Design Manual Neutralization Of Acid Mine Drainage


            United States
            Environmental Protection
            Agency
Industrial Environmental Research EPA-600/2-83-001
Laboratory           January 1983
Cincinnati OH 45268
                                                    'J
            Research and Development
&EPA     Design  Manual:

            Neutralization  of
            Acid Mine  Drainage

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EPA-600/2-83-001
January 1983
DESIGN MANUAL
NEUTRALIZATION OF
ACID MINE DRAINAGE
U.S. ENVIRONMENTAL PROTECTION AGENCY

Office of Research and Development

Industrial Environmental Research Laboratory
U.S,
-------
DISCLAIMER
This report has been reviewed by the Industrial Environmental Research Lab-
oratory, Cincinnati, U.S. Environmental Protection Agency, and approved for
publication. Approval does not signify that the contents necessarily reflect
the views and policies of the U.S. Environmental Protection Agency, nor does
mention of trade names or commercial products constitute endorsement or
recommendation for use.
ii
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FOREWORD
When energy and material resources are extracted, processed, converted, and
used, the related pollutional impacts on our environment and even on our
health often require that new and increasingly more efficient pollution con-
trol methods be used. The Industrial Environmental Research Laboratory -
Cincinnati (lERL-Ci) assists in developing and demonstrating new and improved
methodologies that will meet these needs both efficiently and economically.

This report provides specific design suggestions for neutralization systems
for acid mine drainage treatment. It details step-by-step procedures,
advantages and disadvantages, arid costs for a variety of mine drainage treat-
ment options. It will be of primary use for industry and consultants and
will be of interest to academia and regulatory agencies. For further infor-
mation, please contact the Noriferrous Metals and Minerals Branch, Energy
Pollution Control Division.
David G. Stephen
Director
Industrial Environmental Research Laboratory
iii
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ABSTRACT
This manual was prepared to assist designers and operators of mine drainage
treatment plants in the selection of processes, equipment, and procedures.
Included is a review of the most popular neutralizing agents and the methods
used to handle, prepare, and feed these alkalis. Also, a detailed engineer-
ing explanation of the various processes applicable to treatment are pre-
sented.

Examples of two treatment facility designs are included, delineating general
equipment specifications and cost breakdowns.

The practical methods of sludge dewatering and disposal are explained, along
with modes of operation to improve solids content of the final volume. Tech-
niques for lagooning and closure of such facilities are also discussed.

Concluding the manual is a cost curve for the installation of treatment
plants of various sizes. This curve will allow designers to derive an esti-
mated budget number for capital expenditures.

This report was submitted in fulfillment of Contract No. 68-03-2599 by Penn
Environmental Consultants, Inc., under the sponsorship of the U.S. Environ-
mental Protection Agency. It covers the period from September 1978 to May
1981.
iv
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CONTENTS

Chapter

DISCLAIMER ii
FOREWORD 111
ABSTRACT 1V
CONTENTS v
LIST OF FIGURES vm
LIST OF TABLES xi
LIST OF CONVERSIONS xm

1 INTRODUCTION 1

1.1 Background 1
1.2 Review of AMD Chemistry 2
1.3 Acidity 4
1.4 Ion Solubility and pH 4
1.5 References 5

2 GENERAL TREATMENT CONSIDERATIONS 7

2.1 Acid Mine Drainage Treatment Systems 7
2.2 References 12

3 CHEMICAL TREATMENT 13

3.1 Introduction 13
3.2 Lime 13
3.3 Limestone 46
3.4 Caustic Soda 49
3.5 Soda Ash 56
3.6 References 58

4 MIXING 59

4.1 Introduction 59
4.2 Types of Mixers 59
4.3 Baffles 63
4.4 Shafts and Drives 64
4.5 Energy Requirements 64
4.6 Flocculant Mixing 67
4.7 Summary 69
4.8 References 69
4.9 Other Selected Readings 70
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CONTENTS (continued)

Chapter page

5 IRON OXIDATION 71

5.1 Introduction 71
5.2 Aeration Systems 71
5.3 Chemical Oxidation 84
5.4 Biological Oxidation 89
5.5 Oxidation Rate Test Procedure 93
5.6 References 94

6 SEDIMENTATION 96

6.1 General Characteristics of Mine Drainage Sludge 96
6.2 Settling Unit Design 100
6.3 Recommended Procedure for a Treatability Settling
Test 121
6.4 References 123
6.5 Other Selected Readings 124

7 SLUDGE DEWATERING AND DISPOSAL 125

7.1 Introduction 125
7.2 Mine Drainage Sludge 125
7.3 Methods of Mine Drainage Sludge Dewatering and
Disposal 130
7.4 High-Density Sludge Process 144
7.5 Summary 148
7.6 References 149

8 ELECTRICAL REQUIREMENTS AND INSTRUMENTATION 152

8.1 Introduction 152
8.2 Electrical Power 152
8.3 Motors and Electrical Controls 153
8.4 Instrumentation 153
8.5 Level Controls 155

9 REVERSE OSMOSIS 157

9.1 Introduction 157
9.2 Operational Considerations 157
9.3 References 163
9.4 Other Selected Readings 163
vi
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CONTENTS (continued)
Chapter

10
11
12
13
ION EXCHANGE

10.1 Introduction
10.2 Sul-biSul Process
10.3 Modified Desal Process
10.4 Two Resin Process
10.5 References
10.6 Other Selected Readings

CHEMICAL SOFTENING

11.1 Introduction
11.2 Lime-Soda Softening Process
11.3 Alumina-Lime-Soda Process
11.4 References

CALCULATIONS AND PROCEDURES FOR DESIGN OF A MINE
DRAINAGE TREATMENT PLANT

12.1 Introduction
12.2 Design Example I
12.3 Design Example II

COSTS

13.1 Introduction
13.2 Cost Breakdown of Design Example I
13.3 Cost Breakdown of Design Example II
Page

164

164
165
167
170
174
175

176

176
176
180
183
184

184
184
203

217

217
218
222
Appendix PHYSICAL AND CHEMICAL PROPERTIES OF LIME

A.I Specifications on Lime
A.2 Solubility of Calcium Hydroxide
A,3 Calculating Weights of Slurry
A.4 Solubility of Magnesium Hydroxide
A.5 Heats of Reaction at 25°C
225

225
230
231
231
231
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FIGURES
Number Page
1-1 Theoretical Solubilities of Selected Ions 6
2-1 Process Flow Sheet for Treatment of Acid Mine
Drainage 8
2-2 Conventional Lime Neutralization Process 9
2-3 High-Density Sludge Treatment Process 11
3-1 Typical Quicklime and Slaker Installation 18
3-2 Hydrated Lime and Slurry Installation 19
3-3 Standard and Offset Hopper Bottoms 21
3-4 Lime Feeders 25
3-5 Oscillating Hopper Feeder 27
3-6 Belt Feeder 27
3-7 Paste Slaker with Classifier for Grit Removal 31
3-8 Paste Slaker with Vibrating Screen for Grit Removal 32
3-9 Detention Slaker 33
3-10 Dipper Wheel and Slurry Feeder 42
3-11 Slurry Feed with pH Control Loop 44
3-12 Slurry Feed by Flow Proportioning 45
3-13 Caustic Soda Treatment System 50
3-14 Portable Caustic Soda Feed Arrangement 51
3-15 Freezing Points of Caustic Soda Solutions 54
3-16 Flume Chemical Feeder 55
3-17 Soda Ash Prill Hopper 57
3-18 Soda Ash Vibrating Feeders 57
viii
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FIGURES (continued)
Number Page
4-1 Comparison of Axial and Radial Flow Patterns 60
4-2 Off-Center, Top-Entering Propeller Positions 61
4-3 Typical Radial Turbine Impellers 62
4-4 Characteristics of a Mixing Tank and Standard Turbine 63
5-1 Solubility of Ferric and Ferrous Iron at Various pH 72
5-2 Typical Floating Mechanical Aerator 76
5-3 Typical Splash Block Placement Pattern 81
5-4 Hydrogen Peroxide Feeding System 87
5-5 Rotating Biological Contactor 90
5-6 Design Procedure for a Four-Stage Rotating
Biological Contactor Configuration 92
6-1 Treatability Test Settling Curves 98
6-2 Type 1 Settling 99
6-3 Zones in a Horizontal Continuous Flow Sedimentation
Basin 101
6-4 Zones in a Circular Center Feed, Horizontal
Continuous Flow Sedimentation Basin 101
6-5 Nondistributed Short-Circuiting Influent 105
6-6 Distributed Influent 105
6-7 Surface-Baffled Pond 106
6-8 Combination Surface- and Submerged-Baffled Pond 106
6-9 Conventional Clarifier 111
6-10 Upflow Flocculator Clarifier 113
6-11 Cable Thickener 118
6-12 Tilted-Plate Gravity Settler 120
ix
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FIGURES (continued)
Number Page
7-1 Manually Operated Filter Press 137
7-2 Top View of Drying Bed Construction 141
7-3 Cross-Sectional View of the Drying Bed Construction 142
7-4 Sludge Drying Bed Sizing Requirements 143
7-5 High-Density Sludge Neutralization Process 146
7-6 Sludge Density vs. Ferrous Iron Percentage for the
HDS Process 147
8-1 Flash Mix Tank with pH Probe Installed 155
10-1 Sul-biSul Process Continuous Ion Exchange Flow Sheet 168
10-2 Modified Desal Process Flow Diagram 171
10-3 Two Resin Ion Exchange System 172
11-1 Unit Processes of Lime-Soda Softening 177
11-2 Stages of the Alumina-Lime-Soda Process 181
12-1 Equalization Basin, Design Example I 187
12-2 Settling Basin, Design Example I 200
12-3 Equalization Basin, Design Example II 206
13-1 Installed Costs vs. Plant Flow 219
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TABLES
Number Page
1-1 Point-Source Discharge Limitations for Acid Mine
Drainage 2
1-2 Alkali Comparison for Treatment of Acid Mine
Drainage 5
3-1 Properties of Lime Slurries 36
3-2 Density of Aqueous Sodium Hydroxide Solutions 53
4-1 Diameter for Radial Turbines in Water 65
4-2 Diameter for Axial Turbines in Water 66
4-3 Radial Turbine Proximity and Liquid Properties
Factors 68
4-4 Axial Turbine Proximity and Liquid Properties
Factors 68
5-1 Cross-Sectional Area of Flow in Circular Pipe 79
5-2 Aeration Detention Time Safety Factors 82
5-3 Physical Properties of Hydrogen Peroxide 86
5-4 Hydrogen Peroxide Costs 88
6-1 Recommended Minimum Crown Widths 104
6-2 Rise Rates for Existing Mine Drainage Clarifiers 115
6-3 Suggested Optimum Design Parameters for Clarifier
Operation 116
7-1 Chemical Analyses of Sludges 126
7-2 Sludge Dewaterability Variables 130
7-3 Vacuum Filtration Operational Variables 134
7-4 Operational Cycles 135
7-5 Mine Water Neutralized Sludge Solids Filtration Rates 136
7-6 Pressure Filtration Cake Data 138
xi
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TABLES (continued)
Number Page
7-7 Pressure Filtration - Norton Treatment Plant Sludge 139
7-8 Summary of Centrifugation Test 145
7-9 Coal Mine Drainage Dewatering Methods 149
9-1 Anticipated Permeate Water Quality 162
10-1 Projected Raw and Finished Water Quality Sul-biSul
Process at Smith Township, Pa. 167
10-2 Typical Water Analysis Hawk Run AMD Treatment Plant 170
10-3 Summary of Ion Exchange System Chemical Analyses 173
11-1 Typical Blended Raw Water Characteristics 179
11-2 Estimated Costs of the Alumina-Lime-Soda Process 183
12-1 Raw Water Quality and Effluent Limitations 185
12-2 Raw Water and Effluent Quality Limitations 204
A-l Typical Analyses of Commercial Quicklimes 226
A-2 pH of Calcium Hydroxide Solutions at 25°C 226
A-3 Properties of Theoretically Pure Lime Components 227
A-4 Gravimetric Percentages of Critical Constituents of
Limes 228
A-5 Properties of Typical Commercial Lime Products 229
A-6 Solubility of Calcium Hydroxide in Water 230
XII
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MULTIPLY (ENGLISH UNITS)
English Unit
acres
acre-feet
British Thermal Units
British Thermal Units/
pound
cubic feet
cubic feet
cubic feet/minute
cubic feet/second
cubic inches
cubic yards
degrees Fahrenheit
feet
foot-pounds
flask of mercury
gallons
LIST OF CONVERSIONS
by
Abbreviation
Conversion
Abbreviation
ac
ac-ft
BTU
BTU/lb
fts
ft3
ft3/min
ftVs
in3
yd3
°F
ft
ft-lb
(76.5 Ib)
gal
0.405
1,233.5
0.252
0.555
0.028
28.32
0.028
1.7
16.39
0.76456
0.555 (°F-32)a
0.3048
0.13825
34.73a
0.003785
ha
m3
kg cal
kg cal /kg
m3
1
m3/min
m3/min
cm3
m3
°C
m
kg-m
kg Hg
m3
TO OBTAIN (METRIC UNITS)

Metric Unit

hectares
cubic meters
ki 1ogram-calori es

ki 1ogram-calori es/ki1ogram
cubic meters
liters
cubic meters/minute
cubic meters/minute
cubic centimeters
cubic meters
degrees Celsius
meters
kilogram-meters
kilograms of mercury
cubic meters
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MULTIPLY (ENGLISH UNITS)

English Unit
gallons
gallons/day
gallons/minute
horsepower
inches
inches of mercury
miles (statute)
million gallons/day
ounces (troy)
pounds
pounds/cubic foot
pounds/gallon
pounds/square inch
(gauge)
pounds/square inch
(gauge)
LIST OF CONVERSIONS (continued)
by
Abbreviation
Conversion
Abbreviation
gal
gal/d
gal /mi n
,hp
in
in Hg
mi
Mgal/d
troy oz
Ib
lb/ft3
Ib/gal
3.785
0.003785
0.0631
0.7457
2.54
0.03342
1.609
3,785a
31.10348
0.454
16.02
119.8
1
m3/d
1/s
kW
cm
atm
km
m3/d
g
kg
kg/m3
g/i
Ib/in2g (0.06805 Ib/in2g)a atm
Ib/in2g
5.1715
cm Hg
TO OBTAIN (METRIC UNITS)
Metric Unit
liters
cubic meters/day
liters/second
kilowatts
centimeters
atmospheres
kilometers
cubic meters/day
grams
kilograms
kilograms/cubic meter
grams/liter

atmospheres (absolute)

centimeters of mercury
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MULTIPLY (ENGLISH UNITS)
English Unit
LIST OF CONVERSIONS (continued)

by
Abbreviation
Conversion
Abbreviation
TO OBTAIN (METRIC UNITS)
Metric Unit
pounds/square inch
(gauge)
square feet
square inches
tons (short)
x tons (long)
yards
Ib/in2g
ft*
in2
ton
long ton
yd
0.0703
0.0929
6.452
0.907
1.016
0.9144
kg/ cm 2
m2
cm2
kkg
kkg
m
kilograms/square
centimeter
square meters
square centimeters
metric tons (1,000
grams)
metric tons (1,000
grams)
meters

kilo-
kilo-

Actual conversion; not a multiplier.
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CHAPTER 1

INTRODUCTION
1.1 Background

Acid mine drainages (AMD) are not new problems to the coal industry or to
many other mining industries for that matter. One need drive only a few
miles through the coal areas of Kentucky, West Virginia, and Pennsylvania,
for example, to see the yellow-stained bottoms of streams that contain few
living organisms.

Acid discharges existed long before the mining of coal began. The acid is
formed when pyrite, which is iron sulfide, is exposed to oxygen and water.
The pyrite oxidizes to form a weak solution of sulfuric acid. As the solu-
tion of sulfuric acid passes over the varieties of rock strata surrounding
the pyrite, it dissolves such metals as iron, aluminum, manganese, calcium,
magnesium, sodium, and possibly some trace metals such as arsenic, selenium,
beryllium, nickel, zinc, and others.

Unfortunately, pyrite occurs naturally in close proximity to the coal seams.
Thus, the mining of coal exposes vast quantities of pyritic material to water
and oxygen and greatly accelerates the natural oxidation processes, resulting
in the significant production of acid mine drainage.

Regulation of the concentration of certain chemical parameters in the dis-
charges is now a way of life for the mine operator. Current new-source dis-
charge guidelines, as proposed in the January 13, 1981, Federal Register
(Vol. 46, No. 8), are shown in Table 1-1 for point-source acid mine drainage
discharges. Depending upon the receiving stream's quality and flow, the
limits may be further restricted by State permits.

Iron, aluminum, and manganese are acid-soluble, so merely neutralizing the
water (increasing the pH) will precipitate these ions. This is not so easy
as it sounds, however, because several factors complicate the precipitation.
First, iron can exist in two forms in acid mine drainage; i.e., ferrous
(unoxidized) and ferric (oxidized). The ferric (Fe3+) form will begin to
precipitate around pH 4.0, forming ferric hydroxide or more complicated
oxy-hydroxides; this is the yellowboy common to stream beds in coal country.
The ferrous (Fe2-*} form begins to precipitate at about pH 8.0 and forms a
blue-green hydroxide. In fact, an easy test for significant ferrous iron
concentration in AMD is to sprinkle lime into the drainage, and if the blue-
green hue develops as the lime dissolves in the water, ferrous iron is pres-
ent. It is usually advantageous to oxidize the ferrous iron to the ferric
state rather than to rely upon ferrous precipitation at high pH's. This

1
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TABLE 1-1

POINT-SOURCE DISCHARGE LIMITATIONS
FOR ACID MINE DRAINAGE
(NEW-SOURCE PERFORMANCE STANDARDS)
Effluent Limitations
Average of Daily Values
Effluent Maximum for Any for 30 Consecutive Days
Characteristic 1 Day Shall Not Exceed
mg/1 mg/1
Iron, total 7.0 3.5

Manganese, total 4.0 2.0

Total suspended solids (TSS) 70.0 35.0

pH Within the range
6.0-9.0
oxidation is accomplished by increasing the pH above 7.0 and inducing air
into the water to provide oxygen. The oxidation rate of ferrous iron is
strongly pH-dependent and proceeds extremely slowly below pH 6.0.

Ferric iron and aluminum both begin to precipitate around pH 5.0. At ex-
tremely high pH's (above pH 10.0), the aluminum may tend to redissolve. A pH
above 8.0 is necessary to precipitate manganese to achieve required effluent
levels.

Although increasing the pH can remove all of the elements from solution as
required by the regulations, many of the floes (precipitates) that form, and
especially iron floes, are quite light and tend to remain suspended rather
than settle. It is probably more difficult in treatment situations to remove
the precipitates (floes or sludge) from suspension than to increase the pH,
oxidize, and precipitate them in the first place. The settling process is
monitored by the total suspended solids (TSS) regulation in the guideline
limitations (Table 1-1). All the regulated floes eventually settle; however,
the trick is to settle them before they reach the outlet of the settling
basin or clarifier and thus be in compliance with the discharge regulations.

1.2 Review of AMD Chemistry

A review of the basic chemical reactions involved in the production of AMD is
helpful in understanding the rationale involved in the design of the unit
processes and the overall treatment scheme.
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Pyrite (FeSz) and marcasite (also Fe$2) are both naturally occurring minerals
associated with coal-bearing strata. Pyrite and marcasite differ in crystal-
line structure but have identical iron disulfide composition. Although both
are AMD-producers, pyrite is more abundant and is thus commonly credited for
AMD generation. Pyrite occurs in several forms, based upon the geological
conditions at the time of formation. The framboidal form is the prolific
acid producer and is characterized by a cluster of spheres of agglomerated
minute pyrite crystals, approximately 25 ym (micrometers) diameter.

Iron disulfides (both pyrite and marcasite) react with water and oxygen to
form ferrous sulfate and to release 2 mol (moles) of hydrogen ions (acid)
(1).

7 +2 -2 +
FeS2 +2 02 + H20 = Fe + 2S04 + 2H (1)

The ferrous iron will eventually use 1 mol of H and oxidize to the more
stable ferric form according to:

Fe+2+ %02 + H+ = Fe +3 + %H20 (2)

Above approximately pH 4.0, the ferric ion will hydrolyze (take on water) and
precipitate, freeing 3 mol of H+.

Fe+3+ 3H20 = Fe(OH)3 + 3H+ (3)

Overall, 1 mol of pyrite will produce 4 mol of hydrogen ions: 2 mol from the
initial oxidation of pyrite and 2 mol (net) from the combined iron oxidation
of ferrous to ferric and the subsequent hydrolysis to ferric hydroxide. The
4 mol of H+ are equivalent to 2 mol of sulfuric acid
Thus, the acidity in the AMD entering a treatment plant is developed not only
from the pyrite but also from the hydrolysis of the metals as they oxidize,
hydrolyze, and precipitate. This fact is important to treatment plant design
and operation because it governs the amount of alkali ultimately required for
neutralization. For example, Equations 2 and 3 indicate that a net 2 mol of
H+ (equivalent to 1 mol of H^OO are produced from oxidizing and precipitat-
ing 1 mol of ferrous iron. Using an inlet ferrous iron concentration of 100
mg/1, 62 mg/1 (0.6 lb/1,000 gal) of hydrated lime would ideally be required
to neutralize the additional acidity from the ferrous iron oxidation/precipi-
tation. Analytical methods for determining acidity in AMD should, and nor-
mally do, include the addition of a strong oxidant such as hydrogen peroxide
to oxidize all the metals, thus producing a measurement of total acidity,
which includes acidity resulting from oxidation of all metals and the
hydrolysis/precipitation of those metals insoluble at the end point of the
acidity titration (normally between pH 7.3 and 8.3).
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This total acidity value is normally expressed in terms of milligrams per
liter (mg/1) of calcium carbonate (CaC03), meaning ideally that if the number
of milligrams of CaC03 specified were added to 1 1 of the AMD, the pH of the
AMD would ultimately increase to the titration end point (normally between pH
7.3 and 8.3). It is important to recognize that the operator may, and prob-
ably will, operate the system at a different pH level than the end point of
the acidity titration; thus, the theoretical alkali requirement for treatment
may differ from that predicted by the acidity determinations.

1.3 Acidity

Acidity is a major factor to be considered in system design because it
strongly influences the choice of alkali. To calculate the theoretical
"ideal" amount of alkali required to neutralize a given acidity concentra-
tion, the alkalis must be converted to an equivalent basis; i.e., expressed
as calcium carbonate. The equivalent weight is the molecular weight divided
by the valence of the dissociated ions. For example, the molecular weight of
CaC03 is 100.09 and the valence of Ca is 2; therefore, the equivalent weight
is 100.09 T 2 = 50 (rounded). Similarly, sodium hydroxide (NaOH) has a
molecular weight of 39.99, the valence of Na is 1, and the equivalent weight
is 39.99 * 1 = 40. Soda ash (Na2C03) dissociates into 2(Na+1) and (COJ2;
thus, the equivalent weight is 105.99 v 2 = 53. The lower the equivalent
weight, the more powerful the alkali, because less reagent is necessary to
provide an equivalent neutralization capability. For example, if 50 mg of
CaC03 were required to neutralize 1 1 of AMD with an acidity of 50 mg/1, only
28 mg of quicklime or CaO (see Table 1-2) would be necessary to provide an
equal amount of neutralization.

Once the ideal alkali requirement has been determined, it is necessary to
compensate for the generally inefficient utilization of the reagent in the
actual treatment process. For example, hydrated lime utilization efficien-
cies are typically near 70%, thus requiring 1.4 times the ideal theoretical
amount to accomplish neutralization; limestone efficiencies are below 50%,
thus requiring over two times the ideal quantity; and, sodium hydroxide
efficiencies are above 90%, requiring only 1.1 times the theoretical reagent
quantity.

1.4 Ion Solubility and pH

The metal ions normally present in AMD are typically relatively insoluble in
alkaline environments, and therefore can be precipitated as hydroxides by in-
creasing the pH. The theoretical solubilities, determined by measurements of
each individual ion dissolved in distilled water, are illustrated in Figure
1-1. Actual mine waters involve complex interactions and result in shifts of
the curves shown in Figure 1-1; however, the general trends remain the same.
Manganese, for example, can generally be precipitated at pH's slightly above
8.0, probably because of coprecipitation with iron. In the infrequent situa-
tion that manganese cannot be removed within the pH 6.0-9.0 requirements of
the New-Source Performance Standards (Table 1-1), the regulations allow ele-
vating the pH slightly above 9.0 to achieve satisfactory manganese removal.
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TABLE 1-2
ALKALI COMPARISON FOR
TREATMENT OF ACID MINE DRAINAGE
A1 kali
Molecular Equivalent
Formula wt wt
Factor to
Convert to
CaCO
Equivalence
Calcium neutralizers
Hydrated lime
(calcium hydroxide)
Quicklime
(calcium oxide)
Limestone
(calcium carbonate)
Magnesium neutralizers
Dolomitic lime
(magnesium hydroxide)
Sodium neutralizers
Caustic soda
(sodium hydroxide)
Soda ash
(sodium carbonate)
Ca(OH)2 74.10

CaO 56.08

CaC03 100.08
Mg(OH), 58.30 29.15
NaOH
39.99
37.05
28.04
50.04
1.35
1.78
1.00
Na2C03 105.99
39.99

53.00
1.72
1.25

0.94
1.5 References
1. Singer, P.C., and W. Stumm. Oxygenation of Ferrous Iron. Federal Water
Pollution Control Administration Research Series 14010, Cincinnati,
Ohio, June 1969.
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CONCENTRATION, mg/l
fD
3-
fD
O

fD
S
Irt
O
cn
TJ
X
)
fD
10
fD

fD
O
ft
fD
Q.
ro
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CHAPTER 2

GENERAL TREATMENT CONSIDERATIONS

2.1 Acid Mine Drainage Treatment Systems

As previously described, AMD is a dilute solution of sulfuric acid and iron
sulfate with iron in the ferrous and/or ferric form. Its treatment consists
of neutralization with a suitable alkali, oxidation of ferrous iron to the
insoluble ferric form, and removal of the resulting metal precipitants by a
sedimentation process. There is one basic process system for the treatment
of AMD; however, there are many options available to the designer when evalu-
ating each of the unit processes or subprocesses within the overall system.
Figure 2-1 shows many of these options within the overall process flow sheet.
Consequently, this manual has been developed in a format that discusses each
unit operation (chemical feeding, mixing, sedimentation) of the process sepa-
rately. A discussion of the various overall processes follows.

2.1.1 Conventional Lime Neutralization Process (2, 3, 4)

In the conventional lime neutralization process, each of the five basic
treatment steps follows in normal sequence; i.e., equalization, neutraliza-
tion (mixing), aeration, sedimentation, and sludge disposal. Flow is once-
through and gravity systems are usually employed. A flow sheet for the typi-
cal system is shown in Figure 2-2.

To simplify the controls needed in the system and to minimize operator at-
tendance, a constant flow with only small variations in quality is desirable.
To accomplish this goal, the mine drainage is collected in large holding or
equalization basins, or in large sumps withirt active portions of the mine.
Such holding basins should have a capacity for storage of 2-3 days flow
during shutdown periods. Normally, 12-24 hours flow is maintained in the
holding basin to equalize flow and quality to the treatment facility. From
the holding basin, the mine drainage either flows by gravity or is pumped to
the treatment plant. Since most mines are in rural areas, both the holding
and settling basins are usually surface impoundments of earthen construction.
(Earthen pond design is discussed in Chapter 6.)

Lime is used as the alkali in practically all large-volume AMD treatment
plants. The selection between quicklime and the hydrate is determined by
availability, cost, or personal preference. This process is discussed in
detail in Chapter 3, as are the several methods by which the lime can be fed
into the mixing or neutralization unit. The use of other alkalis is also
discussed in Chapter 3.
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ACID MINE DRAINAGE
COLLECTION/STORAGE
1. IN MINE
2. SURFACE IMPOUNDMENT
ALKALI SELECTION
1. QUICKLIME
2. HYDRATED LIME
3. LIMESTONE
4. SODA ASH
5. CAUSTIC SODA
ALKALI STORAGE
AND FEEDING
1. DIRECT FEED
2. SOLUTION FEED
3. SLURRY FEED
MIXING
1. MECHANICAL
2.TURBULENT
3. NATURAL DISSOLUTION
oo
ROUTINE PROCESS
OPTIONAL
IRON OXIDATION
1. AERATION
2. CHEMICAL OXIDANTS
3. BIOLOGICAL OXIDATION

COAGULANT
ADDITION
I
SEDIMENTATION
1. CLARIFIERS
2. SEPARATORS
3. SETTLING PONDS
4. IMPOUNDMENTS
EFFLUENT
SLUDGE DISPOSAL
1. DEEP MINE
2. LAGOONING
3. FILTRATION
4. DRYING
5. CENTRIFUGATION
Figure 2-1. Process flow sheet for treatment of acid mine drainage.
-------
MINE
PUMP
LIME SLURW SYSTEM
SLUDGE REQUIRING
DISPOSAL .
DISCHARGE
Figure 2-2. Conventional lime neutralization process.
Aeration is a straightforward process for oxidizing ferrous iron to the less
soluble ferric form. Ferrous iron is much more soluble than the ferric form,
with minimum solubility occurring in the pH range of 9.3-12.0 (1). Ferric
iron, on the other hand, is much less soluble and begins to precipitate as a
hydroxide at a pH of 4.0, with minimum solubility occurring about pH 8.0.
Obviously there is an economic advantage in removing iron in the ferric form
at the lower pH. Less lime is required for neutralization to the pH level
needed to maintain minimum iron solubility (8.0 vs. 12.0).

The forced oxidation of ferrous iron is usually included in AMD treatment.
This oxidation is pH-dependent, with the reaction proceeding rapidly at a pH
above 8.0. With the pH requirement satisified, iron oxidation becomes de-
pendent upon the availability of oxygen. This oxidation reaction is ex-
pressed by Equation 2, previously discussed in Chapter 1. The theoretical
oxygen requirement is one unit weight for each seven weights of ferrous iron
to be oxidized.

It is important to point out that most existing aeration units do not have
sufficient hydraulic detention capacity; this may be one major cause of
effluent compliance problems with iron. Chapter 5 discusses the design re-
quirements for iron oxidation and the aeration equipment available to accom-
plish this task. Other methods for iron oxidation are also included in Chap-
ter 5. It is worth noting that in a few cases there are advantages in pre-
-------
cipitating iron as ferrous hydroxide because of peculiarities in water chem-
istry.

Once the drainage has been neutralized and the ferrous iron oxidized, if
necessary, the subsequent step in the treatment process is sedimentation.
Settling of the iron hydroxide and other suspended solids is commonly accom-
plished in earthen settling basins. These must have at least 12 hours of
clear water detention above the sludge storage zone to meet minimum design
requirements in Pennsylvania (5). Other design considerations are discussed
in Chapter 6. When using small settling basins with 12-48 hours detention
and minimal sludge storage capacity, two units operated in parallel are
highly recommended to allow sufficient time for sludge removal without inter-
rupting the treatment process. If treatment plant site conditions allow,
large impoundments that provide many years of sludge storage can be advanta-
geous. Chapter 6 includes design information for other sedimentation units;
e.g., mechanical clarifiers, thickeners, and tilted-plate separators.

A necessary part of the treatment process is the need for adequate planning
for sludge handling and disposal. This will be a significant part of both
the construction and operating costs of the system. Without proper planning
and process selection in this area, day-to-day treatment plant operation can
become a considerable and overly expensive problem.

Even though many mine drainage treatment facilities have been in operation
for 10 years or so, the handling and disposal of the sludge produced contin-
ues to be a problem. The simplest method for final disposal is to pump the
sludge into abandoned deep mines. Although this practice is widespread, the
overall environmental effects of this disposal method have yet to be deter-
mined. Unfortunately, this practice cannot be used at all facilities.
Another method that has been used successfully is lagooning, where the sludge
thickens naturally. Eventually the sludge must be disposed of in a more sat-
isfactory manner, such as burial in a surface mine reclamation project.

There are other methods for sludge dewatering that are used infrequently but
can be considered by the designers. Among these are drying beds, vacuum and
pressure filtration, and centrifugation. Each of these is discussed in Chap-
ter 7, Sludge Dewatering and Disposal.

2.1.2 High-Density Sludge Process (4, 6)

Variations on the conventional lime neutralization process previously dis-
cussed are the sludge recirculation processes that can .be utilized to achieve
better reactivity of the lime and produce smaller volumes of sludge contain-
ing higher solids. One such procedure is the High-Density Sludge Process
developed in 1970 by the Bethlehem Steel Corporation. This process uses lime
for neutralization and can produce a dense sludge that reduces its volume
significantly more than the conventional process. The process is based on a
high sludge recirculation rate within the system, where the optimum ratio of
solids recirculated to solids removed is in the range of 20:1 to 30:1. The
sludge is returned to a reactor vessel where the lime slurry is added. This
point of alkali introduction is peculiar to the Bethlehem system. The slurry

10
-------
is then mixed with the AMD in a neutralization reactor, where aeration is
provided for oxidation of ferrous iron. The process flow sheet is shown in
Figure 2-3.
LIME

STORAGE
WATER
SLUDGE
REACTION
•AMD
NEUTRAL
EFFLUENT
—\
NEUTRALIZATION
AND OXIDATION
AIR-
SOLIDS-LIQUID K
SEPARATION /
RECYCLE SLUDGE
WASTE SLUDGE
15-40% SOLIDS
Figure 2-3. High-density sludge treatment process.
Removal of the solids is accomplished in large mechanical thickeners. The
achievement of high sludge solids, reported by Bethlehem to be as great as
50%, is dependent upon the ferrous-to-ferric iron ratio. If ferric iron
dominates this ratio, sludge densities may be limited to 20% solids. The
user is cautioned that certain aspects of this system may be covered by
patents issued to the Bethlehem Steel Corporation, but other variations on
the sludge recirculation processes are common practice in AMD treatment.
2.1.3 Other Treatment Processes

There can be numerous variations on the conventional process for treatment of
AMD. Where acidity is the main problem and the flow is low, other alkalis
such as soda ash, caustic soda, or limestone can be used. Portable caustic
soda treatment units are common in surface mine operations. Limestone has
been used in several applications for in-place treatment. These applications
are discussed in more detail in Chapter 3.

