26. American Water Works Association, Inc., Water Quality
and Treatment, 1971.
27. Persona 1 correspondence with Roma Bhattacharya, Product
Specialist Heat Transfer Dept., Bell & Gossett and
Wallace Eannace Asso ates, Inc.,Bell & Gossetts'
New York Representatives.
28. Personal correspondence with William R. Ryan, Sales
Manager Water Products & Environmental Systems,
Westinghouse Electric Corp. Power System, Heat
Transfer Division.
29. Personal correspondence with Anthony J. Giuriceo, Sales
Engineer, Mesco Tectonics, Inc.
30. Hittman Associates, Inc., HIT-524 (draft to be Published)
"Vertical Tube Distillation Desalting State-of-the Art
1971."
31. Metzler, D.F. and F.O. Bogedain, "The Cost of Water
Quality Goals," National Symposium ort Cost of Water
Pollution Control, April 1972,
32.. Federal Water Pollution Control Act Amendments, 1972,
Public Law 92-500, Title Three.
33. Office of Saline Water, "Draft Environmental Impact
Statement for Diablo Canyon Prototype Desalting Plant,"
April 1972.
34- Water Desalination Report, Richard Arlen Smith, Publisher-
Editor, P.O. Box 35-K, Tracey's Landing, Md. 20869.
Vol. Ill, No. 36.
102
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SECTION XIV
LIST OP PATENTS AND PUBLICATIONS
No inventions, patents or publications have resulted from
the performance under or in conjunction with this contract,
103
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SECTION XV
GLOSSARY OF TERMS, ABBREVIATIONS
TERMS
Coagulation
BOD
COD
Floe
Hindered settling
Solid Flows
Aerobic
Anaerobic
Mesophilic
Thermophilic
The process of agglomeration of small
particles into larger particles through
agitation with or without the aid of
chemicals.
Biochemical Oxygen Demand.
Chemical Oxygen Demand.
A particle formed by smaller particles
through coagulation, usually promoted
by chemical addition.
Settling of particles in a liquid medium
wherein particles do not behave as a
single particles because of the inter-
action of other near field particles.
Downward passage of solids in a thicken-
ing unit process.
Processes taking place in the presence of
oxygen
Processes taking place in the absence of
oxygen
Pertains to a group of microorganisms
that thrive in a temperature range of
about 30 to 40°C.
Pertains to a group of microorganisms
that thrive in a temperature range of
about 40 to 50°C.
104
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Digestion
Pathogen
Disinfection
A process where complex organic com-
pounds are decomposed into methane and
carbon dioxide gases by facilitative
anaerobic microorganisms. Digestion
can also be accomplished by aerobic
bacteria, to produce carbon dioxide
and ammonia
A microorganism that produces disease.
The elimination of pathogenic and other
microorganisms by chemical addition or
other means.
Tertiary treatment
Processes that are added to secondary
waste treatment facilities to improve
the quality of the effluent.
Abbreviation
\
ASDA
AEC, USAEC
Btu
BWR
csteam
10 CFR 100
COD
CW
DEC
ENR
EPA
gpcd
gpd
gpm
"Hg, "Hg abs
HP
HTGR
Meaning
New York State Atomic and Space Develop-
ment Authority
United States Atomic Energy Commission
British thermal unit (s)
Boiling water reactor
Cost of steam, cents per 10^ Btu
Title 10, U.S. Code of Federal Regula-
tions, Part 100
Chemical oxygen demand, ppm
Cooling water
New York State Department of Environ-
mental Conservation
Engineering News Record Construction
Cost Index
United States Environmental Protection
Agency
Gallons per capita per day
Gallons per day
Gallons per minute
Absolute steam pressure, inches of
mercury
High pressure (steam), > 100 psia
High temperature gas-cooled reactor
105
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Kw, Kwe
LILCO
LWR
MBtu
MGD
MP
MSF
Mw, Mwe
Mwt
NSSS
O&M
QC
PHS, USPHS
psia, psi (abs)
PWR
VTE
Electrical kilowatts
Long Island Lighting Company
Light water r.eactor
Million Btu
Millions of gallons per day
Medium pressure (steam, > 50 psia)
Multistage flash distillation system
Electrical megawatts
Thermal megawatts
Nuclear steam supply system
Operations and maintenance (costs)
Quality control
United states Public Health Service
Absolute pressure, pounds per square inch
Pressurized water reactor
Vertical tube evaporator
106
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APPENDIX A
TEMPERATURE EFFECTS ON WASTEWATER TREATMENT PROCESSES
CONTENTS
Section
SUMMARY OF FINDINGS 112
A-I INTRODUCTION .* 113
A-II PHYSICAL-CHEMICAL PROCESSES 115
Grit Removal 115
Sedimentation 119
Coagulation 121
Gravity Filtration 122
Flotation 128
A-III SLUDGE HANDLING PROCESSES 130
Thickening 130
Dewatering 132
A-IV BIOLOGICAL PROCESSES 141
Biological Waste Treatment 141
Stabilization Ponds 145
Aerated Basins 149
Activated Sludge 153
Trickling Filters 158
Rotating Disks 158
Anaerobic Digestion 160
Aerobic Digestion 164
A-V DISINFECTION 167
Chlorination 167
A-VI ADVANCED TREATMENT PROCESSES 170
Ultra-High Rate Filtration 170
Organic Carbon Removal 170
Nitrogen Removal. 173
Phosphorus Removal 183
A-VII REFERENCES 184
A-VIII GLOSSARY 194
107
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FIGURES
Number
1 Effect of Temperature on the Efficiency of
Grit Removal, Unit Size Held Constant .118
2 Effect of Temperature on Grit Chamber Size
To Achieve a Constant 75% Grit Removal
Efficiency -118
3 Effect of Temperature on the Suspended Solids
Removal in Sedimentation Basins, Unit Size
Held Constant -120
4 Effect of Temperature on Sedimentation Basin
Surface Area To Achieve a Constant 50% Sus-
pended Solids Removal 120
5 Effect of Temperature on the Alum Dosage
Required To Produce Satisfactory Flocculation. .123
6 Effect of Temperature on the Time Required
for Floe Formation 123
7 Effect of Temperature on Head Loss through
Stratified Bed Gravity Filters 125
8 Effect of Temperature on Filter Capacity To
Produce a Constant Head Loss in Stratified
Bed Gravity Filters 125
9 Effect of Temperature on Required Backwash
Rate for Gravity Filters 127
10 Effect of Temperature on Filter Efficiency,
Constant Filter Size 127
11 Effect of Temperature on Mass Loading Rate
to Thickener To Produce a Constant Underflow
Concentration 133
12 Effect of Temperature on Thickener Size To
Produce a Constant Underflow Concentration 133
108
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FIGURES
(continued)
Number
Page
13 & 14 Effect of Temperature on the Compaction
Characteristics of Sludge Produced by the
Pure Oxygen Activated Sludge Process 134
15 Effect of Temperature on Vacuum Filter
Yield, Constant Solids Loading 137
16 Effect of Temperature on the Required
Vacuum Filter Size To Produce a Constant
Filter Yield at a Constant Solids Loading
Rate 137
17 Effect of Temperature on the Capacity of a
Constant Size Centrifuge 140
18 Effect of Temperature on Centrifuge Size
To Achieve a Constant Throughout Capacity 140
19 Effect of Temperature on the BOD Removal
Efficiency of a Constant Unit Size Stabili-
zation Pond 148
20 Effect of Temperature on the Required Size
of a Stabilization Pond To Produce a Con-
stant BOD Removal Efficiency 148
21 Effect of Temperature on the BOD Removal
Efficiency of a Constant Size Aerated Basin... 151
22 Effect of Temperature on the Required Size
of an Aerated Basin To Achieve a Constant
BOD Removal Efficiency 151
23 Optimum Temperatures for the Activated
Sludge Process Reported in the Literature .... 155
24 Effect of Temperature on BOD Removal Effi-
ciency of a Constant Size Completely Mixed
Activated Sludge Unit I57
25 Effect of Temperature on the Required Acti-
vated Sludge Unit Size To Produce a Constant
BOD Removal Efficiency I57
109
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FIGURES
(continued)
Number
Page
26 Effect of Temperature on the BOD Removal
Efficiency of Trickling Filters, Unit Size
Constant 159
27 Effect of Temperature on Anaerobic Digestor
Size Required To Produce a Constant Vola-
tile Solids Destruction at Various Solids
Concentration 162
28 Effect of Temperature on Gas Production in
Anaerobic Digesters of Constant Size 163
29 Effect of Temperature on Volatile Solids
Destruction Efficiency in Constant Size
Anaerobic Digesters 163
30 Effect of Temperature on Volatile Solids
Destruction Efficiency in Constant Size
Aerobic Digesters 166
31 Effect of Temperature on the Time Required
To Produce a Constant Percent Kill with a
Constant Chlorine Residual 168
32 Effect of Temperature on the Chlorine
Required To Produce a Constant Percent Kill .. 168
33 Effect of Temperature on the Adsorption
Efficiency of Activated Carbon with a Con-
stant ABS Application 172
34 Effect of Temperature on Adsorption Capa-
city of Activated Carbon with a Constant
ABS Application 182
35 & 36 Effect of Temperature on the Rate of Nitri-
fication in the Activated Sludge Process 175
37 Effect of Temperature on Nitrification
Loading at Optimum Rate 176
38 Effect of Temperature on Denitrification
Rate in the Activated Sludge Process 177
110
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FIGURES
(continued)
Number , Page
39 Effect of Temperature on Denitrification
Loading Rate 179
40 Effect of Temperature on Denitrification
Efficiency for Constant Size Denitrifying
Columns 180
41 Effect of Temperature on Ammonia Stripping
Efficiency for Constant Size Unit 181
42 Effect of Temperature on Ammonia Stripping
Unit Size for Constant Efficiencies 182
111
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SUMMARY OF FINDINGS
The effects of heat enrichment on the efficiency of common
waste treatment unit operations were investigated based
upon a review of generally available literature. The table
below presents a summary of the units investigated and the
change in size and in efficiency for these operations when
operated at elevated temperatures. The basis of this table
is an increase of temperature of from 20°C to 30°C.
SUMMARY OF UNITS SIZE AND EFFICIENCY
CHANGES DUE TO 10°C TEMPERATURE INCREASE
(20°C to 30°C)
Unit Operation
Change Due to Temperature Increase
% Change in % Change in Unit
Unit Size Efficiency
Grit Chamber
Primary Clarifier
Aeration Basin
Trickling Filter
Stabilization Pond
Aerated Basin
Final Clarifier
Chlorine Contact Tank
Thickener
Anaerobic Digestor
Vacuum Filter
Centrifuge
Filtration (Strat.)
Rapid Sand
Backwash Rate
Activated Carbon
Foam Separation
Nitrification
Denitrification (A.S.)
Ammonia Stripping
Anaerobic Column
(Nitrogen Removal)
Pure 02 Activated Sludge
Coagulation
16
20
10
68
48
52
20
28
20
38
14
20
19
-8
27
92
50
5
5
13
13
12
30
15
8
8.2
9
25
20
60
-16
29
-5
65
47
14
28
114
50
Note: Minus sign indicates an increase in unit size or a
deoTease in efficiency. No sign indicates a
decrease in unit size or an increase in efficiency,
112
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SECTION A-I
INTRODUCTION
The effect of heat addition on the unit operations commonly
found in a waste treatment facility has been investigated
and is reported in this appendix. These investigations are
based on a review of generally available literature. The
unit operations evaluated include the following major
operations:
1. Sedimentation
2. Solids thickening
3. Anaerobic digestion of solids
4. Anaerobic reduction of carbon and nitrogen
5. Aerobic bio-oxidation of carbon
6. Aerobic bio-oxidation of ammonia
7. Anaerobic bio-reduction of nitrate
8. Solids dewatering
9. Thermal processing of solids
10. Solids drying
11. Effluent disinfection
For discussion purposes, the unit operations are divided
into the major categories of:
Physical-chemical processes
Sludge handling processes
Biological processes
Disinfection
Advanced treatment processes
It is recognized that assignment of some processes in a
specific category is somewhat arbitrary.
Graphical presentations showing possible process loadings
and/or efficiency variations as a function of temperature
were developed where possible. The sources of the correla-
tions are presented in this appendix. These illustrations
do not firmly establish process design criteria but, rather,
qualitatively describe the effect of temperature on process
performance. For most of the unit operations evaluated,
further laboratory-scale or pilot-scale investigations
would be required to firmly establish design relationships.
The sewage temperature at the inlet to the sewage treatment
plant is influenced by many factors such as ambient
113
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temperature, length of interceptors, and source of the
wastewater. Though variations about a mean value will oc-
cur throughout the annual cycle, a temperature of 20°C is
selected as a basis for comparing the temperature effects,
Most of the illustrations show values relative to 20°C.
114
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SECTION A-II
PHYSICAL-CHEMICAL PROCESSES
Increasing the temperature has a significant effect upon
the physical-chemical processes that are commonly employed
in waste treatment. The physical processes - grit removal,
clarification, thickening, sludge dewatering, and flotation
- involve a separation of solid material from water. Since
this physical separation process is mainly a function of
the fluid viscosity, decreasing the viscosity of a fluid by
increasing the temperature can increase the efficiency of a
separation process.
Chemical reaction rates are enhanced by elevated tempera-
tures. Chemicals are used in wastewater treatment plants
primarily to aid in the liquid-solid separation processes.
However, certain processes (disinfection, phosphorus re-
moval, and biological denitrif ication) employ chemicals for
a specific function other than as a separation aid. These
processes will be discussed in later sections of this ap-
pendix. This section will discuss the physical-chemical
processes that primarily effect a liquid-solids separation.
Grit Removal
Grit removal in large wastewater treatment plants is gener-
ally accomplished in grit chambers which are designed to
separate inert solids from the flowing medium. These inert
solids are removed to alleviate excessive wear on the mech-
anical equipment (pumps) and to minimize possible interfer-
ence with wastewater treatment processes following the grit
chambers .
When a discrete particle settles in a quiescent medium, the
particle will accelerate until the frictional resistance or
drag force equals the gravitational force. Thereafter, the
particle will settle at a constant speed. For spherical
particles, the terminal settling velocity is as follows
[5, 6]:
vs = Iv
where: Vs = Terminal settling velocity
g = Gravitational constant
115
-------
CD = Drag coefficient
Ss = Specific gravity
d = Diameter of sphere
The drag coefficient (CD) is a function of the Reynolds
number. For a Reynolds number (R)<10\ the value of CD is
as follows:
CB - TT + A + °'34 <*2)
For a Reynolds number less than 0.5, the drag coefficient
can be described as:
CD = if (#3)
Equation (#1) similarly reduces to the following when the
Reynolds number is less than 0.5:
where: v = Coefficient of kinematic viscosity
This relationship (#4), known as Stoke's law, describes the
terminal settling velocity derived for spherical particles
that settle discretely. Discrete sedimentation assumes
that the particles settle without colliding or interacting
with any other particles. From Equation (#4) , the influ-
ence of viscosity and hence temperature is readily
apparent.
Of the three unit operations which separate solids from the
waste, namely, grit removal, primary clarification, and
secondary clarification, the settling phenomenon in a grit
chamber most closely agrees with discrete particle settling
described by Equation (#1) or (#4). Grit chambers are gen-
erally designed to remove inert particles that have a size
greater than 2 x 10"2 cm and a specific gravity of approxi-
mately 2.6.
For a continuous flow tank with turbulence, the works of
Hazen and/or Dobbins can be used to compute solids removal
efficiencies. Hazen's real tank theory [6] states that the
116
-------
efficiency of a settling basin is a function of the parti-
cle settling velocity, the surface area of the basin, rate
of flow, and the hydraulic characteristics of the basin.
The Hazen relationship is:
Removal of Suspended Solids = 1 - |~l + 2^§.l~1/n. . . (#5)
where: n = Coefficient that identifies basin performance
n = 1/3 for good performance
n = 1/2 for poor performance
Vs = Terminal settling velocity
Q = Hydraulic flow rate
A = Surface area
The theoretical increase in removal efficiency and decrease
in grit chamber length with rising temperatures have been
computed, using Hazen 's model (#5), as shown on Figures 1
and 2. Figure 1 shows the relative suspended solids re-
maining at various temperatures. A base removal of 75% at
20°C in a good performance chamber is used. At 30°C, ap-
proximately 20% of the suspended solids remaining at 20°C
would be removed. This corresponds to an overall suspended
solids removal of approximately 80% in a grit chamber sized
for 75% removal at 20°C.
Figure 2 illustrates the relative change in grit chamber
size with variation in temperature and constant removal ef-
ficiency. The width of the grit chamber is kept constant
and a chamber with good performance characteristics is
used. At 30°C, grit chamber length can be reduced by ap-
proximately 20% over an equivalent chamber designed at
20°C.
The composition of settleable and non-settleable solids in
the wastewater is an important factor in analyzing the per-
formance of grit chambers. Some experimental data show
that the actual removal efficiencies in grit chambers are
lower than those predicted from Hazen 's real tank theory
[6]. The actual performance of a grit chamber and other
physical separation units is therefore very much a function
of raw waste characteristics.
117
-------
GRIT CHAMBER
CURVE BASED ON CONSTANT UNIT SIZE
AND HAZEiyfe GOOD PERFORMANCE BASIN
FOR 75% REMOVAL OF KT2 cm DIA.
2.65 DENSITY PARTICLES AT 20°C
% REMOVAL
NOTE:
CURVE BASED ON CONSTANT
WIDTH CHAMBER, AND HAZEN'S
GOOD PERFORMANCE BASIN
FOR 75% REMOVAL OF
2.65 DENSITY
30 40
TEMPERATURE, °C
FIGURE i EFFECT OF TEMPERATURE ON THE EFFICIENCY OF
GRIT REMOVAL, UNIT SIZE HELD CONSTANT,
FIGURE 2 EFFECT OF TEMPERATURE ON GRIT CHAMBER SIZE TO
ACHIEVE A CONSTANT 75% GRIT REMOVAL EFFICIENCY,
118
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Sedimentation
Generally, the gross settleable solids present in waste-
water are removed by gravity settling in primary sedimenta-
tion basins. This process also provides a reduction of the
waste load to subsequent treatment units. The settling
phenomenon in primary basins can be described as a mixture
of discrete and flocculent particle sedimentation. The
particles removed are of smaller size and have a lower spe-
cific gravity than the grit particles removed in the grit
chambers.
As an illustration of the theoretical effect of temperature
on the settling of discrete particles in a sedimentation
basin, Figures 3 and 4 are presented. These curves apply
Hazen's real tank model in rectangular basins operating at
an overflow rate of 800 gpd/SF. A suspended solids removal
efficiency of 50% at 20°C is utilized for these "good per-
formance" basins.
Increase in wastewater temperature from 20°C to 30°C will
result in an increase in suspended solids removal from 50%
to 56% in a constant size sedimentation basin. To achieve
an equivalent suspended solids removal of 50% at the ele-
vated temperature, a basin approximately 20% smaller than
would be required at 20°C is needed. These changes in tank
size and efficiency are based on theoretical considerations
of discrete particle settling.
For flocculent particles, the settling velocity in a sedi-
mentation basin is variable. During the settling process,
the particles coalesce in the basin, thereby affecting the
size and density of the aggregate. The net effect is an
increase in velocity as the particles collide. Since the
settling velocity is not constant and depends upon the
flocculent nature of the solids, the removal efficiency de-
pends not only on the surface area and flow rate, but also
on the detention time. As the temperature increases, the
settling velocity of the flocculated solids increases. Ad-
ditionally, temperature is believed to have a beneficial
effect on the flocculating characteristics of the solids
[8]. Pilot-scale testing of flocculent settling at various
temperatures would provide information on removal efficien-
cies to be expected in secondary clarifiers.
Many factors reduce the efficiency of a prototype sedimen-
tation basin. Various currents such as wind-induced sur-
face currents, convection currents, density currents,
119
-------
SEDIMENTATION BASIN
HAZENS REAL TANK MODEL - GOOD PERFORMANCE BASIN
CONSTANT UNIT SIZE
50% SUSPENDED SOLIDS
REMOVAL® 20°C
CONSTANT SUSPENDED SOLIDS
REMOVAL EFFICIENCY OF 50%
30 40
TEMPERATURE, °C
50
60
FIGURE "5 EFFECT OF TEMPERATURE ON THE SUSPENDED SOLIDS
REMOVAL IN SEDIMENTATION BASINS/ UNIT SIZE HELD CONSTANT,
FIGURE 4 EFFECT OF TEMPERATURE ON SEDIMENTATION BASIN
SURFACE AREA TO ACHIEVE A CONSTANT 50% SUSPENDED SOLIDS
REMOVAL,
120
-------
and velocity distribution tend to contribute to short-
circuiting of flow through the basin. Thermal gradients
may exist within an uncovered basin. These gradients can
have a deleterious effect on the basin flow pattern and re-
movals. The effects of thermally induced gradients have to
be evaluated on at least a pilot scale. For the purposes
of analysis, the reductions in size presented on Figure 4
will be utilized for both primary and secondary clarifiers.
