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).
                              12

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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

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                    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

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                       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

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                      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

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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

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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

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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

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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

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                                         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

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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

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                     SOUTHOIX) P?C°NIC-/ BAYV.EW
                                            NOMTH BRIDeEmMPTON
                                                AREA
Figure 5.   Reference Site Topography
                       29

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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

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    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

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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

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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

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N
                                                              \
                                                                \
                                         1-Reactor Containment
                                         2-Turbine Building
                                         3-Distillation Plant
                                         4-Wastewater  Plant
            Figure 7. Site  Plan - Integrated Facility
                                36

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                         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

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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

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                          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

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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

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                         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

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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

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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

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                          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

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                         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
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                                               GRIT CHAMBER
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                                                                                                                                                 -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
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-------
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

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                                     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

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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

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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

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                         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

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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

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                           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
-------
                          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.

                             89

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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

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                     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

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                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
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12-t
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to
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-
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CURVE B/
RATE OF
POROSITY
COEFFICIE
A CLEAN
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3 MGAD. BED
OF 0.35, SPH
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9
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DEPTH OF 4
EROSITY OF C
: ABILITY OF £
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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

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z *>
z
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CO
UJ
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uj 50
s
" 4(1
CO
<
»30
*
0
<
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10






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-^







^






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— RA




^^^
TE = 3<
FOR
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^^
)^M( i + aoeox)
d = 0.40mm
X = 50
ROM LAWRENCE16



K> 20 30 40 50
FIGURE 9


****"

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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

-------
            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

-------
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

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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

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               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

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                      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
                       REFERENCES
 1.  Carter,  R.  & Karassik,  I.,  "Basic Factors in Centrif-
     ugal Pump Application," Water & Sewage Works, 21RP477.

 2.  Perry, J.H.  & Perry,  R.H.,  Engineering Manual,  McGraw-
     Hill, 1959.

 3.  Rudolfs,  W.  et al.,  "Property of Sludge Which Effects
     Its Discharge Through 24-inch Pipe,"  Sewage Works
     Journal,  Vol. 12,  1,  61, January 1940.

 4.  Brooke,  M.,  "Flow  Formula  for Slurries," Chemical
     Engineering, January  1958.

 5.  Eckenfelder, W.W., Jr.  & O'Connor,  D.J., Biological
     Waste Treatment, Permagon  Press, New  York,  1961.

 6.  Fair, G.M.  et al., Water and Wastewater Engineering,
     John Wiley,  Vol. 2,  1968.

 7.  Chasick,  A.H. & Theodore,  B.B., "Using Graded Sand to
     Test Grit Removal  Apparatus," JWPCF,  Vol.  36, 7,  884,
     July 1964.

 8.  Renn, C.E.,  "Warm  Water Effect on Municipal Supplies,"
     JAWWA, Vol.  49, 405,  1957.

 9.  EPA, Process Design Manual for Suspended Solids
     Removal,  EPA Technology Transfer, 1971.

10.  Velz, C.J.,  "Influence  of  Temperature on Coagulation,"
     Civil Engineering, Vol. 4,  345, 1934.

11.  Camp, D.A.,  Root & Bhoota,  B.V., "Effect of Tempera-
     ture on  Rate of Floe  Formation," JAWWA, Vol.  32,  1913,
     1940.

12.  Willcomb, G.E., "Floe Formation and Mixing Basin
     Practice,"  JAWWA,  Vol.  24,  1416, September 1932.

13.  Parsons,  W.A., Chemical Treatment of  Sewage & Indus-
     trial Wastes, National  Lime Association, 1965.
                            184

-------
14.   Stukenberg,  J.R., "Biological-Chemical Wastewater
     Treatment,"  JWPCF, Vol. 43, 9, 1791, 1971.

15.   Ives,  K.J.  & Sholji, I., "Research on Variables
     Affecting Filtration," Proc. ASCE, J. San. Eng. Div.,
     Vol.  91,  SA4, 1, 1965.	

16.   Lawrence, W.C., "Varying Wash Water Rates with Changes
     of Water  Temperature," JAWWA, Vol. 22, 208, February
     1930.                   	