In addition to the treatment of AMD to achieve a desired effluent quality for

11
-------
discharge, other methods of treatment are available to produce a product
water of higher quality. These include reverse osmosis, ion exchange, and
chemical softening; discussions of these methods are presented in Chapters 9,
10, and 11, respectively. Each of these processes can possibly be utilized
to produce water acceptable for human consumption.

2.2 References

1. Singer, P.C., and W. Stumm. Oxygenation of Ferrous Iron. Federal Water
Pollution Control Administration Research Series 14010, 1969.

2. Dorr Oliver, Inc. Operation Yellowboy-Mine Drainage Treatment Plants
and Cost Evaluation. Report to the Pennsylvania Department of Mines and
Mineral Industries, Coal Research Board, 1966.

3. Holland, C.T., J.L. Corsaro, and D.O. Ladish. Factors in The Design of
an Acid Mine Drainage Treatment Plant. Second Symposium on Coal Mine
Drainage Research, Mellon Institute, Pittsburgh, Pennsylvania, 1968.

4. Skelly and Loy and Penn Environmental Consultants, Inc. Processes,
Procedures, and Methods to Control Pollution from Mining Activities.
EPA-430/9-73-011, Washington, D.C., 1973.

5. Bureau of Water Quality Management. Mine Drainage Manual. 2nd ed.
Department of Environmental Resources, Publication No. 12, Harrisburg,
Pennsylvania, September 1973.

6. Haines, G.F., and P.O. Kostenbader. High-Density Sludge Process for
Treating Acid Mine Drainage. Third Symposium on Coal Mine Drainage
Research, Mellon Institute, Pittsburgh, Pennsylvania, 1970.
12
-------
CHAPTER 3

CHEMICAL TREATMENT
3.1 Introduction

The chemical treatment of acid mine drainage has involved the use of practi-
cally every available neutralizing agent in either pilot or full-scale opera-
tion. Information concerning the application and results associated with the
use of these many reagents is beyond the scope of this manual. Only the more
practical and commonly used alkalis are included. Chemical and physical data
for lime and its solutions are in Appendix A. Presented in the following
chapter are cost comparisons and descriptions of the process systems for the
handling, storage, and feeding of each alkali. Those alkalis included are
quicklime, hydrated lime, limestone, caustic soda, and soda ash.

More than 90% of all facilities treating acid mine drainage utilize a form of
lime. The main factors affecting the selection of any alkali are cost, suit-
ability, reactivity, availability, ease of use, and sludge volume. It is
important for the designer to evaluate these factors fully because each
alkali requires significantly different equipment, which limits further
changeover to another type. In addition, the alkali selected may affect the
design of other processes in the overall system.

3.2 Lime

Lime is a general term that, by definition, encompasses only burned forms of
limestone. The two forms of particular interest in AMD treatment are quick-
lime and hydrated lime. Carbonates, such as limestone, are frequently but
erroneously referred to as "lime." For this and other reasons, limestone is
described separately.

3.2.1 Quicklime

Quicklime (CaO) is a product resulting from the calcination of limestone.
Limestone basically consists of 50%-90% calcium carbonate (CaCOs). When
limestone is burned in a kiln at a temperature of about 1,000°C (1,835°F),
carbon dioxide (COa) gas is driven off and calcium oxide (CaO) or quicklime
is produced. On the basis of its chemical analysis, quicklime may be divided
into three classes:

1. high calcium quicklime - containing less than 5% magnesium oxide;

2. magnesium quicklime - containing 5%-35% magnesium oxide; and

13
-------
3. dolomitic quicklime - containing 35%-40% magnesium oxide.

Quicklime is available in a number of standard sizes, but the following
describes those most applicable to mine drainage treatment:

Ground lime - the product resulting from grinding the larger sized
material and/or screening off the fine size. A typical size is the
majority passing a #8 sieve and 40%-60% passing a #100 sieve.

Pulverized lime - the product resulting from a more intense grinding
than is used to produce ground lime. A typical size is the majority
passing a #20 sieve and 851-95% passing a #100 sieve.

Quicklime is most often obtained in either bulk carloads or 18.2-ffg (20-ton)
pneumatic trucks, and then transferred to a storage silo. To be used effi-
ciently in mine drainage neutralization, the quicklime must be slaked (see
3.2.7, Quicklime Slaking Systems). The slaking process must be carefully
controlled and requires daily attention. Quicklime, combined with a good
slaking operation, offers a low unit cost per gram of acidity neutralized.

The primary disadvantages of a quicklime system are high capital investment
for a slaker, grit removal, close operational control, and the danger to
personnel of possible severe burns.

3.2.2 Hydrated Lime

Hydrated lime (Ca(OH)2) is the most commonly used alkali for neutralizing
acid mine drainage in existing treatment plants. It is preferred when lime
consumption rates are low or when the cost for a slaking system is prohibi-
tive.

Hydrated lime is air-classified to produce the fineness necessary to meet
user requirements. Normal grades of hydrate used for chemical purposes will
have 75%-95% passing a #200 sieve. Due to air classification, the commercial
hydrate produced is purer than the quicklime because most of the impurities
are rejected.

Hydrated lime is packaged in paper bags weighing 22.7 kg (50 Ib) net. It is
also available in bulk. Physical and chemical properties of hydrated lime
can be found at the end of this chapter.

3.2.3 Lime Handling

Hydrated lime can be purchased in various quantities. Where daily require-
ments are small, less than 138 kg (300 Ib) bagged lime may be preferred. The
handling and storage operations are relatively simple, usually involving man-
ual labor or mechanical equipment, depending on the volume involved.

Most often bulk lime, either quick or hydrated, is more efficient and econom-
ical to use. In such installations, the lime is delivered by truck and con-

14
-------
veyed by mechanical or pneumatic s
The factors determining the type of
ics of this chapter.
stems into weathertight bins or silos.
lime to use are discussed in the econom-
3.2.3.1 Bagged Lime

Bagged lime is delivered loose or palletized in truck or box car, and gener-
ally handled by hand truck or forklift to storage. Unloading conveyors may
be preferred for loose bags, particularly if there is a long distance between
r - .. ., . . i . •- _1T_i_« __!_!_• i C'.-l.l'.PJ.
the unloading point and the storage
area. For palletized shipments, forklift
trucks are utilized to move the linje to storage or point of use, thus elim-
inating much manual labor. To facilitate bag dumping, the hopper located
above the feeder can be fitted with

Bagged lime should be stored in dry
aged in multiwall paper bags, but
a bag ripper and screen.

areas. Hydrated lime is normally pack-
exposure to moisture will permeate the
liner and cause caking. Bags of hydrated lime may be stacked 20 high without
damaging the bottom bags. In dry storage, hydrate may be stored for periods
of up to 1 year without encountering serious deterioration. Care should be
exercised to use the material in the order it is received, rather than main-
taining an inactive reserve that may not be consumed for several months or
years. I

3.2.3.2 Bulk Lime

Considerable savings can be realized by using bulk lime instead of bagged
lime, not only in initial cost but
ination of losses from broken bags
also in reduced labor involved in handl-
ing. In additon, there are other advantages including faster loading, elim-
and spillage, better housekeeping because
modern handling systems are completely enclosed, and less dust hazard to
employees.

Delivery of bulk lime to the treatment plant can be accomplished by a variety
of truck or rail-car equipment. Truck transportation, particularly the
blower truck, is generally the fastest, most common, and most economical way
to handle bulk shipments.
i
The pneumatic truck has become popular because of its simplicity and speed of
delivery and unloading. The lime lis blown from the truck directly to silo
storage via a 10-cm (4-in) pipeline that eliminates mechanical conveyors.
The only extra equipment required is a safety release valve and dust collec-
tor mounted atop the silo to exhaust the conveying air. The release valve is
important to accommodate the large [volume of air, up to 36.4m (1,300 ft ),
in the blower truck when the tank is empty. The valve can be a hinged man-
hole cover or a simple 20-cm (8-
in) pipe with weighted, gasketed cover.
Today, the bag-type dust collector'
truck unloading. For hydrated lime
0.06 m3 (2.0 ft3) of air is recommended
cloth area are needed for a large
is used to filter air expelled during
normally 0.09 m2 (1.0 ft2) cloth area/
Approximately 34.8 m2 (375 ft2) of
tJruck rotary blower of 21 m3 (750 ft3)/min
15
-------
capacity. Additional cloth area of 9.3-27.9 m2 (100-300 ft2) may be justi-
fied, however, to accommodate the final cleanout period. These recommenda-
tions are conservative, and more commonly, vendors use an air-to-cloth ratio
of 3:1. If more extensive dust collection is required, the design should
tend to be conservative.

Blower trucks are available with compartment tank capacity varying from 19.6
to 36.4 m3 (700-1,300 ft3). (The latter delivers up to 18.2 kkg (20 tons) of
hydrate and 21.8 kkg (24 tons) of pebble lime.) Air is provided by a
trailer-mounted positive displacement rotary blower, which furnishes up to 21
m3 (750 ft3) of air/min. It is operated from a power takeoff or by a sepa-
rate engine.

The pipe used for transferring lime to storage can be ordinary black iron
pipe, but galvanized pipe is highly recommended because of its resistance to
rust. All bends should be made with a minimum radius of 0.9-1.2 m (3-4 ft)
to reduce wear and resistance to flow. The intake end of the pipe should be
mounted vertically, with the bottom 1.2 m (4 ft) above ground level, and
equipped with a quick-connect coupling for the rubber blowing hose on the
truck.

The largest size of pebble lime that can be pumped efficiently from blower
trucks is 3.2 cm (1.25 in), although a top size of 2.5 cm (1.0 in) is pre-
ferred. Pebble lime may be blown as much as 30.5 m (100 ft) vertically and
45.7 m (150 ft) in a combined vertical and horizontal run. For greater dis-
tances, the unloading time becomes excessively long.

Hydrated lime, however, can be blown readily to 91.4 m (300 ft) in a combined
vertical and horizontal run. If lime has to be blown both vertically and
horizontally, it is much easier to blow vertically first, then horizontally,
rather than vice versa. Otherwise, there will be excessive wear at the 90°
elbow because of the resistance provided by the column of lime in the verti-
cal pipe section, and unloading will be delayed. It is advisable to use only
one 90° turn in the piping system, or none if possible.

When hauled by rail, most bulk lime is shipped in covered hopper cars, the
largest capable of hauling 90 kkg (100 tons) of quicklime or 45 kkg (50 tons)
of hydra ted lime. These cars have two to four compartments, each with a
bottom discharge gate. Generally, the lime is discharged to an under-track
hopper, then taken by screw conveyor or bucket elevator to plant storage, or
an adapter is fastened to the discharge gate by clamps for pneumatic unload-
ing. Air vibrators are usually attached to the hopper bottoms to facilitate
unloading, particularly for hydrate. For protection against rain when not in
use, the under-track hopper should be covered.

3.2.4 Storage of Lime

Since quicklime and hydrated lime are not corrosive, conventional steel or
concrete bins and silos can be used for storage. The storage units, however,
must be watertight. Of all the storage units used for lime, the steel silo
with cone bottom is the most popular.

16
-------
The variety of bin or silo designs for lime storage includes rectangular,
square, hexagonal, and circular. The first three occupy less plant space be-
cause they utilize common walls and are generally easy to clean, but they
have the disadvantage of causing material retention in the corners. With
round silos, this accumulation or bridging is uncommon, although there is
greater tendency for lime to arch than in rectangular bins. In either case,
the storage units should always be designed with a hopper or conical base to
facilitate discharge.

The decision to install one or more large storage units versus several small
ones will depend on the individual plant and on such factors as daily lime
requirements, type of delivery (rail or truck), and operation (continuous or
intermittent). In any event, the total storage capacity should be at least
twice the minimum truck or rail delivery to guarantee a lime supply while
awaiting an order. Most mine drainage treatment plants have a steady lime
demand and operate continuously. It may be prudent to provide at least 7
days storage capacity, and preferably 2-3 weeks.

The importance of ample storage capacity cannot be overemphasized, especially
when it is the cheapest cost per unit volume in the overall treatment system.
Then, too, by having large storage capacity, future plant expansion can be
readily accommodated.

In designing a specific size of silo or bin, it is advisable to use a bulk
density of 481 kg/m3 (30 lb/ft3) for hydrated lime, and 881 kg/m3 (55 lb/ft3)
for quicklime. These density figures are subject to variation, depending on
lime source, type of lime, particle size (pebble vs. pulverized), and grada-
tion. With both products, these values are conservative; hence, bins will
have additional capacity for denser limes. Figures 3-1 and 3-2 show typical
silo installations.

3.2.5 Flowability

The flowability of lime varies from good, for pebble and granulated quick-
lime, to erratic, for pulverized quicklime and hydrated lime. Lime tends to
absorb moisture readily, forming an adherent soft cake that can cause arch-
ing or bridging in storage. Hydrated lime, in particular, also tends to
"rathole," because of its fluffy texture and possibly electrostatic charges.
Then, after collapsing, the hydrate may become fluidized and flood the dis-
charge.

Because of the inherent problems with lime flowability, several units have
been developed to provide a uniform dens.ity from storage to the feeder.
These include special considerations in bin construction, the use of external
vibrators, internal antipacking and antiarching devices, and live bin bot-
toms.

There are various opinions as to the best bin or silo shape and height-to-
diameter ratio for facilitating flow. It appears that a tall, slender struc-
ture is preferred to the short, squat one, with a height-to-diameter ratio in
the range of 2.5:1 to 4:1 the most desirable. Good flowability is promoted

17
-------
DUST FILTER
LADDER a CAGE-
TRANSFER
PLATFOR
LADDER 8 CAGE
SLURRY HOPPER-
RAILING
STORAGE BIN
BIN SIGNAL
DUST a
VAPOR
REMOVER
j^ LIME SLAKER
• BIN ACTIVATOR
(OPTIONAL)

-VOLUMETRIC OR WEIGH BELT

FEEDER

-GRIT SCREEN
PANEL

-CONTROL PANEL

-SLURRY PUMP

Figure 3-1. Typical quicklime and slaker installation.

18
-------
DUST FILTER^

LADDER a CAGE -

,::t

\r
TRANSFER PLATFORM^
,_
LADDER a CAGE Hi
SLURRY HOPPER -
SERVICE PANEL —
^xsr

ii
ii
. ii
i
i
i
i
i
I
i
i
i
1
i
i
i
i
ii «,„

.i7\
• i V

. v\
-
1
1
:
= J r** —
. M-l
4
F -

— BLOWER
S

STORAGE BIN
BIN SIGNAL ^_
TD

\
\
i-r3

\
^tz
*"i

i
*

s^.j
7
_n.
fe
-n--

K
1-
~-l
1
1
"-*.
/
/
/
/
/
/
/
/
it
1

£EP
^
- — -_
r~TL
i i !
^-—RAILING

f BIN ACTIVATOR CONE
oCRtW FEEDER
FILL PIPE
-CONTROL PANEL
SLURRY PUMP
'" :-.*'• •oS'.iV'.'c'.'o-'•*'•''*'"'•' ;° ••'.'•"!•''""I
Figure 3-2. Hydrated lime and slurry installation.
19
-------
by having the discharge area as large as possible in relation to bin cross-
sectional area.

With quicklime or hydrate, it is advisable for hopper bottoms to have a mini-
mum slope of 60° from the horizontal. One way of accomplishing a steeper
slope is to use an offset hopper (Figure 3-3), which serves to increase flow-
ability by decreasing the weight of material going through the opening.

The need for such a steep slope is based on the relatively high angles of
repose of 50°-55° (average) for quicklime and 15°-80° for hydrated lime. In
the case of quicklime, the angle of repose is affected by the particle size,
shape, and gradation (particularly percent of fines), and type of lime. The
wide variation in angle of repose for hydrated lime is the result of the
variation in particle size and gradation, moisture content, degree of aera-
tion, or presence of electrostatic charges. The lower value of 15° would
apply to highly aerated hydrate, which literally flows like water; in con-
trast, high moisture content and electrostatic charges produce the 80° value.

Generally, with the recommended 60° slope, quicklime will flow readily with-
out the need for external aids, whereas hydrated lime and pulverized quick-
lime will require one or more means to promote uniform discharge from bins.

3.2.5.1 Vibrators or Bin Activators

The simplest device for improving flowability is an electromagnetic vibrator
attached to the outside of the hopper face. This type is more suitable for
quicklime than hydrate. Vibrators, however, can be used for hydrate if the
unit is cycled to produce 1-2 seconds of vibration every 5-10 seconds. With
quicklime, the vibrator can be operated continuously during discharge. For
best results, the vibrator should be bolted directly to the conical hopper
face, one-fourth or less of the distance from the discharge to the top of the
cone. Vibrators should be operated only while the hopper is open to flow;
this prevents packing. If noise is a problem, the electromechanical vibrator
is recommended over the electromagnetic unit. Both types are available in a
variety of sizes to fit the individual hopper size and metal plate thickness.

Air pads and jets are other means of inducing material flow. They are not,
however, recommended for lime. They are best suited for highly dense materi-
als such as iron ore. Vibratory bin activators are usually preferred over
air pads for the lime systems used in treating acid mine drainage. The pri-
mary problem with air pads is that they reduce the bulk density of the lime,
which interferes with feeder accuracy and, more importantly, causes flooding.

3.2.5.2 Live Bin Bottoms

Several companies manufacture "live bin bottoms," which are fitted to new or
existing bin structures. These units are vibrated continuously during un-
loading by a gyrator exerting a horizontal force of up to 18,160 kg (40,000
Ib). This promotes a steady flow of material with a uniform density to the
feeder.

20
-------
DUST COLLECTOR

1.2m (4ft)
M1N. RAD.—/
10.2 cm
(4 in) PIPE
60 CONE
QUICK
COUPLING-
HIGH LEVEL
INDICATOR
DUST COLUECTO

90°BEND .
1.2m (4ft) MIN.RAD.
10.2 cm (4 in) PIPE-
LOW LEVEL
INDICATOR
VIBRATOR
ORATOR
QUICK COUPLING -
/SAFETY VALVE
UGH LEVEL
INDICATOR
IOW LEVEL
t±
TO FEEDER |
Figure 3-3. Standard and offset hopper bottoms.
-------
A device called the "bin activator" incorporates an internal dome-shaped
baffle plate above the discharge, which exerts a 45,400-kg (100,000-1b)
vertical thrust into the bin. Vibrations from the baffle plate penetrate the
overlying material, forcing it to move freely to the periphery of the plate,
then down to the discharge. Bridging and ratholing are virtually eliminated.
The bin activator is available in 0.6- to 3.0-m (2- to 10-ft) diameter sizes
and is usually sized one-half the total diameter of a circular silo.

3.2.6 Lime Feeders

Existing AMD treatment processes utilize dry lime in a liquid suspension or
slurry before introducing it into the raw water. Today, however, designers
are beginning to feed dry hydrated lime directly to the acid water. This
eliminates a slurry feed system. Dry lime feed has been used when the drain-
age streams are small and mildly acidic. These conditions require little
lime (less than 0.1 kg/1,000 1), and past designers have installed such sys-
tems because of economics. The primary reason is that dry hydrated lime can
be employed this way with an acceptable degree of efficiency.

For mine drainage requiring larger amounts of lime per unit flow, 0.36-0.48
kg/1,000 1 (3-4 lb/1,000 gal), special design considerations must be pro-
vided. An example would be a larger aeration or flash mixer to insure com-
plete mix and utilization of the lime. This method of lime feeding is fur-
ther explained in Chapter 12.

The lime feeder selection for a treatment plant depends largely on the type
and size of lime specified and daily lime requirements (generally figured in
kg/hr (Ib/hr) or m3/hr (ft3/hr)). Regardless of whether a volumetric or
gravimetric feeder is used, one should be selected that provides sufficient
flexibility, protection against exposure to lime dust, low maintenance, few
moving parts, and ease of cleaning. Some feeders are designed especially for
granular materials such as pebble lime, others for powdery materials such as
hydrate.

A machine that requires frequent maintenance is undesirable in an efficient,
continuous, automated process. Easily cloggable mechanisms, such as plungers
and plates using small orifices or clearances, have been replaced by the
oscillating hopper, belt conveyor, and vibrating feeder. These usually have
large clearances at all points, so larger grains and small pieces of paper or
wire pass the machines. Many operators take the precaution of placing a
screen across the hopper opening to remove paper and wire during loading.
This reduces the chances of blockage, and helps to maintain a dry hydrated
lime of constant density within the feeder hopper.

All feeders should be capable of confining dust for the sake of cleanliness,
health, economy, and compliance with increasingly stringent air pollution and
OSHA regulations. Any feeding mechanism that cannot be easily and conve-
niently enclosed may create a dust problem. Fortunately, most feeds are de-
signed with this in mind. Consideration also should be given to necessary
shutdowns for repair, maintenance, and adjustments to the feed. Moving parts
of mechanical equipment must be lubricated regularly, and electrical equip-

22
-------
ment must be replaced from time to time.

3.2.6.1 Dry Feeders

The efficiency of an acid mine drainage treatment plant depends on the speed
and accuracy with which the lime is handled and fed to the process. To main-
tain the proper function of the neutralization process, it is essential that
an uninterrupted flow of lime be maintained upon demand. The importance of
operating lime feeders properly cannot be overly stressed.

The dry lime feeder consists of two parts:

1. a feeder hopper, usually with a throat at the bottom through which
the lime falls by gravity;

2. a feeding element (i.e., screw) that can be adjusted to give differ-
ent rates of material. The feed may be by volume or by weight;
i.e., volumetric or gravimetric. A volumetric feeder will deliver a
constant volume, but the weight may vary.

With dry feeders, the decision to use volumetric or gravimetric will depend
upon desired accuracy and economics. The accuracy of gravimetric feeders is
not needed for treating acid mine drainage. Volumetric feeders, which may
have an error of 7%-15% by weight, are more than sufficient. This error is
dependent upon the flowability and uniform density of the lime. In most lime
feeding systems, the lime requirements are controlled more by the method of
feeding the solution than by the accuracy of the feeder itself. This is
probably the reason for the popularity of volumetric feeders in AMD plants.

The type of lime used in the feeder is an important factor. A feeder that
may introduce pebble lime with accuracy may not do so with hydrated lime:
hydrated lime is more difficult to feed.

a. Feeder Hopper

A dry feeder will operate only as well as the hopper that charges material to
it. Therefore, it is important that the material flows freely and uniformly
from the hopper. Several techniques have already been discussed under bulk
silo storage. The following discussion presents additional information use-
ful to the designer and treatment plant operator.

Most feeders are equipped with a standard hopper, the size of which depends
on the capacity of the feeder. Feeder hoppers are universally designed so
that the slope angle of the hopper bottom is greater than the angle of repose
of the lime to be fed from it. The slope angle, although varying slightly
from model to model, is usually about 60°. Hoppers for feeders are usually
conical or rectangular (see Figure 3-4 A). No matter what shape, hydrated
lime may arch or bridge.

When the material arches in an unagitated hopper, the operator is powerless

23
-------
to do anything about it except resort to such primitive methods as pounding
the side of the hopper, or going on top of the hopper and agitating the mass
with a stick or paddle. None are effective or acceptable methods. Measures
should be taken to prevent the arching from ever occurring, although no
method is absolutely foolproof.

b. Volumetric Feeders

As mentioned earlier, volumetric feeders supply a constant, preset delivery
of material by volume and do not recognize changes in material density. Con-
sequently, this type of feeder must be calibrated by trial and error at the
outset, then readjusted periodically if the lime changes in density.

There are more than a dozen types of volumetric feeders on the market. Only
five, however, are suited to feeding lime in the quantities required for
treatment of acid mine drainage. Of these, the screw-type conveyors are most
commonly used for pebble lime, and the rotary paddle or star-type feeder for
hydrated lime. The following sections discuss these feeders and present
details on specific models.

c. Screw Feeder

The screw conveyor feeder (Figure 3-4 B) delivers a constant stream of mate-
rial from the hopper. For solid screw-type conveyors, usually constant-pitch
flights are employed, but tapered or variable-pitch flights can also be used.
At the same time, the load on the screw should not be so excessive as to pre-
vent turning. The capacity of a screw feeder can readily be changed by vary-
ing the speed of the shaft. It is highly recommended that a variable-speed
screw feeder be installed. This provides the operator with better control
and flexibility for seasonal variations of lime demand.

A recent development in volumetric screw feeding is the vibrating screw. The
feeder assembly is subjected to continuous, controlled gyratory vibration,
which insures that each flight of the screw conveyor will be filled to the
maximum and that each will be completely emptied at the discharge tube.
There are two models: (1) a heavy-duty feeder, with feed rates from as low
as 0.0078 m3/hr (0.28 ft3/hr) for a 2.5-cm (1-in) screw size to as high as
16.8 m3/hr (600 fts/hr) for the 15.2-cm (6-in) screw; and (2) the live bin
screw feeder, with rates varying from 1.0 x 1Q-1* m 3/hr (0.0037 ft3/hr) for a
0.64-cm (0.25-in) screw to 5.6 m 3/hr (200 ft3/hr) for a 10.2-cm (4-in) screw.
With the latter unit, the entire bin conveyor assembly is subjected to gyra-
tory vibration to insure undiminished flow from hopper to discharge. The
manufacturer claims acceptable feeder accuracy.

d. Oscillating Hopper

The oscillating hopper feeder (Figure 3-5) consists of an oscillating hopper
that swivels on the end of the main hopper. The material completely fills
both hoppers and rests on the tray beneath. As the oscillating hopper moves

24
-------
CO
S-
d)
-a

LO
CM
-------
back and forth, the scraper, which rests on the fixed tray below, is moved
first to the left and then to the right. As it moves, it pushes a ribbon!ike
layer of material off the tray. The capacity is fixed by the length of the
stroke, which may be varied by means of a micrometer screw. Further adjust-
ment is possible by changing the clearance between the hopper and the tray,
which can be raised or lowered. This type of feeder is one of the most wide-
ly used in small installations.

e. Belt Feeder

With a belt feeder, the material enters the feed section from an overhead
hopper, falls on the feed belt, and passes beneath a vertical gate. For a
given belt speed, the position of the gate determines the volume of material
passing through the feeder (see Figure 3-6). With one particular feeder, the
gate is manually positioned by a cam, and the setting indicated by dial and
pointer. A 22.8-cm (9-in) belt feeder is available with a 16.8-m3/hr (600-
ft3/hr) capacity. The feed rate is adjustable over a basic range of 10 to 1
(100 to 1 optional). This particular belt feeder can accommodate pebble lime
of maximum 3.8-cm (1.5-in) size.

f. Rotary Paddle

The rotary paddle feeder is particularly effective for fine materials that
tend to flood, such as hydrated lime. The paddle or vane is located beneath
the hopper discharge, with the feed varied by means of a sliding gate and/or
variable-speed drive for the paddle shaft. The pocket feeder, also called
the star or revolving-door feeder, where the paddle is tightly housed, per-
mits delivery against vacuum or pressure. One such star feeder made espe-
cially for hydrated lime and soda ash incorporates tight-fitting neoprene
blades that control floodable materials. Star feeders work best when coupled
with a live bottom bin.

g. Vibrating Feeder

The vibrating feeder obtains motion by means of an electromagnet anchored to
the feeding trough, which in turn is mounted on flexible leaf springs. The
magnet, energized by a pulsating current, pulls the trough sharply down and
back, then the leaf springs return it up and forward to its original posi-
tion. This action is repeated 3,600 times/min (when operating on 60-cycle
AC), producing a smooth, steady flow of material.

3.2.6.2 Prevention of Feeder Flooding

Feeder flooding is closely related to arching and is caused by a sudden
breaking of an arch or otherwise clogged state. The theories on why this
happens vary; however, one theory states that the material entrains air,
becomes fluidized, and floods (literally gushes), thereby inundating the
feeder. This tendency is much more pronounced in light fluffy materials,

26
-------
OSCILLATING HOPPER
SCRAPER
Figure 3-5. Oscillating hopper feeder.
FEED SECTION
GATE
,FEED BELT
STATIONARY DECK
I
Figure 3-6. Belt feeder.
27
-------
such as hydrated lime, which flow freely and are sometimes difficult to con-
trol. This fact seems to add credence to the entrainment theory; for this
reason, the practice of blowing air into a hopper to break an arch can be
considered an invitation to feeder flooding.

Flooding often occurs when new lime is added to an almost empty hopper. The
effect of the lime dropping through the hopper and probably entraining air
may cause flooding. The silo hopper should always be equipped with either a
bin gate valve or a feeder flood shutoff. Most vendors will supply flood
shutoffs at a slight extra charge. The primary purpose of the slide valve,
which is always positioned between the feeder and bin activator, is to pre-
vent the uncontrolled dispersion of lime when filling an empty silo. Many
times, several feet of lime empty onto the slurry or slaker room floor be-
cause the designer overlooked this item.

3.2.7 Quicklime Slaking Systems

The term slaking refers to the combination of varying proportions of water
and quicklime, which yields a milk of lime (lime slurry) or a viscous lime
paste with some degree of consistency. For maximum efficiency when using
quicklime, it is desirable to slake the lime at or near optimum conditions.
Since many limes have different slaking characteristics, the optimum condi-
tions are usually determined by trial and error. Most lime suppliers can
simplify this determination by providing data or recommendations on the
slaking behavior of their quicklime. The way the quicklime is slaked can
mean the difference between an efficient and a wasteful operation.

Briefly, the variables exerting a profound effect on slaked hydrate quality
are:

1. Reactivity of the quicklime - The reactivity of the quicklime de-
pends upon whether the quicklime is hard-, soft-, or medium-burned.

2. Particle size and gradation of quicklime - No matter what grade of
quicklime is used (lump, pebble, ground, pulverized, or run-of-the-
mill), the finer sizes slake most rapidly because of more available
wetting area.

3. Optimum amount of water - Too much or too little water will slow the
reaction and reduce efficiency.

4. Temperature of water - Slaking water that is too cold or possibly
too hot(steam) for the particular slaking conditions also slows
the reaction.

5. Distribution of water - An even flow of water introduced into the
slaking chamber is highly recommended.

6. Agitation - Too vigorous or insufficient agitation of quicklime and
water will result in a poor and hazardous operation.

28
-------
The heat of the slaking reaction is important (average 88°C (191°F)), and the
reaction can be artificially accelerated by using hot water. By such mea-
sures, it may be possible to increase the slaking rate of a medium reactive
lime to the approximate behavior of a high reactive lime. When applying this
method to a reactive lime, it is possible to obtain extremely rapid, almost
instantaneous slaking, so the lime and water literally explode on contact.
Striving for such explosive slaking as this, however, is inadvisable. A com-
plete slaking time of 5-10 minutes at a rapid and uniform rate is consider-
ably more desirable.

Two extreme conditions should be avoided. If excessive quantities of slaking
water are used, particularly cold water, "drowning" occurs. The surface of
the quicklime particle hydrates quickly, but the mass of hydrate formed
impedes the penetration of the water into the center of the particle, delay-
ing explosion of the particle into microparticles. The rise in temperature
is stifled and slaking delayed, resulting in coarser hydrate particles and
incomplete hydration. The other extreme is adding insufficient water to the
lime, causing the hydrate to be "burned" because of excessive temperatures,
121°-260°C (250°-500°F), instead of the desired temperature just below boil-
ing, 88°C (191°F). Much of the hydration water is lost as steam; thus, a
considerable number of nonhydrated particles remain. Also, the heat can be
so intense that paint on the equipment can blister or ignite, and lime parti
cles initially hydrated can be dehydrated.

3.2.7.1 Continuous Slaking

Because of the obvious economy and efficiency in slaking under optimum or
near-optimum conditions, manually operated batch slaking has been largely
replaced by continuous slakers. There are two basic types of slakers: (1)
the detention type that produces a lime slurry or creamy suspension; and (2)
paste slakers that produce only a paste or putty. Both types are generally
equipped with dilution tanks so any desired concentration of lime slurry can
be made.

There is some variance in the proportion of water and lime used, depending on
the characteristics of the lime and the type of slaker. With detention
slakers, the lime-to-water (weight) ratio averages 3:1 to 4:1 for high cal-
cium quicklimes and 2:1 for paste slakers. Most detention slakers will
handle a quicklime of 5 cm (2 in) top size down to pulverized forms, but
paste slakers require nothing larger than 19 mm (0.75 in). The trend today
is toward the use of smaller sizes of fairly restricted gradations of 19 mm x
9.5 mm (0.75 in x 0.38 in) or smaller granular and pebble sizes. The prin-
ciple of either type of slaker is the same; i.e., to interact quicklime and
water with sufficient contact time to achieve complete hydration while con-
tinuously discharging a slaked lime from the vessel.