Coagulation
A wide variety of organic and inorganic solids in waste-
water will not be removed by sedimentation unless agglom-
erated into larger particles. These particles are
stabilized or kept separate by electrical and physical
forces. Since natural destabilization forces are generally
not sufficient to allow efficient solids removal, chemical
coagulants are employed to stabilize and agglomerate the
solids. For each combination of coagulant and wastewater,
there is an optimum dosage of coagulant and an optimum pH
range for coagulation.
Little experimental information is available on the effect
of temperature on coagulation [6] . Velz [10] cites detri-
mental effects of higher temperatures on coagulation and
removal of color. Higher alum dosages were required at
elevated temperatures to achieve an equivalent effluent
color concentration. He also related temperature, dosage,
and time of appearance of first floe, all illustrating the
disadvantages of high-temperature coagulation. However,
these studies were performed at constantly changing pH
values which were considered of minor importance at that
time.
Camp [11] found that coagulant dosage and temperature could
change optimum pH values markedly. With pH adjustment,
Camp found shorter settling times at higher temperatures.
Renn [8] also found shorter floe formation times at higher
temperatures and that the floe was generally found to be
coarser in nature. Willcomb [12] explains that, as a re-
sult of the increased viscosity of the water and the sur-
face tension change in the floe at low temperatures,
coalescing tendencies are resisted. Prolonged agitation
would then be necessary to enlarge floe sizes to settleable
proportions. Parsons [13] presented ranges of removal
efficiency for coagulation of domestic sewage. These effi-
ciencies range from 65% to 85% for removal of suspended
solids and from 45% to 75% for removal of BOD.
121
-------
Coagulation of effluent from secondary biological treatment
facilities has also been evaluated. Stukenberg [14]
achieved better BOD and COD removals at warmer operational
temperatures.
The optimum conditions for flocculation are determined by
three variables: the chemical dose, the pH, and the tem-
perature. Renn [8] determined that the isoelectric point,
or optimum pH value for coagulation, varies with tempera-
ture. It has been found that, at the optimum pH value, the
required coagulant dosage decreases as the temperature in-
creases. Additionally, the time of floe formation decreases
as temperature increases.
As an illustrative example of the data reported on the ef-
fect of heat on coagulation, Figures 5 and 6 are presented.
The data presented are taken from studies performed on mu-
nicipal water supplies. Temperatures evaluated in these
studies range from approximately 5°C to 28°C. Figure 5 in-
dicates a rapid increase in required alum dosage to achieve
satisfactory flocculation at temperatures lower than 10°C.
Figure 6 presents the time required to form a good floe at
optimum pH and with varying chemical dosage. The figure re-
veals that chemical dosage is the most significant parameter
affecting floe formation in the temperature range studied.
At a chemical dosage of 2 ppm at 20°C, approximately 20 min-
utes were required for good floe formation. This floe for-
mation time decreased to about 25 minutes at 28°C with the
same chemical dosage. However, at a chemical dosage of 4.3
ppm, floe formation times remained constant.
;
Studies of coagulation at temperatures above 28°C were not
found. Additionally, those studies that have been reviewed
do not indicate the quantity or cost of pH adjustment. Wet
testing on specific wastewaters is necessary.
Gravity Filtration
With the exception of gravity sedimentation, deep bed fil-
tration is the most widely used unit process for liquid-
solids separation. Recently, it has been employed in
physical-chemical systems for polishing effluent prior to
discharge [9].
The temperature at which filtration takes place has a large
influence on the process. Filter bed variables affected by
temperature are:
122
-------
COAGULATION
o
o
Q
O
o
u. 5
O —
S*
| | 20
21
O
UJ
10
NOTE: INFLUENT
TURBIDITY- 300 PPM
FIGURE 5
STANLEY $RUTHRAUFF DATA8
5
UJ
u
o:
1.5 <
UJ
(E
CAMR , ET. A I. DATA
II
_L
_L
FIGURE 6
10 20
TEMPERATURE, °C
30
FIGURE 5 EFFECT OF TEMPERATURE ON THE ALUM DOSAGE RE-
QUIRED TO PRODUCE SATISFACTORY FLOCCULATION,
FIGURE 6 EFFECT OF TEMPERATURE ON THE TIME REQUIRED
FOR FLOC FORMATION,
123
-------
(a) Depth of media
(b) Rate of filtration
(c) Expansion of bed on backwash
(d) Efficiency of filtration
(e) Head losses
Head losses through the filter are directly proportional to
viscosity and/ therefore, are reduced by temperature
increase.
A relationship for head loss in a clean stratified bed sand
filter was developed by Kozeny and modified by Fair and
Hatch [6] to:
vv
g
d-f)
Pi
n
£
i=l di2
(#6)
where: y- = Head loss in ft/ft length
k = Coefficient of permeability
v = Kinematic viscosity
v = Liquid velocity through filter
f = Porosity of filter media
¥ = Spheroscity of filter media
Pi = Fraction analyzed sand
di = Average diameter of sieved sand
Based on this equation, an illustrative example of the the-
oretical effect of temperature on a £lean unstratified sand
filter is presented on Figure 7. An increase in water tem-
perature from 20°C to 30°C will result in a decrease in
head loss of approximately 20%. A temperature increase to
60°C will result in a 50% reduction in head loss through
the sand bed. The same equation (#6) can be used to com-
pute the change in filter area required for a constant head
loss and flow rate. Figure 8 presents the relative capac-
ity of a sand filter with constant head and variable tem-
perature. An increase of water temperature from 20°C to
30°C will require a sand filter of approximately 20% less
area or depth than at 20°C.
124
-------
GRAVITY FILTRATION - STRATIFIED BED
c/>
en
o
Q
<
1.3-
1.2-
LO-
ui
.6-
.5-
.4-
12-t
1.0- O
CM
O
x 4.0
X
o
z
UJ
h-
L_
1-
u.
to
3
o
UJ
X
in
-
\
-
i
CURVE B/
RATE OF
POROSITY
COEFFICIE
A CLEAN
-v^-T
'•^
\SED ON KOZ
3 MGAD. BED
OF 0.35, SPH
NT OF PERME
FILTER BED
9
^^
i
ENY EQUATIO
DEPTH OF 4
EROSITY OF C
: ABILITY OF £
^ (fi;
•
i
FIGURE 7
N FOR FLOW
2 IN.,
>, FOR
T,"^"
1
O .9-
5 -8H
UJ
> -7H
UJ
tr
1500
or
u.
CM
.5-
.4-
.3-
1000
0
2
500
FIGURE 8
FLOW BASED ON KOZENY EQUATION
AS ABOVE WITH FLOW = 50 MGD, AT
CONSTANT HEAD LOSS OF 1 FT.
X
20
30
__J
40 50
TEMPERATURE^
60
70
FIGURE 7 EFFECT OF TEMPERATURE ON HEAD LOSS THROUGH
STRATIFIED BED GRAVITY FILTERS,
FIGURE 8 EFFECT OF TEMPERATURE ON FILTER CAPACITY
TO PRODUCE A CONSTANT HEAD LOSS IN STRATIFIED BED
GRAVITY FILTERS,
125
-------
The backwash water rate required to clean a sand filter has
been found to vary with temperature [16] . A relationship,
based on Hazen's formula, for backwash rate is:
R = SOd1'5 (1 + 0.060x) (t QQ30) ..... (#7)
where: R = Backwash rate (in./min)
d = Effective sand size (mm)
x = % bed expansion (expressed as a whole number)
t = Wash water temperature (°F)
An illustrative example of the theoretical variation in
backwash rate with temperature is presented on Figure 9.
Also indicated on this figure are the limited data reported
by Lawrence [16] . The portion of the curve above 25°C is
shown as a dashed line to show that it is an extrapolation
of the basic equation. A temperature increase from 20 °C to
30°C will result in a requirement of 20% higher backwash
rate for an equal bed expansion.
The spatial distribution of solids or turbidity was experi-
mentally determined by Ives [15] to be described by the
following:
= Ac ........................... (#7a)
where: c = Concentration of suspension
1 = Depth of filter medium
A = Filter coefficient
The filter coefficient (A) is a function of the amount of
solids that are deposited in the filter. The coefficient
(A) in Equation (#7a) is obtained for a clean filter opera-
ting with an initial filter coefficient of AO. The equation
which describes the variation of the filter coefficient is
as follows [15] :
X , X + ccr -
126
-------
GRAVITY FILTRATION
Ul
I- is-
Sli7:
I 1.5-
CO I ^_
1.2-
4
UJ
z *>
z
^ 10
CO
UJ
| 60
uj 50
s
" 4(1
CO
<
»30
*
0
<
(0 20
10
^
^"
-^
^
* ^^
— RA
^^^
TE = 3<
FOR
DATA Fl
^*»
•--
^^
)^M( i + aoeox)
d = 0.40mm
X = 50
ROM LAWRENCE16
K> 20 30 40 50
FIGURE 9
****"
(t+j
8C
*»
10)
1
60 °C
UJ
g 4-
u.
u.
UJ
CC 3
2-
ui
a:
FOR f = 0.4
AFTER 1 HR AT 20 cm DEPTH
v = 2.0 GPM / FT2
. d = 0.065 cm
CURVE DEVELOPMENT
BASED ON CURVES
EXPERIMENTALLY
DEVELOPED BY IVES15
40 50 60
TEMPERATURE , °C
FIGURE 9 EFFECT OF TEMPERATURE ON REQUIRED BACKWASH RATE
FOR GRAVITY FILTERS.
FIGURE 10 EFFECT OF TEMPERATURE ON FILTER EFFICIENCY,
CONSTANT FILTER SIZE,
127
-------
where: A0 = Initial (clean) filter coefficient
a = Volume of deposit per unit filter volume
<|> = A second filter constant
f = Porosity of filter medium
The filter coefficient is a direct measure of the filter
efficiency. An illustrative example of the theoretical
effect of elevated temperature on filter efficiency is pre-
sented on Figure 10. For this figure, A0, c and were de-
termined at various temperatures using data presented by
Ives [15] . For the conditions cited on Figure 10, A was
then computed as shown. A temperature increase from 20°C
to 30°C will result in a 50% increase in the filter coef-
ficient and therefore the filter efficiency. Full-scale
evaluation of the effects of heat addition to sand filters
has to be performed for temperatures above 25°C to verify
the theoretical relationship presented in this discussion.
Flotation
Flotation is generally confined to a process in which air
is dissolved into the wastewater. The mixture is held
under pressure to ensure adequate solution of the air into
the liquid. The mixture is then released to atmospheric
pressure upon which the air in the form of small bubbles is
released. These bubbles either become enmeshed in or
attached to the suspended material.
The performance of the unit depends upon having sufficient
air bubbles to float the suspended material. Effluent
quality and/or solids concentration in the float are related
to an air/solids ratio which is defined as the pounds of air
released per pound of initial suspended solids [17]. The
rise rate or velocity of the solids is also directly influ-
enced by the air/solids ratio. Parameters important for
flotation processes include: (1) pressure, (2) solids con-
centration, (3) detention time, (4) type and quality of
waste and sludge and its volatile content, (5) solids and
hydraulic loading rates, (6) temperature, (7) recycle ratio,
(8) air to solids ratio, and (9) use of chemical aids [18].
Flotation processes include: (1) dispersed air-flotation,
(2) dissolved air-vacuum flotation, (3) dissolved air-
pressure flotation, (4) biological flotation. In
128
-------
application, the first two methods may be employed for
wastewater treatment while the last two are for sludge
thickening operations.
At higher wastewater temperatures, the decrease in liquid
viscosity will lessen the resistance to liquid-solid sep-
aration. However, the solubility of air in water is in-
versely proportional to the temperature and, to maintain
comparable air/solids ratios at higher temperatures, the
pressure will have to be correspondingly increased. The
actual effect of elevated temperature upon sizing a speci-
fic unit would have to be determined in laboratory and
pilot studies.
Sludge thickening by heat flotation has been investigated
by Malina [20] and Laboon [21]. In these processes, as
the sludges are heated the solid material tends to rise as
a mat, resulting from a lifting effect caused by the release
of absorbed gases in the form of small bubbles. The appli-
cation of these processes depends upon the amount of en-
trained gases in the sludge, the temperature to which the
sludge is raised, the point of heat addition, and the quan-
tity of sludge to be thickened.
Because many primary variables exist for flotation processes
as indicated, and because some of these variables are inter-
related or affected by prior processes in a treatment sys-
tem, when a flotation unit is selected for operation at high
temperature, pilot plant studies are suggested for effective
design of the unit.
129
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SECTION A-III
SLUDGE HANDLING PROCESSES
Sludge handling processes have been investigated and re-
ported in great depth by Burd [18]. Some 450 references
are cited in his work. For the current study, we have lim-
ited our investigation to the major unit processes. The
reader is directed to the Burd report for discussion of
specific unit processes not present in this report. The
material which follows has been arranged in two categories:
(1) thickening and (2) dewatering. Generally, the condi-
tioning and dewatering unit operations respond to tempera-
ture increases in ways which incorporate many of the
previously developed temperature relationships. Where this
is the case, general reference will be made to the previous
development.
Thickening
Primary sludge and secondary sludge (waste-activated sludge)
require further concentration prior to digestion or dewater-
ing. A gravity-type or flotation-type thickener is gener-
ally employed for this application. The solids settle or
rise at different velocities depending upon the solids con-
centration and temperature. Previous discussions have
dealt with Stoke's law and the change in discrete particle
settling velocity with increases in temperature (see Grit
Removal).
A relationship has been developed to describe the hindered
settling velocity of flocculent sludges as a function of
sludge characteristics, concentration, and the discrete
particle settling velocity [32]. The relationship is de-
scribed by the following equation:
Vi = Vo (1 " C)n (#9)
where: V. = Critical settling velocity of the sludge at
concentration C
VQ = Discrete settling velocity of a particle
n = Empirical value depending upon the sludge
characteristic
130
-------
C = Concentration of solids at velocity V
Solids are transported to the bottom of a gravity thickener
by two mechanisms : their subsidence due to gravity and the
bulk downward transport due to sludge withdrawal from the
bottom of the thickener. The rate at which solids of con-
centration Cj[ pass downward in the thickener can be de-
scribed by the following [32] :
G = CiVi + C±U ....................... (#10)
where: G = Solids flux, expressed in Ibs/day-SF
V. = Settling velocity of the sludge at concentra-
tion CL
U = Average downward velocity caused by removal of
the sludge from the bottom of the tank
The term C^Vi depends upon the settling characteristics of
the sludge, while GJ.U is a variable controlled by the opera-
tion. Characteristically, the solids flux achieves a mini-
mum value which provides the basis of a design to ensure
sufficient thickener surface area to meet the area require-
ments of this minimum solids flux.
A more familiar form of this relationship is presented by
Eckenf elder [25] , based on a material balance between the
influent, the underflow, and the effluent, as:
(1/C. - 1/C )
O.A. = * - ±-
where: U.A. = Unit area (SF/lb solids/day)
C. = Solids concentration at settling velocity
1 Ib/CF
C = Underflow solids concentration, Ib/CF
u
The unit area is inversely proportional to the particle
settling velocity, which is affected by viscosity (and
therefore temperature) changes.
An illustrative example of the theoretical effect of temper-
ature increase on thickener requirements is presented on
131
-------
Figures 11 and 12. Figure 11 presents the relative mass
loading (Ibs/SF/day) as a function of temperature. The
mass loading is the reciprocal of unit area as described in
Equation (#11). Figure 11 describes the relative increase
in allowable solids application rate to a thickener of
known size operating at a constant underflow concentration.
A temperature increase from 20°C to 30°C would allow an in-
crease of 27% in the mass loading on a thickener for the
same underflow solids concentration. Figure 12 presents
the relative unit area required as a function of tempera-
ture. A temperature increase from 20°C to 30°C would per-
mit a 21% reduction in the size required at 20°C to achieve
the same underflow solids concentration.
The relationships presented on Figures 11 and 12 assume
that the sludge settling characteristics [21] and sludge
blanket behavior (gasification) would not change signifi-
cantly by increases in temperature. Experimental evalua-
tions are required to determine what these effects might
be.
Data on the compaction characteristics of pure oxygen acti-
vated sludge have been presented by Stamberg [125]. These
data are presented on Figures 13 and 14. Figure 13 pre-
sents the change in initial batch flux (Ib/SF/day) with
initial mixed liquor concentration and temperature. At
4,000 mg/1 MLSS, an increase in temperature from 10°C to
29°C resulted in a 64% increase in the flux rate (55 lb/SF/
day to 90 Ib/SF/day). Figure 12 presents the change in
initial settling velocity (ft/hr) with initial mixed liquor
concentration and temperature. At 6,000 mg/1 MLSS, an in-
crease in temperature from 23°C to 27°C resulted in a 71%
increase in the initial settling velocity (7 ft/hr to
12 ft/hr).
Dewatering
The dewatering of waste sludges is a cumbersome and costly
part of the wastewater treatment plant operation. The main
objectives of these processes are to reduce the sludge vol-
ume for ease and economy of disposal. Commonly used sludge
dewatering processes include (1) vacuum filtration,
(2) pressure filtration, and (3) centrifugation. Discussion
of the effect of elevated temperatures on these processes
follows.
132
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GRAVITY THICKENING
CONSTANT UNDERFLOW
CONCENTRATION
TEMPERATURE,°C
FIGURE II EFFECT OF TEMPERATURE ON MASS LOADING RATE
TO THICKENER TO PRODUCE A CONSTANT UNDERFLOW CONCENTRATION.
FIGURE 12 EFFECT OF TEMPERATURE ON THICKENER SIZE TO
PRODUCE A CONSTANT UNDERFLOW CONCENTRATION,
133
-------
PURE OXYGEN ACTIVATED SLUDGE
COMPACTION CHARACTERISTICS
too
DATA REPORTED BY STAMBERG 125
Z 34 6 8 10 15 20 30 40
INITIAL MIXED LIQUOR CONCENTRATION (gm/l)
FIGURE 13 & 14 EFFECT OF TEMPERATURE ON THE COMPACTION
CHARACTERISTICS OF SLUDGE PRODUCED BY THE PURE OXYGEN
ACTIVATED SLUDGE PROCESS.
134
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Vaouum Filtration
Vacuum filtration is commonly carried out on slowly rotating
drum filters in a continuous operation. Variables affecting
the performance of vacuum filters include: solids concen-
tration, sludge type, temperature, sludge and filtrate
(liquid) viscosity, and chemical composition.
*
The rate of filtration of sludges has been formulated ac-
cording to Poiseville's and D'Arey's laws by Carmen and
Coakley, as follows [25]:
dv ,„,-.
= (#12)
' U (rcV + RmA)
where: V = Volume of filtrate
t = Cycle time (approximates form time in continu-
ous drum filters)
P = Vacuum
A = Filtration area
y = Filtrate viscosity
r = Specific resistance
c = Weight of solids per unit volume of filtrate
Rm is the initial resistance of the filter medium and can
usually be neglected as compared with the resistance devel-
oped by the filter cake. This equation can be modified to
express filter loading rate (neglecting the initial resist-
ance of the filter medium) as follows:
1/2
where: R = r x 107 sec2/gm
P = Vacuum, psi
C = Solids deposited per unit volume filtrate, gm/ml
135
-------
y = Filtrate viscosity, centipoises
t = Form time, min
Since the filter loading rate is inversely proportional to
the square root of the viscosity, the loading rate increases
with increasing temperature. The specific resistance will
similarly decrease at elevated temperatures. However, this
effect has not been quantified since the reduction in spe-
cific resistance depends upon the sludge characteristics.
The theoretical effect of elevated temperatures can be esti-
mated using Equation (#13). Figure 15 presents the theoret-
ical filter yield as a function of temperature. Figure 16
presents the relative filter area required for a constant
loading rate as a function of temperature. For a 10°C tem-
perature rise from 20°C to 30°C, the filter area required
drops 10%.
Pressure Filtration
Like vacuum filtration, a porous medium is used in leaf fil-
ters to separate solids from liquids [18]. (Leaf filters
are the most commonly used pressure filter.) As sludges are
forced onto the medium under pressure, the solids are cap-
tured in the medium pores and build up on the medium surface.
In general, when the liquid phase is highly viscous, or when
the solids are so fine that vacuum filtration is too slow,
pressure filtration provides a convenient solution to the
separation problem.
The temperature effect on pressure filtration should gener-
ally exhibit the same relationship presented for vacuum
filtration.
Centrifugation
Applications of centrifuging are washing, dewatering, clas-
sification, clarification, or more usually a combination of
these.
The most effective dewatering centrifuges are horizontal,
cylindrical, conical, solid-bowl machines.