17.   Vrablik,  E.R., "Fundamental Principles of Dissolved
     Air Flotation of Industrial Wastes," Purdue Conference,
     XIV,  1959.                            	

18.   Burd,  R.S.,  "A Study of Sludge Handling and Disposal,"
     FWPCA, Pub.  WP-20-4, 1968.

19.   ASCE,  WPCF,  Sewage Treatment Plant Design, 1959.

20.   Malina, J.F., "Temperature Effects on the Dewatering
     of Activated Sludge," Technical Report to U.S.  PHS,
     EHE-69-02, CRWR-35.

21.   Laboon, J.F., "Experimental Studies on the Concentra-
     tion of Raw Sludge," S&IW, Vol. 24, 4, 423, April 1952.

22.   Gassett,  R.B. et al., "Influent of Various Parameters
     on Foam Separation," JWPCF, Vol. 37, 4, 470, April
     1965.

23.   Eldib, I.A.  & Skororoski, R.M., "Renovation of Waste-
     water by Foam & Emulsion Fractionation," ESSO Res. &
     Eng.  Co., Report No. PERD4M-63, 1963.

24.   Rubin, E. et al., "Contaminant Removal from Sewage
     Plant Effluents by Foaming," EWQA, AWTR-5, 1963.

25.   Eckenfelder, W.W., Jr., Industrial Waste Pollution
     Control,  McGraw-Hill, 1969.

26.   Jenkins,  D.  & McGauhey, P.H., "Broad Characterization
     of the Improvement of Sewage Effluents by Foam Separa-
     tion," San.  Eng. Res. Lab., U. of Calif., No. 65-5,
     1965.
                             185

-------
27.  Miller, J.K.P.  & Legatski,  L.K.,  "Investigation of
     High Pressure Foam Wastewater Treatment Processes,"
     Water Poll.  Cont. Res.  Ser.,  17030ES 04/70,  1970.

28.  Kennedy, R.R.,  "Thickening  by Elutriation and Chemical
     Coagulation," in "Sludge Concentration, Filtration and
     Incineration,"  Continued Education Series No. 113,
     U. of Michigan, Ann Arbor,  1964,  as quoted in Ref. 18.

29.  Lamb, C., "Heat Treatment as  an Aid to Sludge Dewater-
     ing - Ten Years' Full-Scale Operation," Journal Inst.
     of Sew. Purif., Part I, pp. 5-15, 1951.

30.  Jepson, C. and Klein, L., "Heat Treatment of Sewage
     Sludge," Journal Inst.  of Sew. Purif., Part I, pp.
     36-45, 1951.

31.  Teletzke, G.H., "Wet Air Oxidation of Sewage Sludge,"
     A Paper Presented at the 16th University of Kansas
     Sanitary Engr.  Conf., 1966.

32.  Dick, R. , "Role of Activated  Sludge Final Settling
     Tanks," ASCE, JSED, April 1970.

33.  Vesiland, P.A., "Estimation of Sludge Centrifuge
     Performance," ASCE, JSED, Vol. 96, SA3, 1970.

34.  Alberton, O.E.  & Guidi, E.E., "Centrifugation of Waste
     Sludges," JWPCF, Vol. 41, 4,  607, April 1969.

35.  Landis, D.M., "Process Control of Centrifuge Opera-
     tions," Chem. Eng. Prog., Vol. 66, 1, 51, January  1970.

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
     Manure," Proc.  18th Ind. Waste Con., Purdue Univ.,
     p. 515, 1963.

39.  Corey, R.C., Principles and Practices of Incineration,
     Wiley, 1969.
                              186

-------
40.  Herman, E.R. & Gloyna, E.F.r "Waste Stabilization
     Ponds, Formulation of Design Equations," J. Sewage &
     Ind. Wastes, August 1958.

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

-------
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

-------
 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

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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

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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

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                    mixed liquor suspended solids
Btu                 British thermal units
                       196

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                      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

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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

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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

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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

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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

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     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

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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

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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

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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

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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
-------