The major difference between these two types is that the paste slaker oper-
ates more easily at higher temperatures, 88°-99°C (190°-210°F), than the
detention slaker, because more heat is generated from the lower water-to-1ime
ratio. To achieve the same desired temperature in a detention slaker, it is
often necessary to augment the natural heat of hydration by one or more of

29
-------
the following methods: (1) using hot water for slaking; (2) using a heat
exchanger in a dual tank system to capture some of the heat of hydration for
use in the first slaking tank; (3) using heavy insulation around the slaking
compartment to reduce heat losses; (4) using a longer retention time to com-
plete hydration. Paste slakers require only 5-10 minutes for complete hydra-
tion as opposed to 20-30 minutes for the detention slaker.

Proponents of the detention-type slaker claim that this unit can slake poorer
quality, slower slaking limes more efficiently, offering more flexibility in
accommodating the whole spectrum of limes. Another possible advantage for
the detention type is the prospect of longer life and less downtime and main-
tenance.

Two popular continuous slaker systems are illustrated in Figures 3-7 and 3-8.
These units provide two separate compartments: one for initially slaking to
35%-40% solids (putty consistency), and the other for dilution that produces
a workable slurry or milk-of-lime solution (5%-10%). Each of these compart-
ments is equipped with rotating paddles for agitation and complete mixing.
Also, the basic quicklime slaking unit provides a shaker or oscillating
screen for slurry degritting, a hood for vapor and dust removal, and the
necessary pumps and piping for transporting the slurry to the point of appli-
cation.

Another widely used slaker is shown in Figure 3-9. This continuous deten-
tion-type slaker provides dual propeller-type mixers, one for each compart-
ment, and a thermostatically controlled slaking temperature. This unit pro-
duces a fairly consistent slurry and provides flexibility for an operator's
needs in AMD treatment. As standard procedure, a thick pastel ike slurry is
made in the first compartment, then diluted and retained for complete hydra-
tion in the second. The desired slurry (solids percent) is controlled by a
preset water-to-quicklime ratio.

The last compartment has a separator trough equipped with a screw raking unit
on an incline for the removal and discharge of grit. This process enables
the grit to be loaded directly into a disposal or transporting container.

Most slakers are available in three and four compartments with a wide range
of capacities, 3.2-45.4 kkg/24 hr (3.5-50 tons/24 hr). Also, they are
equipped with heat exchangers for heat conservation and temperature control.

As mentioned previously, a slaking temperature ranging from 82°-88°C (180°-
190°F) is important, and it can be maintained automatically by a thermostat-
ically controlled water valve. Generally, the valves respond to a 3°C (5°F)
temperature change by adding varying amounts of water. Often, modern slakers
are insulated to retain the heat of hydration, which enhances operation
during seasonal weather variations.

When quicklime is selected as the neutralizing agent, grit removal must be
undertaken to preserve pump and equipment (tanks, pipes, valves) life. Even
the highest quality quicklimes contain 1.535-3.0% grit. The grit is composed
of silica, alumina, carbonate core, and insoluble calcium compounds. Based
on a lime usage of 18.2 kkg/d (20 tons/d), approximately 273-545 kg (600-

30
-------
QUICKLIME
WATER FOR GRIT WASHING-

TORQUE CONTROLLED WATER VALVE •,

JET SPRAY

SLAKING SECTION'
GRIT REMOVAL SECTION

SLURRY DISCHARGE SECTION

LIME SLURRY DISCHARGE

CLASSIFIER
GRIT DISCHARGE
LIQUID LEVEL
GRIT CONVEYOR
Figure 3-7. Paste slaker with classifier for grit removal.
-------
AGITATOR DRIVE
CHEMICAL MOTOR -
(ENCLOSED!
AGITATOR BLADES
I HARDENED STL.)
-VAPOR CONDENSOR
AND SEPARATOR WITH
REMOVABLE COVER AND
BAFFLES
WATER FLUSHED SEAL
GO
ro
ROTOMETER WITH
BYPASS PIPING
SHIPPED ON SLAKER
GRIT SCREEN DRIVE
CHEMICAL MOTOR
(ENCLOSED)
GRIT WASH JETS

IBRATING GRIT SCREEN
HEAT EXCHANGER
WITH BAFFLES
INVELOPES COMPLETE
TANK
SLURRY TANK
DISCHARGE OUTLET-
Figure 3-8. Paste slaker with vibrating screen for grit removal
-------
-s
to
CO
I
o
tt>
rt-
n>
o
ro
-------
1,200 Ib) of grit will require daily disposal.

The designer should be aware of the utilities required for a lime slaker.
Sometimes three-phase power or pneumatic air is needed. Slaking water is
most important with a clean supply always preferred. Past practice, however,
has been to use plant effluent, although in some cases the raw drainage is
used. This practice must be exercised with caution and the dissolved ion
chemistry carefully reviewed to prevent gypsum formation or retarding reac-
tion.

Research to show the importance of water quality in efficient lime slaking
proved that water of or near potable quality is most desirable. In particu-
lar, it was demonstrated that wastewater or recycle-process water containing
sulfites and sulfates retarded the slaking process. Not only was more time
needed to complete slaking, but the quality of the resulting lime slurry was
impaired. The lime hydrate particle size became much larger and the surface
area smaller, which, in turn, retarded the neutralization reaction with
acids. In fact, some of the lime did not hydrate and was wasted. The main
explanation is that lime precipitates the sulfite and sulfate ions that coat
the unreacted calcium oxide, preventing complete water penetration into the
particles.

It was also discovered that the wastewater or recycle-process water can be
used after slaking to dilute the thick lime slurry to the desired consist-
ency. The effect of the sulfite and sulfate ions on the quality of the
diluted lime slurry was negligible. The chloride ion in reasonable amounts
(500 mg/1) did not appear to exert any deleterious effect on slaking, but
higher concentrations retarded the reaction.

3.2.7.2 Degritting of Quicklime Slurries

Even the highest quality quicklimes have a grit content ranging between 1.5%
and 3.0% of the weight. Included in the grit, along with the carbonate core,
are insoluble silicates, aluminates, sulfates, ferrites, and all impurities
in the limestone before the lime was calcined. When the grit is ejected from
the slaker, it resembles a mass of wet sand with particles ranging from 6.4
mm (0.25 in) in diameter to #100 mesh.

Degritting improves lime quality and reduces abrasion and wear on equipment.
In extreme cases, cast iron centrifugal pumps have been worn out within a
month when pumping a slurry that has not been degritted. With efficient
degritting, the same equipment can operate for years without maintenance.

Degritting is performed in the dilution or stabilization tank adjacent to the
slaking chamber. As the slurry or paste passes over a weir into the dilution
chamber, it is dispersed and diluted by water sprays. The heavier grit
particles settle rapidly to the bottom and are removed automatically by
rakes. In some slakers, there is a classifier in the bottom of the dilution
tank where the grit is washed and the washwater recovered.

The washed grit is then disposed of manually or automatically. Slakers with

34
-------
a capacity of 227 kg/hr (500 Ib/hr) or more usually employ automatic educ-
tors. Meanwhile, enough turbulence is maintained so that nearly all of the
slaked lime in the diluted slurry remains in suspension and is piped to
storage.

3.2.7.3 Quicklime Slurry Concentrations

Designer preference as to the concentration of lime solids in the milk-of-
lime slurry used in their process varies from 5% to 20%. Most designers will
dilute the slurry to at least a 10% lime solids concentration before it
leaves the slaker. There is no fixed rule; however, it is recommended that
slurry feed systems be designed at a 10% lime-to-water weight ratio because a
more concentrated slurry causes additional maintenance and operational prob-
lems. In any event, the slurry feed rate should never exceed more than 10%
of the drainage flow. The concentration of lime can be checked for specific
gravity with a hydrometer and by using the data in Table 3-1.

The use of automatic pH control systems for the feeding of lime solutions
into the flash mix tank has become common in recent years. Therefore, the
strength of the slurry is not as important as the volume of slurry being fed.
Regardless of the approach used in controlling the lime feed, it is necessary
to maintain the lime in suspension and prevent settling to maintain a uniform
feed.

Also, it is important to recognize that percent lime solids is not equivalent
to percent calcium oxide. The percent lime solids refers strictly to the
hydroxide (Ca(OH)2) and not the oxide (CaO). Table 3-1 shows the equivalent
of Ca(OH)2 as CaO. To determine what concentration of lime solids exists,
the following simple equation is used:

% Solution = k9 of Ca(OH)2 x 100 (4)
kg of Ca(OH)2 + (1.0 kg/1 x No. of 1 of water)

% Solution = 1b of Ca(QH)2 x 100
Ib of Ca(OH)2 + (8.345 Ib/gal x No. of gal of water)

3.2.8 Slurried Lime Transfer

Usually the diluted paste or slurry ready for use must be transferred a short
distance to the mixing or neutralization tank. It is here that scaling can
become a serious problem. Lime is slightly soluble at best; a saturated
solution is 1.7 g/1 (0.14 Ib/gal) at 0°-10°C (32°-50°F). As the temperature
rises, the solubility of lime decreases until at 90°-100°C (194°-212°F) it is
only 0.55 g/1 (0.005 Ib/gal). Thus, it is economically essential to convey
lime in a much more concentrated form such as a suspension. Because lime
slurry has a pH greater than 12.0, the water that carries it undergoes a
softening action and precipitates fresh calcium carbonate as a dense, hard

35
-------
co
01
TABLE 3-1

PROPERTIES OF LIME SLURRIES
Approx.
% Solids
Ca(OH)2
(by wt)
1.5
3.2
4.8
6.3
7.8
9.4
10.8
12.2
13.7
15.2
16.4
18.0
19.2
20.4
21.8
23.1
24.4
25.6
26.9
28.0
29.2
30.4
31.5

Specific
Gravity
§
15°C
1.010
1.020
1.030
1.040
1.050
1.060
1.070
1.080
1.090
1.100
1.110
1.120
1.130
1.140
1.150
1.160
1.170
1.180
1.190
1.200
1.210
1.220
1.230

Pounds
per
gallon
8.41
8.50
8.58
8.66
8.75
8.83
8.91
8.99
9.08
9.16
9.25
9.33
9.41
9.50
9.58
9.66
9.75
9.85
9.91
10.00
10.08
10.16
10.24

Grams
CaO per
liter
11.7
24.4
37.1
49.8
62.5
75.2
87.9
100.0
113
126
138
152
164
177
190
203
216
229
242
255
268
281
294

Grams
Ca(OH)2
per liter
15.46
32.24
49.02
65.81
82.59
99.37
116.15
132.14
149.32
166.50
182.35
200.85
216.71
233.89
251.07
268.24
285.42
302.60
319.78
336.96
354.14
371.31
388.49

Pounds
CaO per
gallon
0.097
0.203
0.309
0.415
0.520
0.626
0.732
0.833
0.941
1.05
1.15
1.27
1.37
1.47
1.58
1.69
1.80
1.91
2.02
2.12
2.23
2.34
2.45
(continued)

Pounds
Ca(OH)2
per gallon
0.128
0.268
0.408
0.548
0.686
0.826
0.966
1.10
1.24
1.39
1.52
1.68
1.81
1.94
2.09
2.23
2.38
2.52
2.67
2.80
2.94
3.09
3.23

Pounds
Slurry per
cubic foot
62.94
63.62
64.22
64.82
65.49
66.09
66.69
67.29
67.96
68.56
69.23
69.83
70.43
71.10
71.73
72.35
72.97
73.73
74.17
74.85
75.44
76.04
76.64

Weight
Ratio
Water
to CaO
85.70:1
40.81:1
26.77:1
19.87:1
15.83:1
13.11:1
11.17:1
9.79:1
8.65:1
7.72:1
7.04:1
6.35:1
5.87:1
5.46:1
5.06:1
4.72:1
4.42:1
4.15:1
3.91:1
3.72:1
3.52:1
3.34:1
3.18:1

Weight
Ratio
Water
to Ca(OHh
64.70:1
30.72:1
20.03:1
14.80:1
11.76:1
9.69:1
8.22:1
7.17:1
6.32:1
5.59:1
5.09:1
4.55:1
4.20:1
3.90:1
3.58:1
3.33:1
3.10:1
2.91:1
2.71:1
2.57:1
2.43:1
2.29:1
2.17:1

-------
TABLE 3-1 (continued)
GO
Approx.
% Solids
Ca(OH),
(by wt)
32.7
33.8
35.
36.
37.
38,
39.6
40.8
41.
42.
43.
44.
45.
46.
47.
48.
49.
50.
51.
52.
52.8
53.9
.7
.7
.7
.7
.7
.6
.7
.4
.5
.6
.5
.2
Specific
Gravity
0
15°C
1.240
.250
.260
.270
.280
1.290
1.300
1.310
1.320
1.330
1.340
.350
.360
,370
.380
.390
1.400
1.410
1.420
1.430
1.440
1.450

Pounds
per
gallon
10.33
10.41
10.49
10.58
10.66
10.74
10.83
10.91
11.00
11.08
11.16
11.25
11.33
11.41
11.50
11.58
11.66
11.75
11.83
11.91
12.00
12.08

Grams
CaO per
liter
307
321
331
343
356
370
382
396
410
422
435
448
460
472
484
496
510
524
538
550
562
575

Grams
Ca(OH)2
per liter
405.67
424.17
437.38
453.24
470.42
488.92
504.77
523.27
541.77
557.63
574.81
591.99
607.84
623.70
639.56
655.41
673.91
692.41
710.91
726.77
742.63
759.81

Pounds
CaO per
gallon
2.56
2.67
2.81
2.92
3.03
3.14
3.25
3.37
3.48
3.58
3.70
3.81
3.92
4.03
4.15
4.25
4.37
4.50
4.61
4.71
4.82
4.93

Pounds
Ca(OH)2
per gallon
3.38
3.52
3.71
3.85
4.00
4.14
4.29
4.45
4.59
4.73
4.88
5.03
5.18
5.32
5.48
5.61
5.77
5.94
6.09
6.22
6.34
6.51

Pounds
Slurry per
cubic foot
77.33
77.92
78.52
79.19
79.79
80.38
81.06
81.66
82.33
82.93
83.53
84.20
84.80
85.40
86.07
86.67
87.27
87.95
88.54
89.14
89.82
90.42
Weight
Ratio
Water
to CaO
3.04:1
2.90:1
2.73:1
2.62:1
2.52:1
2.42:1
2.33:1
2.24:1
2.16:1
2.09:1
2.02:1
1.95:1
1.89:1
1.83:1
1.77:1
1.72:1
1.67:1
1.61:1
1.57:1
1.53:1
1.49:1
1.45:1
Weight
Ratio
Water
to Ca(OH)2
2.20:1
1.96:1
1.83:1
1.75:1
1.67:1
1.59:1
1.52:1
1.45:1
1.40:1
1.34:1
1.29:1
1.24:1
1.19:1
1.14:1
1.10:1
1.06:1
1.02:1
0.98:1
0.94:1
0.91:1
0.89:1
0.86:1
-------
scale. If unattended, scale will build up and clog pipes. Most often, scale
accumulates at the termination of the line to the treatment tank.

There is no foolproof solution to the problem; however, there are a few
corrective measures that will minimize the problem and often prevent exces-
sive maintenance. Those designing new plants should give serious thought to
this problem because it may be possible to preclude scaling by appropriate
design.

Scaling can also be prevented or minimized chemically. Sodium hexametaphos-
phate can be used to soften the slaking or dilution water so that the calcium
carbonate that precipitates does not accumulate or scale in the pipe.

The best method is to locate the feeder so that the slurry flows by gravity
directly into the solution or mixing tank. While this might not always be
possible, designers should strive to accomplish this in the arrangement of
the lime-feeding equipment. Scaling at the end of the slurry pipe can be
avoided by discharging slurry through an air gap into an open solution tank.
The use of heavy-duty, flexible rubber hoses or plastic pipes should be con-
sidered instead of metal pipes. These can be flexed or rapped to loosen any
scale buildup. Using open troughs to convey the lime suspension makes scale
removal simpler.

Much has been written about the components in a slurry feed system such as
piping, valves, pumps, and configurations. The designer, however, should be
aware of "practical engineering." That is, design with materials suitable
for the application and with operator/maintenance personnel in mind; provide
quick-disconnect, easy-to-assemble fittings and valves; avoid confined spaces
for suspected problem areas; and enhance flow patterns or directions by
proper use of tees, wyes, and elbows. Check valves should not be used, be-
cause slurry systems tend to cause them to plug and fail.

For example, a handy flushout within the piping network can be provided by
placing a wye on its back, with the branch having a removal plug. This
simple configuration provides easy flushing and cleaning without disassem-
bling the piping.

Also, pipe reducers should be avoided, because hydraulic conditions induced
by this fitting can cause "dewatering" (loss of fluidness) and compaction of
the lime slurry.

Much has been written about the type of piping material to use. Slurry feed
systems in the past have employed heavy-duty plastic pipe (Schedule 80),
flexible rubber hoses, and either stainless or galvanized steel pipe. The
final choice of a particular pipe material depends largely on personal pref-
erence, economics, anticipated problems, type of slaking water, and ease of
assembly.

Perhaps the best piping network for a slurry system was seen in Pennsylvania
at the Slippery Rock Acid Mine Drainage Plant constructed in 1967 under
"Project Scarlift." It utilizes clear braided or reinforced tygon tubing
connected to stationary fittings (pump inlets and outlets) with radiator

38
-------
clamps. This system allows constant flow observation, which immensely en-
hances location of blockages, along with easy breakdown for cleaning. Other
advantages to this system are the ability to incorporate continuous long
radius bends (ideal for slurry flows), and the ability of the pipe to flex
under blows from a rubber mallet to loosen buildups.

The primary disadvantages are limitations in size, difficult procurement, and
relative expense. The material cost can be overwhelmingly justified, how-
ever, by the savings in labor from maintenance.

A flushing system is not imperative in every situation. It is, however,
highly recommended for any system undergoing frequent shutdowns (i.e., for
any AMD plant that does not operate daily or operates only in wet weather).

Ideally, a minimum flow velocity of 1.0-1.2 m/s (3-4 ft/s) should be main-
tained within the slurry loop. Consequently, pipe siziag must be designed
and not arbitrarily selected; piping should not be oversized because of
anticipated future expansion. This common mistake can cause initial opera-
tional and maintenance problems amounting to a surprising cost.

The types of valves used in the slurry feed system determine, to a large
extent, the degree of operation or maintenance. Among the many types avail-
able, pinch valves have proven most successful. The nature of their opera-
tion provides a self-cleaning mechanism. Pinch valves will close and perform
efficiently, despite solid accumulations in the tube, and easily release any
dewatered lime deposits upon opening.

Pinch valves are not foolproof, however, and can cause operational headaches
if not properly sized (1), The tendency to use pinch valves the same size as
the slurry piping has resulted in oversizing and poor valve operation. Port
sleeves within a pinch valve are available in various sizes adaptable to
operational flow conditions and should be designed properly.

Ball or plug valves can be used in the slurry piping system for either fully
opened or fully closed conditions. The use of these valves for throttling
implies obvious disadvantages and results in problems. These valves are best
employed near pumps or on bottom nozzles of tanks.

A tight control on the metering devices (lime feeder, water valve) is impera-
tive for the bleed-feed slurry system to work. Extreme variations in slurry
concentrations (specific gravities) produce an inferior system with oscillat-
ing response times. A process designed to meet the demands of the control
system by varying the slurry concentration generally WILL NOT work.

3.2.9 Slurry Tanks

Storage tanks for lime slurries can come in a variety of shapes and sizes.
Generally, materials of construction can be whatever suits structural re-
quirements and resists the high pH of the slurry. The configuration is
usually dictated by the space available or assigned. Storage tank layout
should tend to eliminate short-circuits in flow. This will help prevent

39
-------
unstabllized slurry from entering the treatment system. To reiterate, pri-
mary attention should be given to simplicity of transport between slurry
preparation and storage tanks.

Tanks exceeding 9,500 1 (2,500 gal) should be fitted with baffles, set 90°
apart, to prevent vortex formation. Baffle widths equal to one-twelfth the
tank diameter are sufficient in most cases (see Chapter 4, Mixing).

Slurry storage tanks with adequate agitation must be provided. Slurries
resulting from proper lime slaking require relatively low-energy agitation to
maintain the suspension. Hydrated lime particle size is small and its set-
tling rate is slow. Mixer horsepower and impeller size and shape should be
designed to keep impurities as well as lime in suspension. Otherwise, grit
particles and precipitated salts might become troublesome sediment.

Agitator requirements are influenced by the tank size and shape. When avail-
able space allows some latitude in tank configuration, consult the agitator
supplier for recommendations. These can direct the designer toward the most
efficient application of his equipment.

3.2.10 Slurry Feed Control

At most plants, the treatment of AMD need not involve a sophisticated slurry
feed control system. Usually the drainage flow is constant both in quantity
and quality because of the beneficial effect of an equalization basin. The
system is best controlled by the pH of the flow leaving the neutralization
tank. This pH signal can control the dry lime feeder, a lime solution feed-
er, or a control valve. All three systems work well; but they are discussed
here in order of preference.

Where drainage flow is fairly constant and the quality does not vary signif-
icantly, the dry lime feeder system can be adjusted so it operates constant-
ly, and the lime slurry continuously overflows from the stabilization tank
into the neutralization tank. This control method works well if ferrous iron
concentrations are less than 100 mg/1. A 2.0- or 3.0-pH unit control range
(usually 7.0-9.0) can be set on the controller, and any variance outside
these limits should sound an alarm or stop the system until adjustments can
be made. Periodic adjustments to the feeder or makeup water flow may be
necessary to control pH trends. More than likely, this will be the result of
gradual changes in the AMD quality or variations in the lime quality.

Where slakers discharge continuously to the stabilization tank, the chances
are better that control can be maintained by the slurry feed rate. Response
time is less, but slakers do not change their output at once when input is
changed. Paste slakers perform better than slurry slakers in these circum-
stances because they maintain a constant paste concentration. Changes in
output occur as the paste flow rate and the small dilution compartment con-
centration are changed. A slurry shaker must change the concentration of the
entire slaker before a rate change is completed. The time for completion of
such a concentration change may be too long and the system will shut down.

40
-------
The importance of the slurry preparation method cannot be overstressed.
Where short response time is required, meeting the demands of the control
system by varying the slurry concentration WILL NOT work. The preparation of
a constant slurry concentration and variation of the rate of application will
work. The possible exception is a paste slaker or a hydrated lime feeder
with varying feed control and small dilution tank from which the lime is
sluiced directly to the neutralization tank.

The second type of feed system is a versatile type of solution feeder that
has been used successfully for most types of lime suspensions. It is the
dipper wheel type of feeder, sometimes called the Archimedes Wheel. This
feeder, usually comprised of eight dippers that hold 0.5 1 each, is rotated
by a variable-speed drive. When submerged in a tank with a constant level,
the dippers fill and discharge into an outlet trough. A mechanical counter
driven by the shaft indicates dipper revolutions. It can be completely
automated pneumatically, electrically, and mechanically with a ratio control-
ler, or with the variable-speed drive and remote pH control (see Figure
3-10).

Another system, designed primarily for feeding lime slurries, is regulated by
a pH probe and controller. It is supplied as part of a compact, complete
package neutralization treatment system for permanent or portable installa-
tions. The whole unit measures 2.6 m x 1.6 m x 1.4 m (8.5 ft x 5.25 ft x 4.5
ft). It comprises a slurry mixing tank of either 0.6- or 0.9-m3 (160- or
240-gal) capacity with a flash mixer; a pump, if gravity flow of the slurry
is not possible; and a highly automated pH sensor at the discharge end of the
unit, which regulates the flow of slurry into the treatment process. This
system is based on using hydrated lime, either introduced manually by feeding
22.7-kg (50-lb) bags of lime into the mixing tank or metered from an overhead
dry feeder. The lime is mixed into a uniform concentrated slurry.

The slurry flows by gravity or is pumped to a rotating cup wheel, where it is
picked up and discharged to a funnel leading to the treatment tank. This
funnel, located inside the wheel, is provided with an adjustable hood that
controls the lime slurry feed. Control settings are provided by a controller
and pH probe.

The feed rate is claimed to be infinitely variable over a 1-100 range, and
the largest capacity model can feed up to 227 kg/hr (500 Ib/hr) of hydrated
lime in slurry form. To provide for continuous pH (and feed) control, a
spare pH electrode is maintained in operating condition and used interchange-
ably.

The third type of lime slurry feeding involves proportioning the slurry flow
in response to the pH signal. This most workable system involves pumping
lime slurry through a continuous pipe loop from the stabilization tank and
returning it to that tank. This pipeline should be sized so there is a sig-
nificant head on the slurry circulating pump. The slurry is fed from a
branch line and controlled by a pinch valve or another type of throttling
valve. As previously mentioned, pinch valves are usually controlled by pneu-
matic systems. This system is best applied when there will be frequent vari-
ations in flow or quality, and when strict pH control is required.

41
-------
Figure 3-10. Dipper wheel and slurry feeder.

42
-------
3.2.11 Pumps

Pumps for lime slurries generally fall into two categories: controlled vol-
ume and centrifugal. Controlled volume pumps are usually reciprocating dia-
phragm or progressive cavity types. The diaphragm metering pump has limited
capacity and is not normally used in mine drainage treatment. Progressive
cavity pumps are available in much higher capacities, but they wear faster
than a centrifugal pump under similar operating conditions. An important
consideration is maintaining adequate pipeline velocities with the modulating
flow produced by these pumps. If the pumps do not provide adequate pipeline
velocities at minimum rates, a recirculated slurry loop should be provided.

Centrifugal pumps and control valves are the obvious choices for a wide range
of slurry flows. They are inexpensive, and standard designs incorporate flow
patterns that lend themselves to easy slurry transfer. Two common configura-
tions have been used with transfer pumps and different control valve opera-
tions (see Figures 3-11 and 3-12). Both methods work satisfactorily; there-
fore, they become a designer preference.

Manufacturers of centrifugal slurry pumps regard lime suspensions as fairly
mild among slurries. These pumps are usually cast iron with replaceable
liners and semiopen impellers. The designer should consider the tendency of
lime to dewater under certain conditions of velocity and turbulence. Pumps
should be easy to disassemble for maintenance and cleaning. Speed of rota-
tion should not exceed 1,750 r/min (revolutions per minute), and should be as
low as hydraulic requirements allow.

Although there appears to be a lack of agreement as to the kilowatts (horse-
power) needed to pump lime slurries, the following formula is generally used:

... Q x U x H . Q x W x H ,^
kW = 76 x E hp - 550 x E (5)

where kW = power in kilowatts hp = power in horsepower

Q = liquid flow in m3/s Q = ft3/s

W = specific weight of lime W = lb/ft3
slurry in kg/m3

H = head in m H = ft

E = hydraulic efficiency of pump E = hydraulic efficiency of
in decimal fraction pump in decimal fraction

76 refers to kg-m 550 refers to ft-lb

For large pumps, the efficiency is in the 70%-80% range; but for small pumps,
40%-50% is generally used.

Pump shaft seals should be specially designed for slurries and should not be

43
-------
AIR SUPPLY
-1-
QUICKLIME
STORAGE
BJ—^ % (PHC) (PHR)
ACKPRESSURE
VALVE
iSLAKER
» .
f— _~ _ — «, ^ «
1
1
00-
--t — 1

t
' SLURRY
- T*-
rii i
U' 1

GRIT
DISCH
CONTROL
FLUSHING
SLURRY
LOOP
PROCESS
REACTOR
SLURRY PUMP
STABILIZATION
&
STORAGE
Figure 3-11. Slurry feed with pH control loop,

44
-------
QUICKLIME
STORAGE
BACKPRESSURE
VALVE
1
1
CXJ-
--I 1

*
' SLURRY
I -A-
rti i
U ' I
.1,1 n
1 IU
f

STABILIZATION
PROCESS
REACTOR
FLUSHING
CONTROL
GRIT
DISCH.
SLURRY PUMP
&
STORAGE
Figure 3-12. Slurry feed by flow proportioning.

45
-------
water-flushed. Introduction of water into the slurry will reactivate
scaling. This usually will not occur in the pump, but starts to form shortly
downstream. In any event, it can be avoided by using dry seals.

Pumps and piping should be sized for minimum required velocities in the re-
circulated slurry loop plus the process requirements. The specific gravity
of the slurry should be taken into consideration in these calculations as
well as those for valve sizing.

Standby pumps should be included in accordance with requirements for process
reliability. If the process operation is critical to facility performance,
full standby can be justified. Standby pumps, as well as other piping,
should be placed in a manner that will minimize or avoid the settling of
solids in the static sections.
3.2.12 Miscellaneous Considerations

Two of the most frequent questions presented to designers about lime slurry
handling relate to friction loss and viscosity in pumps. For straight pipe-
lines up to 152.4 m (500 ft) in length, the friction loss is relatively neg-
ligible. With several elbows and turns, there will be a greater friction
loss, particularly with lime slurry containing some grit. In designing a
small pipeline, it is common practice to increase the pump horsepower about
10% to allow for friction loss.

Viscosity of lime slurry is highly variable and depends upon such diverse
factors as type of lime used (hydrate vs. quicklime), particle size (surface
area), reactivity, slaking water, lime ratio, initial temperature of slaking
water, slurry concentration, and temperature. Nevertheless, the viscosity of
a usable slurry (10% solids) can be assumed the same as water.

3.3 Limestone

There has been considerable research, development, and demonstration in the
use of limestone for the neutralization and treatment of acid mine drainage.
On the surface, there appears to be an economic advantage with limestone,
which is available at about 30% of the cost of quicklime or hydrated lime.
As will be explained, there are inherent disadvantages for using limestone in
this kind of treatment process. As a result, there are very few if any
operating facilities using limestone to treat acid mine drainage.

The dolomitic (magnesium) and high-calcium form of limestone are available in
a variety of sizes. Commercially available sizes range from 76.2 mm (3 in)
to 0.074 mm (200 mesh).

Although there are two types of limestone, the high-calcium form has under-
gone the most study as a neutralizing agent because of better reactivity.
The rate of reaction for dolomitic limestone was studied equally but found to
be ineffective. Subsequent discussions about limestone refer only to the
high-calcium form.

46
-------
High-calcium limestone usually has the following chemical composition (1, 2):
Calcium Limestone
(Rock Dust)
% Composition
Calcium Oxide (CaO) 53.0 - 56.0
Magnesium Oxide (MgO) 0.12 - 3.11
Calcium Carbonate (CaC03) 92.66 - 98.6
Silica Dioxide (Si02) 0.1 - 2.89
The reaction of limestone with an acid is accompanied by the liberation of
carbon dioxide (C02), expressed by the following equations.
Limestone + Strong Acid ->- Gypsum + Carbonic Acid
CaC03 + H2SOif •> CaSO^ + H2C03 (6)
3CaC03 + Fe2(SO.l)3 + 6H20 + 2CaS04 + 2Fe(OH)3 + 3H2C03 (7)
3CaC03 + A12 (SO^Ja + 6H20 + 3CaSO^ + 2A1(OH)3 + 3H2C03 (8)
The effervescent reaction, or release of carbon dioxide, from limestone neu-
tralization is not so obvious with acid mine drainage as with stronger acids,
but it does occur.
To utilize limestone effectively as a neutralizing agent, certain quality
criteria must be maintained. An early study by Bituminous Coal Research
(BCR) involving two years of investigation concluded that the effectiveness
of limestone as a neutralizing agent depends upon the following criteria (3):
1. minimum particle size, preferably a minus 325 mesh (approximately
0.044 mm);
2. high calcium content, approaching pure calcium carbonate;
3. low magnesium content, thus eliminating dolomites (CaMg(COa)2) and
magnesites (MgC03);
4. high specific (surface) area.
There are many different quality limestones available; however, finding a
limestone that meets all these criteria can be difficult.
3.3.1 Treatment of Ferrous Iron Drainage
An EPA study by Wilmoth (1) indicates that limestone treatment of ferrous
iron streams is possible, but economically undesirable. The most significant
47
-------
factors inhibiting full-scale use of powdered limestone for AMD treatment are
its slow reactivity and its inability to increase the pH above 7.0 (maximum
obtainable was 7.4). A pH above 7.0 could be achieved only through effluent
recycle and increased reaction time, which enlarges the process units.

At pH 7.0, the slow rate of oxidation of ferrous iron is a deterrent. Exces-
sive aeration times of several hours (normally 30 minutes or less at pH 8.5)
are needed to oxidize the ferrous iron completely. In addition to increasing
the normal costs for the aeration unit, the longer detention periods have a
tendency to destroy the floe particles required for good settling, which re-
sults in an effluent with high suspended solids and turbidity.

According to Wilmoth, the final process scheme that produced a satisfactory
effluent incorporated a 20% sludge recycle rate to maintain a pH near 7.4; 30
minutes detention within the flash mixer; 4-6 hours aeration; and finally,
coagulant addition for reflocculation and effective settling of the sludge
and limestone particles.

The reagent cost alone for this process approached $0.15/3.8 m3 (1,000 gal),
which essentially nullifies any economic advantage in using limestone over
other lime forms. The additional capital investment for enlarged process
units (i.e., flash mixer, aerators, and aeration tank) and power costs total-
ly invalidate the practicality of the limestone treatment system for high
ferrous iron drainages.