Centrifuges separate solids from liquids, through sedimen-
tation and centrifugal force. Typically, sludge is fed
through the center of the unit through a screw conveyor,
136
-------
VACUUM FILTRATION
1.6
- 1.4
1.2
UJ
>
<
_i
UJ
1.0
.8
.6
FILTER RESISTANCE ASSUMED
NEGLIBLE, C, P $ t ARE CONSTANT
1.4
u "-2
cr
a: 1.0
UJ
b
il .8
UJ
>
UJ
IT
FIGURE 16
FOR CONSTANT
LOADING RATE
10
20 30
TEMPERATURE , °C
40
FIGURE 15 EFFECT OF TEMPERATURE ON VACUUM FILTER YIELD,
CONSTANT SOLIDS LOADING.
FIGURE 16 EFFECT OF TEMPERATURE ON THE REQUIRED VACUUM
FILTER SIZE TO PRODUCE A CONSTANT FILTER YIELD AT A
CONSTANT SOLIDS LOADING RATE,
137
-------
mounted inside a rotating conical bowl. Sludge leaving the
feed tube is accelerated and is distributed to the periphery
of the bowl where it is settled and compacted by centrifugal
force. It is then conveyed by the screw to an inclined
"beach" area where it is further dewatered and discharged.
Separated liquid is discharged continuously over adjustable
weirs at the opposite end of the bowl.
The factors that determine the success or failure of cen-
trifugation are (1) cake dryness and (2) solids recovery.
Guidi [34] summarized the effect of the various parameters
on these two factors as follows:
Process
Variable
To improve
recovery
To improve
cake solids
Feed
Rate
decrease
increase
Feed
Consistency
increase
decrease
Temperature
increase
increase
Flocculents
increase
decrease
Thus, we see that increasing temperature increases both cake
dryness and solids recovery. The paramount variable though
is the design of the unit itself [33].
As the temperature of the liquid carrier medium increases,
the viscosity and density decrease, thus increasing the set-
tling rate of the solids. The reduction in the moisture
viscosity aids in dewatering of the solids, producing a
drier cake.
The throughput capacity, Q, of a settling centrifuge at the
"cutoff point" (50% of feed particles removed and 50%
passed) can be described by:
(#14)
y
with:
and:
ID ' ~
where: g = Acceleration of gravity, 981 cm/sec2
c = Volume of liquid in bowl, cm3
Q = 2 Vge
Vg =
-------
s = Effective thickness of liquid layer in which
settling is occurring, cm
p = Density of particle
Pi = Density of liquid
d = Size of particle
y = Absolute viscosity of liquid medium
r = Radius of curvature of path
u = Angular velocity
Thus it can be seen that the throughput capacity for the
cutoff point is inversely proportional to the viscosity and
therefore directly proportional to the temperature of the
liquid. An illustrative example of the theoretical effect
of heat addition on centrifuge capacity is presented on
Figures 17 and 18. In Figure 17, the machine and sludge
operating variables are kept constant to allow determina-
tion of the theoretical effect of temperature. An increase
in temperature from 20 °C to 30°C is computed to approxi-
mately a 25% increase in the throughput capacity of a cen-
trifuge. In Figure 18, the throughput rate is kept constant
while the radius of curvature (r) is allowed to vary with
temperature. An increase in temperature from 20°C to 30°C
will result in a 20% reduction in the unit's radius of cur-
vature. This would result in the use of a smaller diameter
centrifuge for a given application.
Temperature increases may have the effect of reducing the
strength of the solids and make them more difficult to con-
vey. Ease of conveyance of a solid necessitates tests at
different temperatures.
The relationships presented are necessarily theoretical
since no large-scale systematic evaluation of heat effects
on centrifugation has been performed.
139
-------
CENTRIFUGATION
260
FOR CONSTANT UNIT SIZE
AND OPERATION FOR 50%
SOLIDS RETAINED
FOR CONSTANT
THROUGH PUT CAPACITY
40 50
TEMPERATURE, °C
FIGURE 17 EFFECT OF TEMPERATURE ON THE CAPACITY OF A
CONSTANT SIZE CENTRIFUGE,
FIGURE 18 EFFECT OF TEMPERATURE ON CENTRIFUGE SIZE TO
ACHIEVE A CONSTANT THRUPUT CAPACITY,
140
-------
SECTION A-IV
BIOLOGICAL PROCESSES
Biological processes are used in treatment plants to remove
organic material from the wastewater and to reduce the
quantity of biological sludge produced in the removal
process.
Biological Waste Treatment
Biological waste treatment is used to reduce the organic
materials present in a waste. It is a process wherein ac-
tive bacteria are admixed with a waste. Under suitable
environmental conditions, the bacteria reduce the waste to
a more stable form. When the reaction proceeds in the
presence of sufficient dissolved oxygen, the system is
aerobic and the final decomposition products are carbon
dioxide and water. Two basic phenomena occur when organic
matter is removed by microorganisms: oxygen is consumed
by the organisms for energy and new cell mass is synthesized.
The organisms also undergo progressive auto-oxidation of
their cellular mass. These reactions can be illustrated by
the following general equations:
organic matter + O2 + NH3 + cells •* new cells + C02 + H20
and
cells + 02 -»• C02 + H20 + NH3
In the design of wastewater treatment facilities, the rate
at which these reactions occur, the amount of oxygen and
nutrient required, and the quantity of biological sludge
produced in the reaction must be determined.
Wastes can contain suspended, colloidal, and dissolved or-
ganics. The organic matter is measured by the biochemical
oxygen demand (BOD) or by the chemical oxygen demand (COD).
The BOD may be defined as the amount of oxygen required by
suitable organisms in the stabilization of a given quantity
of organic matter. Theoretically, an infinite time "re-
quired for complete biological oxidation of organic matter,
but for practical purposes, the reaction may be considered
141
-------
complete in twenty days . The conventional BOD test is a
measure of the quantity of oxygen utilized in the first
five days of oxidation, under standard conditions, and is
designated as BOD 5. The quantity of oxygen required to
satisfy the twenty day demand is usually referred to as
ultimate BODU. The COD is a measure of the ultimate BOD.
However, in the COD determinations, organic matter is con-
verted to carbon dioxide and water regardless of the bio-
logical assimilability of the substances. In the analysis
of data, it must be remembered that some materials which
are chemically oxidized will not be biologically oxidized.
Biological waste treatment, then, essentially consists of
controlling environmental factors to enable a mixed culture
of microorganisms to utilize the organic matter in the
waste as a food source for reproduction (synthesis) and
energy (assimilation) . In aerobic treatment systems, orga-
nisms are generally suspended in a liquid medium with the
waste to be treated. In trickling filters, organisms are
fixed to a solid medium and the waste is trickled over the
medium. Dissolved oxygen is required by the culture and
sufficient time is allowed for the organisms to utilize the
organics as a food source.
The suspended and colloidal organic matter measured as BOD
undergoes an initial reduction by adsorption to the organ-
isms . Thereafter BOD removal is assumed to be reduced in
accordance with kinetics of the first order. The removal
reaction is usually expressed:
- f| = KL ............................ (#17)
which can be written in the form:
(#18)
-rr = Rate of change of BOD with respect to time
L = BOD remaining
Sa = Quantity of microorganisms present
K2 = BOD removal rate
142
-------
The BOD removal can be expressed for more complex reactions
that are retardant in nature as:
• • + KnL .......... (#19)
The reaction constant in the BOD removal equations is tem-
perature dependent. It is possible to relate the effect of
temperature on BOD removal by the following relationship:
Kt = K20e- ...................... (#20)
where: Kt = BOD removal coefficient at temperature T (°C)
K2o= B0D removal coefficient at 20°C
T = Temperature in treatment system (°C)
6 = Temperature coefficient (1.020 - 1.080)
Increases in system temperature generally increase substrate
removal rates in biological treatment processes. The opti-
mum temperature for biological reactions depends upon the
type of process that is considered. For aerobic mesophilic
systems (e.g., trickling filters, activated sludge and aer-
obic lagoons) , the optimum temperature has a range between
30°C and 35°C. For aerobic thermophilic reactions (e.g.,
composting) , the optimum temperature is approximately 52°C.
For anaerobic mesophilic digestion, the optimum temperature
is 37 °C, while thermophilic digesters are operated at tem-
peratures of 52 °C.
In general, the system will approach some minimum BOD value
rather than~ zero concentration due to an equilibrium between
the bacteria and their liquor. The magnitude of the initial
removal is a function primarily of sludge concentration, ac-
climatization, and waste composition. The rate of reaction
is a function of temperature, nutrient level, concentration
of waste, and sludge composition.
The growth of biological solids may be considered in the fol-
lowing steps: first, a lag period in which the culture
adapts from its previous environment to the present; second,
a period of maximum growth under conditions where unlimited
food is available; third, a period of declining growth where
food availability finally becomes a limiting condition and
143
-------
the sludge consumes previously stored food; and finally, an
endogenous phase where, under severely limited food condi-
tions, cells die and are, in turn, consumed so that mass
population is reduced. The final sludge mass is always
more than the initial, since certain non-disposable mate-
rials are generated during synthesis. The oxygen utiliza-
tion rate per unit weight of sludge is low at first, but
quickly reaching a maximum. As the competition for food
becomes more acute, the rate decreases until an endogenous
level of demand is reached. A knowledge of the parameters
governing these reactions is necessary in the design of a
biological system to treat any organic waste.
A general substrate removal equation can be derived from a
materials balance around a completely mixed suspended growth
system.
INPUT - OUTPUT - REACTION = CHANGE
W(t) - QC - Vf (Sa,C) = V .......... (#21)
where: W(t) = Influent quantity (QCQ)
C = Effluent concentration
Q = Flow
V = Volume of reactor
f(SaiC) = Functional form describing biological
reaction
Sa = Active biological population measured as
mixed liquor suspended solids
C0 = Initial concentration
The reaction term can be further defined and the equation
rearranged:
W(t) „ rn _ dc C /JIO-M
TF ~ K»oat, — -Tj7 """ r— (wH)
where: K2 = Biological reaction rate
n = Order of biological reaction defining
dependence on substrate concentration
144
-------
t0 = Detention time, V/Q
This general equation can now be used to develop the steady-
state equation which assumes no change with respect to time
and would be representative of a biological system operating
at equilibrium under a constant organic load. The general
steady-state equation reduces to:
(C — C)
—V K2SaCn =0 (#23)
In each of the foregoing equations, the term n, designating
the order of the biological reaction, has been included.
Kinetics describing substrate removal can generally be de-
fined as:
n=l First order kinetics generally applied to BOD
and COD removal. The substrate removal is
directly proportional to the substrate
concentration.
n-0 Zero order kinetics generally applied to the
removal of specific compounds such as linear
alkyl sulfonate and phenols. The substrate
removal proceeds at a fixed rate independent
of the concentration.
n*-l Retardant kinetics generally applied to bac-
teriostatic or inhibitory compounds such as
formaldehyde. The substrate removal is in-
versely proportional to the concentration;
removal decreases as concentration increases.
A short discussion of the specific biological processes re-
viewed, the pertinent literature findings, and the tempera-
ture model developed follows.
Stabilization Ponds
Waste stabilization ponds have undergone sufficient study
and development to be classified as one of the major types
of wastewater treatment systems. The design of a waste sta-
bilization pond depends upon the treatment objective. A
pond may be designed to receive untreated wastewaters, pri-
mary treatment plant effluents, secondary biological treat-
ment plant effluents, or excess activated sludge [40].
145
-------
Stabilization ponds rely on natural reaeration to apply
oxygen to the biological populations present. Large level
areas are required as well as temperate climates. The cost
of stabilization ponds can be up to 50% less than equiva-
lent activated sludge biological treatment [41].
Stabilization ponds can be generally divided into three
classifications: aerobic, anaerobic, and facultative ponds.
The classification depends on the organic loading and there-
fore dissolved oxygen content. Factors which affect a sta-
bilization pond's efficiency include detention time, depth,
organic loading, temperature, visible light energy, and the
efficiency of conversion of light energy into chemical
energy.
Fair et al.[6] suggest that the effluent BOD from a single
stabilization pond can be described by a first-order equa-
tion as follows:
Y =
(#24)
where: Y0 = Influent BOD5
Y = Effluent BOD5
K0 = BOD removal rate constant, I/day
tfl = Detention time, days
Gloyna [40, 41] indicates that for single pond the ratios
of the reaction rates are equal to the ratios of the deten-
tion times and are a function of temperature. This rela-
tionship is:
"35
ICm
'35
(#25)
where: k = Reaction rate constants for various tempera-
tures , I/day
t = Reaction times
T = Temperature
146
-------
Laboratory data obtained by Gloyna [41] at 9°, 20°, 24°,
and 35°C showed that 8 ranged between 1.072 and 1.085 and
K35 =1.2 for a synthetic non-settleable sewage.
As pond temperature increases, the equation shows that the
detention time (pond volume) requirements will decrease
accordingly until the temperature reaches 35°C. The rela-
tionship between pond capacity and temperature is valid only
for temperatures ranging between 3°C and 35°C. The lower
limit is due to retardation of bacterial and algal activity
as the temperature approaches the freezing point, while the
upper limit is imposed by thermal inactivation of most types
of algae.
A pond can function very well when the entire contents are
not oxygenated photosynthetically. The biological degrada-
tion rate in ponds is temperature dependent. Practical de-
sign criteria necessitate careful selection of reaction
rates and minimum temperatures, as well as the common con-
siderations of light intensities, food, etc. For many do-
mestic wastes, the following empirical relationship is
suggested by Gloyna [41]:
V = CQLa [6(35-T)] f-f1 (#26)
where: V = Pond volume (ac-ft)
Q = Influent flow (gpd)
La = Ultimate influent BOD (mg/1)
9 = Temperature coefficient (1.072 - 1.085)
T = Average temperature of the coldest month (°C)
C = 10.7 x 10~8 (used where temperature fluctua-
tions are large and designs are based on a
depth of 5 ft and one extra foot for solids
storage)
f = Algal toxicity factor = 1 for domestic wastes
f1 = Sulfide correction = 1 for SO^ concentrations
of less than 500 mg/1
An illustrative example showing the theoretical effect of
temperature on BOD removal efficiency and required pond
volume is presented on Figures 19 and 20. Figure 19 shows
the decrease in BOD removal efficiency expected in a sta-
bilization pond designed to achieve 85% removal in 7.3 days
147
-------
STABILIZATION POND
100
90 -
NOTE: BASED ON
td =7.3 DAY,
REMOVAL AT
0 = 1.072
20 30
TEMPERATURE, °C
FIGURE 19 EFFECT OF TEMPERATURE ON THE BOD REMOVAL
EFFICIENCY OF A CONSTANT UNIT SIZE STABILIZATION POND,
FIGURE 20 EFFECT OF TEMPERATURE ON THE REQUIRED SIZE
OF A STABILIZATION POND TO PRODUCE A CONSTANT BOD
REMOVAL EFFICIENCY.
148
-------
at 30°C. Equations (#20) and (#24) and a 9 of 1.072 are
employed to develop the model shown. An increase in pond
temperature from 20°C to 30°C will result in an increase in
BOD removal to 85%.
Figure 20 shows the relative pond volume required to achieve
a specified removal efficiency as temperature changes.
Equation (#25) and a 6 of 1.072 are employed to develop the
model shown. All variables, except time and temperature,
are kept constant in the pond evaluated. A reference tem-
perature of 20°C is used for Figure 20. A decrease in tem-
perature from 20°C to 10°C would require a pond of twice the
size needed at 20°C to achieve an equivalent BOD removal,
while by increasing the temperature from 20°C to 30°C a 40%
reduction in size is possible.
Temperature variations have significant effects on stabili-
zation ponds, as Figures 19 and 20 show. Large capital cost
savings can be realized if pond temperatures can be main-
tained at uniformly high levels. For this reason, stabili-
zation ponds are found mostly in the southern and south-
western areas of the United States. In colder northern
climates, weather and land costs mitigate against their use.
Aerated Basins
An aerated basin (lagoon) differs from a stabilization basin
in that it is usually deeper, in the order of 10 to 14 ft,
and that the majority of the dissolved oxygen required for
the biological processes is supplied by mechanical devices.
These mechanical devices also provide the mixing in the sys-
tem. The turbulence levels maintained within the basin
should be sufficient to ensure a uniform dissolved oxygen
concentration in the basin, but in many cases are not suffi-
cient to maintain all suspended solids in solution. As a
result, certain solids settle to the bottom of the pond,
where they undergo anaerobic decomposition with the subse-
quent return of materials to the basin contents.
The principal drawback of the aerated lagoon system is the
relatively high concentration of suspended solids leaving
the system. Many recent lagoons have included baffle sec-
tions which allow the solids to settle and re-enter the
mixed portion of the pond [25]. Other designs have favored
clarifiers or clarification ponds following these lagoons.
When internal baffling systems or external clarifier systems
with sludge return are employed, aerated lagoon systems be-
come similar to activated sludge systems.
The aerated lagoon has many desirable features. Among these
are ease of operation and maintenance, equalization of the
149
-------
waste, and the ability to dissipate heat when desirable.
The principal disadvantages of the process are the rela-
tively large land areas required, the inability to signifi-
cantly modify the process, the effluent solids loss, and
the overall sensitivity of the process efficiency to changes
in ambient temperature.
The rate of BOD removal may be defined by taking a materials
balance around the system as described in Equation (#24).
The form of the first-order equation used to describe BOD
removal in aerated lagoons is [25]:
L0
(#27)
where: Kt = The product of k and the solids level
Le = Effluent BOD concentration, mg/1
LQ = Influent BOD concentration, mg/1
This equation ignores the effect of solids sedimentation or
resuspensions which are assumed to be in equilibrium.
Temperature affects the rate of biological oxidation in aer-
ated lagoons as in other biological systems. The tempera-
ture relationship normally used to describe this phenomenon
is [25] :
Kt = K209(T~20) (#28)
where: 0 = Temperature coefficient
Kt & K2Q = Biological reaction rates
Aerated lagoons have been found to be quite sensitive to
temperature, with 6 of 1.06 and 1.09 being reported.
An illustrative example of the effect of temperature on aer-
ated lagoon performance has been developed based on data
presented by Eckenfelder [25]. Figures 21 and 22 present
the results of this example. Figure 21 presents the change
in BOD removal efficiency with temperature. A base condi-
tion of 85% BOD removal in 3.5 days at 20°C with 9 = 1.08
150
-------
AERATED BASIN
90
80
70
#
50
«o
o
8
20
10
3.0
2.O
1.0
FIGURE 2l
NOTE: BASED ON 85%
BOD REMOVAL AT
9 = 1.08
% REMOVAL =
l+Kt
J_
FIGURE 22
NOTE: CONSTANT REMOVAL
EFFICIENCY OF 85%
6= 1.08
10 20 30
TEMPERATURE, °C
40
50
60
FIGURE 21 EFFECT OF TEMPERATURE ON THE BOD REMOVAL
EFFICIENCY OF A CONSTANT SIZE AERATED BASIN,
FIGURE 22 EFFECT OF TEMPERATURE ON THE REQUIRED SIZE
OF AN AERATED BASIN TO ACHIEVE A CONSTANT BOD REMOVAL
EFFICIENCY,
151
-------
is used for the example. An increase in temperature from
20°C to 30°C will result in an increase in BOD removal
efficiency to 92%. Figure 22 presents the relative deten-
tion time required to achieve 85% BOD removal as the tem-
perature changes. An increase in temperature from 20°C to
30°C will require a basin approximately 40% smaller than
that required to achieve 85% BOD removal at 20°C.
The temperature of a waste will change during treatment in
a lagoon system. The temperature which will be maintained
in the lagoon system will depend on the heat balance be-
tween the influent wastewater and the ambient air tempera-
ture. Heat is lost through evaporation/ convection, and
radiation, and is gained by solar radiation. The total
heat loss from the lagoon may be defined by the following
relationship [42]:
H = He + Hc + Hr - Hs (#29)
where: H = Net heat loss
He = Heat loss by evaporation
Hc = Heat loss by convection
Hr = Heat loss by radiation
Hg = Heat gain by solar radiation
The heat loss due to evaporation, He, is expressed by the
relationship [42]:
He = 0.00722 HVC (1 - 0.lW)(Vw-Va) ...(#30)
where: Hv = Latent heat of vaporization, Btu/hr-SF
C = Constant characteristic of the lagoon
W = Mean wind velocity, mph
Vw = Vapor pressure at the liquid surface
Va = Vapor pressure at the atmosphere
The heat loss by convection, Hc/ is computed from the rela-
tionship:
152
-------
Hc = (0.8 + 0.32W/2) (Tw - Ta) (#31)
where: Tw = Lagoon temperature (°F)
Ta = Air temperature (°P)
The heat loss by radiation, Hr/ can be expressed by the re-
lationship :
Hr = 1.0 (Tw - Ta) (#32)
The net heat gain by solar radiation has not been defined
at this time from available data and is neglected in these
calculations.
Activated Sludge
The activated sludge process may be defined as a system in
which flocculated biological growths are continuously cir-
culated and contacted with organic wastes in the presence
of oxygen. The oxygen and mixing in the system are supplied
by diffused air, mechanical aeration, or a combination of
the two. The process involves an aeration step followed by
a solids-liquid separation step. The separated solids are
returned for admixture with the waste in the aeration phase.
Many modifications of the process are employed to obtain the
desired degree of treatment. They range from high-rate
processes with aeration detention times of 2 hrs to low-rate
systems with aeration times of 24 hrs or more, depending on
waste strength.
There are several models which relate BOD removal efficiency
with temperature. However, these models are applicable only
in the mesophilic temperature range. The basic form of the
equation describing the BOD removal efficiency in a com-
pletely mixed, activated sludge process has been presented
by Eckenfelder [5, 25]:
KTSatLe
ET = = (#33)
1 La
where: ET = BOD removal efficiency at temperature T°C
KT = BOD reaction coefficient at temperature T°C
Sa = Mass of biological volatile solids
153
-------
t = Detention time in aeration tank
Le = Effluent BOD concentration
La = Influent BOD concentration
The van't Hoff-Arrhenius relationship for the effect of tem-
perature (in a certain range) on purification rates can be
used for analysis, provided that the experimental determina-
tion of the 9 value for the sewage at different temperature
ranges is obtained. This relationship is expressed in the
following equation [5, 25]:
= K2Q e(T~20) (#34)
where: K = BOD reaction coefficient at 20°C
0 = Temperature coefficient
T = Temperature in °C
The optimum temperatures for the biological oxidation of
various types of waste have been reported in the 30°C to
52°C range. A summary of these investigations is presented
on Figure 23. For each investigation reported, the tempera-
ture range and optimum temperature are indicated.
The elevated temperature used in the activated sludge proc-
ess depends upon the optimum value for biological reaction
obtained experimentally for the specific wastewater in-
volved. Based on the information presented on Figure 23,
the optimum temperature for mesophilic biological systems
appears to be in the 30°C to 37°C range. A temperature co-
efficient (0) for activated sludge systems has been re-
ported to average 1.02 at temperatures up to 30°C [25].
There are limited data published for biological oxidation
of wastes in the thermophilic range. A thermophilic bio-
logical population does exist which is capable of BOD re-
movals at temperatures up to 65°C [25]. However, this high
temperature operation requires very stringent control over
system temperature since these cultures are very
temperature-sensitive.
Many factors affect the complex process of organic removal
by microorganisms and it is not surprising to see different
154
-------
ACTIVATED SLUDGE STUDIES
OPTIMUM AND RANGE OF TEMPERATURE
INVESTIGATION
ECKENFELDER?5..
GEHM43..
HUNTER?.5.....
GOT A AS.4.?
RUDOLFS4.?....
GELLMAN50.....
MALZ5.1
DOUGHERTY52
HOWLAND5?....
QL$ M64
L,
i
t32
t ,
30
P2
I37
30
i
»36
i36
136
I"
, t
RANGE OF
If
_ j
in J
OPTIMUM
TEMR
TEMP STUDIE
«i
M
O
^s 3
M
fO
u
10
20 30 40
TEMPERATURE, °C
50
60
FIGURE 23 OPTIMUM TEMPERATURES FOR THE ACTIVATED SLUDGE PROCESS REPORTED IN THE LITERATURE,
-------
investigators reported different results. One study [48]
shows an efficiency decrease with rising temperature in the
30°C to 40°C range, followed by increasing efficiency until
reaching an optimum value between 50°C and 55°C. This
phenomenon can be explained by the fact that after the op-
timum mesophilic temperature range, the mesophilic bacterial
population dies off, followed by growth of a thermophilic
bacterial population. Other data [25, 46] show a decrease
of efficiency after the mesophilic optimum is reached, which
continues through the thermophilic range. A pilot study
[49] of paper mill wastewater treatment reports similar ef-
ficiency at optimum mesophilic and thermophilic tempera-
tures. It has been concluded that the activated sludge
process can be operated successfully at feed temperatures
as high as 52°C.
As an illustration of the effect of temperature on the acti-
vated sludge process, a model is proposed based on Equation
(#33) . Model performance is described by Equation (#33) up
to a temperature of 30°C based on a base condition of 80%
BOD removal at 20°C and a 0 = 1.02. At temperatures above
30°C and below 50°C, model performance remains constant at
the BOD removal level achieved at 30°C. This model is pro-
posed in light of the range of optimum performances, shown
on Figure 23, most of which were from laboratory-scale
evaluations. Experimentation in the 30°C plus range (up to
the thermophilic range) is required to develop a specific
temperature model.
Figures 24 and 25 present the anticipated effect of tempera-
ture on the relative removal efficiency and detention time.
Figure 24 indicates that an increase in temperature from
20°C to 30°C will result in an additional 20% increase in
removal of the remaining substrate. In the example used,
this would correspond to an 84% BOD removal at 30°C in an
activated sludge system designed for 80% BOD removal at
20°C. Figure 25 presents the relative detention times re-
quired to achieve a constant BOD removal or the relative %
substrate remaining. Also presented on Figure 25 are
Rudolf's data [48] for the temperature range of 30°C to
50°C. These data substantiate the proposed model (constant
removal efficiency) in that temperature range. In this
model, a temperature increase from 20°C to 30°C would result
in a system with 30% less detention time to achieve the same
BOD removal.
In order to keep BOD removal at optimum rate, sufficient
amounts of dissolved oxygen are needed in the activated
156
-------
ACTIVATED SLUDGE
UJ
o
15
1.0
(O
£S05
o
z
UJ
o 1.5
u.
u.
UJ
_l
<
1 1.0
UJ
< 05
_i
UJ
tr
0
-
-
i
[COMPLETELY MIXED MODEL
FOR EM =80
^
1
% e 0=1.02
^-"
1
PROJECTED MODEL
BASED ON E
NOTE:
CONSTANT
TIME
ET= KTS«
T L
i
WPIRICAL DATA
DETENTION
itL.
a
'
FIGURE 24
1
-
-
-
1
COMPLETLY MJXED MODEL
0 = 1.05
1
\ "
~^*^^^ i
1
m PROJECT
BASED ON EM
r-DATA FRO
t-J A —i
NOTE:
EXTENSIO
CONSTANT
EFFICIEN
ALSO API
TO % SU
REMAINS
I
iD MODEL ^
PIRICAL DATA
M REF. 48
k 4
N FOR
REMOVAL
CY
'LICABLE
BSTRATE
IG
FIGURE 25
1
10
20 30 40
TEMPERATURE °C
50
6O
FIGURE 24 EFFECT OF TEMPERATURE ON BOD REMOVAL
EFFICIENCY OF A CONSTANT SIZE COMPLETELY MIXED
ACTIVATED SLUDGE UNIT.
FIGURE 25 EFFECT OF TEMPERATURE ON THE REQUIRED
ACTIVATED SLUDGE UNIT SIZE TO PRODUCE A CONSTANT BOD
REMOVAL EFFICIENCY,
157
-------
sludge process. Dissolved oxygen saturation varies with
temperature. At elevated wastewater temperatures, the oxy-
gen utilization rate of microorganisms increases and the
saturation of dissolved oxygen decreases. The driving force
(difference between saturation and actual concentration) for
oxygen transfer from air to wastewater decreases with in-
creasing temperature. A check of the theoretical oxygen re-
quirements of an activated sludge system operating over a
temperature range of 10°C to 50°C has been made. The dif-
ference in aeration requirements over the temperature range
of 10°C to 50°C is within 10%. This would be within the
design safety factor normally used when selecting oxygena-
tion equipment.
Trickling Filters
A trickling filter is a packed bed of medium (e.g., plastic,
stone) covered with microbial slime. As the wastes pass
over the slime layer, the organic material present in the
waste is reacted upon by microbial action.
Similar to the activated sludge process, the van't Hoff-
Arrhenius relationship can be applied for predicting the
temperature effect on trickling filter performance. Howland
determined a 6 = 1.035 from his experimental data.
An illustrative example of the effect of temperature on
trickling filter performance is presented on Figure 26.
Rowland's filter equation and temperature coefficient are
used in this example. A base condition of 75% BOD removal
at 20°C and a constant detention time are also used. Figure
26 indicates that a temperature increase from 20°C to 30°C
will result in a 15% increase in relative removal efficiency
from 75% BOD removal at 20°C to 86% BOD removal at 30°C.
Rotating Disks
Biological fixed-film rotating disk (BFFRD) is a treatment
process which involves the use of biological films attached
to the rotating disk. The experimental results reveal that
the BFFRD is an efficient treatment process for organic re-
moval, ammonia removal, and resisting organic shock load-
ings [60]. The limiting factors in substrate removal are
the dissolved oxygen content, diffusion, or substrate con-
centration. The process variables include influent loading,
flow rate, detention time, temperature, number of stages of
disk, surface area, submerged depth, speed, and direction.
158
-------
TRICKLING FILTER
L8
>-
U
z
UJ
o
U. 14
U.
UJ
UJ
>
6
uJ UQ
or
0.6
<
-
-
_
i
NOTE:
BASED ON
FILTER E
. E20=75«><
CONSTANT
6 = 1.035
^^^^
i
1
ROWLAND'S
OUATION
>
DETENTION
Vi
^
i
TWO STAGE
r — |
TIME
L^-"
(^
FIGURE 26
i
1
1
1
D 10 20 30 40 50
TEMPERATURE, °C
FIGURE 26 EFFECT OF TEMPERATURE ON THE BOD REMOVAL
EFFICIENCY OF TRICKLING FILTERS, UNIT SIZE CONSTANT,
159
-------
Mathematical models describing the BOD removal efficiency
are reported [60]. Temperature is one of the variables in-
cluded in these models. Experimental data indicate that
BOD removals increase with increasing temperature. However,
experimental data are limited to the mesophilic range of
20°C to 30°C. A temperature coefficient for BOD removal
rate of 9 = 1.025 has been determined for a single-stage
BFFRD.
Anaerobic Digestion
Anaerobic digestion is commonly used to stabilize the waste
sludge generated during municipal waste treatment. Princi-
pally, the sludges are comprised of primary sludge and ex-
cess sludge from either the activated sludge or the
trickling filter operations.
Digestion produces a sludge more amenable to dewatering,
disposal, lagooning, dilution, and other disposal methods
[19]. Digestion occurs in a mixed culture of microorganisms
where particular species are most active in different
stages. The decomposition is accompanied by gasification,
liquefaction, stabilization, colloidal structure breakdown,
and the release of moisture. Since the digestion process
is not complete, byproducts of intermediate metabolism in-
clude organic acids, ammonia, methane, hydrogen sulfide,
carbon dioxide, and carbonates [61]. Volatile solids reduc-
tion of about 70% is commonly achieved by anaerobic diges-
tion [18] in the mesophilic range with approximately 24
days' detention time.
One of the more important factors controlling the rate and
completeness of digestion is temperature [62]. Numerous
studies [19,63,64,65,66,67] describe the effect of tempera-
ture on these design parameters and models. There are two
distinct ranges for operation of anaerobic digestors. Low-
range temperatures are called mesophilic and cover the
range of about 88°F to 103°F (31°C to 40°C) . The thermo-
philic temperatures range from approximately 116°F to 132°F
(46°C to 55°C). There is some disagreement concerning the
exact maximum and minimum temperature, of these ranges, as
the following table shows:
160
-------
Thermophilic* Mesophilic*
Opt. Max. Min. Opt. Max. Min. Reference Date
108 68 1934
128 98 65 1937
122 86 66 1948
133 122 108 90 69 1953
100 70 1956
128 130 115 98 100 80 19 1959
135 122 104 84 5 1961
* Alt temperatures in °F.
For the purpose of this discussion, 128°F (53°C) and 98°F
(37°C) will be used for the optimum temperature for thermo-
philic and mesophilic operation, respectively. In each
range a corresponding bacteria population is responsible
for digestion.
Thermophilic operation offers various advantages over meso-
hilic operation. Thermophilic temperatures result in sig-
nificantly smaller digestion periods [19]. The capacity of
the digestor is directly proportional to the time of diges-
tion, thus indicating proportional decrease in required ca-
pacity in the thermophilic ranges. Reduction in required
digestor capacity of 33% to 50% is suggested by A.S.C.E.
[19] for thermophilic operations, independent of concentra-
tion and storage. A capacity temperature relationship
based on the work of Fair and Moore [65] is presented on
Figure 27. Loadings for digesters run at thermophilic
ranges were found to be 1.6 times the maximum loading in
the mesophilic range for the same density of sludge [71] .
Digesters run at thermophilic temperatures (128°F) are re-
ported to be stable and resistant to upset [72]. However,
because of operating difficulties few plants utilize the
process. Doubling of solids loading of the digestorwas
found to have little effect at the™°P^ i^^^f^uc_
except to reduce gas production and volatile solids reduc
tion a small amount [72].
Gas oroduction at thermophilic temperatures is approxi-
mate?^ Sgher than «?.~?^i%f^ratSSeS8S5l
also found that the gas components are essentially tne
161
-------
ANAEROBIC DIGESTION
FIGURE 27
NOTE:
BASED ON FAIR t MOORE
DATA65 FOR LOW RATE
DIGESTION
AS PRESENTED IN REF. 19
2% SOLIDS IN
SLUDGE ADDED
30 40
TEMPERATURE, °C
FIGURE 27 EFFECT OF TEMPERATURE ON ANAEROBIC DIGESTOR
SIZE REQUIRED TO PRODUCE A CONSTANT VOLATILE SOLIDS
DESTRUCTION AT VARIOUS SOLIDS CONCENTRATION,
162
-------
ANAEROBIC DIGESTION
BASED ON DATA REPORTED
BY GOLUEKE62
BASED ON DATA
BY GOLUEKE62
40 45
TEMPERATURE, °C
FIGURE 28 EFFECT OF TEMPERATURE ON GAS PRODUCTION
IN ANAEROBIC DIGESTORS OF CONSTANT SIZE.
FIGURE 29 EFFECT OF TEMPERATURE ON VOLATILE SOLIDS
DESTRUCTION EFFICIENCY IN CONSTANT SIZE ANAEROBIC
DIGESTORS,
163
-------
same, while some investigators found an increase in methane
from thermophilic digesters. Maly [67] found a greater
organic nitrogen conversion to ammonia at thermophilic
ranges, while Garber [72] found less total nitrogen in the
digested sludge.
The efficiency of the digestor in destruction of volatile
matter decreases by approximately 5% at thermophilic tem-
peratures [62]. Figure 29 presents the data developed by
Golueke in laboratory-scale studies of the effect of diges-
tor temperature on the destruction of volatile matter.
Improved sludge handling characteristics for thermophilic
sludges were found by Garber [72] and supported by Golueke
[62]. These changes in characteristics were found to be:
(1) the average particle size was larger; (2) the protein-
aceous material was more completely digested; and (3) the
sludge had less total nitrogen. Digested sludge concentra-
tions increased from 3.64% to 4.85% and average vacuum fil-
ter yields increased from 1.7 to 6.3 Ibs/SF/hr [72].
Chemical dosage (ferric chloride) was also reduced from
6.5% to 3.4% in Garber's study.
Popova [73] found that most pathogenic microbes (except
viruses) are destroyed at 50°C and that all viable eggs of
helminths were gone at a digestion temperature of 51°C.
Rawn [74] found no problems with odor in his studies when
live steam was used to heat digesters to the thermophilic
range.
Digestor supernatants can impose a high BOD and solids load
on other treatment plant processes and the effluent
receiving water. Fischer [71] found that the supernatant
from thermophilic digestors contained more solids, being
high in colloidal and non-settleable solids which were dif-
ficult to remove even at high coagulant doses of lime and
ferric chloride. Two-stage digestion, a mesophilic first
stage followed by a thermophilic second stage, resulted in
a supernatant of better quality. Golueke [62] states that
Fischer's findings of poor supernatant quality could have
been due to inadequately adapted digestors which produced
poor digestion.
Aerobic Digestion
Aerobic digestion can be described as a process where micro-
organisms obtain energy by auto-digestion of the cell
164
-------
protoplasm and the biologically degradable organic matter
in the sludge cells is oxidized to carbon dioxide, water,
and ammonia [75]. Aerobic digestion produces a biologi-
cally stable sludge suitable for a variety of further de-
watering and disposal operations [18].
Volatile solids reduction is dependent upon detention time
with a sharp increase in volatile solids reduction as the
detention time is extended to about 12 days [76] .
Lawton [76] and Drier [77] found that temperature has an
appreciable effect at short detention times and that the
effects of temperature decreased at longer detention times
or lower loading rates. Figure 30 presents the relation-
ships developed by Drier [77] in his study of temperature
effects on aerobic digestion. These results are based on
bench-scale studies of mixed primary and waste-activated
sludge. Continuous flow reactors, with feed sludge at 3.2%
solids concentrations and approximately 70% to 80% volatile
solids, were used by Drier.
Drier [77] also found that pH and alkalinity in the aerobic
digestor rose with increasing temperature. Lawton [76]
found that long detention times were required to produce a
digested sludge with good settling and dewatering charac-
teristics. Loehr [78] and Woodley [79] found that thermo-
philic (52°C) aerobic digesters were less efficient than
those operating in the mesophilic (35°C) range.
165
-------
AEROBIC DIGESTION
50
40
o
o
o
UJ
o
V)
30
.025 Lbs/FT3/DAY (60-DAY)
-0.048 Lbs/FT3/DAY (30-DAY)
-0.145 Lbs/FT3/DAY*(IO-DAY )
FIGURE 30
DATA REPORTED BY
DRIER
77
o
>
20
10
-0.29 Lbs/FT3/DAY (5-DAY)
I
VOLATILE SOLIDS LOADINGS
10
20 30
TEMPERATURE, °C
50
FIGURE 30 EFFECT OF TEMPERATURE ON VOLATILE SOLIDS DES-
TRUCTION EFFICIENCY IN CONSTANT SIZE AEROBIC DIGESTORS.
166
-------
SECTION A-V
DISINFECTION
Chlorination
Chlorination has long been considered to have the greatest
practical potential of all disinfection systems for freeing
sewage of pathogens [80] . Chemical disinfection theoreti-
cally proceeds in two steps: (1) penetration of the cell
wall, and (2) reaction with the cell enzymes [6].
The rate of disinfection is generally considered to be
determined by the rate of diffusion of the disinfectant
through the cell wall or the rate of reaction with an
enzyme. The van't Hoff-Arrhenius relationship can be used
to describe temperature effects [6] .
A convenient form for this relationship is:
(T2 - Ti)
IOCT =
10g
t2 4.56
where: Tx & T2 = Two absolute temperatures (°K) for which
the rates are to be compared
t: & t2 = Times required for equal percentages of
kill at fixed concentrations of
disinfectant
E = Activation energy (calories) and is a
constant characteristic of the reaction
(for aqueous chlorine, E = 8,200 @ pH
7.0)
For T2 - TI = 10, the ratio ti/t2 (called Qio) is approxi-
mately related to E at normal water temperature as follows
[6] :
log Qio = log (t!/t2) = E/39,000 ..... (#36)
For aqueous chlorine at a wastewater pH of 7.0, E = 8,200
and Qio = 1.65. An illustrative example of the theoretical
temperature effect on the contact time required to achieve
a constant percent kill using aqueous chlorine is presented
on Figure 31. A temperature increase from 20°C to 30°C
167
-------
CHLORINATION
o
UJ
u.
z
Ul
cc
16
1.4
1.2
1.0
08
0.6
0.4
0.2
0
NOTE: CONSTANT% KILL
AQUEOUS CHLORINE
E = 82OO
t, . E(Tg-T.) _
4.56 T, T2
O
I-
Z
UJ
o
z
o
o
UJ
i
o
UJ
tr
1.6
1.4
1.2
1.0
Q8
0.6
0.4
0.2
0
(
\
\
^
\
.
\
N
NO
\
TE: C
A
C
F
1 FIGURE 32
ONSTANT CONTACT TIME
iQUEOUS CHLORINE
ONSTANT % KILL
"OR 99 % KILL OF E COLI
Coa6tp =0.24
N,^
^1
^^
b
) 10 20 30 40 50
TEMPERATURE, °C
FIGURE 31 EFFECT OF TEMPERATURE ON THE
PERCENT KILL "ITH a
168
-------
will result in approximately a 35% reduction in the required
contact time for disinfection.
A relationship generally used to describe the observed dis-
infection efficiency as a function of disinfectant concen-
tration is [6] :
C tp = Constant ...................... (#37)
where: C = Concentration of disinfectant
= Time required to effect a constant % kill
n = A coefficient of dilution or a measure of the
order of the reaction
For a 99% kill of E . coli with aqueous chlorine as HOC1,
Berg [6] has found the following relationship:
• 0.86
t = 0.24 (#38)
Using this equation and the information on Figure 31, a
relationship between required chlorine dosage and tempera-
ture can be developed. Figure 32 presents the relationship
between chlorine dosage and temperature for a 99% kill of
E. coli with aqueous chlorine in a constant contact time.
An increase in temperature from 20°C to 30°C will result in
a 42% reduction in the amount of chlorine required for
disinfection.