Perhaps a better design criteria (3) for limestone treatment of AMD would be:

Ferrous Iron
Concentration Response to Limestone Treatment
0-50 Effective treatment may be achieved without pre- or
postneutralization iron oxidation.

50-100 May be effectively treated but requires postneutrali-
zation aeration and significant reaction-retention
time. Preneutralization oxidation can reduce ferrous
iron concentrations.

> 100 Potential treatment is uncertain with experience to
date, unless combined with preneutralization ferrous
iron oxidation to achieve the above ferrous levels.

Although only a guide, these statements are "cautionary" and imply skepti-
cism. More recent studies supplement these statements further by discour-
aging the use of limestone with water having more than 100 mg/1 ferrous iron.

Lovell also emphasizes many significant design parameters that must be evalu-
ated before selecting a limestone neutralization process (4). Those high-
lighted are:

1. specifications for limestone grade, size, and hardness;

48
-------
2. mode of operation (mixer or rotary mill);

3. gas exchange capabilities (oxidation and C02 removal);

4. supplementary reagent requirements (lime, polymers);

5. operating pH and aeration requirements for ferrous iron oxidation;

6. ratio of recycle volume;

7. sludge settleability.

3.3.2 Treatment of Ferric Iron Drainage

To this point, only the feasibility of successful limestone treatment of mine
drainage containing ferrous iron has been discussed. If iron is mostly in
the ferric form (at least a 4:1 ferric-to-ferrous ratio), treatment with
limestone appears feasible. Wilmoth and Hill, however, could only realize a
32% limestone utilization efficiency in their studies on Grassy Run drainage
at Norton, W. Va. (5). Consequently, it would take three times the amount of
limestone to neutralize this water. At the time of the study (1970), lime-
stone cost less than one-third that of lime, which placed it at an economic
disadvantage. Today, with the price (including delivery) for limestone at
$24/Mg ($22/ton) and hydrated lime at $72/Mg ($65/ton), the treatment eco-
nomics are more favorable.

In summary, the use of limestone for the treatment of AMD has decisive dis-
advantages, both economic and functional. Limestone cannot be used to treat
drainages containing manganese because the maximum obtainable pH is not
sufficient to precipitate the manganese.

3.4 Caustic Soda

Traditionally, caustic soda (NaOH) has been employed as a neutralizing agent
for low-flow, mildly acidic drainages located in remote areas. Generally
associated with surface mine drainage problems, it is extensively used for
neutralization in the panhandle region near West Virginia, Ohio, and Pennsyl-
vania.

The conventional process usually consists of a horizontally mounted 38-m3
(10,000-gal) storage tank for the caustic soda, and a flume-type chemical
feeder (Figure 3-13). Gravity serves as the source of power, with a constant
head valve in the chemical feeder controlling a constant feed.

In surface mining operations, caustic is frequently used to neutralize acid
water as it is being pumped from the pit. For such a temporary or portable
need, caustic can be siphoned into the suction side of a portable mine water
pump as shown in Figure 3-14. The flow of caustic is controlled by adjust-
ment of the gate valve on the storage tank outlet. The discharge line of the
pump should be of sufficient length to provide 1-2 minutes retention for ade-

49
-------
INLET-

STEAM
COIL
MANWAY
-SUPPORT

CAUSTIC SODA STORAGE TANK
COAL
STORAGE
PILE
CHEMICAL FEED
CONTROL PUMP
(IF REQUIRED)
OPTION
pH CONTROL-
SET AT 8.5
OPERATING RANGE
^ 7.0-9.0
EXISTING SETTLING POND
50% Caustic Feed Options

1. Variable chemical feed pump with pH probe controller.

2. Solenoid valve with pH probe controller (gravity).

3. Constant chemical feed pump with electric pinch valve and pH probe
control .
Figure 3-13. Caustic soda treatment system.

50
-------
CAUSTIC SODA (NaOH)
STORAGE TANK
" GATE
VALVE
FLEXIBLE SUCTION LINE-
PLASTIC
I"SUPPLY LINE
JL
-GASOLINE MOTOR
CONNECTION TO
PUMP SUCTION
PORTABLE
6" PUMP
(TYPICAL)
PIT WATER
\
SETTLING BASIN
Figure 3-14. Portable caustic soda feed arrangement.
-------
quate mixing of the caustic with the mine drainage.

The decisive limiting factor to the use of caustic soda in AMD treatment is
its cost. Caustic soda costs average about $287/dry kkg ($260/dry ton).
Since it is usually bought as a liquid (50% concentration), the purchaser
pays for hauling 50% water, which further increases this cost. In cold wea-
ther, caustic is purchased as a 20% solution, which allows for the lowest
temperature before freezing occurs.

Aside from the dangers in handling caustic, the chemical characteristics of
concentrated solutions require design consideration. Table 3-2 illustrates
some of the properties of sodium hydroxide solutions. This knowledge is
extremely helpful when one also knows the theoretical alkali requirements. A
close approximation for the dilution of the concentrated solution (50%) can
be computed and made directly at the storage area upon delivery, following
all appropriate safety precautions.

Figure 3-15 presents the freezing points for various solution strengths. The
50% caustic soda solution freezes at 12°C (54°F), and is quite viscous (bare-
ly pumpable) at a slightly higher temperature of 15°C (60°F). Ideally, an
18% solution offers the lowest freezing temperature; however, higher concen-
trations have been used in coal-producing states with minimal freezing prob-
lems in the winter.

Despite its detrimental properties, sodium hydroxide will produce an excel-
lent effluent quality. A properly treated drainage will be low in suspended
solids and turbidity, and will have an iron content within limitations.
Also, caustic soda is nearly 100% reactive. Generally, sodium hydroxide
sludges are fluffy and more voluminous than lime sludges; however, they dis-
play acceptable settling properties (5).

In his experiments at the Crown Mine drainage field site, Kennedy encountered
several problems related to winter weather that must be overcome for caustic
soda to be reliable (6). One of these was acquiring a good water supply to
dilute the sodium hydroxide. This can be a severe problem. An illustration
of the type of chemical feeder used by Kennedy is shown in Figure 3-16.
Devices such as these are available from several manufacturers with installed
costs in the area of $10,000.

The settling ponds in many of these facilites are merely depressions in re-
claimed strip land. The volume of sludge associated with this type of treat-
ment will vary from 4% to 6% of the flow, depending upon the acidity and iron
content of the drainage.

The use of sodium hydroxide for treatment of high sulfate drainages (greater
than 2,500 mg/1) has merit because there is no gypsum (CaSO^) formation. The
effluent, however, will have a higher dissolved solids content because of the
total solubility of sodium sulfate.

In summary, caustic soda is an excellent neutralizing agent capable of pro-
ducing an effluent within discharge limitations. The high cost of this
reagent, along with its undesirable handling properties, limits its use.

52
-------
TABLE 3-2

DENSITY OF AQUEOUS
SODIUM HYDROXIDE SOLUTIONS AT 20°/4° C
VALUES GIVEN IN THE INTERNATIONAL CRITICAL TABLES

Weight of NaOH in Solution
Specific
Gravity
1.0095
1.0207
1.0318
1.0428
1.0538
1.0648
1.0753
1.0869
1.0979
1.1089
1.1309
1.1530
1.1751
1.1972
1.2191
1.2411
1.2629
1.2848
1.3054
1.3279
1.3490
1.3696
1.3900
1.4101
1.4300
1.4494
1.4685
1.4873
1.5065
1.5253
Grams per
liter
10.10
20.41
30.95
41.71
52.69
63.89
75.31
86.95
98.81
110.9
135.7
161.4
188.0
215.5
243.8
273.0
303.1
334.0
365.8
398.4
431.7
465.7
500.4
535.8
572.0
608.7
646.1
684.2
723.1
762.7
Pounds per
U.S. gallon
0.08425
0.1704
0.2583
0.3481
0.4397
0.5332
0.6285
0.7256
0.8246
0.9254
1.133
1.347
1.569
1.798
2.035
2.279
2.529
2.788
3.053
3.325
3.603
3.886
4.176
4.472
4.774
5.080
5.392
5.710
6.035
6.365
Pounds per
cubic foot
0.6302
1.274
1.932
2.604
3.289
3.989
4.701
5.428
6.169
6.923
8.472
10.08
11.74
13.45
15.22
17.05
18.92
20.85
22.84
24.87
26.95
29.07
31.24
33.45
35.71
38.00
40.34
42.71
45.14
47.61
Percent
NaOH
1
2
3
4
5
6
7
8
9
10
12
14
16
18
20
22
24
26
28
30
32
34
36
38
40
42
44
46
48
50
Degrees
Baume
1.4
2.9
4.5
6.0
7.4
8.8
10.2
11.6
12.9
14.2
16.8
19.2
21.6
23.9
26.1
28.2
30.2
32.1
34.0
35.8
37.5
39.1
40.7
42.2
43.6
45.0
46.3
47.5
48.8
49.9
Degrees
Twaddel 1
1.90
4.14
6.36
8.56
10.76
12.96
15.16
17.38
19.53
21.78
26.18
30.60
35.02
39.44
43.82
48.22
52.58
56.96
61.28
65.58
69.80
73.92
78.00
82.02
86.00
89.88
93.70
97.46
101.30
105.06
53
-------
160
120
AREA SOLID AND SOLUTION PHASES
en
80
u.
o
I
£

I
Q. 40
0)
NaOH
NaOH
NaOH
NaOH
NaOH
-40
1
2
3
4
5
6
7
8
9
10
11
12
13
14
Ice + Solution
Ice + NaOH -7H,0
NaOH • 7H20 + Solution
• 5H2O-t- Solution
-7H,0 + NaOH • 5H20
-4H,0 + Solution
-4H20
- 3 VaH,0 + Solution
NaOH-4HaO + NaOH-3V2H20
NaOH-SVaHjO + NaOH- 2H,0
- 2HS0 + Solution
-HjO + Solution
NaOH-2HiO+NaOH-H,0
NaOH + Solution
NaOH-H,
60
70
80
90
100
Percent NaOH
Figure 3-15. Freezing points of caustic soda solutions.
-------
CHEMICAL
FEED
RESERVOIR
NaOH SOLUTION
STORAGE
FEED ADJUSTMENT
ORIFICE AND METERING ROD
FLUME
FLOW
STILLING WELL
(MAINSTREAM FLOW
INDICATOR)
Figure 3-16. Flume chemical feeder.

55
-------
Sodium hydroxide has been successfully used in remote locations to treat
mildly acid or small flow drainages. The automatic feeding methods always
present the danger of overtreatment in incidences of a feeder malfunction.
Winter weather protection must be provided to minimize freezing problems.
Finally, sludge produced by treatment with this chemical will have good
settling properties, but it tends to be fluffy and more susceptible to wash-
out. In most cases, the settling pond should represent the final disposal
site for sludge.

3.5 Soda Ash

Soda ash (Na2C03) has rarely, if ever, been used as the principal neutraliz-
ing agent in large-flow mine drainage treatment facilities. Due to its high
cost of $309/Mg ($280/ton) and limited availability, soda ash is usually
used only for treatment of low-flow drainages that contain little ferrous
iron such as would occur in surface mines. Its selection for such applica-
tions is more for convenience than cost efficiency.

Soda ash is produced in solid pellet form called briquettes or "prills."
Production of these is limited and the availability is scarce. Effective
treatment has been demonstrated by simply immersing these prills in a wire
basket in the flowing drainage. The rate of dissolution will enable such a
system to be effective for a 24-hour period. The volume of drainage flow and
its acidity will define the number of prills to be used each day. This is
best determined by experimentation.

Devices are available for feeding soda ash prills over a longer period. One
method consists of a storage hopper mounted over a basket (see Figure 3-17).
The hopper usually accommodates one or two bags up to 90 kg (200 Ib) of
prills. The prills will tumble through the hopper opening at the bottom and
keep the basket full. While this is a simple system, it is subject to clog-
ging at the hopper opening as the soda ash absorbs moisture, causing the
prills to expand.

This method of treatment is controlled by the rate of dissolution of the
prills. Depending upon the acidity in the drainage, the final treated water
may not be within the desired pH range. It should also be recognized that
drainages with significant iron concentrations may cause the prills to be
coated, rendering them ineffective. Multiple systems could be employed in
series or parallel if one unit cannot adequately treat the flow or acidity
involved.

In practice today, a typical soda ash small-scale feeding operation is shown
in Figure 3-18. The cost for such a feeder would be about $5,000.

A more effective method of using soda ash is in conjunction with a dissolver,
if electric power is available. Flaked soda ash can be fed into a batch tank
(comparable to a slurry tank) equipped with a mixer to dissolve the soda ash
in a water solution. The tank size would depend upon the amount of alkali
(soda ash) needed, the time period involved, and the concentration of solu-
tion used. The soda ash solution would then be fed into the AMD stream by

56
-------
. *
Figure 3-17. Soda ash prill hopper.
Figure 3-18. Soda ash vibrating feeders.
57
-------
gravity with flow controlled by a valve, or pumped at a constant to variable
rate depending on the degree of sophistication used (refer to Figure 3-13).
This system offers better control and reliability in treatment; however, it
does require a power source and additional equipment.

Wilmoth and Hill, in their studies on neutralizing ferric iron mine drainages
(approximately 125 mg/1) with soda ash, were able to produce an effluent of
satisfactory discharge quality, but treatment costs were impractical (5).
Based on their results and today's prices, soda ash treatment would be the
most expensive method for treating mine drainage and appears economically
undesirable.

In laboratory tests with several mine drainages, Lovell found soda ash con-
sumption ranged from 150% to 310% more than the theoretical requirement for
neutralization (4). He explicitly reports a 56% use-efficiency in treating
AMD, which means that twice the theoretical soda ash requirement will be
required.

Both researchers found that the sludge formed by soda ash neutralization
settled well and compacted to densities comparable to those obtained wth
quicklime and hydrated lime.

3.6 References

1. Wilmoth, R.C. Limestone and Lime Neutralization of Ferrous Iron Acid
Mine Drainage. EPA 600/2-77-101, May 1977.

2. Murry, J.A., et al. Journal of American Ceramic Society. 37(7):238-
323, 1954.

3. Bituminous Coal Research. Studies on Limestone Treatment of Acid Mine
Drainage. Water Pollution Control Research Series, January 1970.

4. Lovell, H.L. An Appraisal of Neutralization Processes to Treat Coal
Mine Drainage. EPA 670/2-73-093, November 1973, p. 81.

5. Wilmoth, R.C., and R.D. Hill. Neutralization of High Ferric Iron Acid
Mine Drainage. Federal Water Quality Administration, August 1970.

6. Kennedy, J.L. Sodium Hydroxide Treatment of Acid Mine Drainage. Crown
Mine Drainage Field Site, February 1973.
58
-------
CHAPTER 4

MIXING
4.1 Introduction
Unit processes that essentially mix are incorporated into every acid mine
drainage treatment facility. Mixing involves the uniform dispersion or
suspension of liquid or solid particles into another liquid or solid media.
Impeller design, rotational speed, and horsepower requirements vary according
to the mixing application and the properties of the reagents.

Rapid mixing is usually employed in AMD treatment for the addition of the
neutralizing agent to the raw water. Mixing is also required for preparation
of lime slurries and polymer solutions. Gentle or slow mixing is associated
with flocculation, large lime slurry storage tanks, and sludge thickening.

The purpose of this chapter is to make the designer aware of mixing technol-
ogy and the techniques employed in sizing a mixer. The basic engineering
principles involved with the proper sizing and design of the mixing vessel
are also presented.

4.2 Types of Mixers

Mixers are classified into two broad categories according to the flow pat-
terns they produce. Axial and radial patterns are general terms describing
the kinds of mixing devices that impel or move the entraining media. Axial
flow includes propellers, axial turbines, fan turbines, and pitched paddles.
The axial and fan turbines are the most common. These mixing devices induce
flow patterns parallel to the drive shaft. Radial flow impellers, on the
other hand, induce currents perpendicular to the drive shaft and include
radial turbines and paddles. Figure 4-1 illustrates the simple difference
between the impellers that produce these two distinct flow patterns.

As shown in Figure 4-1, the basic difference between the two types of mixers
is that the axial turbine blades are pitched. Consequently, each impeller
type has its own characteristic flow patterns, making each more suitable for
certain mixing applications. In this manual, emphasis will be on axial
impellers because most mixing applications in AMD treatment require this
type.
59
-------
FLOW
TOTh

PARALLE
iE AXIS -
/"

V
V
-
V
^_
"^ -^

i
1
1

-i
J

4
J
i-

y
^*

\
\

' .
,:
•* X
AXIAL MIXING

-BAFFLES
TYPICAL FLOW PATTERN IN
BAFFLED TANK WITH AXIAL TURBINE
POSITIONED ON CENTER
AXIAL TURBINE
V

RADIAL MIXING
FLOW PERPENDICULAR
TO THE AXIS
TYPICAL FLOW PATTERN IN BAFFLES
BAFFLED TANK WITH RADIAL TURBINE
POSITIONED ON CENTER
RADIAL TURBINE
Figure 4-1. Comparison of axial and radial flow patterns.

60
-------
4.2.1 Propeller Mixers

Propeller mixers are used primarily for rapid mixing. Axial flow propeller
mixers can be portable or fixed-mounted, depending on the mixer size and its
application. Portable mixers are usually mounted on the side of the mixing
vessel. Generally, angular, top-entering propeller mixers range in size from
0.37 to 2.24 kW (0.5 to 3.0 hp), although many designs limit the size to 0.75
kW (1.0 hp) and a maximum shaft length of 1.83 m (6.0 ft).

Portable mixers are usually mounted angularly off-center to obtain a top-to-
bottom mixing pattern. This prevents a swirling pattern from developing and
also eliminates the need for tank baffles. Typically, the maximum vessel
volume used with angular mixers 2.24 kW (3.0 hp) or less is 3.785 m3 (1,000
gal), but can vary with the application. The mixer shaft should enter at a
15° angle from vertical and, if possible, at a point off the tank centerline
as shown in Figure 4-2.
\
•
PROPELLER TURNING
COUNTERCLOCKWISE
LOOKING DOWN ON SHAFT
Figure 4-2. Off-center, top-entering propeller positions.
Fixed-mounted, right-angle drive propeller mixers should be positioned ver-
tically, on-center, in a baffled tank.

Propellers are usually no larger than 46 cm (18 in) in diameter regardless of
the vessel size. In deep tanks with on-center, top-mounted, right-angle, or
vertical mixers, multiple propellers can be placed on a single shaft, usually
aiming the liquid in the same direction.
61
-------
The flow patterns produced by propeller and other axial mixers drive the
water down to the tank bottom, then horizontally until turning upwards at the
wall, and eventually return it to the suction side of the propeller (oval
loop), as illustrated in Figure 4-1.

There are two basic speed ranges applicable to both portable and fixed-
mounted propeller mixers. Direct-drive propellers (high speed) rotate at
either 1,150 or 1,750 r/min, while gear-driven propellers (low speed) rotate
within the 350-420 r/min range (1). The faster speeds provide a high level
of shear with a low draft capacity, making it suitable for flash mixing of
chemicals. The low speeds provide less fluid shear with a large draft capac-
ity, making them more suitable for solids suspension applications, such as
slurry makeup tanks less than 11.355 m3 (3,000 gal) in size (2).

4.2.2 Turbine Mixers

Turbine mixers can be axial or radial flow, depending on the impeller design.
Axial impellers are pitched-blade or fan turbines, while radial impellers are
flat, curved, or with a spiral backswept blade as illustrated in Figure 4-3.
The curved and spiral backswept impellers are used only for high-viscosity
applications, which are not often encountered in AMD treatment unless mixing
large quantities of sodium hydroxide or soda ash.
Figure 4-3.
Typical radial turbine impellers: (1) flat-blade turbine,
(2) spiral backswept turbine, (3) curved-blade turbine.

62
-------
Axial turbines are used for most large-scale mixing applications involving
liquid-solid suspensions such as mixing large, stored volumes of a lime
slurry. Turbine mixers are usually fixed-mounted, vertically, in fully
baffled tanks. This configuration gives good top-to-bottom fluid circulation
throughout the mixing vessel. Turbine impeller diameters are usually one-
third of the tank diameter, but can range between 30% and 40%.

The standard number of blades on a turbine impeller is either 6 or 8, but
there can be anywhere from 4 to 16. The turbine impeller is mounted about 1
turbine diameter above the tank bottom. Turbine units are always mounted
vertically. Their power range can be provided between 0.75 and 373 kW (1 and
500 hp). Common turbine impeller speeds are available between 50 and 150 r/
min, but can be obtained in a range from 15 to 420 r/min (1).

When mounted in a tank with a liquid depth equal to or greater than its diam-
eter, one turbine impeller is sufficient. If the liquid depth-to-tank diam-
eter ratio exceeds 1:3, two turbine impellers must be used. The upper tur-
bine should be located 0.5-1.0 turbine diameter below the liquid surface (E).
The lower turbine should be mounted 1 turbine diameter above, the tank bottom
(C) as shown in Figure 4-4.

n '
^-
r---|
r-r
n
n
t
>.j
— 14
I 1- " -1 t
1 ' D
C

bw ;
ba

—be
ba = Baffle Bottom Clearance
be = Baffle Clearance
bw = Baffle Width
C = Turbine Bottom Clearance
D = Turbine Diameter
E = Upper Turbine Depth
n = Rotation Speed
T = Tank Diameter
W = Turbine Blade Height
Z = Liquid Depth
Figure 4-4. Characteristics of a mixing tank and standard turbine.
4.3 Baffles

Baffles are used with turbine impellers and on-center vertically mounted
propeller mixers. The standard width for baffles (bw) used with turbine
63
-------
mixers is one-twelfth the tank diameter, while the baffle width used with
propeller mixers is one-eighteenth the tank diameter (1). The standard num-
ber of baffles is four, mounted 90° apart. To prevent the formation of dead
spots, the baffles should be mounted with a clearance between the baffles and
the side wall (be). This space should be 10%-15% of the baffle width; how-
ever, this clearance ranges between 2.54 and 7.6 cm (1 and 3 in). For tanks
with flat or slightly conical bottoms, the baffles should end a minimum dis-
tance of one-half baffle width above the tank bottom (ba) (3). Also, baffles
should extend at least 15.2 cm (6 in) above the liquid level.

4.4 Shafts and Drives

The small, high-speed turbine impellers, used with portable mixers, are con-
nected directly to the drive motor. Slow-speed turbines have a shaft coupled
directly to a speed reducer, in which case torsional bending stresses are
transmitted to the reduction gears. The shaft can be isolated from the speed
reducer by independent bearings. This provides a flexible coupling that does
not transmit stresses to the speed reduction gears.

Shaft lengths are a function of the tank depth. Short shafts and those up to
4.57 m (15 ft) in length usually need no support, but longer shafts usually
require foot bearings. These should be avoided if at all possible. Closed
tanks require mechanical seals or stuffing boxes.

4.5 Energy Requirements

A general rule often used by designers for sizing mixer motors is 0.2 kW/m3(l
hp/1,000 gal) of tank volume (2). Using the ideal turbine design as a start-
ing point, the mixer speed (n) and exact turbine diameter (D) can be deter-
mined for the given slurry concentration in the tank. This procedure in-
volves applying correction factors to the turbine diameter for differences
between the ideal turbine and the one being designed in turbine depth and the
specific gravity of the slurry (3).

The power dissipated during mixing is a function of the turbine diameter, its
shape and speed, and the specific gravity of the fluid. Increasing the diam-
eter, speed, or specific gravity requires an increase in the motor size
needed to impart the same mixing to the fluid, unless another property is
correspondingly reduced. Thus, turbine diameter or speed must be reduced for
•operation in a fluid whose specific gravity is greater than 1.0 if the motor
size is to remain unchanged.

To begin this sizing procedure, the designer should estimate the motor size
needed at 0.2 kW/m3 (1 hp/1,000 gal) of mixing tank capacity and the turbine
impeller diameter as one-third the tank diameter. The turbine should be
located 1 turbine diameter above the tank bottom. The following procedure is
then applied to determine the turbine speed and the corrected turbine diam-
eter. Tables 4-1 and 4-2, relating ideal turbine diameter, speed, and motor
horsepower, have been compiled for radial and axial turbines with one and two
blades, operating in water. The procedure for axial turbines is the same as

64
-------
01
TABLE 4-1

DIAMETER (IN) FOR RADIAL TURBINES IN WATER
(SINGLE AND DUAL TURBINES)
r/mina

A on
'tf.O
350
280
230
190
175
155
125
115
100
84
68
45
37
30

15
16
17
18
21
23
24
27
30
39
44
50
2
DT

13
14
15
16
19
21
22
24
27
35
40
46

16
17
18
20
23
25
26
29
33
42
47
54
3
DT

14
15
16
17
21
23
24
26
30
38
42
48

18
19
21
22
26
28
29
32
36
46
53
59
5
DT

16
17
18
19
23
25
26
29
33
42
47
54
7%
ST

19
21
23
24
28
30
32
34
38
51
56
66
DT

17
19
21
22
25
27
28
32
35
46
50
58
10
ST

20
22
24
26
29
31
33
36
41
54
60
70
DT

18
20
22
23
26
28
30
34
38
48
54
62
15
ST
16
19
21
24
26
28
31
34
36
39
45
58
66
76
DT
17
19
22
24
25
28
31
33
36
41
51
59
66
Motor horsepower
20
ST
17
20
22
25
27
29
33
36
38
42
48
62
70
80
DT
18
20
23
25
27
30
33
35
39
44
54
64
72
25
ST
18
21
24
27
29
30
34
37
40
44
51
64
72
82
DT
19
21
24
26
28
32
34
36
40
46
57
66
74
30
ST
19
22
25
28
30
32
36
39
41
46
53
68
74
86
DT
20
22
25
27
29
33
35
37
41
47
60
68
76
40
ST
20
23
26
29
31
33
39
42
44
49
54
70
76
90
DT
21
23
27
29
30
34
37
40
43
50
62
72
82
50
ST
21
24
27
31
33
35
40
43
46
51
57
74
82

Dl
22
25
28
30
32
36
39
41
46
52
66
76

aRotational speed
-------
en
TABLE 4-2

DIAMETER (IN) FOR AXIAL TURBINES IN WATER
(SINGLE AND DUAL TURBINES)
r/min

420
350
280
230
190
175
155
125
115
100
84
68
56
45
37
30

ST
11
13
15
18
19
20
21
24
26
28
31
35
39
45
51
57
2
DT
10
11
13
15
17
18
19
20
22
25
27
31
35
40
44
50

ST
12
14
16
19
20
21
23
26
28
30
33
38
43
49
56
62
3
DT
11
13
14
16
18
19
20
23
24
26
29
33
38
43
49
55

ST
14
15
18
20
22
24
26
29
31
33
37
42
48
55
62
68
5
DT
12
14
16
18
19
20
22
26
26
29
32
37
42
48
54
60
7h 10
ST
15
17
19
21
25
26
28
32
32
36
39
44
51
59
66
74
DT
13
15
17
19
21
22
25
27
29
32
35
40
44
52
58
66
ST
16
18
20
23
26
27
29
33
35
38
42
48
55
62
70
78
DT
14
16
18
20
22
24
26
29
31
33
37
42
48
55
62
70
15
ST
17
19
22
25
28
30
32
36
38
41
45
52
59
68
76
86
DT
15
17
19
21
25
26
27
32
33
36
40
45
51
60
68
74
Motor horsepower
20
ST
18
20
24
26
30
32
33
38
40
44
48
55
64
72
80
90
DT
16
18
20
23
26
27
29
33
35
38
42
48
55
64
70
80
25
ST
19
21
25
28
31
32
35
39
42
46
50
57
66
74
84

DT
17
19
21
24
27
28
31
35
37
40
44
50
56
66
74

30
ST
20
22
26
29
32
33
37
41
44
48
52
59
68
78
88

DT
18
19
22
26
28
30
32
36
38
41
45
52
60
68
76

40
ST
21
24
27
31
34
36
38
44
47
50
56
64
72
82
__

DT
19
20
24
26
30
32
33
38
40
44
49
56
64
72
--

50
ST
22
26
28
32
36
38
40
45
49
52
58
66
74
86
--

DT
20
21
25
28
31
32
35
40
42
46
50
58
66
76
__

aRotational speed
-------
that for radial turbines.

1. Using the estimated horsepower and turbine diameter values, read the
required turbine speed from Table 4-1 or 4-2.

2. Choose all applicable correction factors from Table 4-3 or 4-4, and
multiply them to obtain a Power Correction Factor (PCF).

3. Refer to Table 4-4 for the turbine diameter correction needed, de-
pending on the PCF. If the PCF value is between 0.95 and 1.10, the
turbine diameter does not need to be corrected. Otherwise, round
the turbine diameter to the nearest whole centimeter (inch). If un-
sure which way to round the number, round down to the lower diam-
eter.

4. If the diameter has been changed, check the new turbine diameter-to-
tank diameter ratio. If it is less than the desired value of 0.33,
reduce the turbine speed to the next lower value and select the
corresponding turbine diameter. Repeat steps 2 through 4 with these
new values.

This procedure has been outlined for the designer primarily for awareness of
basic factors when purchasing a mixer or mixing tank from a vendor. If the
designer, however, chooses to fabricate a tank for a given application, he
can exercise this option.

4.6 Flocculant Mixing

Flocculation refers to the physical and chemical process in which suspended
solids form larger, faster settling particles. This process immediately pre-
cedes sedimentation. Flocculation is accomplished in AMD treatment by the
addition of chemical polymers that act as bridging agents between the sus-
pended particles. The polymer solution is prepared in a mix tank and then
added to neutralized mine drainage before the clarifier or sedimentation
basin. The flocculant reaction is achieved through gentle mixing, which can
be done in a vessel equipped with a flocculation chamber. The most common
method is by normal turbulence in the transfer pipe to the clarifier. The
latter method, however, cannot always be assumed to be adequate or efficient.
Many flocculants require a specific reaction time to form the necessary
bridging for improved settling. Thus, this common simple method can be
wasteful and expensive.

The polymer makeup tank should be stirred by a low-speed, portable propeller
mixer (less than 400 r/min) to avoid destroying the polymer molecules. A
constant supply of polymer solution is usually fed from a storage tank by a
volumetric (positive displacement) pump. Centrifugal pumps should not be
used because they can destroy the polymer molecules by shear force. The tank
size depends upon the polymer feed rate, the feed concentration, and the
solution makeup schedule. Care must be taken to avoid formation of gummy,
clogging masses of polymer during the solution makeup.

67
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TABLE 4-3

RADIAL TURBINE PROXIMITY AND
LIQUID PROPERTIES FACTORS
Condition Different Factor
from Base Case Single Turbine
C/D = 0.4-0.7
C/D = Less than 0.4
In unbaffled square or
rectangular tank
Specific gravity other
Viscosity 1-10,000 cP

0.95
0.85
0.85
than 1.0 sp gr
(centi poise) 1.0
TABLE 4-4
AXIAL TURBINE PROXIMITY AND
LIQUID PROPERTIES FACTORS
Dual Turbine
1.0
0.92
0.85
sp gr
1.0

Condition Different Factor
From Base
C/D = 0.6-0.9
C/D = 0.3-0.6
C/D = Less than 0.3
Cover (Z-C) = 0.5D to
In unbaffled square or
rectangular tank
Specific gravity other
Viscosity, cP
Case Single Turbine
1.05
1.10
Not used
l.OD
0.85
than 1.0 sp gr
200 1.0
500 1.1
1,000 1.2
2,500 1.3
5,000 1.5
10,000 1.8
Dual Turbine
1.0
1.05
Not used
0.95
0.85
sp gr
1.0
1.1
1.2
1.3
1.5
1.8
68
-------
Rapid mixing of the polymer solution with mine drainage flow is desired to
insure adequate polymer dispersion. A small flash mix tank providing a de-
tention time of 10-30 seconds, with a mixer providing 26-53 kW/1,000 1/s
(2.2-4.4 hp/1,000 gal/min) of flow, can be used for this purpose, as well as
in-line mixers (4). Common practice is to inject polymer into a pipe or
flow-splitter box where the turbulence provides mixing. This is acceptable
as long as the turbulence and residence time provides complete mixing.

Gentle agitation is desired to promote floe growth. Mechanical flocculators
equipped with slowly rotating paddles (less than 60 r/min) serve this pur-
pose. Standard designs are available that can provide 30-45 minutes for floe
formation (4). Again, these units have not found much use in AMD treatment
plants, because designers have relied upon process turbulence to enhance floe
formation. Flocculator-clarifiers combine the two unit processes, elimina-
ting the need for a separate flocculation unit; however, these units are
rarely used.

4.7 Summary

There are basically three applications of mixing in AMD treatment. Mixers
are used in lime slakers and lime slurry storage tanks, flash mix tanks, and
flocculation tanks. The recommended designs for each of these are summarized
below.

Slurry mix tanks exist in a wide range of sizes. Smaller tanks, less than
11.355 m3 (3,000 gal), are equipped with either angular-mounted, portable
propeller mixers or with fixed-mounted propeller mixers with baffles. Large
tanks, greater than 11.355 m3 (3,000 gal), are usually equipped with top-
mounted turbine mixers and also contain baffles. Mixer horsepower require-
ments can be estimated from the general rule of 0.2 kW/m3 (1 hp/1,000 gal).