169
-------
SECTION A-VI
ADVANCED TREATMENT PROCESSES
The unit operations discussed in this section have been
classified as advanced treatment processes. Generally,
these processes are not commonly employed in "conventional"
secondary wastewater treatment. Therefore, they are clas-
sified as advanced which is arbitrary in some cases.
Many of these processes could be included solely or as
part of a system to provide tertiary treatment after con-
ventional secondary biological treatment. For most of the
processes reviewed, insufficient information is available
to allow development of a model describing temperature
effects on the process.
Ultra-High Rate Filtration
The ultra-high rate (UHR) filtration process involves the
filtration of wastewaters through a multi-media bed up to
20 ft in depth at application rates of from 12 to 30 gpm/SF
[93]. Chemical coagulation prior to filtration may be em-
ployed to enhance removal efficiencies and decrease the
head loss through the filter.
UHR has been applied as a tertiary treatment step to achieve
higher degrees of removal of suspended solids, BOD, and
phosphates. Experimental tests have been reported using
alum and polymer additions to achieve suspended solids, BOD,
and total phosphate removal efficiencies of 99%, 97%, and
98%, respectively [93]. The average effluent concentrations
of suspended solids, BOD, and total phosphates were reported
as two, four, and less than one mg/1, respectively. Further
studies on the removal of other contaminants are required
to establish the total effectiveness of the UHR filtration
process.
The temperature effect on the UHR filtration process is
expected to exhibit a similar relationship to that presented
for the gravity filtration process.
Organic Carbon Removal
Activated carbon is used in wastewater treatment to remove
soluble organic compounds from solution. The wastewater is
170
-------
normally pumped through packed bed or expanded bed acti-
vated carbon columns to remove residual refractory organic
compounds from biological secondary treatment effluents.
Recently activated carbon adsorption has also been used for
the treatment of effluents from physical-chemical processes.
Carbon adsorption is a surface phenomenon where molecules
are adsorbed due to the attraction between surface charges
within the carbon pores and the adsorbate. The overall
rate of adsorption is generally considered to be limited by
the rate of diffusion of molecules into the carbon pores.
Therefore, Equation (|35), the van't Hoff-Arrhenius equa-
tion, can be used to describe the effect of temperature on
the overall rate of adsorption.
For liquid adsorption systems, there is no precise method
for selecting and predicting the performance of activated
carbon types founded on their basic properties or those of
the adsorbing material. Design data must be obtained from
pilot plant testing for the particular application under
consideration.
As an illustration of the effect of temperature on carbon
adsorption, the work of Morris and Weber [94] is used to
develop the relationship presented in Figures 33 and 34.
A series of carbon adsorption test runs were performed at
various temperatures on water containing alkyl benzene sul-
fonate (ABS). Figure 33 presents the change in the rela-
tive adsorption rate as temperature increases. The van't
Hoff-Arrhenius equation is presented along with data re-
ported by Morris and Weber [94J. A temperature increase
from 20°C to 30°C will result in a 25% increase in the rate
of ABS adsorption on activated carbon.
Figure 34 presents the change in relative adsorption capa-
city of activated carbon for ABS as temperature increases.
The Langmuir equation and data presented by Morris and
Weber [94] for ABS are used to develop the relationship
shown. A temperature increase from 20°C to 30°C will result
in an 8% reduction in the relative adsorption capacity of
the activated carbon. A specific set of design conditions
would be needed to demonstrate the total value of increasing
the temperature in an activated carbon system. Where high
volumes of water are to be treated, an increase in adsorp-
tion rate and therefore reduction in liquid detention time
would be desirable.
171
-------
ACTIVATED CARBON ADSORPTION
UJ
a.
o
<
ui
ui
or
1.0
0.9
as
2.0
IB
1.6
1.4
L2
I.O
OB
Ofi
"
X
I
x
r
X
x
4
V
X
>
t
f
| FIGURE 33"
/
A
\
*
NOTE: ABS REMOVAL
DATA FROM REE 94
Lo
k,
n I
E =
• 4 HR TEST
+ 8 HR TEST
A 10 HR TEST
_ E(T2-T,)
4.5
4300
6T, T,
2
IO
20
30
40
50 °C
\
\J
•
\
| FIGURE 34
NOTE: EQUILIBRIUM ABS
CONCENTRATION = 2O/*mol/l
BASED ON LAN6MUIRS EQUATION
AND DATA FROM REF. 94
X = Xmbc/(l + bc)
X
S^
\
\
>»>.
^sj
^^
IO 20 30 4O
TEMPERATURE °C
5O
FIGURE 33 EFFECT OF TEMPERATURE ON THE ADSORPTION
EFFICIENCY OF ACTIVATED CARBON WITH A CONSTANT ABS
APPLICATION.
FIGURE 34 EFFECT OF TEMPERATURE ON ADSORPTION
ACTIVATED CARBON WITH A CONSTANT ABS
172
-------
Nitrogen Removal
Physical, chemical, and biological processes have been used
to remove nitrogen compounds from wastewaters. Present
water quality criteria may require 90% nitrogen removal
and/or a specified nitrogen content in the effluent from
wastewater treatment plants.
The principal forms of nitrogen present in sewage are ammo-
nia and organic nitrogen. Organic nitrogen is generally
present in suspended form. Therefore, those processes
which effect suspended solids removal also effect a removal
of organic nitrogen. Ammonia nitrogen can be removed by
air stripping or by conversion to another more oxidized
form of nitrogen (nitrite or nitrate). The oxidized nitro-
gen forms can be reduced to nitrogen gas by bacterial
action.
In addition to sedimentation, air stripping, and the biolog-
ical processes, other more sophisticated nitrogen removal
processes may be employed including electrodialysis, reverse
osmosis, ion exchange, and distillation. Nitrogen removal
processes discussed herein are:
Activated sludge
Anaerobic columnar filters
Ammonia stripping
Activated Sludge
Nitrogen removal via air activated sludge process is a two-
stage phenomenon: nitrification (the oxidation of nitrogen
forms to nitrite and nitrate) and denitrification (the re-
duction of nitrite and nitrate to nitrogen gas) .
Nitrification occurs in two steps. The first is oxidation
of ammonia nitrogen to nitrite and the second is oxidation
of nitrite to nitrate [95] .
The pH range for'the oxidation of ammonia to nitrite is 7.5
to 9.0 and the range for oxidation of nitrite to nitrate
8.0 to 9.0. The growth rate of Nitrosomonas has been de-
scribed by the following [95, 96]:
= 0.18 e- <#39)
173
-------
where: 1^ = Growth rate
T = Temperature (°C)
Experimental evaluations of the rate of nitrification [96,
97,98,99] show variations away from this formula. Figure
35 presents the data of Mulbarger [97] and a relationship
developed by Metcalf and Eddy [98]. The relative rate of
nitrification is presented as a function of temperature.
Above 25°C, Mulbarger's data show a wide divergence from
Equation (#39). Figure 36 presents the data of Metcalf and
Eddy relating the rate of ammonia nitrogen nitrification to
temperature at various pH values. Based on the information
presented on these two figures, it is apparent that nitri-
fication is severely limited by low wastewater temperatures.
Figure 37 presents the permissible nitrification tank load-
ings based on the work of Sawyer at Marlboro, Massachusetts
[98]. The effect of lower wastewater temperatures on nitri-
fication rate can be offset somewhat by increasing the mixed
liquor volatile suspended solids, as shown on Figure 37.
Where wastewater temperatures drop below 65°F (18°C), such
as in northern climates, the required sludge age may be
excessively high for operation of a single carbonaceous re-
moval and nitrification stage, necessitating sequential
staging [98] . However, with heat enrichment, a single
sludge system for carbonaceous removal and nitrification
may be sufficient and thereby provide a definite cost
savings.
Nitrites and nitrates are biologically reduced to nitrogen
gas in the denitrification step by a wide variety of common
facultative bacteria. Denitrification can occur through the
endogenous respiration of the biomass or through the addi-
tion of an organic carbon source to increase the denitrifi-
cation rate and reduce the required residence time. Various
organic compounds have been used as a carbon source, but
methanol has been found to be the least expensive.
No workable relationship for describing the effect of tem-
perature on the rate of denitrification in activated sludge
has been developed, though data show increase of rates with
higher temperature [97, 100]. Figure 38 presents
Mulbarger's [97] data for denitrification rates as they
vary with temperature. A rise in temperature from 10°C to
174
-------
NITRIFICATION
(ACTIVATED SLUDGE)
BASED ON REF 96
KM= 0.18
M 8 E EQUATION
REF. 98
30
1
o:
UJ
a.
.35
.30
UJ
u.
QL
CO
CO
.25
.20
E
o>
10
E
o>
.15
.10
.05
50% OPTIMUM RATE pH
T
FIGURE 36
OPTIMUM RATE pH 8.4
75% OPTIMUM RATE pH
7.5. 9.3
10
15
20 25 30 35
TEMPERATURE, °C
45 50
FIGURE 35 & 36 EFFECT OF TEMPERATURE ON THE RATE OF
NITRIFICATION IN THE ACTIVATED SLUDGE PROCESS,
175
-------
PERMISSIBLE NITRIFICATION TANK LOADINGS
DATA FROM REE 98
o
o
o
I
10
i
z
NOTE:
PEAK LOAD
FACTOR IS 1.5
TEMPERATURE,°C
FIGURE 37 EFFECT OF TEMPERATURE ON NITRIFICATION
LOADING AT OPTIMUM RATE,
176
-------
DENITRIFICATION IN ACTIVATED SLUDGE
NOTE:
DATA FROM REF. 97
10 20
TEMPERATURE, °C
FIGURE 38 EFFECT OF TEMPERATURE ON DENITRIFICATION
RATE IN THE ACTIVATED SLUDGE PROCESS,
177
-------
20°C resulted in a 300% increase in the rate of denitrifi-
cation. Figure 39 presents the relationships developed by
Sawyer [98] in the Marlboro, Massachusetts studies.
Anaerobio Columnar Filters
Denitrification of fully nitrified waste streams may be
achieved by the anaerobic filtration process. Using meth-
anol as the hydrogen acceptor, denitrifying organisms
within the adhering filter surface slime growth will reduce
nitrate nitrogen to molecular nitrogen gas.
Recent experiments have shown the denitrification capabili-
ties of the anaerobic filtration process [101]. However,
mathematical models of the anaerobic filtration process
have not as yet been confirmed through sufficient
application.
Temperature effects were found to be small. Increasing
temperature has been shown [102] to increase slightly the
denitrification rate and decrease slightly the required
methanol doses. Figure 40 presents the effects of temper-
ature on nitrate nitrogen removal for various methanol
doses.
The columnar contacting denitrifying system provides con-
struction cost savings over suspended growth reactors and
has the added advantage of suspended solids removal and
ease of process control [103].
Ammonia Stripping
Air stripping is used to remove ammonia nitrogen from waste-
waters . Ammonia stripping efficiency increased considerably
with increased temperature [84]. A mathematical model has
been developed which relates ammonia stripping rates with
temperature [104]. Ammonia removals experienced at Lake
Tahoe increased in excess of 10% for a temperature rise from
15°C to 20°C. Figures 41 and 42 present the Lake Tahoe data
on ammonia stripping [84]. Figure 41 presents the relative
ammonia removal rate as a function of temperature. Figure
42 presents the required stripping tower depth, to achieve
a stated percent ammonia removal, as a function of
temperature.
178
-------
PERMISSIBLE DENITRIFICATION TANK LOADINGS
DATA FROM REF. 98
180
160
140
120
o
O
O
O
O
V,
m
1500 ma/I MLVSS
IOOO mg / I MLVSS
NOTE'
PEAK LOAD FACTOR IS
15 20 25
TEMPERATURE,^
FIGURE 39 EFFECT OF TEMPERATURE ON DENITRIFICATION
LOADING RATE.
179
-------
100
ANAEROBIC
COLUMNAR DENITRfFICATION
BASED ON DATA REPORTED
IN REFERENCE 102
NOTE: PARTS METHANOL / PART No3-N ON
WEIGHT BASIS GIVEN IN PARENTHESIS
10 15 20
TEMPERATURE, *C
25
30
FIGURE 40 EFFECT OF TEMPERATURE ON DENITRIFICATION
EFFICIENCY FOR CONSTANT SIZE DENITRIFYING COLUMNS,
180
-------
AMMONIA STRIPPING
DATA FROM REF. 84
1.0
I
UJ
a:
z
o
<
UJ
UJ
a:
10
FIGURE 41
NOTE:
TOWER DEPTH=
20 FEET
FOR 2.0 6PM/ft.2
AND 48 Oca ft.
olr/gal.
12 14 16 18
INFLUENT TEMPERATURE , °C
20
FIGURE 41 EFFECT OF TEMPERATURE ON AMMONIA STRIPPING
EFFICIENCY FOR CONSTANT SIZE UNIT,
181
-------
AMMONIA STRIPPING
DATA FROM REF. 84
Q.
UJ
a
cr
UJ
o
35
90
25
20
15
10
10
FIGURE 42
I
•PROJECTED SOUTH TAHOE DATA
75% NH3
REMOVAL:
NOTE: FOR 2.Qgpm/ft.2
AND 480cu.ft. air/gal.
12 14 16 18
INFLUENT TEMPERATURE, °C
20
\
22
FIGURE 42 EFFECT OF TEMPERATURE ON AMMONIA STRIPPING
UNIT SIZE FOR CONSTANT EFFICIENCIES,
182
-------
Phosphorus Removal
Phosphorus removal may be achieved by "luxury" uptake in
biological systems or by chemical precipitation. Other
less commonly employed processes removing phosphorus in-
clude: electrodialysis, reverse osmosis, ion exchange, and
distillation.
"Luxury" phosphorus removal in the activated sludge is
achieved by incorporation of phosphorus in the synthesized
cell mass beyond that amount required by the cell for syn-
thesis on new cell material. Phosphorus may also be re-
moved from recycled sludge by subjecting the sludge to
controlled anaerobic or low pH conditions [105].
"Luxury" uptake of phosphorus in biological systems has not
been refined to effectively predict its occurrence and ex-
tent based on design criteria and operational parameters
[106]. However, it is predicted that phosphorus removal
would increase at higher temperatures due to increased bio-
logical and chemical activities.
Chemical precipitation employing lime, alum, and ferric
salts, or a combination thereof have been successfully em-
ployed for high phosphorus removals [105]. A survey of the
solubilities of conventionally used chemical additives and
resulting precipitates formed shows that the chemical addi-
tives are highly soluble and increase their solubility only
slightly with increasing temperatures up to 60°C and that
the precipitates formed are insoluble over this temperature
range. Chemical precipitation of phosphates at higher tem-
peratures should be more efficient due to enhanced physical
separation and increased chemical reaction rates [108].
183
-------
SECTION A-VII
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ugal Pump Application," Water & Sewage Works, 21RP477.
2. Perry, J.H. & Perry, R.H., Engineering Manual, McGraw-
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3. Rudolfs, W. et al., "Property of Sludge Which Effects
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4. Brooke, M., "Flow Formula for Slurries," Chemical
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5. Eckenfelder, W.W., Jr. & O'Connor, D.J., Biological
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6. Fair, G.M. et al., Water and Wastewater Engineering,
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184
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14. Stukenberg, J.R., "Biological-Chemical Wastewater
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185
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27. Miller, J.K.P. & Legatski, L.K., "Investigation of
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31. Teletzke, G.H., "Wet Air Oxidation of Sewage Sludge,"
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33. Vesiland, P.A., "Estimation of Sludge Centrifuge
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34. Alberton, O.E. & Guidi, E.E., "Centrifugation of Waste
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35. Landis, D.M., "Process Control of Centrifuge Opera-
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36. Lubow, L.A., "Drying of Sludge on Heated Sludge Beds,"
Sewage Works Journal, Vol. 13, 5, 1011, May 1941.
37. Swanick, J.D. & Baskerville, R.C., "Sludge Dewatering
on Drying Beds," Paper Presented at the London Inter-
national Engineering Exhibition, April 1965.
38. Wiley, J.S., "Utilization and Disposal of Poultry
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p. 515, 1963.
39. Corey, R.C., Principles and Practices of Incineration,
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186
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40. Herman, E.R. & Gloyna, E.F.r "Waste Stabilization
Ponds, Formulation of Design Equations," J. Sewage &
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41. Gloyna, E.F., Waste Stabilization Ponds, Vanderbilt,
November 1971. ~~~~
42. Barnhart, E.L., Aerated Lagoon Design, Vanderbilt,
November 1971.
43. Gehm, H.W., "Activated Sludge at High Temperatures and
High pH," Biological Treatment of Sewage and Industrial
Wastes, McCabe & Eckenfelder Eds., Reinhold, New York,
1956.
44. Kshirsagar, S.R., "Effect of Temperature on Sewage
Treatment," Instn. of Engineers (India), Vol. 44, 10,
3, June 1964.
45. Hunter, W.P. et al., "Temperature & Retention Time
Relationships in Activated Sludge," Proc. 21st Ind.
Waste Conf., Purdue Univ., EES #121, 1966.
46. Gotaas, H.B., "Effect of Temperature on Biochemical
Oxidation of Sewage," SWJ, Vol. 20, 3, 441, May 1948.
47. Carpenter, W.L., "Temperature Relationships in Aerobic
Treatment and Disposal of Pulp and Paper Wastes,"
JWPCF, Vol. 40, 5, 737, May 1968.
48. Rudolf, W. et al, "White Water Treatment, Aeration
with Non-Flocculent Growths," JWPCF, Vol. 25, 1, 70,
January 1953.
49. National Council for Stream Improvement, "Temperature
Relationships in Aerobic Treatment and Disposal of
Pulp and Paper Waste," Tech. Bulletin #191.
50. Gellman, I., "Influence of High Temperature on Acti-
vated Sludge," Johns Hopkins Univ., Res. Project,
unpublished paper.
51. Maly, F., "Investigation of Biological Purification
of Sewage on Aerobic Thermophilic Basis," M.S. Thesis
for Reinisch-Westfalisch Technische Hochschule,
Aachen (W. Ger.), Br. W.P. Abs., Vol. 40, I9b5.
187
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52. Dougherty, M.H., "Elevated Temperature Effect on Citrus
Waste Activated Sludge," JWPCF, Vol. 30, 10, 1263,
October 1958.
53. Rowland, W.E., "Effect of Temperature on Sewage Treat-
ment Processes," Sewage and Ind. Wastes, Vol. 25, 2,
161, February 1953.
54. Quirk, Lawler & Matusky Engineers, Technical Files
#105, 175.
55. Velz, C.J., "A Basic Law for the Performance of
Biological Beds," SWJ, Vol. 20, 4, 607, July 1948.
56. Committee Report, "Sewage Treatment at Military
Installations," SWJ, Vol. 18, 5, p. 791, September
1946.
57. Rankin, R.S., "Evaluation of the Performance of Bio-
filtration Plants," ASCE, JSED, Vol. 120, 823, 1955.
58. Rowland, W.E., "Flow Over Porous Media as in a
Trickling Filter," Proc. 12th Ind. Waste Conf., Purdue
Univ., 435, 1957.
59. Caller, S.W. & Gotaas, H.B., "Analysis of Biological
Filter Variables," ASCE, JSED, Vol. 90, SA6, 59, 1964.
60. Weng, C.N., "Biological Fixed-Film Rotating Disks for
Wastewater Treatment," Ph.D. Dissertation, New York
University, New York, 1972.
61. Dague, R.L., "Anaerobic Activated Sludge," JWPCF,
Vol. 38, 2, 220, February 1966.
62. Golueke, C.G., "Temperature Effects on Anaerobic
Digestion of Sludge," Sew. & Ind. Wastes, Vol. 30,
1225, 1958.
63. Lawrence, A.L., & McCarty, P.L., "Kinetics of Methane
Fermentation," JWPCF, Vol. 41, 2, Rl, February 1969.
64. Fair, G.M. & Moore, E.W., "Heat and Energy Relation-
ships in Digestion of Sewage Solids," Sewage Works
Journal, Vol. 4, 755, March 1932.
188
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65. Fair, G.M. & Moore, E.W., "Observations on the Diges-
tion of a Sewage Sludge Over a Wide Range of Tempera-
ture," Sewage Works Journal. Vol. 9, 3, 1937.
66. Henkelkian, H. et al., "Effect of Change of Tempera-
ture on Thermophilic Digestion," Sewage Works Journal,
Vol. 20, 5, 806, May 1968.
67. Maly, J. & Fadrus, H., "Influence of Temperature on
Anaerobic Digestion," JWPCF, Vol. 43, 4, 641, April
68. Fair, G.M. & Moore, E.W. , "Time and Rate of Sludge
Digestion and Their Variation with Temperature,"
Sewage Works Journal, Vol. 6, 3, 1934.
69. Coulter, R.G., "Environment for Anaerobic Destruction
of Organic Material," Public Works, Vol. 84, 5, May
1953
70. Imhoff, K. & Fair, G.M., Sewage Treatment, Wiley &
Sons, New York, 1956.
71. Fischer, A.J. et al., "Plant Scale Tests on Thermo-
philic Digestion," SWJ, V. 17, 4, 718, 1945.