Flash mix tanks are sized to provide a detention time between 10 and 30
seconds. The mixer employed is almost always an angular-mounted, portable
propeller mixer sized accordingly, but limited to 2.24 kW (3.0 hp).

To enhance chemical coagulation and improve settling performance, mechanical
flocculators can be used. A flocculation detention time of 3-5 minutes
should be provided before settling. Several types of flocculators are avail-
able, including vertical and horizontal paddle and turbine units. Floccula-
tor suppliers should be contacted to provide detailed information on the
basin and flocculation unit.

In conclusion, this chapter has presented the more basic, common sense fac-
tors influencing mixer and mixing tank design. The designer should be aware
that good and complete mixing is not always easily accomplished.

4.8 References

1. Perry, R.H., C.H. Chilton, and S.D. Kirkpatrick. Chemical Engineers'
Handbook. 4th ed. McGraw-Hill, New York, 1963.

69
-------
2. Beals, J.L. Mechanics of Lime Slurries. Proceedings: 37th Interna-
tional Water Conference, Engineers Society of Western Pennsylvania,
Pittsburgh, Pennsylvania, October 1976.

3. Casto, L.V. Practical Tips on Designing Turbine Mixer Systems. Chemi-
cal Engineering, 79(1), January 10, 1972.

4. The American Water Works Association, Inc. Water Quality and Treatment.
3rd ed. McGraw-Hill, New York, 1971.

4.9 Other Selected Readings

McCabe, W.L., and J.C. Smith. Unit Operations of Chemical Engineering. 3rd
ed. McGraw-Hill, New York, 1976.

Metcalf and Eddy, Inc. Wastewater Engineering. McGraw-Hill, New York, 1972.
70
-------
CHAPTER 5

IRON OXIDATION
5.1 Introduction
Acid mine drainage often contains significant concentrations of iron result-
ing from the oxidation of pyritic minerals associated with coal. The mining
process exposes these pyrites or iron sulfides to the atmosphere and to mois-
ture, thus causing their oxidation to soluble ferrous sulfate salt and sul-
furic acid. These salts readily dissolve in water forming mine drainage. If
there is an overabundance of acid salts to the alkalinity in the water, the
mine drainage will be acidic.

As mine drainage is formed, iron is first present in the ferrous (Fe"^) form.
Ferric iron is much less soluble than ferrous and can be precipitated as a
hydroxide to effluent quality levels below the minimum allowable pH of 6.0
(1, 2). As shown in Figure 5-1, minimum ferric solubility occurs at a pH of
8.0, while ferrous does not reach minimum solubility until the pH approaches
11.0 (2, 3). At the maximum allowable discharge pH of 9.0, ferrous iron is
soluble to about 4 rng/1, which is in excess of the discharge limitations for
new sources. Therefore, in most mine drainage treatment systems, such as any
of the chemical neutralization processes, it is imperative to oxidize any
ferrous iron to the ferric form so it can be effectively removed at a lower
system pH. The methods available to accomplish this oxidation are natural or
mechanical aeration, chemical oxidation, and biological systems.

5.2 Aeration Systems

Ferrous iron, when exposed to oxygen, will oxidize to ferric iron at a rate
determined by the ferrous iron concentration, the dissolved oxygen concentra-
tion, and the pH of the solution. At pH values greater than 6.0, the reac-
tion occurs according to the following rate equation:

- k (Fe~) (02) (OH-)2
The reaction is first order with respect to the ferrous iron and the dis-
solved oxygen concentrations. This means that the oxidation rate decreases
as the concentration of either decreases. The reaction rate is second order
with respect to the hydroxyl ion (OH") concentration for pH values greater
than 6.0. Thus, the reaction rate increases 100 times for each one-unit rise

71
-------
PO
O)
E
tl
LL
1000-
500-
200-

100-
10-

3-
J-
o
1-

0.1-

0.01

0.001-
OPERATING
SOLUBLE
RANGE
FERRIC
HYDROXIDE
Fe (OH).
INSOLUBLE
EPA LIMITATION
(30-DAY AVERAGE )
i
8
FERROUS
HYDROXIDE
SOLUBLE
12
14
,H
Figure 5-1. Solubility of ferric and ferrous iron at various pH.
-------
in pH above pH 6.0. The rate of ferrous iron oxidation can be classified as
extremely slow (days) at a pH of less than 3.0, slow in the pH 3.0-6.0 range,
moderate to fast in the pH 6.0-8.0 range, and rapid above this point (3).

Ferric iron will precipitate as ferric hydroxide sufficiently to meet efflu-
ent limits at a pH of 5.0. At this pH, unfortunately, the oxidation rate for
ferrous iron is slow. The reaction does not increase to an acceptable rate
(minutes) until the pH is 8.0 or greater. When iron is mostly in the ferrous
form, aeration processes are most efficiently operated within a pH range of
8.0-9.0. In this pH range, the oxidation reaction takes place in a matter of
minutes, and the controlling parameter for the design of the aeration unit
becomes a function of the oxygen transfer efficiency and not the chemical re-
action of oxygen and iron. The aerator should be sized to provide dissolved
oxygen saturation in the aeration basin, assuming maximum oxygen transfer.

5.2.1 Oxygen Requirements

The amount of iron to be oxidized in a given time will determine the capacity
of the aeration system. If the system cannot meet this oxygen requirement,
then oxidation will be incomplete. The rate of oxidation increases as the
dissolved oxygen concentration in water increases to its saturation point.
Any additional aeration capacity beyond saturation does not benefit the
ferrous iron oxidation rate.

The rate of oxygen transfer into water depends on the initial oxygen deficit;
i.e., the lower the initial dissolved oxygen concentration, the easier it is
to dissolve oxygen by aeration. Conversely, the ferrous iron oxidation rate
depends on the dissolved oxygen concentration, with the maximum rate occur-
ring at saturation. A compromise between these two mechanisms must be met to
optimize the aeration process. In water with a pH greater than 8.0, the oxi-
dation rate is rapid enough that maintaining dissolved oxygen near saturation
is unnecessary. If aeration is performed at a pH less than 8.0, then a
fairly high dissolved oxygen level must be maintained for efficient ferrous
iron oxidation (4).

The chemical equations for the oxidation of ferrous iron to ferric and its
subsequent hydrolysis are as follows:

Fe+2 + %02 + H+ -> Fe+3 + %H20 (10)

Fe+3 + 3H20 -»• Fe(OH)3 + 3H+ (11)

The stoichiometric relationship of this equation indicates that 1 kg of
oxygen will oxidize 7 kg of ferrous iron under ideal conditions. During this
oxidation and hydrolysis reaction, 1 mol of acidity (as H2S04) is formed for
each mole of ferrous iron that is oxidized. Therefore, sufficient excess
alkalinity must be added during neutralization to compensate for the acidity
formed and to maintain optimum pH conditions during aeration for ferrous iron

73
-------
oxidation.

Whatever type aeration system is used, it must meet the oxygen demand for
ferrous iron oxidation. The theoretical oxygen demand for any mine water can
be calculated by Equation 12 (4).

02 = Qw x Fe x 5.16 x 10"^ (12)

where 02 = Theoretical 02 demand (kg 02/hr)

Qw = Acid mine drainage flow rate (1/s)
+ 2
Fe = Fe initial concentration (mg/1)

02 = Qw x Fe x 7.14 x 10~5

where 02 = Theoretical 02 demand (Ib 02/hr)

Qw = Acid mine drainage flow rate (gal/min)

Fe = Fe 2 initial concentration (mg/1)

Oxygen makes up about 21% of air by volume. Only a fraction of the oxygen
that conies in contact with the water is actually absorbed. This fraction,
expressed as the oxygen transfer efficiency, differs for each aeration system
under actual operation conditions. The total air needed to supply the theo-
retical oxygen demand, taking into account the above considerations, can be
calculated by Equation 13 (4).

6.324 x 02
ga - £

where Qa = Total air demand (standard m3/nrin)

02 = Theoretical oxygen demand (kg/hr of oxygen required)

E = Oxygen transfer efficiency (as %)

101.36 x 0.
Qa = £ -

where Qa = Total air demand (standard ft3/min)

74
-------
0 = Theoretical oxygen demand (Ib/hr of oxygen required)

E = Oxygen transfer efficiency (as %}

Oxygen transfer efficiencies (E) range from 3% to 25%, depending upon the
type and size of aerator used and the depth of submergence. Manufacturers of
aeration equipment should be consulted for the transfer efficiency of their
specific aeration system. A 10% transfer efficiency is generally used when
the exact value is not available. The maximum air requirements will be at
the highest anticipated air temperature when the density is least.

In addition to providing the required oxygen for ferrous iron oxidation, the
aeration system must be capable of keeping the ferric hydroxide solids and
unreacted reagent in suspension. If there is not sufficient mixing, these
solids will settle to the bottom of the aeration basin. Settled solids, in
effect, reduce the aeration volume and aeration time, leading to incomplete
ferrous iron oxidation. The aerator must be sized to meet both oxygen and
mixing requirements. In mine drainage with very high iron concentrations,
the power required for oxygen transfer is usually sufficient to meet the mix-
ing requirements. In most other cases, the aerator size may be determined by
the mixing requirements.

Aeration is the conventional method used for iron oxidation. The aeration
processes, which solubilize atmospheric oxygen in mine drainage, can be
classified into four categories; i.e., mechanical surface aeration, diffused
air aeration, submerged turbine aeration, and cascade aeration.

5.2.1.1 Mechanical Surface Aeration

Mechanical surface aeration introduces atmospheric oxygen into water by
rotating blades positioned below the static water level in an aeration basin.
The turbulence created by the aerator disperses air bubbles and keeps the
iron floe in suspension. Oxygen is absorbed by the water at the air-water
interface following the classical laws of gas absorption. The dissolved
oxygen then reacts with the ferrous iron to complete the reaction.

Mechanical surface aerators (Figure 5-2) are the more popular choice by the
designers of mine drainage treatment plant aeration systems. They can be
either slow-speed turbines or high-shear, high-speed, axial flow turbines.
Both types pull water from beneath the rotor blades and spray it across the
water surface. The oxygen transfer occurs during this splashing. Mechanical
aerators are either floating or structurally supported. The oxygen transfer
efficiency of surface aerators is generally the highest of the aeration
systems considered, usually in the 1.8- to 2.1-kg 02/kW-hr (3.0-3.5 Ib 02/hp-
hr) range (5). Floating aerators perform best in circular basins. These
basins should be limited to a depth of 3.0 m (10 ft) to eliminate the need
for a draft tube or submerged piping. Slow-speed aerators are generally used
because more water is pumped and a better efficiency is obtained for the same
horsepower than with a high-speed aerator. Mechanical aerators can be used
in both square and circular basin configurations.

75
-------
MOTOR
DIFFUSER HEAD
n
UUL>
^POLYURETHANE
FILLER — ^
INTAKE VOLUTE-
PROPELLER
Figure 5-2. Typical floating mechanical aerator.
5.2.1.2 Diffused Air Aeration

Diffused air aeration systems have rarely been used for the treatment of acid
mine drainage. Although diffused systems are used extensively in sewage
treatment, gypsum and iron or aluminum precipitates encountered with AMD tend
to make this system unsuitable for this application.

Diffused air systems introduce air, supplied by a blower, to the water
through diffusers placed near the tank bottom. Oxygen transfer occurs as the
air bubbles rise to the surface. The diffuser devices deliver air through a
porous medium, such as carborundum, nylon, or saran. These have a better
oxygen transfer efficiency than coarse bubble diffusers, which deliver air
through perforated pipes. This advantage is offset, though, by the cleaning
and replacement costs caused by clogging of the porous medium with precipi-
tates in the water. A gypsum problem resulting from the treatment of AMD
would prohibit the use of a diffused air system. Coarse bubble devices

76
-------
generally have a lower oxygen transfer efficiency and a higher capital cost
than mechanical aerators, making them impractical for AMD treatment.

The air and oxygen requirements for diffused aeration have already been
specified (see Equations 12 and 13). The manufacturer will specify the opti-
mum configuration and spacing for the desired oxygen transfer rate and mixing
needs. Diffused air aeration basins can be constructed of concrete, are
usually rectangular in shape, and operate at depths of 3.7-4.6 m (12-15 ft).
The diffusers work most efficiently when placed to one side in a single 1-i»«
along the horizontal axis of the basin (4). The basin design and diftuser
placement are critical for good mixing and oxygen transfer.

5.2.1.3 Submerged Turbine Aeration

Submerged turbine aerators combine both of the previous methods. An air
sparger is located near the bottom of the tank and above it is one or more
turbine impellers. The shearing action of the rotating impeller produces
small bubbles necessary for good oxygen transfer efficiency.

Submerged turbine aerators offer few benefits over mechanical surface aera-
tors. The higher, degree of oxygen transfer is of little benefit to ferrous
iron oxidation. These are less efficient than surface aerators and require
more power to meet the oxygen demand. Standard oxygen transfer efficiencies
are 1.0-1.3 kg 02/kW-hr (1.5-2.0 Ib 02/hp-hr) for single impeller turbines,
and 1.7-2.1 kg 02/kW-hr (2.5-3.0 Ib 02/hp-hr) for dual impeller turbines (6).
Submerged piping is also required for these turbine units. The submerged
turbine units are fixed-mounted and operate well in winter conditions.

5.2.1.4 Cascade Aeration

Aeration by gravity or cascading is a practical, inexpensive, and popular
method employed for oxidizing low ferrous iron concentrations in mine drain-
age (less than 50 mg/1). The most common type of cascade aerator is an open
trough, usually a half-round pipe, lined with splash blocks to induce turbu-
lence and increase aeration.

Other methods of cascade aeration are stairsteps, falls, a combination of
both, and wide, shallow, open flumes.

Assigning a scientific or technical approach to the design of any type of
cascade aerator is difficult and, for the most part, can be classified as a
rough estimate. Many variables, such as mine water temperature, initial
oxygen deficit, oxygen transfer, and air temperature, influence aeration
efficiency. It is difficult to predict efficiency unless actually performed
at full scale.

Therefore, this chapter will present an estimate of the design of an aeration
trough. The designer can then refine the device (i.e., increase length of
the trough, insert more splash blocks) when the unit goes on-line. Also,
this uncertainty will require the designer to provide flexibility and margin

77
-------
for error in the design and erection.

5.2.2 Aeration Trough Design

The maximum rate of iron oxidation occurs at pH 8.5 provided oxygen is avail-
able. Therefore, an operating pH less than ideal will require longer aera-
tion times and usually renders this method impractical for waters with high
ferrous iron concentrations (greater than 100 mg/1).

Detention times for this method are generally less than 5 minutes. The
detention time for any trough can be found using Equations 14 and 15.

V = A (L) (14)

where V = volume of trough m3 (ft3)

A = cross-sectional area of flow m2 (ft2)

L = length of trough m (ft)

D = (15)
where D. = detention time (min)

V = volume of trough m3 (ft3)

Q = m3/min (gal/min)

The difficult part of the equation is computing the cross-sectional area of
the flow in the pipe at a given depth. To alleviate this problem, Table 5-1
has been provided. This table presents the cross-sectional area per linear
meter (foot) of the most commonly used pipes, 15.2-121.9 cm (6-48 in), flow-
ing at various depths. Only the lower depths, those less than half the diam-
eter, are applicable. Volume is merely the cross-sectional area multiplied
by the trough length (see Equation 14).

With the volume of the trough in cubic meters (gallons) known, the detention
time can be easily computed. A rule of thumb among designers has been 0.1
min/mg/1 of ferrous iron at pH 8.5. Therefore, if a mine drainage contains
30 mg/1 of ferrous iron neutralized to 8.5, 3 minutes of trough detention
time with vigorous aeration should be provided. The factor is a good start-
ing point but in no way guarantees complete oxidation. This can be deter-
mined only in the field by analysis of the drainage.

The number of splash blocks, size, and spacing are chosen arbitrarily.

78
-------
TABLE 5-1

CROSS-SECTIONAL AREA OF FLOW IN CIRCULAR PIPEa
Depth
Flow
cm
1.27
1.91
2.54
3.18
3.81
4.45
5.08
5.72
6.35
6.99
7.62
8.26
o QQ
9^53
10.16
10.80
11.43
12.70
13.34
13.97
of
(d)
in
0.50
0.75
1.00
1.25
1.50
1.75
2.00
2.25
2.50
2.75
3.00
3.25
3.50
3.75
4.00
4.25
4.50
4.75
5.00
5.25
5.50
Diameter of
cm in cm in cm
15.2 6.0 20.3 8.0 25.4
.001 .008 .001 .009 .001
.001 .014 .002 .017 .002
.002 .022 .002 .025 .003
.003 .030 .003 .035 .004
.004 .038 .004 .045 .005
.004 .048 .005 .056 .006
.005 .057 .006 .068 .007
!oio

in
10.0
.010
.019
.028
.039
.051
.064
.078
!l07

cm
30.5
.001
.002
.003
.004
.005
.007
.008
!oii
.013
.014

in
12.0
.011
.020
.031
.043
.057
.071
.086
!l!9

Pipe
cm
38.1
.001
.002
.003
.005
.006
.007
.009
.'012
.016
.018
.020
.022

(D)
in cm
15.0 45.7
.013 .001
ri9?
.035 .004
f)AQ
.064 .007
nan
.097 .010
^5
.134 .014
.175 .018
IOC
.218 .022

"
_,_
'

in cm in
18.0 53.3 21.0
.014 .001 .015
.039 .004 .042
.070 .007 .076
.107 .011 .117
.149 .015 .162
.197 .020 .211
.242 .024 .264
.292 .030 .319
.345 .035 .378

'
047 50?

cm in
61.0 24.0
.001 .016
.004 .045
.008 .082
.012 .125
.016 .174
.021 .227
.026 .284
.032 .344
.038 .408
.044 .474
.050 .543
aThe units for the area of flow in pipe are m2 when the diameter is given in cm and ft2 when the diameter
is given in in.
-------
TABLE 5-1 (continued)
Depth of
Flow (d)
in
Diameter of Pipe (D)
CO
o
cm
1.27
2.54
0.5
1.0
3.81 1.5
5.08 2.0
6.35 2,5
7.62 3.0

8.89 3.5
10.16 4.0
11.43 4.5
12.70 5.0

13.97 5.5
15.24 6.0
16.51 6.5
17.78 7.0

19.05 7.5
20.32 8.0
21.59 8.5
22.86 9.0

24.13 9.5
25.40 10.0
26.67 10.5
27.94 11.0

29.21 11.5
30.48 12.0
31.75 12.5
33.02 13.0

34.29 13.5
35.56 14.0
36.83 14.5
38.10 15.0
cm
58.6
.002
.004
.008
.012
.017
.022
.028
.034
.040
.047
.054
.061
.068
—
^ _
—
--

in
27.0
.017
.048
.087
.133
.185
.241
.302
.367
.436
.507
.581
.658
.737
—
__
—
—

cm
76.2
.002
.005
.009
.013
.018
.024
.030
.036
.043
.050
.057
.065
.073
.081
.089
—
—

in
30.0
.018
.050
.092
.141
.195
.255
.320
.389
.462
.538
.617
.699
.783
.870
.960
—
—

cm
83.8
.002
.005
.009
.014
.019
.025
.031
.038
.045
.053
.060
.069
.077
.085
.094
.103
.113

in
33.0
.019
.053
.096
.148
.205
.269
.337
.410
.486
.567
.651
.738
.827
.920
1.015
1.112
1.211

cm
91.9
.002
.005
.009
.014
.020
.026
.033
.040
.047
.055
.063
.072
.081
.090
.099
.109
.118
10,0

in
36.0
.020
.055
.101
.154
.214
.281
.353
.429
.510
.595
.683
.774
.869
.967
1.067
1.170
1.275
1 "3ft9

cm
121.9
.002
.006
.011
.017
.023
.030
-.038
.046
.055
.065
.074
.084
.095
.105
.116
.128
.140
1 5?
i fid

17fi
IRQ

pi c
. £10
ppo

in
48.0
.023
.064
.110
.179
.250
.327
.411
.500
.595
.694
.798
.907
1.019
1.135
1.254
1.377
1.502
1 (\*%. 1

JL • / Oo
1 RQ7
p rioo
? 1 7?
p qi -3
c.olo
p AC.J

cm
137.2
.002
.006
.012
.018
.025
.032
.041
.049
.059
.069
.079
.090
.101
.112
.124
.136
.149
Ifi?
17C

IRQ

?1fi
pon

?dd
pcq
?7d
ppq

in
54.0
.024
.064
.124
.190
.265
.348
.437
.532
.633
.739
.850
.966
1.086
1.210
1.338
1.469
1.604
1 7d?
1 RRd
Loot
p n?Q
p i 71:

p AT;

p cop
p 700
p qdfl
o, inq

cm
152.4
.002
.007
.012
.019
.026
.034
.043
.052
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IRK
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ppq

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pen
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pqi
on?
ooo

.151
in
60.0
.025
.071
.131
.201
.280
.367
.461
.562
.669
.782
.899
1.022
1.149
1.281
1.417
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1 Rd7
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-------
Splash blocks are usually tack-welded to the bottom of the trough (corrugated
metal pipe) and inclined to cast the flow into the air. Figure 5-3 shows a
typical splash block and the placement pattern. This is one example of many
configurations used today. However, any concept for inducing the required
turbulence is suitable; the possibilities are limited only by the imagina-
tion.
3' - 4'
FLOW
I I
ELEVATION
PLAN
Figure 5-3. Typical splash block placement pattern.
The application of gravitational or cascade aeration is limited by the ini-
tial ferrous iron concentration (less than 50 mg/1) and the neutralization pH
(8.5). Any variance from these criteria, especially pH, will affect and
inhibit complete iron oxidation by this method. A good trough velocity is
0.9-1.2 m/s (3-4 ft/s); flow depth should not exceed one-fourth of the pipe
diameter. This will allow adequate freeboard to contain splash and reduce
overflow.

When the trough exceeds 366 m or 1.2 m/s x 300 s (1,200 ft or 4 ft/s x 300 s)
or when ferrous iron concentrations are greater than 50 mg/1, a mechanical
aerator should be considered.

5.2.3 Aeration Basin Design

Proper design of the aeration basin is necessary for the efficient oxidation
of ferrous iron. The sizing of this unit is important to insure that ade-
quate, but not excessive, aeration periods are provided. Attention must also
be given to the basin plan, depth dimensions, and inlet and outlet struc-
81
-------
tures. Aeration basins are most often earthen units lined with riprap or
asphalt, but can also be constructed of concrete or steel.

Mathematical models have been used to predict aeration times for the oxida-
tion of ferrous iron at varying pH ranges (1, 2, 4, 15). Due to the variable
nature of mine drainage and the effects other dissolved ions may impose on
the reaction, laboratory tests are the most reliable way to optimize the
aeration system design. The laboratory tests should follow the method out-
lined at the end of this chapter. To insure plant operation at peak condi-
tions, the tests should be performed on a sample containing the maximum
ferrous iron concentration and at the lowest anticipated operating pH. Also,
the tests should be carried out at the desired operating pH or at several
others if a comparison of results is desired.

The detention time needed for ferrous iron oxidation, as determined by the
laboratory tests, must be multiplied by a safety factor for the design of the
full-sized aeration basin to reduce the possibility of short-circuiting
through the aeration basin. The chances of short-circuiting are greater with
the shorter detention periods. The calculated detention time should be
multiplied by the safety factors listed in Table 5-2.
TABLE 5-2

AERATION DETENTION TIME SAFETY FACTORS

Calculated Detention Time Safety Factor
min
> 16 2.0

11-15 3.0

6-10 4.0

3-5 8.0

1-2 10.0

<1 >10.0
The volume of the aeration basins is easily determined using the mine drain-
age flow rate and the calculated detention time as follows:

V = QDtf (16)

82
-------
where V = Volume, m3 (gal)

Q = Flow, m3/s (gal/min)

D. = Detention Time, s (min)

f = Safety Factor (from Table 5-2)

The design of the aeration basin must be made in conjunction with the se-
lected aerator. The aerator must be positioned in the basin so the entire
volume is aerated and well mixed, thus eliminating dead spots where solids
will accumulate. Aerator manufacturers will recommend the surface area and
depth for complete oxygen dispersion and mixing. Therefore, the designer
must know these limitations. Manufacturers will also specify the aerator
size in kilowatts, kW (horsepower, hp), and the oxygen transfer rate, kg
02/kW-hr (Ib 02/hp-hr). Using this data, an efficient and proper aeration
system can be designed.

The oxygen transferred can be calculated by multiplying the aerator kilowat-
tage (horsepower) by the oxygen transfer rate.

02 supplied = 02 transfer rate x aerator power

kg 02/hr (Ib 02/hr) = kg 02/kW-hr x kW (Ib 02/hp-hr x hp) (17)

The designer must determine whether one or more aerators will meet the oxygen
requirement. This decision will be based on an evaluation of the capital and
operating costs and the ability of the aerator(s) to keep the basin contents
mixed.

The size of tank matters also. A floating aerator will generally be effec-
tive only in a 7.6-m (25-ft) diameter pattern and a 3.1- to 3.7-m (10- to
12-ft) depth. Typical aeration basin depths range between 1.5 and 4.6 m (5
and 15 ft). The depth selected will depend upon the detention required, the
basin surface area, and the mixing characteristics of the aerator(s). In the
case of mechanical surface aerators, draft tubes can be added if the basin
depth becomes excessive. The draft tube is an extension that causes a deeper
intake, increasing the effective operating depth. Turbine and diffused aera-
tion systems have optimum operating depths, and the equipment manufacturer
should be consulted.

When a single surface aerator is used, the ideal basin shape is circular.
This usually eliminates dead spots where solids accumulate. Also, this geo-
metric shape reduces the possibility of short-circuiting, which reduces
aeration time. When using more than one aerator, a rectangular shape may be
preferred. For smaller flows, minimum basin sizing as stipulated by the
manufacturer may overrule earlier decisions on the basin size as calculated
for detention requirements.

83
-------
Design and construction of the aeration tank should follow standard engineer-
ing practice. Steel and concrete construction should follow American Insti-
tute of Steel Construction and American Concrete Institute specifications. If
an earthen basin is to be used, consideration must be given to the soil and
geology, basin location, and construction procedure. Stable slopes should be
established, with typical values as 3:1 for the outside slope and 2:1 for the
inside slope. The soil should be compacted properly to insure stability.
The basin bottom and inside slopes should be lined with clay, concrete,
asphalt, or a synthetic liner. When using clay as a liner, riprap at the
water periphery will prevent erosion caused by wave action.

Regardless of the type of aeration basin, the inlet and outlet structures
must be designed to minimize short-circuiting. They should be located at
opposite ends of the aeration tank, allowing maximum use of the aeration tank
volume. The water level in the tank is determined by the elevation of the
outlet structure. Outlet structures are typically weirs, weir boxes, or open
vertical pipes. A baffle should be placed in front of the outlet structure
to separate the turbulent aeration zone from the outlet zone. The inlet
structure should allow the water to free-fall into the tank, thus eliminating
hydraulic problems. A direct discharge by means of a pipe into the aeration
tank is adequate, because the mixing from the aerator will disperse the in-
fluent water. Provisions for access to both floating and supported aerators
should also be considered.
5.3 Chemical Oxidation

Besides aeration, several chemicals have been utilized for iron oxidation.
This chapter presents those chemicals that have potential use for low ferrous
iron drainages.

5.3.1 Iron Oxidation by Ozone

Ozone (03) is a reactive, gaseous allotrope of oxygen. The only practical
method presently available to produce ozone is by electric discharge in which
a current is passed through a stream of air or oxygen. In 1970, the EPA
sponsored a study of the cost-effectiveness of on-site ozone generation
facilities and larger central plants that would supply ozone via a distribu-
tion system (6). It was concluded that the central plant system was more
economical than individual on-site ozone generators except for low-iron (less
than 50 mg/1), low-flow (less than 946 m3/d) conditions.

Theoretically, 1.0 kg (2.2 Ib) of ozone will oxidize 2.3 kg (5.1 Ib) of fer-
rous iron. The same amount of acid is released during ozone oxidation as
during aeration. Assuming 86% ozone utilization, 1.0 kg (2.2 Ib) of ozone
will oxidize 2.0 kg (4.4 Ib) of ferrous iron (6).

The benefits of ozone treatment over lime aeration are as follows: (1) the
oxidation reaction is efficient and quick, allowing use of smaller reaction
basin; (2) close process control needed for lime treatment is not needed for
ozone treatment; and (3) neutralization to pH 6.0 is all that is required,

84
-------
allowing treatment with limestone as well as other alkalis. The sludge pro-
duced by limestone-ozone combination is denser than lime sludge, reducing
sludge handling requirements.

5.3.2 Iron Oxidation by Hydrogen Peroxide

The oxidation of iron using hydrogen peroxide (H202) merits consideration
where "specialty" conditions exist. These conditions can be generally clas-
sified as follows:

1. alkaline mine drainages (pH greater than 6.0) with low oxygen re-
quirements for iron oxidation;

2. where pH adjustment is made merely for more favorable iron oxida-
tion rates;

3. as a supplemental source where existing facilities become iron-over-
loaded and expansion is impossible.

No steadfast rules can be presented as to when to consider hydrogen peroxide.
An overall economic evaluation of the particular application must be per-
formed. Comparing reagent cost against mechanical aerator operating cost,
which includes capital reinvestment for equipment and amortization, indicates
that the hydrogen peroxide oxidation method was more expensive; however, a
blanket elimination should not be made for every treatment plant.

The general chemical characteristics of this reagent follow, along with the
simple stoichiometric relationship necessary for estimating the theoretical
requirement.

5.3.2.1 Physical Properties, Handling, and Storage of Hydrogen
Peroxide

Hydrogen peroxide is colorless, has a distinctive pungent odor, and is com-
pletely miscible with water. It is prepared in three commercial grades, 35%,
50%, and 70% by weight; the remaining percentage is water (see Table 5-3) (7,
8). Hydrogen peroxide is neither poisonous nor flammable, but it strongly
irritates skin, mucous membrane, and eye tissue; it also decomposes to form
oxygen, which supports combustion. Additives keep decomposition of uncontam-
inated solutions to about 2%/year. Many metals, salts, dust, dirt, oil, and
rust greatly increase the decomposition rate. Thus, proper handling and
storage procedures must be followed.

Standard means of shipment of hydrogen peroxide include 57-1 (15-gal), 114-1
(30-gal), and 208-1 (55-gal) polyethylene-lined drums; 114-1 (30-gal) alumi-
num drums; special tank trucks, 7,570-15,140 1 (2,000-4,000 gal); and tank
cars, 15,140, 22,710, and 30,280 1 (4,000, 6,000, and 8,000 gal). The drums
should be accompanied with special drum rockers, wrenches, and spouts.

Hydrogen peroxide should be stored only in the original containers or in

85
-------
TABLE 5-3
PHYSICAL PROPERTIES OF HYDROGEN PEROXIDE
H202 concentrations, weight %
Active oxygen content, weight %
g H202/l
Specific gravity 20°C/4°C
Ib/gal at 20°C
Boiling point, °C
Boiling point, °F
Freezing point, °C
Freezing point, °F
Viscosity 20°C (cP)
Total vapor pressure at 30°C (mm Hg)
Heat of dilution cal/g mol of H202
at 25°C and 1 atm -84 -178 -381
35
16.5
396
1.133
9.4
108
226
-33
-27
1.11
23.33
50
23.5
600
1.20
10.0
114
237
-52
-62
1.17
18.3
70
32.9
903
1.29
10.8
126
258
-40
-40
1.24
10.1
90
42.3
1248
1.387
11.6
141
286
-11
12
1.26
5
98
46.1
1407
1.436
11.95
149
300
-3
28
1.25
3
-------
properly designed tanks made of properly prepared compatible material. Stor-
age tanks for hydrogen peroxide must be vented and stored in a clean, fire-
proof area. Hydrogen peroxide suppliers should be consulted for detailed
design requirements (see Figure 5-4).
LINE TO USE POINT
WATER FILTER
& METER
FREE LIFT
MANHOLE
COVER

JET
MIXER
ASSEMBLY

STORAGE TANK

FILL LINE

FILL
CONNECTION
WATER AND H202
BALL VALVE
HOSE
FOR WATER ADDITION
WATER
SUPPLY
LINE
FLUSH
VALVE
FLUSHING & SAFETY HOSE
LENGTH AS REQUIRED TO
REACH FILL LOCATION
Figure 5-4. Hydrogen peroxide feeding system.
5.3.2.2 Chemistry

Equation 18 expresses the reaction of ferrous iron with hydrogen peroxide
(H202).
H202 + 2Fe++ + 2H+ •* 2Fe+++ + 2H20
(18)
One mole of hydrogen peroxide oxidizes 2 mol of iron. Expressed as a weight

87
-------
ratio, 0.45 kg (1.0 Ib) of H20? will oxidize 1.5 kg (3.3 Ib) of ferrous iron.
The oxidized iron is then available to react with alkalinity to form insol-
uble ferric hydroxide.

Using Equation 18 and knowing the ferrous iron loading, the theoretical
hydrogen peroxide requirement can easily be calculated.

5.3.2.3 Cost

The capital cost for a portable or permanent feeder installation is about
$5,000. This includes duplicate pumps, a 19,000-1 (5,000-gal) storage tank,
and foundation. The reagent costs freight on board (FOB) are presented in
Table 5-4.
TABLE 5-4

HYDROGEN PEROXIDE COSTS
i/Jsa
Bulk delivery 17,000 1 (4,500 gal) tanker @ 50% 9.7 21.5

Bulk, 1,900-2,300 1 small (500-600 gal) @ 50% 10.9 24

Drums, 208 1 (55 gal) @ 50% 12.25 27

Drums, 114 1 (30 gal) @ 50% 13.64 30

Full strength (100%) 19.5 43
The daily reagent cost can be determined based on the desired method of
delivery.

5.3.3 Summary

The oxidation of iron with hydrogen peroxide can be more convenient than
economical, but merits consideration for certain applications. It can be the
primary source of oxygen, but is best used in conjunction with existing aera-
tion facilities. Proved feasible, hydrogen peroxide offers a viable alterna-
tive for iron oxidation when overloaded facilities cannot be expanded, or
where effluent pH limits can be met without pH adjustment (9).
-------
5.3.4 Other Chemical Oxidants

In- addition to ozone and hydrogen peroxide, chlorine and potassium permanga-
nate are oxidants that could be applied to ferrous iron oxidation. Chlorine
forms acid upon reaction with water, making it undesirable for AMD treatment.
Both chlorine and permanganate are expensive chemicals; consequently, neither
are economical for AMD treatment.
5.4 Biological Oxidation

Research indicates that bacteria capable of oxidizing ferrous iron exist
naturally in most acid mine drainages. These bacteria, Thiobacillus ferro-
oxidans. obtain their carbon and nitrogen requirements from inorganic sources
and their energy from the oxidation of ferrous iron. These bacteria have
been implicated as a catalyst in the formation of acid mine drainage, and
several methods have been proposed to utilize their ability to oxidize fer-
rous iron in mine drainage. Dispersed growth systems have proven unsuccess-
ful; however, fixed growth systems, such as trickling filters and rotating
biological contactors (RBC), are effective in supporting bacterial popula-
tions. Experiments by Lovell have shown that synthetic filter media best
support the bacterial growth necessary for oxidation in a trickling filter,
while rock material is subject to degradation by the acid water, causing
clogging and ponding (10, 11).

01 em and Unz have completed research on ferrous iron oxidation of mine waters
directly from the mine using rotating biological contactors (12, 13, 14).
The RBC is an aerobic treatment device, consisting of a series of four
plastic discs mounted on a horizontal shaft (see Figure 5-5). This assembly
is placed in a trough through which the wastewater flows, submerging slightly
less than half the surface area of the discs. The discs rotate slowly on the
shaft, causing the biological growth on the discs to alternate contact be-
tween air and water. This is the first application of rotating discs to AMD
treatment.

Olem and Unz obtained data from RBC pilot plants utilizing actual acid mine
drainage. They compared performance at two peripheral disc velocities, 0.32
and 0.17 m/s (63 and 34 ft/min), and at five hydraulic loadings ranging be-
tween 110 and 440 1/d/m2 (2.7 and 10.8 gal/d/ft2). They found a linear re-
lationship between ferrous iron removal and stage retention time (time per
individual compartment). This relation held for any set of operating condi-
tions. Also, a faster disc velocity produced better iron oxidation at any
constant hydraulic loading. At a given disc velocity, an increase in hydrau-
lic loading from 110 to 440 1/d/m2 (2.7 to 10.8 gal/d/ft2) resulted in an
increase in effluent ferrous iron concentration, even though the ferrous iron
oxidation rate also increased. The RBC can be expected to produce an efflu-
ent containing less than 10 mg/1 ferrous iron at loading rates up to 88 kg
Fe++ applied/d/1,000 m2of disc surface (18 Ib Fe++/d/l,000 ft ).

More recent tests have been conducted on acid mine drainages that flow over-
land for more than a mile before reaching the treatment facility, thus expos-
ing the water to stream ecology and ambient air temperatures. The data col-

89
-------
1C
o
FEED BUCKET

DRIVE SYSTEM
FEED

CHAMBER

INFLUENT
STAGES OF DISCS
EFFLUENT
Figure 5-5. Rotating biological contactor.
-------
lected during winter operation showed that the effect of low temperatures on
the biological oxidation process was not as great as expected. The removal
efficiency at 0.4°C (32.7°F), the lowest temperature recorded, dropped only
10% below that found at 10°C (50°F), the initial mine water temperature.
This effect can be negated by a lower hydraulic loading.

Although it is impossible to remove ferrous iron to 1.0 mg/1 with the RBC,
this degree of removal by a biological system is probably not necessary.
Olem and Unz restate Lovell's belief that an effluent of 10 mg/1 ferrous iron
is adequate from the RBC. Since the effluent still must be neutralized, it
is assumed that agitation during the neutralization step will provide enough
aeration to oxidize any remaining ferrous iron.

Though the capital cost of the RBC system is greater than a conventional
mechanical aeration system, the attraction of the RBC system is a significant
reduction in energy costs through lower power requirements. The horsepower
required to rotate the discs is substantially less than that required to
power an aerator. The possibility arises that limestone can be used as the
neutralizing agent, resulting in a savings in chemicals. Overall, the two
most costly operational items (power and chemical) can be reduced. With the
conventional aeration process, limestone neutralization is possible only in
water containing less than 100 mg/1 ferrous iron. Also, the RBC effluent pH
need only be raised to 6.0, acceptable for discharge, while pH 8.5 is needed
with the conventional system for efficient iron oxidation.

5.4.1 Design Procedure for a Four-Stage Configuration

Data obtained from the Olem and Unz study can be applied to the design of a
four-stage single-shaft RBC system (12). Equation 19 is used to determine
the required disc surface area.

A = fSOJLl (19)
L

where A = disc surface area (m2)

Fo = initial ferrous iron concentration (mg/1)

Q = flow (1/s)

L = ferrous iron loading (kg/d/1,000 m2) determined from Figure 5-6
120L

where A = disc surface area (ft2)

Fo = initial ferrous iron concentration (mg/1)

91
-------
Q = flow (gal/d)

L = ferrous iron loading (lb/d/1,000 ft2) determined from Figure 5-6
DC<
OC5
U1UJ
u-cc
40-

35-

30-

25-

20-

15-

10-

0
FERROUS IRON LOADING (Ib. Fed I) applied/day/IOOOsq.ft.)
10
15
20
25
30
—i
i
20
I
40
I
80
T
0 20 40 60 80 100 120 140

FERROUS IRON LOADING (kg. Fe(ll) applied/day/IOOOsq.m.)
Ib./day/IOOOsq.ft. = 4.89 kg./day/1000 sq.m.
Figure 5-6.
Design procedure for a four-stage rotating biological
contactor configuration.
The designer chooses the desired effluent ferrous iron concentration from the
RBC system listed on the vertical axis (Figure 5-6). A line is drawn horizon-
tally, intersecting the curve, and dropped vertically from this point to the
horizontal axis. Derived is the maximum ferrous iron loading (L) that will
yield the desired effluent. This value, as well as the initial ferrous iron
concentration (Fo) and the flow (Q), are applied to Equation 19, yielding the
92
-------
necessary disc area. RBC manufacturer data then must be utilized to deter-
mine the number of RBC discs needed to meet this surface area requirement.

5.5 Oxidation Rate Test Procedure

The oxidation of ferrous iron is a first-order reaction with respect to
ferrous iron concentration, meaning that the reaction rate is dependent on
the ferrous iron concentration at any point in time. This reaction rate can
be expressed as
(20)
dt
where Fe is the ferrous iron concentration at any given time and k is the
oxidation rate constant. The integrated form of the equation is

-let
Fe = Feoe Kt (21)

where Fe is the initial ferrous iron concentration and Fe is the ferrous
iron concentration at time t.

This value of the reaction rate constant k, can be determined from a graph of
log Fe vs. time. Ferrous iron oxidation data will yield a straight line when
plotted on semi log paper. The slope of the line multiplied by 2.303 equals
the rate constant, k.

k = slope x 2.303 (22)

Once the rate constant for a particular mine drainage at a given operating pH
is determined, the detention time required for oxidation of the same water
from any initial concentration to any final concentration can be determined
by Equation 23.

(Fe.)
t = 1/k x 2.303 log^-F-V (23)
where Fe. and Fe^ are the initial and final ferrous iron concentrations,
respectively.

The oxidation rate test should be performed on fresh samples of mine drainage
to minimize any natural oxidation of the ferrous iron. A large sample, pref-
erably 4 1 (1 gal) or so, should be used. The sample should be stirred con-

93
-------
tinuously.

Initially, a small sample is taken and properly preserved for analysis to
determine the initial ferrous iron concentration. Lime or caustic soda is
added quickly to raise the pH to the desired operating level. At the same
time, air is supplied through a diffuser such as an aeration stone or disc.
Samples are then removed periodically throughout the test for ferrous iron
analysis. All samples must be preserved properly with hydrochloric acid to
prevent the oxidation from continuing. The pH should be measured frequently
and maintained at the selected level.

A 1-hour test should be sufficient to generate data for determination of the
oxidation rate constant. The test should be repeated once or twice to assure
repetitive results.

5.6 References

1. Lovell, H.L. An Appraisal of Neutralization Process to Treat Coal Mine
Drainage. Technology Series Report, EPA-670-2-73-093, Washington, D.C.,
November 1973.

2. Singer, P.C., and W. Stumm. Oxygenation of Ferrous Iron: The Rate-
Determining Step in the Formation of Acidic Mine Drainage. U.S. Envi-
ronmental Protection Agency Water Pollution Control Research Series
Report, DAST-28, 14010 06/69, Washington, D.C., June 1969.

3. Skelly and Loy, Engineers and Consultants. Processes, Procedures, and
Methods to Control Pollution from Mining Activities. U.S. Environmental
Protection Agency Report 430/9-73-011, Washington, D.C., October 1973.

4. Selmeczi, J.G. The Design of Oxidation Systems for Mine Water Dis-
charges. Fourth Symposium on Coal Mine Drainage Research, Pittsburgh,
Pennsylvania, April 1972.

5. Cheremisinoff, P.N. Aerators for Wastewater Treatment. Pollution
Engineering, March 1974.

6. Seller, M., C. Waide, and M. Steinberg. Treatment of Acid Mine Drainage
by Ozone Oxidation. U.S. Environmental Protection Agency Water Pollu-
tion Control Research Series Report, 14010 FMH, Washington, D.C., Decem-
ber 1970.

7. FMC Corporation. Industrial Wastewater Treatment - A Guidebook to
Hydrogen Peroxide for Industrial Wastes. Philadelphia, Pennsylvania.

8. E.I. duPont de Nemours & Company, Inc. Hydrogen Peroxide Solutions,
Storage and Handling. Wilmington, Delaware, November 1977.

9. Cole, C.A., A.E. Molinski, N. Rieg, and F. Backus. Peroxide Oxidation
of Iron in Coal Mine Drainage. Journal of the Water Pollution Control
Federation, Vol. 49, No. 7, July 1977.

94
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10. Lovell, H.L. Studies in the Treatment of Coal Mine Drainage by Biochem-
ical Iron Oxidation and Limestone Neutralization. Pennsylvania Depart-
ment of Environmental Resources Special Research Report SR-98, February
1974.

11. . Experience with Biochemical-Iron-Oxidation Limestone-Neutraliza-
tion Process. Fourth Symposium on Coal Mine Drainage Research, Pitts-
burgh, Pennsylvania, April 1972.

12. Olem, H., and R.F. Unz. Acid Mine Drainage Treatment with the Rotating
Biological Contactor. Institute for Research on Land and Water Re-
sources Publication 93, University Park, Pennsylvania, September 1976.

13. . Acid Mine Drainage Treatment with Biological Contactors, Biotech-
nology, and Bioengineering. Vol. XIX, 1977.

14. . Microbiology Oxidation of Ferrous Iron in Coal Mine Drainage
Treatment. Sixth Symposium on Coal Mine Drainage Research, Louisville,
Kentucky, October 1974.

15. Wilmoth, R.C., J.L. Kennedy, and R.D. Hill. Observations on Iron Oxi-
dation Rates in Acid Mine Drainage Treatment Plants. Fifth Symposium on
Coal Mine Drainage Treatment Research, Louisville, Kentucky, October
1975.

5.7 Other Selected Readings

Holland, C.T., J.L. Corsaro, and D.J. Ladish. Factors in the Design of an
Acid Mine Drainage Treatment Plant. Second Symposium on Coal Mine Drainage
Research, Pittsburgh, Pennsylvania, May 1968.

Omelia, C.R. Oxygenation of Iron (II) in Continuous Reactors. U.S. Depart-
ment of Interior, Office of Water Resources Research Project A-022-NC, 1969.

Pennsylvania State University. College of Earth and Mineral Sciences. Short
Course on Controlling Water Pollution in Coal Mining, March 1975.

Singer, P.C., and W. Stumm. Kinetics of the Oxidation of Ferrous Iron.
Second Symposium on Coal Mine Drainage Research, Pittsburgh, Pennsylvania,
May 1968.
95
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CHAPTER 6

SEDIMENTATION

6.1 General Characteristics of Mine Drainage Sludge

Mine drainage sludges are generally composed of hydrated ferrous or ferric
oxides, gypsum, hydrated aluminum oxide, unused lime, varying amounts of sul-
fates, calcium carbonates, bi carbonates, and trace amounts of silica, phos-
phate, manganese, copper, titanium, and zinc (1). Hydroxides are formed by
the reaction of metal salts with hydroxyl ions from natural or added alkalin-
ity (lime, caustic).

The sludge characteristics vary with mine drainage quality and neutralization
method. Important sludge characteristics include settleability, density,
dewaterability, particle surface properties, and viscosity (2). Sludge
settleability combines the aspects of sludge settling rate and final sludge
volume. Sludge density is usually reported as percent solids by weight. It
is extremely desirable to produce a dense sludge to reduce sludge volumes and
handling and disposal costs. Sludge dewaterability is defined as the ability
of a sludge to be concentrated into a more manageable and less voluminous
form by centrifuging, filtering, or lagooning. The electrostatic charge, or
particle surface property, is important in determining how well individual
particles will flocculate into larger particles and settle out of suspension.
The viscosity of sludge measures the flowability, sometimes an important
consideration when pumping sludges.

Mine drainage neutralized with lime (hydrated or quicklime) produces sludges
that settle slowly. These hydrate-produced sludges are light, gelatinous,
and very voluminous.

The ferric hydroxide sludges associated with mine drainages have high iron
concentrations and are generally a fluffy mass with very low solids, usually
less than 1%. Sludge production increases as the iron and aluminum concen-
trations increase, averaging 5%-10%, and can be as high as 30% of the total
treated volume. Sludges formed in the pH range of 6.4 to 7.2 have the best
settling properties (2), but the iron oxidation rate is poor. The ideal pH
for iron oxidation is around pH 8.5.

Sludges generated from mine drainage neutralized with limestone have a higher
density than those generated from other lime products. The volume of lime-
stone sludge can be as little as 20% that of sludge neutralized using lime,
and the solids content can be up to 15 times greater. This results in a
smaller volume required for sludge disposal. Sludges produced from highly
acidic waters neutralized with caustic soda (NaOH) have very low densities

96
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and resist compaction despite long storage detention times. Even the low
acidic waters, when neutralized with caustic, will produce large sludge
volumes. Sludges formed from soda ash (sodium carbonate, Na2C03) neutraliza-
tion compact to densities somewhere between those generated by lime and lime-
stone.

Sammarful (3) reported that mine drainage neutralized with carbonates yields
a granular, dense sludge, while hydroxide neutralizing agents produce a
semigelatinous sludge. Figure 6-1 illustrates the relationship of settling
time to sludge volume for various neutralizing agents (4).

Polyelectrolyte addition may be used to improve sludge settling rates. Poly-
electrolytes are water-soluble, high-molecular-weight, organic polymers that
may be cationic (positively charged) or anionic (negatively charged). These
charged polymers adhere to sludge particles and improve settling. Determina-
tions about whether cationic and anionic polyeletrolytes should be used, and
about what polymer dosages are proper, depend on sludge characteristics.
Treatability tests provide a means of determining these variables.

Lovell (1) separated AMD sludge settling rates into three types or classifi-
cations of settling phenomena. Type 1 settling is associated with the neu-
tralization of mine drainage containing high concentrations of iron and alu-
minum. The volume of the sludge (mostly ferric hydroxide) usually amounts to
5%-10% of the average daily treated volume. Type 1 sludge settles rapidly
with a well-developed liquid-solid interface. This most common type of mine
drainage settling phenomenon is illustrated by Figure 6-2 (5). In part I of
the figure, the precipitated sludge is a homogeneous mixture throughout the
sample. In II, stratification begins and the particles in Zone D (bottom)
begin settling onto already settled particles. Water becomes trapped inside
the layers, forming a gelatinous mass. The adjacent upper layer (Zone C) is
a transition zone characterized by a suspended solids concentration lower
than Zone D but greater than Zone B. The supernatant, Zone A, develops as
the liquid-solids separation is completed. Settling continues, as illus-
trated in III, and Zones A and D increase in depth while B and C decrease.
In IV, only two zone areas (A and D) remain, with a large majority of the
solids present in Zone D. At this point, Zone D begins to compress. Compac-
tion forces exerted by individual particle density cause the bridging of floe
to break down as the structure is overcome by its own weight. The weakly
trapped water is displaced by the weight of the sludge as it achieves equi-
librium with the compressive strength of the floe. This volume occupied by
the sludge usually represents the final volume.

Type 2 settling pertains to sludge generated from mine drainage containing
low concentrations of iron and aluminum with a pH of 6.5-7.5 and little or no
acidity. Lime addition is not always required. Since this class of mine
drainage lacks nuclei for flocculation, the sludge (natural suspended solids)
experiences poor settling, and most tends to remain in solution. The parti-
cles that do settle are light and fluffy in character and produce sludges
with only 0.5% solids (6). Satisfactory liquid-solids separation may take
several days, or perhaps not occur at all. If treatment does not require
neutralization, the volume of sludge is totally dependent upon the iron
content, suspended solids, and other possible precipitable elements present

97
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vo
oo
LIMESTONE CoC03
HYDRATED LIME Ca(OH)2
SODIUM CARBONATE Na^CO,
I
I
30 mm.
60 mm.
4hrs. 8hrs. I2hrs.
TIME
16 Irs.
20hrs
24hrs
Figure 6-1. Treatability test settling curves
-------
B
IE
in the water (7).
enhance settling.
Figure 6-2. Type 1 settling.
Polymer or flocculant addition is highly recommended to
Type 3 settling behavior is independent of solids content and pollutional
loadings and is generally associated with limestone or sodium carbonate (soda
ash) neutralization. A two-phase separation system develops, with 90% of the
solids settling rapidly. The solids settle as a cloud of individual parti-
cles without a distinct liquid-solid interface. The sludge volume approaches
the maximum, usually within 10 minutes, yielding a turbid, cloudy super-
natant. The supernatant turbidity can be controlled with the use of a me-
chanical upflow clarifier in which the influent passes through a sludge
blanket. The sludge blanket filters the solids from the drainage, enhancing
removal.
6.1.1 Settling Performance

Settling performance is related to the hydraulic surface loading, which is
calculated by dividing the design flow (1/d) by the surface area of the pond
or clarifier (m2), with the resulting hydraulic surface loading in units of
1/d/m2 (gal/d/ft2). Common values range from 175 to 350 1/d/m2 (500 to 1,000
gal/d/ftz).

Hazen showed that separator performance is related to hydraulic surface load-
ing and the following variables (8):

1. flow turbulence in the basin;

2. velocity distribution throughout the pond;

3. particle interaction;

4. particle resuspension.

Starting at the water surface, a particle must settle the depth (D) of the
separator at a velocity (Vs) such that the settling time is less than or

99
-------
equal to the time the liquid is in the basin (9). Expressed empirically,

Sett1ingPVelocity (Vs) = Detent1on T™e (24)

D (m) _ Volume (L x W x D. m3)
Vs (m/min) Flow (m3/min)
Simplifying,
Vs (m/min) = Flow (m3/min)
1 ' ' Surface Area (W x L, mz)
Therefore, all particles having a settling velocity (Vs) greater than or
equal to the hydraulic surface loading (Flow/Area) are removed.

Actual basin performance is affected by many factors. Inlet and outlet
devices and wind induce turbulent, nonquiescent flow currents that lead to
poor settling and possible short-circuiting. Flocculation of particles into
large agglomerations results in nonuniform settling velocities. Also, influ-
ents with high suspended solids concentrations tend to settle as a mass
rather than as discrete particles.

6.2 Settling Unit Design

Settling units used in mine drainage treatment vary in size, configuration,
and method of solids removal. Earthen settling ponds are the most popular
because of their low capital and operational costs. The use of mechanical
clarifiers or thickeners, however, should be considered because they enable
the operator to exercise more control over the treatment system and improve
sludge densities (thus yielding lower volumes). Mechanical separators will
be discussed in detail later in this chapter.

A settling unit can be divided into four effective zones (see Figures 6-3 and
6-4) (10):

1. the inlet zone, in which the influent enters the unit and is uni-
formly distributed over the cross-sectional area;

2. the settling zone, where the majority of liquid-solid separation
occurs;

3. the sludge zone, used for temporary or permanent storage and compac-
tion of settled solids;

4. the outlet zone, where the supernatant is removed from the unit.

100
-------
INLET
ZONE
SETTLING

ZONE
OUTLET
ZONE
SLUDGE ZONE///77
Figure 6-3. Zones in a horizontal continuous flow
sedimentation basin.
OUTLET ZONE
\
INLET ZONE
-SLUDGE ZONE

-SETTLING ZONE-
Figure 6-4. Zones in a circular center feed, horizontal
continuous flow sedimentation basin.
101
-------
Little if any settling occurs in the inlet and outlet zones, reducing the
effective settling area to Zone 2. An optimum settling unit design is one
that minimizes the effects of the inlet and outlet zones.

6.2.1 Treatability Test

Before sizing and designing any settling pond or clarifier, a treatability
study should be performed to determine the behavior and characteristics of
the sludge along with expected supernatant quality. A simple testing proce-
dure yields much information on the properties of the sludge that are impor-
tant to design. These properties include sludge settling velocity, optimum
pH, best neutralizing agent, dosage rate, and sludge density and volume. A
recommended procedure for a treatability test can be found at the end of this
chapter.

Results obtained from laboratory tests, while useful, cannot be applied
directly to basin design. Sludge settling velocity, for example, which is
extremely important in clarifier design, can be determined, but the value
obtained is usually reduced to account for variables in the system.

Treatability studies also give an indication of sludge volume produced per
unit volume of influent treated. This value is very important in determining
tank volume.
6.2.2 Earthen Ponds

The following criteria should be considered when evaluating a site and con-
structing a settling pond.

6.2.2.1 Site Location

Earthen basins should not be located in swamps, marshes, or floodplains, on
steep slopes, or over abandoned wells or mine workings which might have
fissured rock that would permit seepage. Test borings are necessary when
information on soil conditions is unavailable (11).

6.2.2.2 Soil Conditions

Ponds should contain a 0.6- to 0.9-m (2- to 3-ft) layer of impervious mate-
rial to prevent seepage. Clays and silty clays are excellent for this pur-
pose. Sandy clays are usually satisfactory. Coarse textured sands, sandy
gravel mixtures, and gob materials are highly pervious, and therefore usually
unsuitable. Limestone areas are especially hazardous as pond sites because
of the possible presence of crevices or sinkholes in the limestone (12).

Soil types can be determined by laboratory testing of samples taken from the
field or from local Soil Conservation Districts.
102
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6.2.2.3 Foundation Conditions

Soil conditions must provide stable support for pond enbankment foundations,
as well as the necessary resistance to passage of water. Good materials,
providing stability and being impervious, are mixtures of coarse and fine
textured soils, such as gravel-sand, sand-clay, and sand-silt.

Tight contact between embankment and foundation should be insured to control
seepage properly along the plane of contact (13). One method for achieving
tight contact is to use layers of clays or silty clays in the construction of
the embankments. Synthetic liners or bentonite addition can also be used to
seal embankments.

The basin bottom should not be founded on bedrock or on stony, rocky soils.
The most suitable bottom foundation consists of a thick layer of relatively
impervious consolidated material.

6.2.2.4 Embankment Construction

The embankments forming the sidewalls of the basin should be constructed of
impervious materials similar to those used in the basin's liner. They should
be placed in layers and properly compacted with a sheepsfoot roller. The
side slopes depend on the properties of the fill and on the strength and
stability of the foundation material. Average slopes for inside embankments
are 2.5:1 or 3:1. Outside slopes should be gentle enough to allow a mower to
cut grass safely (3:1 or greater), and should be protected from erosion by
low-growing grass. The recommended minimum crown widths for earth embank-
ments of various heights are shown in Table 6-1 (12). If the crown is to be
used as a roadway, it should be at least 4.3 m (14 ft) wide at any height.

6.2.2.5 Design of Ponds

The settling pond should have an impermeable layer of clay or a'plastic
membrane liner in areas where fill material is unsuitable (permeabilities
greater than 10~6cm/s). A minimum freeboard of 0.61 m (2 ft) should be
maintained in the pond to prevent overflow.

The inside slope of the embankments should be protected ag-ainst erosion
caused by wave action. Placing a 0.61-m (2-ft) wide collar of riprap (aggre-
gate greater than 5 cm (2 in) in diameter) at the expected water level on the
embankment is an effective method of erosion protection. Ponds having wide
fluctuations of water levels require wider collars, 0.91-1.22 m (3-4 ft).
Synthetic collars (liners) may also be used.

Inlet and outlet design is a critical factor in providing quiescent condi-
tions for good settling pond performance. Inflow should be uniformly dis-
tributed over as much of the pond width as possible. This can be accom-
plished with the use of multiple inlets or a continuous width, multiple
V-notch box weir. This decreases the inlet flow velocity, which reduces the
probability of washout or resuspension of solids. Bad inlet design can lead

103
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TABLE 6-1

RECOMMENDED MINIMUM CROWN WIDTHS

Height of Embankment Minimum Crown Width
~m [ft]" "in 1ft)

1.52 (5) 0.912 (3)

1.52 - 3.04 (5 - 10) 2.43 (8)

3.04 - 4.57 (10 - 15) 3.04 (10)

4.57 - 6.1 (15 - 20a) 3.65 (12)

6.1 - 7.62 (20 - 25a) 4.26 (14)
aA 3.04-m (10-ft) wide horizontal bench is required for every 6.1 m (20
vertical ft) of embankment with l%-3% backs!ope.
to short-circuiting, which reduces the detention time and removal efficiency
of the settling pond. It also produces "dead" areas of noncirculating water
in the settling ponds, creating channels of flow within the pond, as shown in
Figures 6-5 and 6-6.

Uniform distribution of influent across the width of the settling pond inhib-
its isolated mounding of settled particles within the basin, thus maximizing
sludge storage volumes.

Similarly, outlet devices should also be multiple or continuous to maintain
low exit velocities. High exit velocities—that is5 those greater than 0.304
m/s (1.0 ft/s) at the effluent—create turbulence that can resuspend settled
solids, causing deterioration of effluent quality.

Baffles, selectively placed in a settling pond, prevent short-circuiting.
Two types used are surface baffles and submerged baffles.

Surface baffling prevents short-circuiting by diverting the flow from a
straight line across the pond to a less direct pattern, thus allowing for
more uniform influent distribution. Surface baffles generally float on the
pond surface and are anchored to the embankments with cables. They project a
few centimeters above the water level, and downward 0.6-0.9 m (2-3 ft), or as
much as one-half the pond depth. They can be constructed of flexible rubber,
PVC, nylon, or wooden planks.

Submerged baffles are used for sludge containment. The baffles extend upward

104
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4W
IO
T T
O
r
WEIR
O
c
Figure 6-5. Nondistributed short-circuiting influent.
EFFLUENT WEIR
SURFACE BAFFLE-
INFLUENT
WEIR-
4W
Figure 6-6. Distributed influent.
105
-------
from the bottom of the pond, and reach within Si few meters (feet) of the sur-
face. They create two separate settling chambers in the basin. The primary,
or influent, chamber collects most of the settled solids. The secondary, or
effluent, chamber acts like a polishing pond and provides additional solids
removal. The separate chambers ease sludge removal.

Submerged baffles are usually used in conjunction with surface baffles. The
combination of the two causes an "S" type flow path, as shown in Figures 6-7
and 6-8.
OUTLET
BAFFLE
INLET
SLUDGE ACCUMULATION
Figure 6-7. Surface-baffled pond.
OUTLET-
BAFFLE
_FLOW
INLET
SUBMERGED BAFFLE
'..-.•.^^WM^^^fM^^^

LE ' Zoi iinftF
SLUDGE ACCUMULATION
Figure 6-8. Combination surface- and submerged-baffled pond.
The effluent weir should not be placed on the leeward side of the pond
because agitation caused by wave action from winds could cause resuspension
of settled solids. The best effluent quality is achieved when the effluent
weir is placed on the pond's windward side.

6.2.3 Types of Sedimentation Ponds

For the purpose of this manual, earthen ponds are separated into two general
categories: settling ponds and impoundments.
106
-------
Settling ponds are small and designed primarily for settling with periodic
sludge removal. The sludge is disposed externally, or recycled to utilize
the unreacted portions of the neutralizing agent and improve final density
(solids content). These ponds require detailed engineering design and close
operational control. There is usually more than one pond, and they are
associated with more complicated treatment operations than impoundments.

Impoundments are large ponds designed for both settling and final sludge
disposal. Sludge accumulates in the settling pond over the life of the mine,
usually 10-20 years. Impoundments require large land areas, extreme depths,
planned methods of handling surface runoff and drainages, and, usually,
extensive construction.

To further define the characteristics of and design parameters for both
basins, the following subsections are offered.

6.2.3.1 Small Earthen Settling Ponds
(Approximately 2 Days Detention Time)

Small earthen settling ponds are designed to provide solids removal with
limited sludge storage capacities (14). Ponds in this category generally
have detention times less than 2 days (48 hours). These ponds require better
initial design, construction, and quality operational control.

The optimum configuration for small earthen ponds is a rectangular shape with
a length-to-width ratio of 4:1 (8). This long, narrow shape minimizes
short-circuiting, and inlet and exit turbulence have minimal effect on the
primary settling zone. Baffles, along with inlet and outlet distribution
devices, are used extensively in these ponds. These ponds may be used either
in series or in parallel in an effort to maximize efficiency and confine
sludge accumulations.

A series operation allows the drainage to flow into the primary settling
pond, which removes the majority of solids and is where sludge removal
devices are usually located. The primary effluent then enters the second
pond for final polishing.

A parallel pond system can operate in two ways. Neutralized water can flow
to only one pond until the sludge accumulations begin carrying over into the
effluent. At this time, the flow is diverted to the second pond until the
sludge in the first is removed. The other method of operation uses both
ponds simultaneously and requires dividing the flow with a splitter box.
This operational mode usually employs sludge removal devices in each pond.
The parallel mode provides a means of removing sludge without interrupting
flow in the basins.

6.2.3.2 Impoundments (Greater Than 2 Days Detention Time)

Impoundments are designed to serve for both solids settling and final sludge
disposal. They have detention times much greater than 2 days and are usually

107
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built as large as possible, without regard to detention time. Impoundments
have an economic advantage over small ponds in labor and operating costs.
Also, they provide the advantage of a large buffering capacity during treat-
ment plant breakdowns, thus producing a more uniform effluent quality (15).
Large impoundments, however, require extensive land area and are the least
realistic conservation approach (16). More importantly, they are almost
impossible to abandon.

Exceptionally large settling impoundments (detention time greater than or
equal to 20 days) assume limnological properties of lakes, experiencing turn-
overs in both fall and spring. During summer, the water stratifies, forming
a warm upper layer of freely circulating water (epilimnion); a middle layer
(metalimnion) containing a rapid temperature drop with depth (thermocline);
and a deep, cold, bottom layer (hypolimnion). As the sun decreases in inten-
sity with the coming of fall, the water temperature and density become uni-
form throughout the impoundment. The slightest wind can now circulate the
water, resuspending the sediments. A reverse effect occurs in the spring as
the water restratifies, resulting in spring turnover (14). This entire
phenomenons however, is rare in larger impoundments.

This problem can be avoided by limiting the depth of the impoundment,
enabling the sun to warm the entire body of water with a minimal temperature
gradient. In many situations, however, this solution is impractical because
impoundments must have the total volume for sludge disposal for the life of
the mine, which can only be provided with excessive depths (greater than 6
m).

Large earthen impoundments are usually constructed by damming a valley or
utilizing an abandoned strip pit. A special permit or state approval is
sometimes required when impoundments obstruct watercourses or change water-
shed drainage patterns. Several states define embankments over certain
heights as dams and require more stringent design and construction practices.
When damming a valley, diversion ditches shuld be provided to route runoff
around the impoundment. Runoff entering the impoundment increases the solids
loading to the pond, decreasing sludge disposal capacity. Also, excessive
amounts of runoff will cause resuspension of settled solids. Most impor-
tantly, eroded sediments can fill a basin quickly, greatly reducing the life
of a facility.