72. Garber, W.F., "Plant Scale Studies in Thermophilic
Digestion," Sewage & Ind. Wastes, Vol. 26, 1203, 1954.
73. Popova, N.W. & Bolotina, O.T., "The Present State of
Town Sewage & The Trend in Research Work in the City
of Moscow," Advances in Water Pollution Research,
Vol. 2, McMillan Co., New York, 1964.
74. Rawn, A.W., "Sludge Digestion Temperature Control with
Live Steam," Water Works and Sewage, Vol. 90, 6, 246,
June 1943.
75. , Drier, P.E. & Obma, C.A., "Aerobic Digestion of
Solids," Walker Proc. Equip. Co. Bulletin No. 26-S-
18194, January 1963.
76. Lawton, G.W. & Norman, J.D., "Aerobic Sludge Digestion
Studies," JWPCF, Vol. 36, 4, 495, April 1964.
77. Drier, D.E., "Aerobic Digestion of Solids," Purdue
Indus. Wastes Conf. 18, p. 123, 1963.
189
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78. Loehr, R.C., "Aerobic Digestion - Factors Affecting
Design/1 Paper Presented at the 9th Great Plains Sewer
Works Design Conference/ March 1965.
79. Woodley, R.A., "A Study of Aerobic Biochemical Oxida-
tion of Primary Sewage Sludge at Mesophilic and Thermo-
philic Temperatures," Unpublished Master's Thesis,
Purdue University, W. Lafayette, Ind., 1961.
80. Rhines, C.E., "Fundamental Principles of Sewage Chlor-
ination," Purdue Industrial Waste Conference, 20, 673,
1965.
81. Tashnival, C.L. et al., "Effect of Temperature on
B.O.D. of Chlorinated Sewage," Chem. Abs., 66, 7450,
1967.
82. AWWA, Water Quality & Treatment, McGraw-Hill, 1971.
83. Goldstein, L.J. et al, "Continuous Flow Water Pasteur-
izer for Small Supplies," JAWWA, Vol. 52, 247, 1960.
84. Gulp, R.L. et al., Advanced Wastewater Treatment, Van
Nostrand, Reinhold, 1971.
85. Thirumurth, J.D., "Ozone in Water Treatment & Waste-
water Renovation," Water & Sew. Works, Vol. 115,
R-106, 1968.
86. Ogden, M., "Ozonation Today," Ind. Water Eng., Vol. 7,
6, 36, June 1970.
87. Marsh, G.R. & Panula, G.F., "Ozonation in the BOD
Reduction of Raw Domestic Sewage," W&SW, Vol. 112, 10,
October 1965.
88. Stumm, W., "Ozone as a Disinfectant for Water Supply
& Sewage," Presented at Meeting of San. Sect., Boston
Soc. of Civil Engineers, 1957T
89. Evans, F.L. & Ryckman, D.W., "Ozonated Treatment of
Wastes Containing ABS," Proc. of 18th Ind. Waste Conf.,
Purdue Univ., 1963
90. Anon., "O2 & °3~RX for Pollution," Chem. Engr., Vol.
77, 4, 46, February 1970.
190
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91. Mixon, P.O., "Filterability Index & Microscreener
Design," JWPCF, Vol. 42, 11, 1944, November 1970.
92. McCarty, P.L., "Biological Denitrification of Waste-
waters by Addition of Organic Materials," Proc. 24th
Ind. Waste Conf., Purdue Univ., p. 1271, 1969.
93. Nebolsine, R., "High Rate Filtration of Secondary
Treatment Plant Effluent," Paper Presented at the New
York Water Pollution Control Federation Annual MeeF1"
ings, January 1972.
94. Morris, J.C. & Weber, W.J., Jr., Adsorption of Bio-
chemically Resistant Materials from Solution 1, U.S.
Dept. of HEW, AWTR-9, 1964.
95. Eckenfelder, W.W., Jr., "A Design Procedure for
Biological Nitrification and Denitrification," Chemical
Engineering Progress Symposium, Vol. 63, 78, 230, 1967.
96. Downing, A.L. et al., "Population Dynamics," Third
Intl. Conf. on Water Pollution Research, Munich, 1966.
97. Mulbarger, M.C., "Nitrification and Denitrification in
Activated Sludge Systems," JWPCF, Vol. 43, 10, 2059,
October 1971.
98. Metcalf & Eddy, Inc., "Nitrification & Denitrification
Facilities," For EPA Design Seminar for Wastewater
Treatment, February 1972.
99. Wild, H.E., et al., "Factors Affecting Nitrification
Kinetics," JWPCF, Vol. 43, 9, 1845, September 1971.
100. Beer, C. , "Discussion of Evaluation of Anaerobic Deni-
trification Processes," ASCE, JSED, Vol. 96, SA6, 1452,
December 1970.
101. Jerris, J., Personal Communication, February 1972.
102. Gulf South Res. Inc., "Methanol Requirement and Tem-
perature Effects in Wastewater Denitrification," Water
Poll. Res. Ser., 17010DHT09/70, August 1970.
103. Smith, J.M. et al., "Nitrogen Removal from Municipal
Wastewater by Columnar Denitrification," Paper at EPA
Design Seminar for Wastewater Treatment Facilities,
February 1972.
191
-------
104. Roesler, J.F. et al., "Simulation of Ammonia Stripping
from Wastewater," ASCE, JSED, Vol. 97, SA3, 8182,
June 1971.
105. Nesbitt, J.B., "Phosphorus Removal - The State of the
Art," JWPCF, Vol. 41, 5, 701, May 1969.
106. Bunch, R.L., "Factors Influencing Phosphorus Removal
by Biological Treatment," Water-1970, AICE, Vol. 67, 90,
1971.
107. Mulbarger, M.C. & Shefflett, D.G., "Combined Biologi-
cal and Chemical Treatment for Phosphorus Removal,"
Water-1970, AICE, 67, 107, 1971.
108. EPA, Phosphorus Removal Process Design Manual, Tech-
nology Transfer, 1971.
109. Helfgott, T. & Hunter, J.V., "The Removal of Phosphates
from Wastewater by Electrodialysis," Water-1969, AICE,
Vol. 65, 218, 1969.
110. Anon., "High Temperature Electrodialysis Without Heat
Addition," Saline Water Conversion Rep. - 1965, U.S.
Dept. Int., Office of Saline Water, 187, 1964.
111. Brunner, C.S., "Pilot Plant Experience in Demineraliza-
tion of Secondary Effluent Using Electrodialysis,"
JWPCF, Vol. 39, 10, Rl, October 1967.
112. Pruyn, K.T., "Mathematical Model of the Electrodialysis
Process," Water Poll. Contr. Res. Ser., 17090FTAO7/69,
July 1969.~~~
113. Stephan, D.G. & Schaffer, R.B., "Wastewater Treatment
and Renovation Status of Process Development," JWPCF,
Vol. 42, 3, 399, March 1970.
114. Sanders, B.H. & Paris, E.F., "Electrodialysis -
Principles & Applications in the Chemical & Food
Industry," Ionics, Inc., 41A, Watertown, Mass.
115. Anon., "Fouling - Resistant Membranes for Water De-
salting by Electrodialysis," Chem. Eng., Vol. 79, 1,
38, January 1972.
192
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116. Cohen, H., "The Use of Reverse Osmosis in the Treat-
ment of Industrial Effluents," from Developments in
Water Quality Research, Shuval, H.I. (Ed.), Ann Arbor
Science Pub., 1971.
117. Shuval, H.I. (Ed.), Developments in Water Quality
Research, Ann Arbor Science Publishers Inc., 1971.
118. Aerojet Gen. Corp., "Reverse Osmosis Renovation in
Municipal Wastewater," FWQA, 14-12-184, December 1969.
119. Osmonics Inc., Reverse Osmosis Systems, No- 67107, UMC.
120. Schmitt, C.R., "Don't Overlook Ion Exchange," Water &
Wastes Eng., December 1971.
121. Mueller, J., Personal Communication, Manhattan College,
Bronx, New York.
122. Levendusky, J.A., "Progress Report on the Continuous
Ion Exchange Process," Water-1969, AICE, Vol. 65, 1969.
123. Dow Chemical Co., "Nitrate Removal from Wastewaters by
Ion Exchange," Water Poll. Contr. Res. Ser.,
17010FSJ01/71.
124. Envirogenics Co., "Investigation of a New Phosphate
Removal Process," Water Poll. Contr. Res. Ser.,
17010DJA11/70, November 1970.
125. Stamberg, J.B., "EPA Research & Development Activities
with Oxygen Aeration," EPA Design Seminar for Waste-
water Treatment, February 1972.
193
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SECTION A-VIII
GLOSSARY
TERMS
Coagulation
BOD
COD
Floe
Hindered settling
Solids flow
Aerobic
Anaerobic
Mesophilic
Thermophilic
Digestion
The process of agglomeration of small
particles into larger particles through
agitation with or without the aid of
chemicals
Biochemical Oxygen Demand
Chemical Oxygen Demand
A particle formed by smaller particles
through coagulation, usually by chemical
addition
Settling of particles in a liquid medium
that does not behave as a single particle
because of the interaction of other near-
field particles
Rate of downward passage of solids in a
thickening unit process
In the presence of oxygen
In the absence of oxygen
Describes a group of microorganisms that
thrive at a temperature range of about
30°C to 40°C
Describes a group of microorganisms that
thrive at a temperature range of about
40°C to 50°C
Process where complex organic compounds
are converted to methane and carbon diox-
ide gases by anaerobic decomposition by
anaerobic and facultative anaerobic micro-
organisms. Can also be accomplished by
aerobic bacteria, producing carbon dioxide
and ammonia.
194
-------
Pathogen
Disinfection
A microorganism that produces disease
The elimination of pathogenic and other
microorganisms by chemical addition or
other means
Tertiary treatment Processes that are added to secondary
waste treatment facilities to improve
the quality of the effluent
SYMBOLS
°C
°F
gm
mg
cm
1
ml
ft
SF
in.
hr
min
sec
Ibs
ac-ft
mph
gpd
ppm
degrees centigrade
degrees Fahrenheit
grams
milligrams
centimeters
liters
milliliters
feet
square feet
inches
hour
minutes
seconds
pounds
acre-feet
miles per hour
gallons per day
parts per million
195
-------
mixed liquor suspended solids
Btu British thermal units
196
-------
APPENDIX B
ENGINEERING AMD ECONOMIC FACTORS AFFECTING
THE THERMAL HEATING OF WASTEWATER
The concept of operating a wastewater treatment plant at an
elevated temperature as part of an integrated utility com-
plex provides two opportunities for economic benefit. First,
cost savings are possible through a reduction in size of the
more efficient heated plant, and second, heat from the power
or distillation plants which otherwise would be wasted can
be recovered with a potential saving through the elimination
of rejection equipment. Against these savings must be con-
sidered the technical and economic feasibilities associated
with the introduction of waste heat to the wastewater pro-
cess - an evaluation which is dependent upon:
1. The energy requirements of the treatment process.
2. The available energy sources.
3. Mechanisms for transferring energy from an avail-
able source to the wastewater.
Specific considerations were given to the effect of thermal
energy addition on the processes employed in wastewater
treatment; the materials of construction required to obtain
consistently reliable performance without excessive main-
tenance and use of redundant equipment; the heat transfer
coefficients that can be expected for each potential ap-
proach; and the effect of surface fouling on these co-
efficients. All calculations are based on an integrated
wastewater treatment plant utilizing conventional primary
and secondary techniques to treat 50 MGD of raw wastewater
with an annual average (design) inlet temperature of 65 F.
ENERGY REQUIREMENTS
In order to determine the relationship between the cost for
197
-------
a plant to process 50 MGD of wastewater and the temperature
at which the plant is operated, an overall heat balance for
the plant was established, incorporating information devel-
oped with regard to the heat losses through each unit of the
proposed plant for various seasons of the year. As was
described in Section VIII, the aeration tank was determined
to be the unit operation which most affected the overall
plant cost. It was shown that for temperatures in the aera-
tion tank of greater than 86°F, virtually no further savings
in the unit operation cost were achievable. Heat losses in
the system up to and including the aeration tank, were com-
puted to be 7°F, requiring that the incoming wastewater be
heated to 93°F in order to realize the 86°F temperature in
the aeration tank.
The heat required to achieve a temperature of 86°F in the
aeration tank is dependent upon the point of heat addition.
If heat is added prior to the grit chamber and a seven de-
gree temperature loss is assumed in and prior to the aera-
tion tank, the heat required would be 4.85 x 108 Btu/hr. If
the heat is added prior to the primary settler, a reduction
of approximately 0.17 x 10 Btu/hr is possible in the heat
required. After accounting for the heat addition attributed
to the 0.27 MGD of filter supernate at 95°F, heat losses due
to radiation, conduction and convection of the heated waste-
water to the atmosphere and the surrounding area as well as
the cooling effect of the addition of approximately 4 MGD of
thickener supernate at 85°F, an average temperature loss
across the primary settler of one degree is anticipated.
Thus, .an influent of 92°F and an effluent of 91°F is accept-
able to maintain an 86°F average temperature in the aeration
tank.
The third alternative point for heat addition is prior to
the aeration tank, in which case 4.50 x 10^ Btu/hr would
be required.
The energy requirements for winter operation will increase
by 8.7 x 10^ Btu/hr since the wastewater temperature will
be approximately 5°F lower than for the design case; how-
ever, the energy requirements for the summer operation will
decrease by the same amount since the summer wastewater tem-
perature is assumed to be 5°F above the design base.
198
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Four considerations dictate the choice of the point for heat
addition: (1) the cost associated with the quantity of heat
required at each point; (2) the type of equipment used to
transfer the heat; (3) the conditions which could cause
equipment fouling and thus reduce the effective heat transfer
coefficient; and (4) corrosion and/or erosion of the heat
transfer equipment.
AVAILABLE HEAT SOURCES
Having determined the quantity of energy necessary to raise
the temperature of the wastewater to the design temperature,
it is next necessary not only to identify the sources of
available heat in the integrated utility complex, but also
to determine if the energy is not so diffuse as to prohibit
attainment of the necessary wastewater temperature. A num-
ber of sources of energy are available for consideration;
namely,
Prime Steam
The first available heat source which may be considered is
prime steam, defined as steam produced specifically for heat-
ing purposes by fuel combustion or produced in a steam gen-
erator or reboiler associated with a primary or secondary
steam system of a power generating station.
Current fuel costs applicable to intermediate size industrial
boilers are in the range of 60-90£/106 Btu and increasing an-
nually. This cost alone precludes further consideration of a
single purpose energy source for heating of the wastewater.
To this cost would have to be added, of course, the cost of
heat transfer equipment, environmental controls, and in the
case of a nuclear plant, a reboiler to eliminate the possi-
bility of carryover of radioactivity.
Process Steam
The second source of energy is low pressure or process
steam. For this analysis, process steam is assumed to be
available saturated at a temperature of 285°F from a reboiler
driven by a turbogenerator steam supply. Since the satura-
tion pressure of 285°F steam is approximately 53 psi (abs),
199
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either surface heat transfer or direct transfer through injec-
tion of live steam can be utilized. Moreover, a reboiler is
provided in the integrated plant to supply heat for the dis-
tillation process; hence, only the incremental cost increase
for the distillation plant reboiler would have to be con-
sidered. The cost allocation approach which results in the
lowest charges for process steam is achieved by determining
the cost of prime steam produced for turbogenerator use,
then prorating the cost of extracted steam on the basis of
available energy. These cost calculations are described in
detail in Section VI, Nuclear Steam Supply, and result in a
cost of 29C/106 Btu for process steam at 285°F.
Product Water
A potentially attractive approach consists of utilizing
wastewater in lieu of salt water as the cooling media for
the product water. Product water from the base case distil-
lation plant leaves at an average temperature of 104°F and
must be cooled before it can be pumped into a reservoir or
directly into a pipeline. The maximum design temperature
for this type of discharge is 85°F to prevent the introduc-
tion of undesirable concentrations of corrosion products in
the water distribution system. If a design temperature of
74°F to 75°F for the product water were selected, there
would be sufficient heat available to raise the wastewater
stream to a temperature of approximately 93°F from the annual
average inlet temperature of 65°F.
To effect the transfer of heat from the product water to the
wastewater, liquid/liquid plate and shell and tube heat ex-
changers were considered. The capital cost of these heat
exchangers is directly proportional to the surface area
requirements and to the materials of construction.
Assuming that the product water is to be cooled from 104°F to
75°F and the wastewater is to be heated from 65°F to 92°F a
log mean temperature difference (LMTD) of 11 is possible.
Though this appears to be low, it is only slightly lower than
the value calculated for the proposed integrated facility
product water cooler using seawater as the coolant (LMTD of
12). The 92°F maximum wastewater temperature reflects the
addition of heat after the grit chamber, to preclude possible
excessive erosion of the heat exchanger by the raw wastewater,
200
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The transfer of heat by liquid/liquid or steam/liquid
surface heat exchange equipment is governed primarily by the
temperatures at which the exchange takes place and by the
resistances to heat transfer. In__general, the heat trans-
ferred may be expressed at Q = UAAt where Q is the heat ex-
changer duty in Btu/hr, U is the overall coefficient in
Btu/hr/f t2-°F, A is the area of the heat exchange surface,
and At is the log mean temperature difference. For liquid
liquid heat exchangers such as a product water/screened and
degritted wastewater interchanger, the overall resistance to
heat transfer is .caused by the following individual resist-
ance:
(a) Cold fluid film resistance (R^ )
£C
(b) Cold fluid fouling resistance (R )
sc
(c) Metal tube wall resistance (R )
(d) Hot fluid fouling resistance (R - )
sn
(e) Hot fluid film resistance (Rf>.)
The total resistance, R. = R_ +R +R + R -U + R
-------
A slightly more conservative value of 300 Btu/hr-ft2-°F was
selected for the overall heat transfer coefficient for the
product water/seawater heat exchanger used in the cost anal-
ysis.
Estimation of the fouling resistance, RSC' the value for en-
tering cold sewage stream has been attempted in various ex-
perimental tests; the Tubular Exchanger Manufacturer's Asso-
ciation (TEMA) has defined fouling coefficients (reciprocal
of resistance) to be used for various types of coolants, tem-
peratures, and velocities. For example, Kern (B-3) shows in
his tabulation of these values, an average fouling resistance
of 0.006 for Chicago Sanitary Canal coolant water treated for
coolant use. Using this value in the previous example in
place of the sum of Rsc + Rsh yields an overall resistance of
0.00717 or an overall heat transfer coefficient of approxi-
mately 140 Btu/hr-ft2-°F, thus the sensitivity of overall
heat transfer coefficient to fouling resistance is readily
apparent. On the other hand, based upon a computer synthesis
of wastewater composition, a manufacturer estimated the foul-
ing factor to be 0.0015, which would give rise to an overall
heat transfer coefficient of 260 Btu/hr-ft2-°F (B-4).
Qualitative results from studies carried out using Contra
Costa canal water (B-5) indicated that a fouling factor of
approximately 0.0033 existed and that direct use of waste-
water for cooling was attended by biogrowth and other fouling
which had to be controlled by the addition of chlorine and
other chemical additives.
An important approach to decreasing the resistance to heat
transfer is to increase the velocity of the fluids. Three
significant results of this strategy are:
1. The film resistance decreases (or, conversely, the
film heat transfer coefficient increases).
2. Deposits, especially soft ones, are scrubbed off
the surfaces more frequently, making the fouling
resistance lower.
3. Pumping power increases due to increased frictional
loss and the increased pass length required to
maintain contact time.
202
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The first of these results is apparent from Sieder and Tate's
(B-6) equation for heat transfer
0.14
Nu = 0.023 (Re)0-8 (Pr) °-33 LJl)
\ WK/
where ,
Nu = Nusselt number
Re = Reynolds number
Pr = Prandtl number
= Viscosity ratio of fluid at wall to that in bulk
stream
The Nusselt number may be written as " , the Prandtl number
as yCP » and the Reynolds number as _ P. Therefore Sieder
k ~y~
and Tate's equation may be expressed as h = ZV°l8 for given
fluid conditions, where Z is a constant of proportionality.
In the above formulae, the nomenclature used is:
h = film coefficient, Btu/hr-ft^-°F, or reciprocals of
Rfc or Rhf .
D = tube diameter.
k = thermal conductivity of fluid.
y = fluid viscosity.
p = fluid density.
Cp = specific heat.
V = fluid velocity.
Thus, it is possible to increase h by increasing velocity.
Computation shows that doubling the velocity increases h by
1.73 times. Since *~ = Rfc and 5-— = Rfh it can be shown
lif c tn
that each film resistance could be reduced by 73 percent.
Since the film resistance accounts only for a small fraction
203
-------
of the total resistance, the impact of such strategies ap-
pears to be low.
The rate at which deposits, whether biological or chemical,
form and are removed from heat transfer surfaces, as a func-
tion of velocity, has not been clearly established.
In summary, it appears that liquid/liquid heat transfer rates
will be governed primarily by the fouling resistances encoun-
tered in heating wastewater for enhancement of both primary
and secondary treatment processes.