The use of small impoundments (detention time 3-5 days) in parallel, each
with sufficient volume to handle the total flow, can be effective and result
in longer pond life. As sludge accumulation begins to reduce detention time
in the first impoundment, the flow is diverted to the second. The sludge in
the first impoundment is permitted to undergo compaction and drying by de-
canting the supernatant. The sludge dewatered atmospherically does not rehy-
drate and occupies a smaller volume compared to the original gelatinous
precipitate (10). Ideal climatological drying will achieve a 30%-40% solids
content; however, a sludge solids content of 10%-15% is satisfactory.
108
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6.2.3.3 Volume Requirements

Many factors influence the sizing of sedimentation basins. These include
detention time, sludge removal and disposal method, and mode of operation.

Detention time is the basic design parameter when sizing earthen settling
ponds for mine drainage treatment, but it is not the only influencing factor.
Inlet and outlet flow conditions must be considered, along with turbulence
and eddy currents induced by wind. These disturbances reduce settling effi-
ciency and resuspend partially or previously settled solids. Baffles reduce
short-circuiting to a degree, but much less than the total basin volume is
actually utilized. For these reasons, a minimum of 12 hours is required for
sedimentation ponds utilizing periodic sludge disposal. A survey investi-
gating performances of mine drainage treatment plants in southwestern Penn-
sylvania and northern West Virginia showed nearly all settling ponds had
detention times in excess of 12 hours (17). It is common practice to build
the ponds as large as possible, subject to land availability.

The basin volume can be calculated, using the detention time and design flow,
with Equation 25.

V = Qtd (25)

where V = volume of settling pond, without sludge storage m3 (gal)

Q = design flow, m3/min (gal/min)

t. = detention time (min)

Additional capacity in settling ponds must be provided for sludge accumula-
tion. The sludge storage volume can either be approximated from methods
explained previously or estimated conservatively as 5%-10% of the average
daily volume treated.

Sludge storage volume requirements in a settling basin depend upon how often
sludge is removed. Ponds constructed without sludge collection or removal
devices should provide a minimum of 1-month sludge storage volume (11).
Ponds equipped with sludge removal devices should allow sufficient volume for
sludge storage between withdrawal operations. In some cases, basins are
designed with enough volume to hold sludge for the life of the treatment
plant, eliminating the need for any sludge removal. In all cases, the sludge
storage volume must be included in the final volume of the basin.

Settling Volume + Sludge Volume = Pond Volume
109
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6.2.4 Mechanical Clarifiers

Many mine drainage treatment plants employ clarifiers, generally circular,
for liquid-solids separation. These devices are usually preferred when land
area becomes limiting. The purpose of this subsection is to inform the
designer of this alternative and to present criteria and cost data for eco-
nomic and engineering evaluation.

When selecting a clarifier, the following features should be considered (18):

1. supernatant clarity;

2. an underflow of settled sludge having high density and suitable for
secondary dewatering (a 3% solids content in the clarifier underflow
is ideal (average l%-256));

3. provisions for sludge recycle to improve reagent utilization;

4. provisions for sludge removal to a second-stage thickening process
(i.e., lagoons);

5. adequate sludge storage capacity (depth) for blanket formation,
settling, and in situ concentrating;

6. provisions for flocculant addition to improve and control sludge
settling rates.

Many existing mine drainage plants utilize one of the following three types
of liquid-solids separators: (1) conventional clarifiers, (2) upflow solids
contact or flocculator clarifiers, and (3) thickeners.

6.2.4.1 Clarifiers

As best defined, a clarifier is a gravitational liquid-solids separator
having the primary objective of producing a high-quality supernatant (over-
flow) regardless of underflow solids content. Conversely, thickeners are
used for the purpose of concentrating underflow, with secondary emphasis on
supernatant quality. Consequently, most liquid-solid separators employed in
mine drainage plants should be called clarifiers, but because of their large
diameters requiring heavy-duty raking mechanisms, they are termed thickeners.

The first type of separator to be discussed is the conventional clarifier.
More often employed in sewage treatment, this horizontal flow-type clarifier
contains a rotating sludge removal mechanism with scraper blades, sludge
hopper, drive motor and unit, and center feedwell. Clarifiers, as applied to
mine drainage treatment, do not have skimmers. A conventional clarifier is
shown in Figure 6-9.

The conventional clarifier operates much like a simple horizontal settling
basin and utilizes radial flow distribution. Neutralized water enters the
circular center feedwell that distributes the flow uniformly 0.6-0.9 m (2-3

110
-------
BRIDGE
HANDRAILING
0.30m(lft)QROUTl
BAEFUE
\SUPPQRTS
TURNTABLE;
MAX. WATER SURFACE

EFFLUENT WEIR

EFFLUENT
\LAUNDER
038 m(lft-3in)MIN.

TOP OF
^INFLUENT BAFFLE

DRIVE CAGE
CENTER PIER

RAKE ARM TRUSS
SCRAPER BLADES
SLUDGE PIPE
6.99cm
(2-3/4in)
.30m (1ft)

0.6I m(2ft)GROUT
3.8lcm (l-l/2in)BLADE
CLEARANCE
SLUDGE HOPPER
HOPPER SCRAPERS
Figure 6-9. Conventional clarifier.
-------
ft) below the surface. The water flows toward the periphery and the solids
are separated by gravity. A continuous overflow weir is provided for super-
natant outflow.

Sludge accumulations are moved to the collection hopper at the center by a
rotating rake mechanism equipped with collection blades. The designer usual-
ly allows the volume required to cover the raking device as sludge storage.
This depth is dependent upon clarifier diameter and mechanism selected.
Sludge can be drawn off either continuously or intermittently. Higher under-
flow solids contents have been realized with intermittent drawoff.

The second type of clarifier, and perhaps the most popular unit, is the
flocculator or upflow solids contact clarifier, as shown in Figure 6-10.
This unit combines two functions into a single operation: flocculation
(particle formation) and filtration. The flocculation occurs when a polymer
is injected into the neutralized mine drainage in the feedwell. The floccu-
lator paddles gently mix the stream, inducing floe formation. The skirted
bottom on the feedwell extends below the sludge blanket level. This directs
the newly generated floe to rise through the sludge blanket, where suspended
solids undergo removal by adsorptive filtration. When this unit is designed
and operated properly, a definite sludge blanket-supernatant interface can be
seen within the clarifier.

More commonly, this unit is used without the addition of polymers for floccu-
lation. This type of application is termed an upflow solids-contact clari-
fier. The floe formed during the neutralization reaction settles poorly, but
well enough to create a filtration blanket. All principles of operation and
flow patterns are the same as those of the flocculator-clarifier.

This mode of operation has produced high-quality effluents for plants uti-
lizing either hydrated lime or quicklime as the neutralizing agent. The
sizing of these clarifiers has principally been the responsibility of the
vendor, and based upon one special parameter, rise rate. Rise rate is de-
fined as the vertical velocity of the water in meters (feet) per minute or
hour through the clarifier. For example, assume an existing plant has a
daily average flow of 3,785 m3 (1,000,000 gal) and a clarifier 30.48m (100
ft) in diameter. The following calculations are performed using this stan-
dard formula:

VRR = Q/A (26)

where VRR = rise rate, m/min (ft/min)

Q = design flow rate, m3/min (ft3/min)

A = cross-sectional area of clarifier, m2 (ft2)
112
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COLLECTOR
D*'VE UNIT
HANDRA/LINQ
BR/DQE
2-54cmCJ,n)6RouT
0.38
LAUNDER

l'B' S.W.D,
SLUDGE DRAW-OFF
PIPE
fl°cculatc
-------
now n - ' v
0 w x
Metric English

- 3'785 m3 d 1.000.000 gal ft3
d l,440min d 7.45 gal l, 440 min
Q = 2.63 mVmln 93.2 ft3/min

Area: A = (30.48 m) 2 % (100) 2 ^
4 4

A = 729.28 m2 7,854 ft2

Velocity: VRR = Q/A

(Hse rate! 2.63jl!M!L 93.2 ft3/min
I rise rate; = 729>28 m2 7,354 ft2

VRR = 0.0036 m/min 0.012 ft/min

Thus, if a rise rate is assumed, the formula can be rearranged to yield the
appropriate surface area or equivalent diameter. The most commonly used rise
rate (VRR) in sizing a mine drainage clarifier is 0.015 m/min (0.05 ft/min).
In actual practice, however, the values acquired from reliable performance
plants are lower.

Illustrated in Table 6-2, items A-l through A-12, are actual flows and sized
clarifiers in operation. These particular plants produce effluents in com-
pliance with EPA standards and with a high degree of reliance. As shown, the
rise rates, with the exception of A-9, are below the vendor recommendation.
This is not to say that a 0.015-m/rm'n (0.05-ft/min) rise rate will undersize
the clarifier, resulting in poor effluent quality. Industry has elected to
oversize clarifiers in an effort to guarantee performance, and to allow addi-
tional capacity as a safety factor for higher-than-predicted drainage flows.

Based upon these data, the designer can size a clarifier with confidence,
producing good performance when using a rise rate in the range of 0.003-0.009
m/min (0.01-0.03 ft/min).

The depth of the clarifier, be it solids-contact, flocculator, or conven-
tional, is independent of surface area and depends upon the detention time,
sludge blanket thickness, required storage volume, raking mechanism, and
desired underflow percent solids content.

Lovell (18) lists various optimum design parameters, which ultimately deter-
mine depth in addition to size. As shown in Table 6-3, the suggested 1.83-m
(6-ft) minimum sludge depth below the feedwell , plus a 72-hour minimum deten-
tion time for sludge storage, significantly influences the design depth.

Similar to values found in actual practice, 0.003-0.009 m (0.01-0.03 ft/min),
Lovell (18) recommended rise rates of 0.136 1/s/m2 (0.2 gal/min/ft2) equiva-
lent to 0.008 m/min (0.026 ft/min).

114
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TABLE 6-2

RISE RATES FOR EXISTING MINE DRAINAGE CLARIFIERS
AMD Plant
Flow
Mgal/d
Clarifier, dla
m ft
Rise Rate (VRR)
m/mi n
ft/min
A-l 3,785
A-2 15,140
A-3 26,495
A-4 28,387
A-5 16,275
A-6 23,467
A-7 25,738
A-8 3,475
A-9 11,355
A-10 946
A-ll 3,785
A-12 18,925
Vendor
recommendation
1.0
4.0
7.0
7.5
4.3
6.2
6.8
0.92
3.0
0.25
1.0
5.0
18.9
67.0
67.0
57.9
57.9
54.8
54.8
22.8
22.8
7.6
30.5
45.7
62
220
220
190
190
180
180
75
75
25
100
150
0.0091
0.0030
0.0057
0.0073
0.0042
0.0067
0.0076
0.0057
0.0192
0.0143
0.0033
0.0079

0.0152
0.03
0.01
0.017
0.024
0.014
0.022
0.025
0.019
0.063
0.047
0.011
0.026

0.05
At this point, the designer can estimate clarifier size and realize the land
area requirements. Knowing the diameter, a rule of thumb for cost is $5,0007
m ($l,500/ft) of diameter excluding the shell. (See Chapter 13, Cost Esti-
mating, for details.) Commonly, the shell is constructed totally of con-
crete, but steel has been used in combination with it.

6.2.4.2 Thickeners

Thickeners are those separator units that usually receive clarifier underflow
or an influent with a high percent solids. For example, the separators asso-
ciated with coal preparation plants are true thickeners. These units are
equipped with heavy-duty raking mechanisms and receive influent (washwaters)
high in suspended solids. Consequently, most separators used by the coal
industry, such as those designed for mine drainage treatment, have been
termed "thickeners," which may not always be correct.

Nevertheless, thickeners have been used extensively in mine drainage treat-
ment to perform the following double function: (1) to produce an overflow
within federal and state standards, and (2) to store sludge to produce denser
underflows.

Thickeners are sized like clarifiers, but design parameters depend upon the
intended purpose of the unit. If a thickener is the only separator unit
before ultimate disposal, the designer may consider a longer-than-normal
(12-hour) detention time in an attempt to produce a dense underflow. These
115
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TABLE 6-3

SUGGESTED OPTIMUM DESIGN PARAMETERS
FOR CLARIFIER OPERATION
1.

10.
Rising water velocity, 1/s/m2 (gal/min/ft2)

Solids settling rate, m/hr (ft/hr)

Unit area, m2/kkg/d (ft2/ton/d)

Percent solids in underflow

Sludge recycle provision -
percent circulating load

Percent solids to feedwell (combined
raw feed plus recycle sludge)

Flocculant dosage, mg/1 influent

Depth of feedwell below sludge level, m (ft)

Sludge depth below feedwell, m (ft)

Settled sludge retention time, hr
0.136 (0.2) maximum

0.61 (2.0) minimum

30 (300) minimum
(may exceed 1,024 (10,000))

5.0 minimum

5-30

1 minimum

1 - 2

0.61 (2) minimum

1.83 (6) minimum

72 minimum
units are of large diameter, usually greater than 30.48 m (100 ft), depending
upon design flow.

Smaller diameter thickeners are used as secondary separators, receiving
clarifier underflow or settled sludge from earthen basins. Operating under
this condition, the primary purpose of the unit is to thicken or dewater
sludge. Usually, special flocculants or polymers are added to enhance the
process.

Regardless of the design purpose, thickeners should be sized using a rise
rate between 0.003 and 0.009 m/min (0.01 and 0.03 ft/min). This rate is more
conservative than the vendor recommendation of 0.02 m/min (0.066 ft/min), but
can vary with the technique used for sludge witdrawal (periodic or continu-
ous).

Also, thickeners can be sized on the basis of solids loading. Surface area,
expressed as m2/kkg/d (ft2/ton/d) can range from 4.09 to 49.16 m2/kkg/d (40
to 480 ft2/ton/d), depending upon the total process operation. The lower
values are applicable when either polymer or sludge recycle is employed,
while the higher values apply for a once-through system utilizing gravity
separation only. An average design value of 30 m 2/kkg/d (300 ft2/ton/d) is
116
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often used. Only a treatability test simulating actual operating conditions,
however, can produce a valid loading rate.

6.2.4.3 Supplemental Operational and Design Considerations

The operational method of a clarifier or thickener is usually the result of
on-site adjustments that produce the best effluent quality. Every treatment
plant has its own peculiar behavior that deviates from laboratory results.

For example, the raking mechanism of a separator can be operated continuously
or intermittently. Best results have been obtained when the collection rake
operates periodically. This allows the settling sludge total quiescence for
maximum compaction. After a period of time (usually days), the collection
rake is activated and sludge withdrawn.

Plants producing large volumes of sludge may require the collection rake to
operate continuously. In such cases, the designer could consider a slower-
than-normal peripheral speed. Generally, a clarifier collection mechanism
moves at 4.57-6.1 m/min (15-20 ft/min), while thickener rakes rotate slightly
faster, 7.62-9.14 m/min (25-30 ft/min). Even slower peripheral speeds of
1.5-3.0 m/min (5-10 ft/min) can improve sludge densities; their use merits
consideration.

Also, the support structure of the collecting rakes warrants forethought,
especially for those plants anticipating a gypsum (CaSOj problem. Many
types of clarifiers and thickener mechanisms are constructed with various
features to allow easy repair or cleaning. Most of these underwater struc-
tures (trusses) are massive and designed as cantilevers. Thus, 0.454 kg (1
Ib) applied to the rake arm 15.24 m (50 ft) from the center support exerts
6.92 kg-m (50 Ib-ft) of torque. Therefore, small amounts of sludge accumula-
tion will produce a high torque requirement that must be transferred through
the raking mechanisms. Typical raking mechanisms plow through the sludge and
slowly move it to the collection well. Where gypsum is not a problem, these
mechanisms function satisfactorily.

If the chemistry of an acid mine drainage indicates a gypsum problem might
occur (sulfate greater than or equal to 2,500 mg/1), these raking mechanisms
with large underwater surface areas should be avoided. The alternative is a
raking mechanism pulled by cables, as shown in Figure 6-11. The drive or
torque truss is visible above water, and only the cables and,raking blades
are submerged. Obviously, this minimizes underwater surface area available
for precipitating gypsum. Also, the cables are flexible supports from which
thin layers of gypsum can be easily removed.

Mechanical separators function well with a minimum of maintenance. Concrete
structures are preferred; however, steel shells can be used when provided
with corrosion protection. The designer should allow for operator access to
potential problem areas, such as the thickener rakes and sludge withdrawal
line, without having to drain the unit.
117
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MECHANISM SUPPORT TRUSS
SOLUTION LEVEL-
00
TORQUE ARM LENGTH
DRIVE HEAD
UPPER VERTICAL SHAFT
TOP OF
TANK
SINGLE PIPE TORQUE ARM
DRAG CABLES
BOOM SUPPORT
SUSPENSION CABLE
DUAL AXIS
APIVOT PIN
LOWER
VERTICAL
SHAFT
LIFT INDICATOR
RAKE ARM
CENTER SCRAPER
"•RAKE BLADES (DOUBLE SWEEP)
SPIDER PLATE
DISCHARGE CONE
Figure 6-11. Cable thickener.
-------
6.2.4.4 Tilted-Plate Gravity Settlers

The tilted-plate gravity settler is an inclined-plate, shallow-depth settling
device. It performs the same function as a conventional clarifier, but
occupies approximately one-tenth the space. These settlers operate on the
principle that solids removal is independent of depth and dependent upon the
surface area of the settling unit. If the depth of the settling compartment
is reduced to a few centimeters (inches) and a number of units are stacked,
increasing available surface area, the handling capacity of the separator
increases proportionately. These types of separator units use parallel
plates tilted at an angle (55°-60°) so that the settled sludge is self-
draining. The true effective surface area of the unit is calculated by using
the cosine of the inclining angle and projecting the plates to a horizontal
plane multiplied by 80%. Twenty percent of the area is lost to influent
turbulence and collected sludge volume.

Total surface area = (N) (A cos <*) (width) (27)

where N = number of plates

A = length of the plate (hypotenuse)

* = angle of inclination

This compacts the surface area of a clarifier into a much smaller size than
those previously discussed.

The influent is introduced into the unit through a bottomless rectangular
feed box located between sets of tilted plates. The influent flows into the
plates from the side and then upward, exiting at the top of the tank through
flow distribution orifices. The orifices are sized to take a specific pres-
sure drop, insuring uniform flow distribution across the plates. The solids
settle in each compartment and slide downward into the sludge hopper. Fur-
ther concentration of the settled solids is accomplished by compression with
a low-amplitude vibrator pack or with a specially designed version of a
conventional thickener.

Figure 6-12 illustrates the flow pattern and compactness of the unit.
Although to date few units are on-line in AMD treatment plants, tremendous
potential exists for this type of application.

The criteria for sizing or computing the required surface area of the unit is
based on 0.34 1/s/m2 (0.5 gal/min/ft2). This surface loading exceeds our
recommended rise rate of 0.003-0.009 m/min (0.01-0.03 ft/min), but can be
justified because most of these units are preceded by polymer flocculation.

The appropriate size unit is selected by equivalent clarifier diameter. For
example, a process flow requiring a 13.7-m (45-ft) diameter clarifier would
use a Model 2000/55, where 2,000 is the surface area (ft2) and 55° the angle
of plate inclination.

119
-------
-n

-------
The unit's fully assembled approximate cost can be estimated by multiplying
the equivalent clarifier diameter by $5,000/m ($l,500/ft).

The major disadvantage of this type of separator is its susceptibility to
clogging caused by gypsum. Where gypsum precipitation is suspected, this
type of separator is not recommended.

6.3 Recommended Procedure for a Treatability Settling Test

1. Pour 1,000-ml portions of well-mixed, fresh (24 hours old or less) mine
drainage into five 1,500-ml beakers. (If lab equipment is limited, only
one sample at a time can be used.)

2. Place the beakers on a gang stirrer or comparable mixing device and stir
for 5 minutes at 60-80 r/min.

3. Leaving one sample as a "blank" for comparison, add measured amounts of
neutralizing agent to each sample, adjusting the pH levels between 6.0
and 9.0. Typically, if ferrous iron removal is the primary objective, a
pH of 8.0 is strongly recommended. Agitate the mine drainage samples
continuously during alkali additions, and thereafter for approximately 5
minutes at 80-90 r/min to insure thorough mixing.

4. Then, aerate each solution approximately 30 minutes either mechanically
or with diffused air. Remove from each sample an aliquot for suspended
solids analysis.

5. Pour each sample into a 1,000-ml graduated cylinder and allow to settle.

6. At 2-minute intervals, record the sludge interface height in the gradu-
ated cylinders for 30 minutes. Continue recording this height for the
next 60 minutes at 10-minute intervals.

7. Plot the data recorded on a graph of liquid-solids interface height
versus time for each sample tested. The plots should form curves simi-
lar to those shown previously in Figure 6-2.

8. Note the final sludge volume (mm). After acquiring the initial sus-
pended solids concentration (mg/1) (step 4), the density of the final
sludge can be calculated as follows:

ViSi = VsSs (28)
where Vi = volume of sample (1,000 ml)

Si = initial suspended solids concentration before settling (mg/1)

121
-------
Vs = final settled sludge volume (ml)

Ss = solids content of settled sludge

The calculation assumes 100% removal of all solids, which is not true. The
supernatant will have some concentration of suspended solids; however, for a
close approximation of the expected sludge density, the procedure is valid.
Another method to use that ignores sludge density is the volume of sludge
produced per volume of water treated. This ratio is the sludge volume ratio
(14).
Sludge Volume Ratio = <»>
The sludge volume ratio times the quantity of AMD treated per day yields a
close approximation of expected sludge production per day. This method can
be used if a laboratory analysis has been performed on the mine drainage to
be treated. The sludge production can be estimated using the sum of the
suspended solids and ferrous iron concentrations.

Suspended solids concentration and dissolved iron concentration, mg/1

x flow, 1/d x -L 9 = sludge solids produced, kg/d

Assume 15% of the neutralizing agent is waste and settles with the sludge.

Sludge solids produced, kg/d + 0.15 neutralizing agent added, kg/d

= total settled solids, kg/d

Assume an average 1% solids content in the sludge or results from step 8.

Total settled solids, kg/d m Sludge produced> kg/d x 11 x lOllm.
\j«\j i j. ^y •

= sludge produced, m3/d

9. The quality of the supernatant is determined by decanting portions of
each sample tested and analyzing for effluent parameters (iron, pH,
suspended solids, manganese). If the supernatant quality does not meet
the standards of Environmental Protection Agency Effluent Guidelines and
state standards, an extension of the detention time, use of an alterna-
tive neutralizing agent, or addition of polymer settling aids should be
considered. If supernatant quality is obviously undesirable across the

122
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pH range tested, then polymer addition becomes necessary. At this
point, it is best recommended that a polymer salesman be contacted for
the type and dosage of flocculant required.

6.4 References

1. Lovell, H.L. An Appraisal of Neutralization Processes to Treat Coal
Mine Drainage. EPA-670/2-73-093, Cincinnati, Ohio, November 1973.

2. Moss, E.A. Dewatering of Mine Drainage Sludges. Part I, EPA-14010-
FJK, Coal Research Bureau, December 1971.

3. Sammarful, I.C. Evaluation of Common Alkalis in Neutralizing Acid Mine
Water. M.S. thesis, The Pennsylvania State University, University Park,
Pennsylvania, March 1969.

4. Wilmoth, R.C., and R.D. Hill. Neutralization of High Ferrous Iron Acid
Mine Drainage. Federal Water Quality Administration, August 1970.

5. McCabe, W.J., and J.C. Smith. Unit Operations of Chemical Engineering.
3rd ed. McGraw-Hill Book Company, New York, 1976.

6. Barthauer, G.L. "Coal Age." Mine Drainage Treatment - Fact and Fic-
tion, 71(6), 1966.

7. Coal Research Bureau, West Virginia University. Dewatering of Mine
Drainage Sludge. EPA-14010 FJX.

8. Hazen and Sawyer, Engineers. Process Design Manual for Suspended Solids
Removal. EPA 625/l-75-003a, Cincinnati, Ohio, January 1975.

9. Janerus and Lucas. Settling and Thickening Metal Hydroxides With the
Lamella Gravity Settler. Paper presented at Pennsylvania Water Pollu-
tion Control Association, Hershey Technical Conference, Hershey, Penn-
sylvania, June 1977.

10. Aqua-Aerobic Systems, Inc. Clarifiers. Bulletin 302, 1976.

11. Bureau of Water Quality Management. Mine Drainage Manual. 2nd ed.
Publication No. 12, Department of Environmental Resources, Harrisburg,
Pennsylvania, September 1973.

12. Soil Conservation Service. Ponds for Water Supply and Recreation.
Agriculture Handbook No. 387, USDA, 1-71.

13. Krynine, D.P., and W.R. Judd. Principles of Engineering Geology and
Geotechnics. McGraw-Hill Book Company, New York, 1957.

14. Smith, R.L. Ecology and Field Biology. Harper and Row Publishers,
Inc., New York, 1974.
123
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15. Wilmoth, R.C., and J.L. Kennedy. Treatment Options for Acid Mine
Drainage Control. U.S. Environmental Protection Agency, Cincinnati,
Ohio.

16. Lovell, H.L. The Control and Properties of Sludge Produced from the
Treatment of Coal Mine Drainage Water by Neutralization Process. Re-
prints of Papers Presented before Third Symposium Committee to the Ohio
River Valley Water Sanitation Commission, Pittsburgh, Pennsylvania,
1970.

17. Skelly and Loy, Engineers and Consultants. Development Document for
Effluent Limitations Guidelines and Standards of Performance for the
Coal Mining Point Source Category. U.S. Environmental Protection
Agency, Cincinnati, Ohio, November 1974.

18. Lovell, H.L. Design of Coal Mine Drainage Treatment Facilities.
Pennsylvania State University, University Park, Pennsylvania, November
1973.

6.5 Other Selected Readings

Lovell, H.L. Experience with Biochemical Iron-Oxidation Limestone Neutrali-
zation Process. Preprints of Papers Presented before the Fourth Symposium on
Coal Mine Drainage Research, Coal Industry Advisory Committee to the Ohio
River Valley Water Sanitation Commission, Pittsburgh, Pennsylvania, 1972.

Pudlo, G.H. Sludge Volume from Treatment of Acid Mine Drainage. M.S.
thesis, West Virginia University, Morgantown, West Virginia, 1970.

Akers, D.J., Jr., and W.F. Lawrence. Acid Mine Drainage Control Methods.
Report No. 81, West Virginia University Coal Research Bureau, Morgantown,
West Virginia, 1973.
124
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CHAPTER 7

SLUDGE DEWATERING AND DISPOSAL
7.1 Introduction
Neutralization of acid mine drainage creates a sludge that can be costly to
handle, dewater, and ultimately dispose. Today, sludge handling and disposal
presents the most recurrent and demanding problem to the designer or operator
of a mine drainage treatment plant.

Environmental regulations coupled with all their ramifications for sludge
disposal categorize sludge as a secondary or potential pollutant. In design-
ing a mine drainage treatment plant, it is important to realize there are two
effluents; i.e., treated water and sludge.

The volume of sludge from a mine drainage treatment plant varies with the
composition of the untreated drainage and the neutralization method employed.
However, a designer should expect the sludge volume to be from 5% to 10% of
the daily flow through a typical treatment facility. Some facilities have
recorded sludge volumes as high as 33% of the average daily flow (1). There-
fore, the methods of sludge handling are extremely important to the designer.

The purpose of this chapter is to describe the chemical and physical proper-
ties of mine drainage sludge, the practical methods of dewatering, and to
discuss the most commonly employed sludge disposal methods.

7.2 Mine Drainage Sludge

This section provides a description of mine drainage sludge, including its
chemical characteristics and physical properties, such as its settleability,
density, dewaterability, particle characteristics, and viscosity.

7.2.1 Chemical Characteristics

The chemical composition of AMD sludge, or yellowboy, varies with the raw
drainage and method of treatment. Lovell reports that the sludge is gener-
ally composed of hydrated ferrous or ferric oxides, gypsum (calcium sulfate),
hydrated aluminum oxide, varying amounts of sulfates, calcium, carbonates,
bicarbonates, and trace quantities of silica, phosphate, manganese, titanium,
copper, and zinc (2).

Illustrated in Table 7-1 are the primary components and compounds of severa'

125
-------
TABLE 7-1

CHEMICAL ANALYSES OF SLUDGES
Weight, % (dry basis 105°C-24 hr)
Alkali Used Hydrated Lime - Air Oxidation
Bennett's
Mine Water Branch Proctor 1 Proctor 2
Hydrated Lime Hydrated
Bio-oxidation Dolomite
Proctor 2
Proctor 2
Calcined Dolomite
Proctor 1 Proctor 2
no
Component

Al 3.8
Fe 19.5
Ca 6.9
Mg 6.6
SOij 5.7
H20 at 180°C 12.5
Al(OH), 11.1
Fe(OH)3 37.6
CaC03 11.4
MgC03
3MgC03-Mg(OH)2.3H20 25.2
CaS(V2H20 10.7
4.7
17.7
5.8
4.3
6.8
15.8
13.7
34.0
7.5

16.4
12.7
3.1
23.1
5.2
5.1
5.8
14.8
Compound
8.9
44.3
7.0
19.4
10.9
8.0
24.3
4.8
1.3
11.5
Composition
23.1
46.6
0.0
5.1
20.7
2.8
13.0
17.2
3.8
4.4
10.2
5.5
7.4
10.7
11.8
1.6
8.7
4.5
13.5
6.7
9.8
2.3
11.7
8.2
24.9
38.3
--
14.4
8.4
16.0
14.2
25.0
21.0
22.4
3.1
13.2
25.9
14.4
12.0
24.6
4.4
4.8
23.2
5.2
5.8
5.5
14.7
.9
.7
13.
44.
7.2
6.1
15.6
10.4
Total
96.0
84.3
90.5
95.5
94.2
101.7
94.5
97.9
-------
mine drainage sludges generated from different waters and neutralization
processes (3). Since most mine drainages can be classified by broad group-
ings according to similar chemical composition, the data sufficiently repre-
sent the expected chemical properties of mine drainage sludge. The general
consensus among researchers is that the chemical makeup of raw water, along
with the unit operations and type of neutralizing agent employed, greatly
influence sludge formation and associated chemical and physical character-
istics.

7.2.2 Physical Properties

Among the many physical properties of mine drainage sludge, the most impor-
tant are settleability, density, viscosity (sludge flowability), particle
size and surface properties, and dewaterability (3).

7.2.2.1 Settleability

Sludge settleability ranks as the most important property with respect to
design. Sludges that settle poorly require specialized treatment (i.e.,
polymer addition) and greatly influence process design. The various alkalis
produce sludges with differing properties.

Hydrated lime (Ca(OH)2) is the most popular neutralizing agent in treating
mine drainage. Sludges generated from such facilities usually settle well,
but compact poorly, producing high volumes of sludge. High sludge volumes
have also been reported for quicklime (CaO) treatment. Bisceglia stated that
most quicklime products contain 2%-3% unburned limestone or "core." Assuming
an 18.1 kkg/d (20 ton/d) usage, this amounts to 364-545 kg (800-1,200 Ib) of
extra solids (dry) (4). This could represent about 19.0 m3 (5,000 gal) or
more of sludge per day. Also, the compressive settling (compaction) of
quicklime sludge is just slightly greater than that of hydrate sludge.

Limestone produces a sludge with good settleability and compaction, but is
limited because of pH (7.4 maximum) and the chemistry of iron oxidation (5).

The sodium neutralizing agents (sodium carbonate, sodium hydroxide) produce
sludges that are more fluffy and voluminous than the limes. Because they are
completely soluble, there is virtually no waste. On a dry solids basis, less
sludge is actually produced.

7.2.2.2 Density

Sludge densities are generally reported as percent solids by weight. They
are an extremely important parameter in both the initial settling process
(pond or clarifier) and final disposal. Densities can vary tremendously,
from 0.5$ for in situ sludge to 63% for1 a cake from a dewatering process (6).

The designer can expect a sludge density from settling ponds with a continu-
ous discharge to be anywhere between 0.5% and 4.5%. Clarifier underflows
have sludge densities ranging from 1.0% to 7.0%, slightly higher than earthen

127
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ponds. Clarifiers offer greater ease in solids handling.

7.2.2.3 Viscosity (Sludge Flowability)

Little has been reported in the literature pertaining to sludge viscosity
because the percent solids content is usually so low (less than 5%) that it
is readily comparable to water. Nevertheless, sludge flowability is not
always that easy. The ability of a sludge to flow is considerably important,
especially when designing a collection system or cleaning a pond. Lovell
outlines some of the properties and flowability problems encountered when
transferring different sludges (3).

1. One sludge moved satisfactorily from the bottom channels of the
settling pond to the sludge drying basin by gravity prior to its
gelation, although it tended to resist flow from the sloped bottoms
to the channel. The limestone-produced sludge was most difficult to
move because of its density. Low-velocity flows are preferred
during transfer because the tendency toward "ratholing" increases
with flow rate.

2. The mobility of an "aged" settled sludge was poor when covered by
several feet of water. It would not flow freely from the sloped
bottom into the sludge channels and hence through the effluent pipes
by gravity flow or from the pump. This fact indicates that the
angle of repose of the submerged sludge is greater than 30°; thus,
slopes approaching 45° are preferable.