Based on the Chicago and the Contra Costa data, it is appar-
ent that fouling resistances on the order of 0.003 to .005
can be expected with wastewater heating, resulting in over-
all heat transfer coefficients on the order of 140 to 193
f\
Btu/hr-ft -°F. However, based on the professional judgments
of various heat exchangers and condenser manufacturers,
higher values could be anticipated. Therefore, a comprehen-
sive testing program would be in order to determine the ap-
propriate fouling factors, the best method of cleaning and
the required frequency of cleaning. For design purposes in
this study, a wastewater fouling resistance of 0.0037 has
been assumed; and the resulting overall heat transfer coef-
ficient of 156 Btu/hr-ft2-°F was utilized for heat exchanger
design calculation. Although this value may be conservative,
the limited data available does not justify the assumption of
a significantly higher coefficient.
Resistance values for the product water/wastewater exchange
(assuming a velocity through the tubes of 3-4 fps, same as
product water to saltwater case) are as follows:
RfC = .00094
Rsc = .0037
Rw = .00025 (22 Bwg 304 SS 1" OD Tube)
Rsh = .0005
Rfh = -0010
Rt = .00639
= 15S
204
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Turbogenerator Condenser
The incoming was tewater could utilize approximately 7 percent,
or 4.85 x 10B Btu/hr of the turbogenerator heat (6.7 x 10^
.
and thus provide economic benefit for the integrated
facility.
To effect the indirect transfer of heat from the turbogenera-
ator exhaust steam to the wastewater, three types of surface
condenser were considered: (1) standard type shell and tube,
single and multipass condensers with the wastewater on the
tube side; (2) a conceptual design of a box cooler with
manifolded, singlepass, vertical tube sections, with the
steam on the tube side. Vertical tubes are often preferred
when the condensate must be appreciably subcooled below its
condensation temperature; (3) sectionalized condenser. For
these condensers, the cost is directly proportional to the
surface area needed for condensation.
From Fourier's Law, the heat exchange area requirement (for
a constant heat load) is inversely proportional to both the
log-mean temperature difference (LMTD) and the overall heat
transfer coefficient (U) . Thus, cost is proportional to
1/UAt. The major concerns with surface condensers are the
unfavorable effects of fouling and corrosion, due to the use
of wastewater, on the overall heat transfer coefficient. A
decreased overall heat transfer coefficient results in high
surface area requirements and the corrosive nature of waste-
water suggests the use of more expensive construction
materials.
The use of corrosion resistant materials for all heat trans-
fer equipment is required on the bases of operational reli-
ability and assurance of an uncontaminated product. Pub-
lished results wherein carbon steel was used as heat transfer
tubing in heating sanitary canal water show corrosion rates
ranging from 6 mils per year (for canal water with corrosion
inhabitators added) to 50 mils per year for untreated canal
water. These tests were performed in recirculating tubes
(B-7) . Alternative materials of construction for heat ex-
changers include 90/10 and 73/30 Cu-Ni, USS 100 (stainless
409), stainless 316, stainless 304, titanium, monel, and
brass and bronze alloys.
205
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The requirements for reliability discussed above clearly ex-
clude the use of carbon steel for the heat transfer surface.
Other studies (B-8) on polluted feedwaters indicate that the
70/30 Cu-Ni alloy may be superior to the 90/10 material in
the presence of ammonia or hydrogen sulfide. However, fur-
ther corrosion testing in the specific environment of interest
appears to be a definite requirement in order to identify and
measure the long term pitting, uniform corrosion rates and re-
sistance to stress corrosion cracking.
Recent evidence (B-9, B-10, B-ll) suggests that USS 100, Type
409 culvert grade stainless steel could have the necessary
qualities of corrosion resistance, formability, and low cost
to be considered for incorporation in a product water/waste-
water heat exchanges. Preliminary data show the corrosion in
saltwater to be only 2.4 mils for 4 years continuous service
and 0.2 mils penetration for 6 years service in the Delaware
and Monogahela Rivers. However, some deeper isolated pitting
and crevice corrosion was noted. The cost for this material
is 32-33£/lb, as compared to approximately 60-70<:/lb for 304
stainless steel and 9-ll£/lb for carbon steel. Based on the
discussions with fabricators of heat exchangers, the use of
titanium is not prohibitively expensive at this time; how-
ever, the use of titanium or other "exotic" materials does
not at the present time appear to be necessary to heat waste-
water .
The design, operation, and maintenance of the wastewater and
distillation plant is directed toward achieving the minimum
overall costs consistent with such constraints as public
safety and health. High availability by using low reliabil-
ity equipment having poor corrosion resistance, may be
achieved only with a great deal of redundancy in design and
maintenance, or alternatively by using highly reliable,
corrosion resistant, equipment with little redundancy or
maintenance. Experimental analysis is required to determine
the actual fouling and corrosion rates associated with vari-
ous heat transfer surfaces.
The discussion of the relationship between fouling factors
and overall heat transfer coefficients presented earlier in
conjunction with liquid/liquid exchangers generally applies
to liquid/vapor exchangers. However, liquid/vapor heat ex-
changers such as condensers or shell and tube type heat
206
-------
exchangers operating with cold wastewater and exhaust steam
are expected to have somewhat higher overall heat transfer
coefficients than would be achieved in liquid /liquid ex-
changers. Primarily, this is due to the fact that steam con-
densing resistances are lower than liquid film resistances
(B-12) . This advantage is not great since the primary re-
sistance to heat transfer remains on the liquid side. For
example, the overall resistance may be determined from the
following individual resistances:
Rfc = 0-00094 (cold wastewater film resistance)
RSC = 0.0037 (cold wastewater fouling resistance)
RW = 0.00025 (22 BWG 304 SS, 1" OD tube)
Rfh = 0.0003 (condensing film resistance)
Rsh = 0-0002^ (steam fouling resistance)
Rt = 0.00539
U = = • 185 Btu/hr-ft2-°F
The difference between Rfh and Rsjj is the liquid/liquid case
described earlier and the above values for the steam/liquid
case is only .0005, which is only ( ' .. x 100 = 9.3 percent
of the total resistance; consequently, improvements in the
steam side coefficients cannot contribute greatly to the
overall performance.
Again, manufacturers of heat transfer equipment have indi-
cated that the fouling resistance of 0.0037 (based on Contra
Costa experience) , may be conservative, and have suggested
values as low as 0.0005 (B-13) . Such a favorable value would
give an overall resistance of .00219 or an overall heat
transfer coefficient of 456 Btu/hr-ft2-°F. Other manufac-
turers (B-4) have suggested a slightly more conservative
overall heat transfer coefficient of 320 Btu/hr-ft2-°F.
For this application of vapor/liquid exchangers, it is par-
ticularly important to ensure that any leakage is in the
direction of the vapor phase, which could be expected to be
contaminated, albeit slightly, with radioactivity.
207
-------
A relatively high head loss through the sectionalized con-
denser is possible, which coupled with the vacuum pump re-
quirements for the vertical tube unit, are important consid-
erations in using this alternative. However, manufacturers
(B-13) have indicated that a pressure drop through the waste-
water section of a condenser would be approximately 15-20
feet of head.
In summary, heat transfer manufacturers feel the tube fouling
caused by the use of wastewater as a coolant in lieu of sea-
water, assuming that the wastewater has undergone filtration
for the removal of large suspended solids, will require only
minor changes in the condenser surface area requirement and
maintenance and operating procedures with the additional in-
stalled cost involving approximately $650,000-$700,000 (B-13).
However, in order to reflect the absence of operating experi-
ence, this alternative was deferred until such time as ex-
perimental data could be developed to verify these opinions.
Distillation Plant Condenser
If a distillation plant matching the size of a wastewater
treatment plant is utilized in conjunction with the waste-
water treatment and power generating facilities, heat is
available from the distillation plant condenser as a poten-
tial source for wastewater heat. Approximately 5 percent of
the product water produced by the distillation plant comes
from the last effect in the form of low pressure steam that
must be condensed, and has a heat value of 12.5 x 108 Btu/hr.
In order to heat the wastewater to 93°F the temperatures of
the steam formed in the last effect (No. 19) was increased
from 91°F at 1.5" Hg to 100°F at 2" Hg by decreasing the
overall distillation plant At.
The transfer of heat could be accomplished either by a sec-
tionalized surface condenser or by direct injection of steam
into the wastewater. If a sectionalized surface condenser
is used to transfer this heat to the wastewater, the techni-
cal considerations are similar to those discussed in conjunc-
tion with the transfer of heat from the turbogenerator con-
denser. Again, the most important consideration is that of
the extremely stringent control which must be exercised in
design and operation to prevent contamination of the conden-
sate.
208
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Direct transfer of heat from the steam to the wastewater may
be affected by either steam injection or barometric condens-
ers. Usually, steam of sufficiently high pressure (> 15 psia)
is used for direct injection or sparging, requiring the use
of process steam, which has been shown to involve relatively
high costs.
Alternatively, low pressure steam, as is available from the
distillation plant, can be used but power must be supplied to
the system to provide the driving forces necessary for getting
the steam into the water since the water pressure at the bot-
tom of a 10-foot high tank would be 16 times greater than the
steam pressure. This injection could be accomplished by an
eductor or ejector; however, these devices have high head
losses, due to the velocities needed to create a suction of
less than 2.5" Hg, necessary for the steam to be transferred.
Probably no other process equipment gives as much performance
in terms of heat transfer per unit of investment cost, as the
barometric condenser, which
inexpensive, provides direct contact between condensing vapor
and cooling water without resistance of an intervening wall.
Because there is direct contact between the vapor and the
cooling water, barometric condensers are used only where the
condensable materials are not to be recovered. However, this
still leaves a wide field for application where process steam
is to be condensed. For example, barometric condensers are
universally used to condense the steam vaporized from vacuum
pans and from the last effect of multiple-effect evaporators,
They are also used between stages, after the last stage of
multistage steam ejectors, and on vacuum distilling columns
to condense the process steam.
The use of barometric condensers ordinarily is not accompanied
by the deposition of scale or other materials which impede the
transfer of heat. Primarily this is due to the method by
which heat is transferred - warm steam condenses directly on
particles or flowing sheets of water without any intervening
metal surface. Little recorded experience on the the use of
raw, screened and degritted, primary settled, or secondary
settled sewage streams is available since such coolants
ordinarily are not used in power or industrial operations.
However, it is anticipated that such streams can be heated
in this equipment without impairment of the heat transfer
209
-------
function providing certain minimum maintenance procedures are
carried out to prevent accumulation of excessive deposits.
The growth of anaerobic organisms in the barometric condens-
ers is expected since adequate concentration of nutrients and
favorable temperatures exist. These micro-organisms are ex-
pected to be anaerobic since the low absolute pressure at-
tained in the barometric condenser will result in the removal
of most of the dissolved oxygen present in the waste stream.
The degree to which such growth takes place is not known.
Scaling by the deposition of calcium carbonate is not expect-
ed on the basis of Langelier Index values ranging from -0.35
to 0.00 for a sewage plant effluent in Bay Park, Long Island.
With this alternative, the condensate is not recovered but
is recycled with the wastewater. This type of system
has a lower capital and operating cost unan the surface con-
denser approach; however, with this system, the steam that is
added to heat the wastewater requires increased capacity of
the equipment downstream of the heat addition point. As a
result, the cost of the wastewater treatment facility as well
as the distillation plant; increases, both in capital and '
operating costs. Since the total heat requirement of
4.85 x 108 Btu/hr is only half the available latent heat
(12.5 x 108 Btu/hr) from the distillation plant overheads,
two condensers, one for heating the wastewater, by means of
barometric leg condensers and another through which seawater
coolant was passed, would be required. The economics of
this approach are discussed in a subsequent portion of this
Appendix.
OVERALL COST OF HEAT ADDITION
The marginal costs of adding heat to the wastewater have been
calculated for the various alternative methods previously
described. In order to compare these costs on a common basis,
the costs for a nonthermally-enhanced wastewater treatment
plant integrated with a nuclear power plant and wastewater
distillation plant have been calculated. The difference
between this nonthermally-enhanced cost and the thermally-
enhanced cost is the cost of adding heat. This difference
is expressed both as $/yr and C/Kgal product water. It
should be noted that the costs presented here do not take
into account any of the costs associated with the wastewater
210
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treatment equipment itself, but only those costs for the dis-
tillation plant and heat transfer equipment required for
thermal enhancement of the wastewater plant.
As stated above, the base case has been taken as a non-
thermally-enhanced integrated plant. The costs that have
been estimated include the capital plus operating charges for
the distillation plant, steam reboiler, and product post
treatment.
Included in the capital cost estimates are the costs of
size increases in distillation plant, product water cooler
and steam supply (reboiler), and credits through si?e re-
ductions in the distillation plant product water condenser,
the integrated plant intake and outfall structures, and
the pump stations associated with the hypothesised heat ex-
changer and/or barometric condensers. Operating cost calcu-
lations include estimates of labor, electric power, steam,
chemical additions, and other operational and maintenance
charges which are either directly associated with the added
heat transfer equipment or are conventionally treated as a
function of plant design capacities.
The following equations were developed during the course of
this study to determine the cost allocations for the various
unit processes within the plant complex.
Distillation Plant
The distillation plant cost has two major components:
(a) Heat transfer surfaces
(b) Volumetric containment and handling
Based upon various economic and design studies for 50 MGD de-
salination plants that had been prepared for the OSW, approxi-
mately 70 percent of the total plant cost is allocated to
heat transfer surface and the remaining 30 percent to fluid
handling and containment structures.
According to Fourier's Law, it can be shown that the heat
transfer area is inversely proportional to the log mean tem-
perature difference. On the basis of these facts and assump-
tions, the following equation was derived for determining the
cost of the distillation plant as a whole:
211
-------
c = c
new old
At
07F51
0.03 + 0.7
old
new
MGD
new
MGD
old
0.85
For plants with constant At (driving force), the term
(0.3 + 0.7 (At nj/At )0.85j _ 1 A is true kecause for
v old new
the same At, the area does not change. For plants with lower
At's, the area must increase by the ratio of At/o^/ A tnew
because for a lower At and the same heat transfer, A must in-
crease according to Fourier's Law. The costs of the compon-
ents (surface and volume) of the distillation plant are
estimated to scale to the 0.85 exponential power.
C = $59.375 x 106, best current estimate for 47.5 MGD
product plant (B-15)
At = steam to first effect - steam from last effect =
285 - 100 = 185°F
MGD ., .. = 47.5 total product water
old
For the cases using a barometric condenser to inject steam
for wastewater heating:
MGD = MGD . , + W ,
new old exhaust
Where:
W , ,_ = exhaust steam flow to the wastewater plant,
exhaust , . .... _ .... ^
expressed in millions of gallons per day
Steam Reboiler
The reboiler cost is estimated at $2.5 x 106 for the 65°F
reference case. For all other cases, the cost of this piece
of equipment is assumed to vary according to the following
equation:
0.85
212
-------
where:
C __ = §2.5 x 105
old
Qnew = the steam load to the distillation plant, Btu/hr
Qold = the steam load to the distillation plant for the
65°F case = 1.177 x 109 Btu/hr (constant)
Distillation Plant Condenser Credits
The condenser credit is based on the change in area require-
ments for varying steam loads. For the 65°F reference case,
A0 « 200,000 ft2- The cost/ft2 of surface has been estimated
at $15, installed. The following equation was used to obtain
the credit for reduced surface requirements.
Condenser credit 4 c - (15)1 l * W ,°"M ~ U » W»]
sTt; (u0) \
where:
Ac = change in capital cost, $ (This term is negative)
Wc = total steam from the last effect
W , = exhaust steam to the wastewater plant
exhaust
X = latent heat of steam from last effect
c
A~t = log mean temperature difference of the condenser
n = overall heat transfer coefficient for the con-
denser = 550 Btu/hr-ft2-°F (constant)
Intake/Outfall Structure Credit
The following expression was used to obtain the credits for a
smaller intake/outfall structure due to decreased cooling
water requirements:
A I/O cost = C
old
MGD
new i _^
MGD
old
213
-------
where:
A I/O
C
'old
MOD
old
MGD
new
= change in cost for the Intake and Outfall structures
(This term is negative)
= $1.22 x 107
= 1070 MGD (constant)
= new cooling water requirements
Product Water Cooler
For the cases where the last effect temperature was raised * in
order to provide higher temperature steam to the wastewater
plant, the product water temperature was also raised in order
,to keep the distillation plant in heat balance. Thus, a
larger product cooler is needed. For these cases only, the
following expression was used to obtain the debit:
0.7
APC =
W ( At - At . , .
p V new old I Cs
1
where:
A PC
W.
P
A t
new
"old
Uo
Tt
change in cost of the cooler, $ (This term is
positive and is added to the plant capital cost.)
product water flow in Ib/hr = 16.5 x 106 (constant)
product water A t across the cooler - varies ac-
cording to case
product water At across the cooler for the 65°F
case = 20°F (constant)
overall heat transfer coefficient for the cooler
= 300 Btu/hr-ft2-°F (constant)
log mean temperature difference for the product
cooler = 12.35°F (constant)
installed cost of heat transfer surface, $15/ft2
(constant)
214
-------
If a liquid/liquid heat exchanger is used to transfer heat
from the product water to the wastewater, from the cost of
such an exchanger may be subtracted the cost of the product
water cooler in the design plant.
Condenser Seawater Coolant Pump Power Credit
For the cases requiring a smaller condenser, less seawater is
required for cooling purposes and, consequently, less elec-
trical power. The following expression is based upon a 25
psi pressure drop for the seawater, a 75 percent pump effi-
ciency, power cost at 9.1 mills/Kw-hr, and a 365-day opera-
tion.
ACSCPP JMGDnew" MGDold) dO6) (25) (2.31) (0.7457) (8760x9.IxlQ"3)
(1440) (3960) (0.75) (10^)
or
ACSCPP = 0.00080268 (JVJGD - MGD )
new old
where:
ACSCPP = change in power cost, $10 (This term is negative.)
MGD = cooling water requirements for the 65°F case = 221
° MGD (constant)
MGD = cooling water requirements for the smaller condens-
new ers and varies according to case selected
Product Cooler Seawater Pump Power Debit
The equation for the condenser seawater coolant pump is also
used for the product cooler seawater pump except that
MGD =38.2 MGD. The results of these computation are
old
positive and added to the annual operating and maintenance
charges.
Calculated Cost
The dominant component of the cost of direct heat addition
with process steam produced either by a separate fuel source
or in a reboiler driven by extracted steam from the hypothe-
sized power generating facilities is the cost of fuel. Con-
sidering a 95°F - single addition point case, 5.2 x 10b
Btu/hr are required for 50 MGD wastewater feed; i.e., the heat
215
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requirement is approximately 0.25 x 106 Btu/Kgal. A fossil
fuel cost of 60-90C/106 Btu leads to an enhancement cost of
15-23C/Kgal for fuel alone; i.e., without including the cost
of heat transfer equipment, pumping power, O&M, etc. A pro-
cess steam cost of 29C/106 Btu is equivalent to an energy
cost of 7.25<:/Kgal, exclusive of those costs associated with
delivery of the energy. These partial costs, when considered
in the perspective of total conventional wastewater treat-
ment cost of 15-20<:/Kgal, preclude further consideration of
the high temperature steam approach to thermal enhancement
of the wastewater processes.
Three alternative ways of adding heat to the wastewater plant
using barometric leg condensers have been costed for opera-
tion for 86°P, 95°F and 104°F temperature levels. These al-
ternatives include:
(a) distillation plant exhaust steam-decreasing distil-
lation plant At.
(b) distillation plant exhaust steam-constant distilla-
tion plant At.
(c) distillation plant exhaust steam combined with in-
ter-effect extracted steam.
For alternative (a), as the design temperature of the waste-
water treatment plant is increased, the final exhaust temper-
ature from the distillation plant also increases, thus de-
creasing the overall distillation plant At since the steam
temperature to the first distillation effect is constant.
Because of this decreasing A t and because of the increased
throughput (due to the use of the barometric leg) the dis-
tillation plant capital cost will increase. Because of the
increased throughput, the steam to the first effect in-
creases, thus increasing the reboiler size and cost. Since
the barometric condenser uses exhaust steam from the last ef-
fect, the condenser size (and cost) decreases due to the
decreased loading. As the last effect temperature is in-
creased, the product water temperature increases, thus
requiring a larger and more costly product cooler. Associ-
ated with the decreasing condenser and increasing cooler is
a decreasing cooling water load and decreasing intake/outfall
216
-------
structure cost. The cost allocation methods have previously
been described. The costs for this alternative are shown in
Table B-l and line 1 of the composite graph (Figure B-I) .
Alternative (b) is essentially the same as (a) with the ex-
ception that the distillation plant At is held constant by
increasing the steam temperatures to the first effect. All
other increases and decreases in equipment size and cost are
included where applicable. The costs for this alternative
are shown in Table B-II and line 2 of the composite graph.
Alternative (c) combines the use of exhaust steam with steam
that is extracted between distillation plant effects. In do-
ing so, the plant At and final exhaust temperature remain
constant for all levels of heating. As a result, the pro-
duct cooler remains constant for all temperature levels. All
other credits and debits have been estimated as previously
described. The costs associated with this alternative are
shown in Table B-III and line 3 of the composite graph.
All three of the above methods were calculated for two point
heating; that is, heat has been added at the influent to the
primary settling tank and at the influent to the final aera-
tion tank with the warm recycle streams. This concept was
originally examined when heat was assumed to be available at
no cost.
Alternative (a) was also calculated using single point heat-
ing at a location ahead of the grit chamber and the primary
settling tank only, without the warm recycle steam. These
costs are shown in Table B-IV and line 4 of the composite
graph.
As an alternative to the use of a barometric leg condenser
system, the costs associated with a shell and tube liquid-
liquid heat exchanger, transferring product water heat to
the wastewater were examined. In this application, the pro-
duct water is. cooled from 104°F to 75°F, and the wastewater
is heated from 65°F to 92°F (the latter temperature reflects
the addition of heat after the wastewater has passed through
the grit chamber, to preclude possible excessive erosion of
the heat exchanger by the raw wastewater) . Under these con-
ditions, the heat to be exchanged is 4.68 x 1()8 Btu/hr, at a
LMTD of 11-
217
-------
TABLE B-I
BAROMETRIC LEG TWO POINT THERMAL ENHANCEMENT ALTERNATIVE A
Process Parameters and Cost Centers
Process Parameters
Wastewater treatment plant reference,
Temperature , *F
Distillation plant feed rate, MGD
Distillation plant product rate, MGD
Distillation plant feed temperature, °F
Exhaust steam to barometric condensers,
Temperature , °F
Exhaust steam to barometric condensers,
Ib/hr
Extracted steam to barometric condensers,
Temperature , °F
Extracted steam to barometric condensers.
Ib/hr
Extracted steam to barometric condensers,
Temperature , °F
Extracted steam to barometric condensers,
Ib/hr
Capital Costs ($10°)
1. Distillation plant
2. Reboiler
3. Product post-treatment
4. Barometric condensers
5. Barometric pump stations
Subtotal
6. Condenser credit (-)
7. Intake/outfall credit (-)
8. Product cooler debit (+)
Net Total Capital Costs ($106)
Annual Fixed Charges ($106)
1. Distillation plant @ 7.823%
2. Reboiler @ 7.823%
3. Product post-treatment @ 7.823%
4. Barometric condensers @ 7.823%
5. Barometric pump stations @ 7.823%
Subtotal
6. Condenser credit @ 7.823* (-)
7. Intake/outfall credit e 7.823%(-)
8. Product cooler debit @ 7.823% (+)
Net Total Annual Fixed Charges (S106)
Annual Operating and Maintenance Charges (-$10 )
1. Distillation plant
Labor
Electric power
Chemicals
Spare parts and maintenance materials
Steam
2. Reboiler
Operation and maintenance
3. Product post-treatment
Operation and maintenance
4. Barometric condensers and pump stations
Operation_and_maintenance
Electric power
Subtotal
5. Condenser pump power credit (-)
6. Product cooler pump power debit (+)
Net Total Annual Operating and
Maintenance Charges (S106)
Grand Total Annual Charges ($10°)
Cost of Heat Addition ($106/yr)
«/Kgal Product
Case Number
1
65
50
47.5
65
— —
~
— ~
"~
"
"™
59.375
2.500
1.300
--
63.175
__
__
—
63.175
4.645
0.196
0.102
— — ,
»„
4.943
—
—
—
4.943
0.408
0.789
0.694
0.891
2.991
0.025
0.020
~
—
5.818
—
—
5.818
10.761
~
—
2
86
51.16
47.5
88
91
402,500
— —
— —
~~
60.604
2.552
1.300
.126
.206
64 . 788
-0.952
-0.274
0
63.562
4.741
0.200
0.102
0.010
0.016
5.069
-0.074
-0.021
0
4.974
0.418
0.808
0.711
0.913
3.064
0.026
0.020
0.003
0.131
6.094
-0.057
0
6.037
11.011
0.250
1.44
3
95
51.70
47.5
93
100
588,990
™
~~
62.937
2.576
1.300
.129
.208
67.150
-1.724
-0.556
+0.534
65.404
4.924
0.201
0.102
0.010
0.016
5.253
-0.135
-0.043
+0.042
5.117
0.423
0.816
0.719
0.923
3.098
0.026
0.020
0.003
0.131
6.159
-0.116
+0.005
6.048
11.165
0.404
2.33
4
104
52.56
47.5
101
109
889,055
65.384
2.614
1.300
.189
.209
69.696
-2.402
-0.730
+1.135
67.699
5.115
0.204
0.102
0.015
0.016
5.452
-0.188
-0.057
+0.089
5.296
0.430
0.831
0.731
0.939
3.152
0.026
0.020
0.004
0.132
6.265
-0.153
+0.009
6.121
11.417
0.656
3.78
i
218
-------
5.0
4.0
3.0
2.0
1.0
A.
B.
Alternatives
Barometric Condensers - 2 Locations
Curve 1. Decreasing Plant At
Curve 2. Constant Plant At
Curve 3. Intereffect Extraction
Barometric Condensers - Single Location
Curve 4. Decreasing Plant At
30
35
40
Wastewater Plant Reference, Temperature °C
Figure B-I Cost Comparison of Thermal Enhancement Alternative
Methods
219
-------
TABLE B-II
BAROMETRIC LEG TWO POINT THERMAL ENHANCEMENT ALTERNATIVE B
Process Parameters and Cost Centers
Process Parameters
Wastewater treatment plant reference,
Temperature , °F
Distillation plant feed rate, MOD
Distillation plant product rate, MOD
Distillation plant feed temperature, °F
Exhaust steam to barometric condensers.
Temperature , °F
Exhaust steam to barometric condenser*,
Ib/hr
Extracted steam to barometric condensers,
Temperature , °F
Extracted steam to barometric condensers.
Ib/hr
Extracted steam to barometric condensers.
Temperature , °F
Extracted steam to barometric condensers.
Ib/hr
Capital Costs ($106)
1. Distillation plant '
2 . Reboiler
3. Product post-treatment
4. Barometric condensers
5. Barometric pump stations
Subtotal
6. Condenser credit (-}
7. Intake/outfall credit (-)
Net Total Capital Costs ($106)
Annual Fixed Charges ($106)
1. Distillation plant @ 7.823*
2. Reboiler @ 7.823%
3. Product post-treatment @ 7.823*
4. Barometric condensers @ 7.823%
5. Barometric pump stations @ 7.823%
Subtotal
6. Condenser credit @ 7.823% (-)
7. Intake/outfall credit @ 7.823%(-)
8. Product cooler debit @ 7.823% (+)
Net Total Annual Fixed Charges ($106)
Annual Operating and Maintenance Charges ($10 )
1. Distillation plant
Labor
Electric power
Chemicals
Spare parts and maintenance materials
Steam
2 . Reboiler
Operation and maintenance
3. Product post-treatment
Operation and maintenance
4. Barometric condensers and pump stations
Operation and maintenance
Electric power
Subtotal
5. Condenser pump power credit (-)
6. Product cooler pump power debit (+)
Net Total Annual Operating and
Maintenance Charges ($106)
Grand Total Annual Charges ($106)
Cost of Heat Addition ($106/yr)
C/Kgal Product
Case Number
1
65
50
47.5
65
—
—
—
—
—
—
H
fl>
H
•a
f
tfi
a
£
1
co
7
86
51.16
47.5
85
91
402,500
—
__
—
__
JH
»
,_!
3
s
8
£
I
U]
3
95
51.70
47.5
93
100
588,990
—
—
—
—
•
61.174
2.576
1.300
0.129
0,208
65.387
-1.724
-0.556
+0.534
63.641
4.786
0.201
0.102
0.010
0.016
5.091
-0.135
-0.043
+0 . 042
4.979
0.423
0.816
0.719
0.923
3.226
0.026
0.020
0.003
0.131
6.287
-0.116
+0.005
6.176
11.155
0.394
2.27
4
104
52.56
47.5
101
109
889,055
—
~
—
"
62.086
2.614
1.300
0.189
0.209
66. 398
-2.402
-0.730
+1.135
64.401
4.857
0.204
0.102
0.015
0.016
5.170
-0.188
-0.057
+0 . 089
5.038
0.430
0.831
0.731
0.939
3.543
0.026
0.020
0.004
0.132
6.656
-0.153
+0.009
6.S12
11.550
0.789 .
4.55
220
-------
TABLE B-III
BAROMETRIC LEG TWO POINT THERMAL ENHANCEMENT ALTERNATIVE C
Process Parameters and Cost Centers
Process Parameters
Wastewator treatment plant reference
Temperature, °p —
Distillation plant feed rate, MGD
Distillation plant product rate, MGD
Distillation plant feed temperature, °F
Exhaust steam to barometric condensers.
Temperature, °F
Exhaust steam to barometric condensers,
Ib/hr
Extracted steam to barometric condensers,
Temperature, °F
Extracted steam to barometric condensers,
Ib/hr
Extracted steam to barometric condensers,
Temperature, °F
Extracted steam to barometric condensers,
Ib/hr
Case
umber
Capital Costs ($10 )
1. Distillation plant
2. Reboiler
3. Product post-treatment
4. Barometric condensers
5. Barometric pump stations
Subtotal
6. Condenser credit (-)
7. Intake/outfall credit (-)
8. Product cooler debit (+)
62.350
2.581
1.300
0.184
0.266
66.681
-1.608
-0.469
0
65.277
2.610
1.300
0.240
69.694
-2.1S2
-0.638
0
Net Total Capital Costs ($106)
Annual Fixed Charges ($106)
1. Distillation plant @ 7.823*
2. Reboiler § 7.823%
3. Product post-treatment @ 7.823%
4. Barometric condensers @ 7.823%
5. Barometric pump stations @ 7.823%
Subtotal
6. Condenser credit @ 7.823% (-)
7. Intake/outfall credit @ 7.823%(-)
8. Product cooler debit @ 7.823% (+)
64.604
4.878
0.202
0.102
0.014
0.021
66.904
8,107
0,304
0,102
0.019
0.031
Net Total Annual Fixed Charges ($106)
Annual Operating and Maintenance Charges ($10 )
1. Distillation plant
Labor
Electric power
Chemicals
Spare parts and maintenance materials
Steam
2. Reboiler
Operation and maintenance
3. Product post-treatment
Operation and maintenance
4. Barometric condensers and pump stations
Operation and maintenance
Electric power
Subtotal
5. Condenser pump power credit (-)
6. Product cooler pump power debit (+)
, " ——-
Nat Total Annual Operating and
Maintenance Charges ($106)
Grand Total Annual Charges ($106)
Cost of Heat Addition (S106/yr)
*/Kgal Product
5.054
0.424
0.819
0.720
0.925
3.104
0.004
0.174
5.23S
0.439
0.130
0.730
0.937
3.147
0.026
0.020
O.OOS
O.Dli
221
-------
TABLE B-I.V
BAROMETRIC LEG ONE POINT THERMAL ENHANCEMENT ALTERNATIVE A
Process Parameters and Cost Centers
Process Parameters
Wastewater treatment plant reference.
Temperature , °F
Distillation plant feed rate, MOD
Distillation plant product rate, MGD
Distillation plant feed temperature, °F
Exhaust steam to barometric condensers,
Temperature , *F
Exhaust steam to barometric condensers.
Ib/hr
Extracted steam to barometric condensers.
Temperature, "F
Extracted steam to barometric condensers,
Ib/hr
Extracted steam to barometric condensers,
Temperature , °F
Extracted steam to barometric condensers.
Ib/hr
Capital Costs ($106)
1. Distillation plant
2. Reboiler
3. Product post-treatment
4. Barometric condensers
5. Barometric pump stations
Subtotal
6. Condenser credit (-)
7. Intake/outfall credit (-)
8. Product cooler debit (+)
Net Total Capital Costs (S10&)
Annual Fixed Charges ($10&)
1. Distillation plant @ 7.823%
2. Reboiler @ 7.823%
3. Product post-treatment @ 7.823%
4. Barometric condensers @ 7.823%
5. Barometric pump stations @ 7.823%
Subtotal
6. Condenser credit @ 7.823% (-)
7. Intake/outfall credit @ 7.823%(-)
8. Product cooler debit @ 7.823% (+)
Net Total Annual Fixed Charges (S106)
Annual Operating and Maintenance Charges ($10 )
1. Distillation plant
Labor
Electric power
Chemicals
Spare parts and maintenance materials
Steam
2. Reboiler
Operation and maintenance
3. Product post-treatment
Operation and maintenance
4. Barometric condensers and pump stations
Operation and maintenance
Electric power
SuBtotal
5. Condenser pump power credit (-)
6. Product cooler pump power debit (+)
Net Total Annual Operating and
Maintenance Charges (S106)
Grand Total Annual Charges (S106)
Cost of Heat Addition (S106/yr)
«/Kgal Product
Case Number
1
61
SO
47. S
' 65
—
«
—
~
--
~
H
*
"9
n
g
41
i
a
2
86
51.00
47.5
80
91
347,860
~
—
--
--
60.436
2.544
1.300
0.116
0.056
64.452
-0.824
-0.235
0
63.393
4.728
0.199
0.102
0.009
0.004
5.042
-0.064
-0.018
0
4.960
0.417
0.806
0.709
0.910
3.054
0.025
0.020
0.002
O.QI*.
5.999
-0.049
0
5.950
10.910
0.149
0.86
3
95
51.44
47.5
86
100
499,425
-•
— —
--
— —
62.658
2.564
1.300
0.123
0.056
66.701
-1.553
-0.515
+0.534
65.167
4.902
0.201
0.102
0.010
0.004
3.21V
-0.121
-0.040
+0.042
5.100
,
0.420
0.813
0.715
0.918
3.082
0.026
0.020
0.002
0.056
6.052
-0.108
+0.005
5.949
11.049
0.288
1.66
4
104
51.88
47.5
93
109
652,505
~*
"
~~
""
65.073
2.584
1.300
0.141
0.056
69.154
-2.030
-0.650
+1.135
67.609
5.091
0.202
0.102
0.011
0.004
5.410
-0.159
-0.051
+0.089
5.289
0.424
0.820
0.721
0.926
3.109
0.026
0.020
0.002
0.056
6.104
-0.138
+0.009
5.975
11.264
0.503
2.90
222
-------
Potential savings which result from this arrangement include
approximately $500,000 from the reduction in size of the in-
take and outfall structures (Table 10, Section IX), $1,355,000
for the product water cooler in the base case (Case II) and
$1,497,000 in wastewater treatment plant costs due to heat
addition. If these savings of $3,332,000 were to be entirely
offset by the cost of a liquid/liquid heat exchanger, the
maximum permissible area costs for installed equipment as a
function of attainable overall heat transfer coefficients
are shown in Table B-V.
TABLE B-V
BREAKEVEN COSTS FOR A PRODUCT WATER/WASTEWATER
HEAT EXCHANGER AS A FUNCTION OF THE HEAT
TRANSFER COEFFICIENT ATTAINABLE
U Btu/hr-Ft2-°F A Scr Ft $/Ft2 (Installed)
150 2.84 x 105 $11.70
200 2.13 x 105 $15.60
250 1.70 x 105 $19.60
300 1.42 x 105 $23.45
350 1.22 x 105 $27.30
400 1.06 x 105 $31.40
223
-------
REFERENCES
B-l Personal correspondence with Anthony J. Giuriceo,
Sales Engineer, Mesco Tectonics, Inc.
B-2 "Standards of the TEMA", Tubular Exchanger Manufac-
turer's Association, 3rd Edition, New York, New
York, 1952.
B-3 Kern, D.Q., Process Heat Transfer, McGraw-Hill Book
Co., Inc., New York, 1950.
B-4 Personal correspondence with Roma Bhattacharya, Pro-
duct Specialist Heat Transfer Dept., Bell & Gossett
and with Walter C. Bosch, Wallace Eannace Associates,
Inc., Bell & Gossett's, New York Representatives.
B-5 "Central Costa County Sanitary District and Contra
Costa County Water District", Bechtel Corporation,
Contra Costa, California. EPA Grant No. 17080 FSF.
B-6 Sieder, E.N., and G.E. Tate, Industrial Engineering
Chemistry, No. 28, McGraw-Hill Book Co., New York,
1936.
B-7 Weddle, C.L., and H.N. Masri, "Industrial Use of Ren-
ovated Municipal Wastewater", Bechtel Corporation,
San Francisco, California, March 1972.
B-8 Todd, Bailie, and Tuthill, "Desalination-Lower Cost
Water by Proper Materials Selection", Desalination
Conference, Yugoslavia, September 1970.
B-9 Preliminary corrosion performance data on stainless
steels - U.S. Steel Research Laboratory.
B-10 Evaluating Material Performance in 3,000 GPD stain-
less steel desalination test plants - 18 months oper-
ation by H.H. Lawson & T.R. Harkin, Paper No. 25,
1972, Conference of the National Association of Cor-
rosion Engineers.
B-ll R. Smith, Corrosion Resistence of Various Steels in
Liquid Manure Proj. No. 57.001-003 (13) April 1969.
224
-------
B-12 Rohsenow, Warren M., and Harry Choi, Heat, Mass and
Momentum Transfer, Prentice-Hall, Inc., Englewood
Cliffs, New Jersey, 1961.
B-13 Personnal correspondence with William R. Ryan, Sales
Manager Water Products & Environmental Systems,
Westinghouse Electric Corp. Power System, Heat Trans-
fer Division.
225
-------
APPENDIX B
ABBREVIATIONS AND SYMBOLS
ABBREVIATIONS
A
Btu
C
new
Cold
Cp
CS
Cu-Ni
D
Op
h
Hg
hr
k
Kgai
LMTD
MOD
Nu
Pr
Q
°-new
Re
MEANING
Rfh
RSC
Area
British thermal unit
New cost
Old cost
Specific Heat
Cost of heat transfer surface
Copper Nickel
Tube diameter
Fahrenheit
Film coefficient
Mercury
Hour
Thermal conductivity of fluid
A thousand gallons
Log mean temperature difference
Million gallons per day
Nusselt number
Prandtl number
Heat exchange duty in Btu/hr
Steam load to the distillation plant
Reynolds number
Cold fluid film resistance
Hot fluid film resistance
Cold fluid fouling resistance
226
-------
ABBREVIATIONS (Cont'd)
MEANING
Rsh
RW
TEMA
U
USS
V
^exhaust
WP
Hot fluid fouling resistance
Total Resistance
Metal tube wall resistance
Tubular Exchanger Manufacturer's Assoc.
Overall heat transfer coefficient
Btu/hr-ft2-°F
United States Steel
Fluid velocity
Total steam from last evaporator effect
Exhaust steam flow to wastewater
Product water flow in Ib/hr
GREEK SYMBOLS
AC
A CSCPP
A I/O
A PC
Ft
X
c
p
p
Change in capital cost
Change in power cost
Change in cost of intake and outfall
structure
Change in cost of product cooler
Log mean temperature differential
Latest heat
Fluid density
Fluid viscosity
227
-------
TECHNICAL REPORT DATA
(Please read Instructions on the reverse before completing)
1. REPORT NO.
EPA-670/2-74-080
2.
3. RECIPIENT'S ACCESSION-NO.
4. TITLE AND SUBTITLE
STUDY OF AN INTEGRATED POWER, WATER AND WASTEWATER
UTILITY COMPLEX
5. REPORT DATE
December 1974; Issuing Date
6. PERFORMING ORGANIZATION CODE
7. AUTHOR(S)
New York State Atomic and Space Development Authorit)
8. PERFORMING ORGANIZATION REPORT NO.
9. PERFORMING ORG \NIZATION NAME AND ADDRESS
New York State Atomic and Space Development Authorit}
230 Park Avenue
New York, New York 10017
10. PROGRAM ELEMENT NO.
1BB043
11.CONTRACT/G«BNj(rXNO.
17080 HHV
12. SPONSORING AGENCY NAME AND ADDRESS
National Environmental Research Center
Office of Research § Development
U.S. Environmental Protection Agency
Cincinnati, Ohio 45268
13. TYPE OF REPORT AND PERIOD COVERED
Final
14. SPONSORING AGENCY CODE
15. SUPPLEMENTARY NOTES
16. ABSTRACT
This study evaluates, technically and economically, a hew approach to siting
power generation, wastewater treatment and water supply facilities. It is included
that the integrated facility results in more efficient utilization of land and
water resources, produces a net reduction in undesirable process effluents, and
achieves at a reduced cost many of the environmental quality goals sought today.
In particular, the use of waste heat for the beneficiation of wastewater treatment
was determined to be sufficiently promising to merit further investigatory
research. The integrated facility studied will supply 1000 Mw of electric power
at 9.1 mills/Kw-hr, will provide secondary treatment for 50 MGD of wastewater for
15
------- |