3. Mobility of aged sludge on the sloped surface of a drained pond was
satisfactory when movement was initiated by a squeegee or hose water
pressure. It tended to move as a large block. On quiescent com-
pression and aging, the sludge forms a gel that has thixotropic
tendencies (the property exhibited by certain gels of liquefying
when stirred or shaken, then returning to a semisolid form upon
standing). The gelation behavior develops within 48 hours under
sludge and water pressure compression. There was no evidence of
this gelation phenomenon in the settled sludge in the thickener.
Apparently, the slow movement of the sludge by the thickener rake
prevented gelation. An understanding of aging and gel formation is
needed to assist with lagoon design and operation.

A 0.61- and 0.9-m (2- and 3-ft) thick layer of gelled sludge in one
section of a pond was moved into the sludge channel by two men in
4-5 hours using squeegees and low-pressure hoses. Blocks of sludge
were cut with shovels; these would slide down the sloping bottom to
the sludge channels, provided the bottom was kept thoroughly wet by
a small flow of plant water from a hose. Such settled sludge would
normally range between 5% and 15% solids. This procedure was rapid
and minimized sludge dilution. There was no evidence that the
bituminous (asphalt) bottoms significantly enhanced sludge movement.
But it did stabilize the pond and eliminate erosion and mud forma-
tion problems during sludge removal.

128
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The sidewalls of the pond were earthen. Although no significant
erosion problems developed with the compacted clay walls, sludge
removal from these areas was more difficult. An occasional small
stone from the pond walls would block the check valve of the sludge
pump. Bituminous coatings of side and bottom surfaces provide
definite advantages and are recommended. Riprap should never be
employed below water levels in a pond.

4. Transfer of gelled sludge from the settling pond to drying beds by
gravity flow was unsatisfactory. Sludge tends to rathole quickly,
allowing excessive amounts of clarified supernatant to move with the
sludge and complicate subsequent operations.

5. The sludge had adequate flow characteristics in lines and through
pumps when movement began. The sludge transfer lines were flushed
with clear water after sludge transfer ended to prevent scaling or
further gelation. No line blockage was experienced.

6. Sludges produced with hydroxide-type alkalis were the least dense
and did not adhere very tightly to the~pond bottom. By contrast,
settled limestone slurries were dense, sticky, and claylike in
character, and adhered tenaciously to the bottom surface.

The experience and properties outlined above by Lovell alert the designer or
operator to the importance of providing for sludge handling and giving fore-
thought to design.

7.2.2.4 Particle Size and Surface Properties

Both of these properties influence the flocculation of sludge particles. The
ability of sludge floes to agglomerate affects the sludge settling veloci-
ties, and ultimately, the settling basin size.

The surface property of individual floes generally refers to the electro-
static charge on the particle. This charge, which causes the particle to
resist flocculation, can be offset with the addition of coagulants or poly-
mers.

Researchers have conducted numerous studies on the various types of polymers
(3, 7). They found that nonionic (neutral charge) and anioaic (negatively
charged) polymers were most responsive in AMD sludge treatment. The best
flocculant and optimum dosage rate must be established experimentally.
Ideally, polymer dosages should range between 0.5 and 2.0 mg/1; however,
higher rates have been used.

7.2.2.5 Dewaterability

The ease of removing water from mine drainage sludge is referred to as its
dewaterability. Others have defined It as the ease with which sludge can be
Concentrated into a more manageable form (3). Obviously, the objective of

129
-------
any sludge dewatering operation is to reduce moisture, thereby increasing the
solids content, which will minimize disposal volumes. Some basic factors
that influence sludge dewaterability are presented in Table 7-2 (6).
TABLE 7-2

SLUDGE DEWATERABILITY VARIABLES

1. Initial concentration of solids 4. Compressibility

2. Age and temperature 5. Chemical composition

3. Viscosity 6. Physical characteristics
These variables apply mostly to mechanical methods of dewatering. In most
cases mine drainage sludge undergoes gravity dewatering or air drying either
by lagooning or disposal into deep mines.

7.3 Methods of Mine Drainage Sludge Dewatering and Disposal

This section describes sludge dewatering and disposal methods and includes
lagooning, deep mines, underground disposal guidelines, filtration, bed dry-
ing, and centrifugation.

7.3.1 Lagooning

Lagoons offer the easiest and one of the cheapest methods of sludge storage
and dewatering when land is available. Mine drainage treatment plants are
usually located in remote or isolated areas where land is readily available.

Lagooning can serve the following three purposes in mine drainage treatment:
(1) as settling ponds or impoundments (Chapter 6) where sludge is collected
and undergoes preliminary dewatering; (2) as the primary sludge dewatering
unit; or (3) as permanent storage or disposal facilities.

Lagooning has been used in every mode of operation (singular, series, paral-
lel) to enhance sludge consolidation and thickening. Large, single lagoons
(settling impoundments) perform all three functions the least effectively.
Very little dewatering and compaction of the sludge takes place in these
facilities. The continuous presence of water above the sludge zone does not
allow any atmospheric dewatering. The designer should not consider a large,
single lagoon (impoundment) as an effective sludge dewatering facility, al-
though it may have certain advantages as a settling unit, as discussed in
Chapter 6.

130

-------
Settling ponds in series, where the first serves as the primary settling unit
and the second as a polishing pond, are often used in mine drainage treat-
ment. This system offers better sludge control by isolating the majority of
the solids in the first pond. This arrangement has the same disadvantage as
a large single pond where little sludge dewatering occurs. In this arrange-
ment, the primary settling pond is usually equipped with a sludge removal
device, and the sludge is transferred to a disposal lagoon where atmospheric
dewatering can occur.

A dual or parallel arrangement of settling ponds, together with an isolated
dewatering lagoon, may be the optimum system for achieving the most effective
sludge dewatering possible through natural methods. Where land is available,
construction of two settling ponds, each with sufficient volume to treat the
design flow, appears ideal. This system allows alternate use of the ponds so
the inactive pond can undergo first-stage dewatering. Its contents are then
transferred to the final disposal lagoon for further dewatering. In both
ponds, supernatant or surface water must be decanted to allow the sludge full
exposure for natural drying.

Holland et al. experienced good results with this method of drying where a
raw sludge with an initial solids content of near 1% dewatered to 14% within
3 weeks. They further observed that solids can be increased by 8% and can
dewater as much as 20% solids. These are remarkable results and should not
be assumed to be easily achieved.

After most of the water is decanted from a lagoon, shrinkage cracks commonly
occur that will honeycomb the entire pond. A unique property of dewatered
sludge, if not totally submersed, is that it will resist "rewetting," which
enhances the dewatering process.

The problem of ultimate sludge disposal still remains. Closure of sludge
disposal lagoons presents a formidable problem. Sludge lagoons are very
difficult to cover with soil. In such an operation, extreme care should be
taken as the sludge exhibits thixotropic tendencies; i.e., it will tend to
liquefy upon vibration. Dried sludge samples have exhibited relatively low
unconfined compressive strengths in the range of 200-400 lb/ft2. Such a
sludge would be under stress during covering with soil, especially by the
weight and movement of the machinery. Special safety precautions are needed
for operators of the covering equipment.

Spread burial of dried sludge can be more economical if it can be done "in-
house" and the land area for spreading is available. This method takes
advantage of the soil conditioning benefit's of the sludge (i.e., alkalinity,
minerals). Conrad, in a test using sludge to treat a corn field, reported
that the corn grew faster and taller and produced a better yield than the
control plot (8). Controlled amounts of sludge applied to an area, then
covered by soil in layers, can be utilized as a means of sludge disposal and
improving the land. The heavy metal content and other toxic constituents of
the sludge should be evaluated for their long-term effect on the vegetation.
Sludge should not be applied to high-acidity soils (i.e., gob piles) where
the alkalinity will totally leach out and the metals will redissolve.

131
-------
In summary, lagoon drying can substantially reduce sludge volumes when oper-
ated properly. The best results are obtained when undisturbed dewatering is
allowed to occur after free water is decanted. The designer should remember
that deep sludge lagoons may not be satisfactory for ultimate disposal.

7.3.2 Abandoned Deep Mine Disposal

Disposal of mine drainage sludge to abandoned deep mines by means of a bore-
hole is exploited whenever possible. The apparent simplicity of operation is
not always indicative of the practice. As with any disposal method, all the
legal and environmental ramifications must be considered before it becomes
possible. Nevertheless, this means of disposal offers the most economical
and simplest method for sludge disposal.

Steinman, because of restricting terrain at a treatment site, could not con-
struct a lagoon of adequate size for sludge disposal (9). It was decided
that sludge disposal would be to an abandoned deep mine 14.5 km (9 mi) away.
Even under these conditions where extra handling and trucking were required,
it was proven economical and successful, although a thorough inspection of
the proposed underground area by the state regulatory agency was required
before permitting the site.

An example of possible state agency guidelines for underground disposal
follows.

7.3.2.1 Guidelines for Underground Disposal of Sludge from Acid
Mine Drainage Treatment (10)

Whenever underground disposal of sludge from treatment of mine drainage is
proposed, the sludge must have a pH of 7.0 or above, and all of the iron must
be in the ferric form.

Any application submitted that proposes underground disposal of sludge should
contain the following supplemental information:

1. Location

• a. Name of the abandoned mine in which the sludge is to be disposed.
b. Outline of the mine workings on the latest available U.S. Geologi-
cal Survey topographic map.
c. Location of the discharge point of the sludge shown on the latest
topographic map.

2. Mine Hydrology

a. Is there water in the mine? If water is present, what area of the
mine is flooded?
b. Is the water level in the mine rising, falling, or static?
c. Does the flooded mine discharge? If so, supply the following:
(1) Location of discharge point(s).
(2) Quantity of the discharge(s).

132

-------
(3) Elevation of the discharge(s).
(4) What quality controls will be at these discharge points to
determine future variations in quantity and quality?
What is the extent of previous mining next to and below the pro-
posed disposal area?
Does the disposal mine have any connections with other nearby
workings?
(1) List discharges to other mines.
(2) List discharges from other mines.
Is there pumping from the disposal mine or any interconnected
mines? If so, supply information concerning quantity and quality
of pumpage.
3. Quality of Mine Water
Supply the following information concerning the quality of water
in the mine where the sludge will be disposed:
(1) PH.
Acidity or alkalinity.
Iron.
v*y

111
4. Input of Sludge

a. Volume of sludge.
b. Volume in the mine available for sludge disposal.
c. Estimated length of time sludge will be disposed of in the mine.
d. Leachate quality from the sludge (ASTM-A Method).
e. The concentration of sludge.
f. What effect will the sludge have on any discharges from the mine?
(1) Quantity.
(2) Quality.
g. Will any of this sludge be flushed out by movement of the impounded
mine water?

5. Geology

a. Structure of the coal beds.
(1) Strike.
(2) Dip.
b. Are there any faults present in the immediate area?
(1) Location.
(2) Strike and dip.
c. What is the elevation of the local groundwater table? How was this
determined?

6. When the pH of the water in the abandoned mine is above 4.0, the ferric
iron is basically insoluble. At a pH below 4.0, ferric iron is soluble.
When it is proposed to dispose of iron sludge into water with a pH below
4.0, supporting data must be supplied to show that the proposed disposal
will in no way adversely affect any present or future discharges from
the mine pool.

133
-------
Little detailed information can be found in the literature pertaining to deep
mine disposal of drainage sludge. The most common process is to withdraw the
sludge from the settling basin by means of a portable or permanent collection
system and pump it directly to the borehole. No dewatering is considered
before injection. Therefore, the solids content is low (0.5%-2.0%) and high
volumes must be handled.
7.3.3 Vacuum Filtration

Vacuum filtration, which has long been employed in sewage treatment for
sludge dewatering, is readily applicable to mine drainage sludge. The prin-
ciple of operation is simple. The revolving drum is perhaps the most common
type of vacuum filter. It has a series of vacuum cells that run the length
of the drum. The drum, turning at less than 1 r/min, passes through the
sludge where a vacuum of 0.40-0.87 atm (12-26 in of mercury) is applied to
the submerged portion (approximately 25% of the periphery), drawing a cake to
the filter media surface. The filter media, referred to as a "cloth," can be
made from a number of materials, the most common of which is polyethylene.
As the sludge emerges from the reservoir, the vacuum dewaters the cake as it
rotates on the drum. The cake is then removed by a scraper, a blast of air,
or coiled springs.

Some of the operational variables that influence the dewaterability of the
sludge by this method are listed in Table 7-3 (6).
TABLE 7-3

VACUUM FILTRATION OPERATIONAL VARIABLES

1. Amount of vacuum 4. Filter media

2. Amount of drum submergence 5. Sludge conditioning before
filtration
3. Drum speed

Lovell, operating a vacuum filter with the cycle times listed in Table 7-4,
was able to achieve the filtration rates presented in Table 7-5 (3).

Lovell states that filter feed slurries with less than 2% solids consistently
produced the lowest solids filtration rate, 73.9-147.8 kg/m2/d (15-30 lb/ft2/
d). Solids at 5% or greater, however, can have filtration rates of 610-1,616
kg/m2/d (125-331 lb/ft2/d), indicating that vacuum filtration can be economi-
cally feasible (3). Other researchers studying vacuum filtration reported
successful results in producing a filter cake with 24%-35% solids (7, 11).
At this percent solids, the sludge can be easily handled and is acceptable
for landfill disposal.

134
-------
.TABLE 7-4

OPERATIONAL CYCLES

Submergence 16 s 21%

Drying time 50 s 64%

Discharge time 12 s 15%

Total cycle time 78 s 100%

Cycles/hr - 46.1

Corresponding filter drum speed - 0.77 r/min

7.3.4 Pressure Filtration

Pressure filtration is merely an acceleration of the vacuum filtration pro-
cess. Based upon the same principle (a pressure differential across a filter
media), pressure filters can consist of plates or of plates and frames.
Sludge enters the filtration chambers, designed so that liquid (filtrate)
passes through the filter medium while the solids (cake) are held within the
chamber. Sludge does not flow from chamber to chamber (series), but enters
each chamber independently (parallel) so that each area fills with solids at
the same rate, retaining the same quantity of filter cake and passing nearly
identical filtrates. The filtration process continues until the chamber area
is full or a predetermined terminal pressure is reached, completing the
cycle. The sludge feed is stopped and the chambers are opened to remove the
filter cake. The cycle is then repeated. Figure 7-1 illustrates a simple
manually operated filter press.

The rate of sludge filtration depends upon (1) feed pressure, (2) thickness
of filter cake, (3) sludge temperature and viscosity, (4) nature of the cake
solids, and (5) the filter medium.

Pressure filtration of mine drainage sludge or other waste sludges has not
been widely employed in the United States because of high labor, maintenance,
and capital costs. Rummel, in East Germany, investigated pressure filtration
of mine drainage sludge, and found that an influent slurry of 1.2% solids
could only be dewatered to a filter cake of 20%-30% solids after filtration
(12). The best filtration rate reported was 50 l/m2/hr (0.061 gal/m/ft2).
Unfortunately, this particular treatment plant generated 40,000 m3 (10,000
gal) of sludge per day. Thus, it was concluded that the filter output was
insufficient for the application.

Several sludges were dewatered by this method with moderate success. Table
7-6 presents the results of experiments conducted by Akers et al. (13). The
data indicate that good filtration rates could be obtained without floccu-

135
-------
Mine Water Source
TABLE 7-5

MINE WATER NEUTRALIZED SLUDGE SOLIDS FILTRATION RATES
Sludge Filtration Rates. 1b dry so1ids/ft2/24 hr
% Solids — - _
in Filter Fresh Old Pebble Dolomitic
Feed Hydrated Hydrated Dolomitic Hydrated
Slurry Lime Lime Lime Lime
Proctor

Proctor
Proctor

Proctor

Proctor
Bennett1

Bennett1

No. 1

No
No

No
s

. 1
. 2

. 2

. 2 (bio)

. 2 (bio)
branch

branch

Tyler run

2
2

7
1

.62

.53
.70

.19

.48

.26

.18
.96

.72

.58

.03

26
27

62
72

33
38
40
42

.4b
.8

.7
.0

,1
.6C
.1
.9

131
137
39
50
82
81
331

.0
.0
.6
.3
.5
.4
.0

54.

63.
64.

15.
21.
128.
125.

0
53.2
53.8
0
0

6
2
5
3
In this case, hydrated lime, stored in the plant lime storage silo for a
period of 3 months, converted to about 50% CaC03.

The double values reported for most tests represent duplicate determinations
made on each sludge slurry.

No precoat material employed.
lants. Also, it was concluded that pressures lower than 5.1 atm (60 Ib/in2g)
should not be used, because the principal advantage of this dewatering method
is a high filtration rate per unit of filter area requiring high pressure.

The capital and operating costs for a full-scale process are shown in Table
7-7 for a plant being built in 1972. Since then many of these prices have
doubled, thus adding to the unattractiveness of this method.

136
-------
CLEAR FILTRATE
OUTLET
FIXED HEAD SOLIDS COLLECT MOVABLE HEAD
IN FRAMES
PLATE FRAME
MATERIAL ENTERS
UNDER PRESSURE
Figure 7-1. Manually operated filter press,
-------
oo
00
TABLE 7-6

PRESSURE FILTRATION CAKE DATA
Sludge
Used
Norton
Edgell
Edgell
Edgell
Banning
Shannopin
Shannopin
Shannopin
Pressure
atm
5.083
5.083
6.444
7.805
7.805
5.083
6.4444
7.805
Ib/in2g
60
60
80
100
100
60
80
100
Thickness
of
Final Cake
cm
19.0500
1.2700
0.9525
1.4288
2.2225
6.0325
6.9850
7.3025
in
7.5000
0.5000
0.3750
0.5625
0.8750
2.3750
2.7500
2.8750
Filtration Time
to Produce
Cake
mm
179
198
138
175
219
170
180
200
Final
Solids of
Cake
%
20.8
26.2
26.0
26.0
11.8
8.7
12.0
10.9
All sludges were air-blown for 5 minutes after break.
-------
TABLE 7-7

PRESSURE FILTRATION - NORTON TREATMENT PLANT SLUDGE
(PRELIMINARY PRICES - SPRING 1972)

Capital
Equipment Costs Costs/yr Costs/d

1 48-in filter press $21,000.00
1 plate shifter 2,700.00
Feed pump with accessories 900.00
Precoat equipment 5,000.00
Construction and installation:
35% of equipment 10.000.00
Total equipment cost $40,000.00

Equipment depreciation $4,000.00 $ 11.00

Building: 1,200 ft2 at
$10.00/ft2 12,000.00

Building depreciation $ 400.00 1.00

Total capital cost $52,000.00

Operational costs

Maintenance: 6% of
total capital cost $3,120.00 8.50

Electricity: 270 kW-hr/d
at $0.0175 kW-hr 4.70

Labor: 2 men at 24 hr at
$6.00/hr each $288.00

Precoat: Johns-Manville Celite
501, $73.00/ton F.O.B.
Ca. warehouse, $105.40/ton
delivered Morgantown, W. Va. 50.00

Total capital and operational cost $363.20

Assuming 50,000 gallons clarifier underflow/d:
$7.30/1,000 gal sludge dewatered

Assuming 20,000,000 gal acid water/d: $0.02/1,000 gal acid water
139
-------
7.3.5 Porous Bed Drying

Sludge drying beds constructed of graded materials have long been employed as
a method of dewatering. A variety of filter media, such as sand, crushed
limestone, coal, red dog, and gravel, have been used. Water is removed in
this dewatering method by decanting the ponded surface water, by percolation
through the bottom of the bed, and by evaporation.

Figures 7-2 and 7-3 illustrate the drying bed constructed by Grube and Wil-
moth in their experiments to evaluate this method of sludge dewatering (14).
Modeled after drying beds traditionally used for sewage sludge, it is obvious
that detailed construction is required, although others have employed beds
with less material classification. For example, just two layers of materi-
als, crushed limestone beneath coarse sand, can be used, with a minimum depth
of 15.24 cm (6 in) for each medium.

Lovell summarizes the sludge drying bed area requirements in the design graph
shown in Figure 7-4.

Based on data from studies by Lovell, a 3,785 m3/d (1 Mgal/d) plant requires
6,691 m2 (72,000 ft2) of bed area. For all practical purposes, this method
of sludge dewatering could only be feasible for a plant with low sludge pro-
duction. Even then, the climate of the northeastern United States does not
favor this method.

Grube and Wilmoth described the operational difficulties encountered with the
application and removal of sludge from drying beds (14). They concluded that
drying bed operations that allowed sludge freezing rendered this method of
dewatering unacceptable.

Summer operation with lime-neutralized, coagulant-treated sludge provided the
following observations:

1. Approximately 20% by volume of the influent sludge (at 2.5% solids)
was retained on the sand bed as a 20% solids gel!ike mass, while 50%
drained through the sand as a clear effluent. The remaining 30%
was assumed to be lost to the atmosphere by evaporation.

2. The drainage rate through the sludge and sand averaged 26 1/d/m2
(0.6 gal/d/ft2).

3. Sludge solids appeared to stabilize near 20% within 20 days drying
time.

Although this method of sludge dewatering will eventually produce a cake,
removal from the bed remains a problem. Grube and Wilmoth attempted to re-
move the sludge with a rubber-tire highlift, but this proved almost impos-
sible and required rebuilding portions of the bed (14). The recommended
alternative was a trac-mounted loader, which functioned well. The large
volumes of sludge and weather conditions that occur in the areas where mine
drainage treatment plants are most prevalent (eastern United States) make
this method of sludge dewatering impractical.

140
-------
(Q
C
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ro
O

<
3
(O

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D-
O
3
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CO
01
3
to c
ro
00
TJOlN
5°
°3
z*"*
mi

3
3
^c
^Si*

19.2m (63ft) i
-^
15.2m (50ft) , 4m (13ft) _

— • SLOPED 0.3cm/m (2in/50ft)

10cm (4in)

PERFORATED PLASTIC PIPE

CONCRETE
APPROA^

APRON

PERFORATED
TJ
m
o>

Z
-------
!f
**
«»-

*.
A.
A^
.*
•
4.
^.
•'
f,
.'
,:
•A
.,
:
.0
&

\
0.80m (2.6ft) OF SLUDGE CAPACITY
20cm «8in) OF MORTAR- GRADE SAND (AASHO #M45)
f
13cm (Sin) OF "SHOT" GRAVEL
f
30cm (12in) OF MEDIUM STONE (AASHO #56)
1
, 15cm (6in) OF COARSE STONE (AASHO #4)
I 4 (.}
\ / A /
*!&'&/&<'/ / COMPACTED SUB-BASE //^//^///
\ / ^-2.5cm din) OF SAND MOcm (4in) PI
\ /_ - PLASTIC PIPE
30cm (12in)
CLASS B
CONCRETE
BITUMASTIC SEALER
Figure 7-3. Cross-sectional view of the drying bed construction.
-------
co
>-

«*
125-
100-
0.
(9
ta
o
UJ
o
UJ
o
O
UJ
CD
.50
J
DRAINAGE BASIN AREA REQUIREMENT-FT2 PER 1000 GPD CMD TREATED
100 150 200 250 300
_L
350
L
400
J
PPM NEUTRALIZED SOLIDS IN BED INFLUENT
Figure 7-4. Sludge drying bed sizing requirements.
-------
7.3.6 Centrifugation

Centrifugation is the separation of substances of different densities by the
use of centrifugal force. Many types of centrifuges are manufactured, but
all incorporate the following three basic operations: (1) a feed system
which delivers the sludge; (2) a revolving solid bowl or basket to collect
dewatering sludge; and (3) a sludge and skim (effluent) removal system.
Dewatered sludge can be removed either by a scraper blade or a screw con-
veyor.

Akers et al. conducted experiments on several sludges using a solid bowl
centrifuge. Table 7-8 summarizes results of the sludges tested (13). These
results vary drastically for each sludge, indicating that a centrifuge is
effective for dewatering some, but not all, sludges.

7.4 High-Density Sludge Process

Researchers from the Bethlehem Steel Corporation Research Department and
Bethlehem Mines Corporation cooperated in the development of a process de-
signed to provide both improved settling characteristics and increased sludge
solids concentration. The result of their research is the high-density
sludge process (HDS). This process reportedly achieves settled sludge con-
centrations between 15% and 40%, compared to a maximum of 15% from conven-
tional lime neutralization. The resulting sludge storage or disposal volume
is thus reduced by a significant factor.

The major differences between the HDS and the conventional lime neutraliza-
tion processes are shown in Figures 2-2 and 7-5. The HDS process has a
second reaction tank for mixing return sludge and lime slurry. All other
unit processes are the same as conventional lime neutralization. A clarifier
is necessary to provide adequate process control of sludge settling and
thickening and to provide an efficient means of sludge collection and return.

Research conducted on both pilot plant and full-scale operations has pin-
pointed several operating parameters that affect the HDS process. These
parameters are as follows (15):

1. ferrous-to-ferric iron ratio in the acid mine drainage;

2. recirculated sol ids-to-precipitated solids ratio;

3. point of alkalinity addition;

4. neutralization pH;

5. reaction tank detention time.

To elaborate, the following discussion of each is offered.
144

-------
TABLE 7-8

SUMMARY OF CENTRIFUGATION TEST*
Feed Solids (5)
Flow Rate
I/sec
(gal/min)
Bowl Solids
Skim Solids
Shannopin sludge
,27
,08
.09
,04
.24
1.85

Banning sludge

0.64
0.68
0.69
0.72
0.73
0.99

Norton sludge
1.
5.
,54
,52
10.05
3.77
.09
.48
.73
2.
2.
2.
6.12
0.03
0.06
0.09
0.12
0.15
0.03
0.03
0.09
0.12
0.15
0.18
0.03
0.03
0.03
0.03
0.03
0.06
0.09
0.12
0.06
(0.5)
(1.0)
(1.5)
(2.0)
(2.5)
(0.5)
(0.5)
(1.5)
(2.0)
(2.5)
(3.0)
0.5
(0.5
0.5
0.5)
0.5)
(1.0)
(1.5)
(2.0)
(1.0)
33.4
33.6
32.
36.
35.4
26.6
.5
.9
8.8
8.4
8.7
8.6
8.1
11.4
.3
.7
41.
53.
64.1
63.0
44.
50.
55.8
51.8
.9
.5
11.0
8.8
9.0
8.0
7.7
13.8
5.1
3.8
4.6
5.0
5.6
9.2
12
19
17
22
11.6
13.5
12.8
19.2
*Capital and operating costs ranged between $1.80 and $4.50 per 2.785 m3
(1,000 gal) of treated sludge based upon 1972 prices.

7.4.1 Ferrous-to-Ferric Iron Ratio

The percentage of ferrous iron in the acid mine drainage has a significant
effect on the maximum settled solids concentration that can be attained by
the HDS process. Test data have been used to generate a curve representing
the relationship between the sludge density and the ferrous iron percentage
in the influent (Figure 7-6) (2). This curve shows that ferrous iron per-
centages below 70% had relatively little effect on the concentration of
145
-------
AMD
HOLDING
POND
AIR
LIME SLUDGE
REACTION BASIN
NEUTRALIZATION
AND OXIDATION BASIN
^WASTE
SLUDGE
Figure 7-5. High-density sludge neutralization process.
sludge solids produced. For ferrous iron percentages greater than 70%, the
sludge concentration approached 50%. This highest sludge solids concentra-
tion was achieved under laboratory conditions with synthetic mine drainage.

7.4.2 Recirculated Solids-to-Precipitated Solids Ratio

The ratio of recirculated solids to precipitated solids has a direct effect
on the settled sludge density. The trend is for the sludge density to in-
crease as the ratio of recirculated to precipated solids is increased. This
increase in rapid up to a recirculation ratio of 20:1, moderate between 20:1
and 30:1, and small above 30:1. The recommended optimum recirculation range
is 25:1 to 30:1 (15). Operation within this range maximizes sludge density
while minimizing the clarifier area requirement.

7.4.3 Point of Alkalinity Addition

The HDS process operates successfully only when the lime slurry and recycle
sludge are mixed in a reaction tank prior to the addition of acid mine drain-
age and aeration. This represents the most critical step in the process.
Any other process arrangement results in the failure of the process to
146
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20 40 60 80

FERROUS IRON,%
100
Figure 7-6. Sludge density vs. ferrous iron percentage for the HDS process.
achieve the desired solids concentration.
7.4.4 Neutralization pH

The neutralization pH affects the settled sludge concentration, the oxidation
rate, and the settling area requirements. The optimum operating pH was found
to be within the 7.2-7.7 range for their AMD (16). Within this range, maxi-
mum sludge densities are produced, a satisfactory iron oxidation rate is
maintained, and mechanical clarifier area requirements are minimized. Opera-
tion within a pH range of 6.0-6.5 results in reduced lime usage, but the
ferrous iron oxidation rate is unacceptably slow. Operation within the pH
range of 8.2-8.7 reduces the sludge solids concentration from 35% to 20% and
increases the mechanical clarifier area requirement. Operation within the
9.0-9.5 range produces a rubbery sludge that hinders pump and clarifier
operation.
7.4.5 Reaction Tank Detention Time

Separate tanks must be provided for mixing the lime slurry with the return

147
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sludge and for the neutralization-oxidation reaction. A 1-minute detention
time in the sludge reaction tank is adequate to maintain a sludge of maximum
density. The detention time in the neutralization-oxidation tank is a func-
tion of the ferrous iron concentration and the operating pH.

The feasibility of this process has been demonstrated by Bethlehem at a test
plant near Ebensburg, Pa., and at several full-scale facilities. Sludge con-
centrations generally conformed to the predictions made by previous small-
scale tests (17). The HDS process, however, is best suited to treat high
ferrous iron mine drainages. Conventional neutralization process plants with
mechanical clarifiers can be easily converted to the HDS process by the addi-
tion of a sludge line and a sludge reaction tank.

In summary, the high-density sludge process offers a method for improving the
sludge concentration of the settled sludge. Nevertheless, an aura of skepti-
cism exists among designers in the field of mine drainage as to the actual
upper limits of performance. Currently, only one publication has reported a
sludge density of 10%-12% in a full-scale mine drainage treatment plant.

The Bethlehem Corporation holds a patent on the HDS process; thus, their
approval of certain designs is required.

7.5 Summary

The most practical and economical method of sludge disposal is pumping to an
abandoned deep mine. The designer should consider this alternative whenever
possible, but only after acquiring a permit that addresses the environmental
impact of this practice.

Another viable alternative is to thicken the sludge in a lagoon to a form
that is more easily handled. The thickened sludge can ultimately be disposed
of in a reclamation program for an active stripping operation, or mixed with
the refuse tailings from a vacuum filter for a coal preparation plant. Both
cases require a good control program to minimize operational problems. A
tailings-to-sludge ratio for a good spreadable mix varies with the individual
moisture content of the two materials involved. Probably the easiest method-
for determining a workable tailings or soil-to-sludge mixing ratio is by
trial and error. After a short period of time, the disposal operator will
become fairly efficient at judging the proper mixing ratio to achieve compac-
tion and burial. Obviously, if a large amount of sludge requires disposal
and insufficient quantities of dewatered tailings exist for satisfying a
proper ratio, another disposal method must be considered.

Table 7-9 presents a summary of the results and costs for several sludge
dewatering methods taken from studies conducted by Akers (13) and the Coal
Research Bureau (6).

In conclusion, sludge disposal will remain the most formidable problem facing
designers and owners of mine drainage treatment plants. Any method of sludge
handling, dewatering, and disposal will incur cost. The most economical
method is deep mine disposal. It is important, however, to emphasize that

148
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TABLE 7-9

COAL MINE DRAINAGE DEWATERING METHODS
Costs3
Dewatering Method

Vacuum filtration
Precoat vacuum filtration
Porous drying beds
Pressure filtration
Centrifugation
Single lagoon
Drying lagoon
Sludge
% Sol
8.8 -
11.4 -
15.0 -
8.7 -
8.1 -
0.5 -
12.0 -
Cake
ids
30.9
35.1
25.0
26.2
64.1
4.5
20.0
($/3.875 m3 (1,000 gal
dewatered sludge)
3.40
1.40 - 3.50
4.80 - 19.10
1.70 - 7.30
1.80 - 4.50
c
c
?Akers, February 1973
°After Lovell and Grube
Assessment based on land value and ultimate disposal (labor)
the environmental acceptability of this method is unknown at this time. In
fact, the environmental acceptability of all of the ultimate disposal tech-
niques needs to be assessed.

7.6 References

1. Holland, C.T., J.L. Corsaro, and D.J. Ladish. Factors in the Design of
an Acid Mine Drainage Treatment Plant. Second Symposium on Coal Mine
Drainage Research, Mellon Institute, Pittsburgh, Pennsylvania, May 1968.

2. Pennsylvania State University. Short Course on the Design of Coal Mine
Drainage Treatment Facilities. College of Earth and Mineral Sciences,
November 1973.

3. Lovell, H.L. An Appraisal of Neutralization Processes to Treat Coal
Mine Drainage. EPA-670/2-73-093, Environmental Protection Technology
Series, November 1973.
149
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4. Bisceglia, T.B. Hydrated Lime Versus Quicklime for Neutralization of
Waste-Acid Waters. Mercer Lime and Stone Company, Pittsburgh, Pennsyl-
vania, October 1966.

5. Wilmoth, R.C. Limestone and Lime Neutralization of Ferrous Iron Acid
Mine Drainage. EPA-600/2-77-101, Environmental Protection Technology
Series, May 1977.

6. Coal Research Bureau. Dewatering of Mine Drainage Sludge. Water Pollu-
tion Control Research Series, December 1971.

7. Dorr-Oliver, Inc. Operation Yellowboy - Mine Drainage Plan. Bethlehem
Mines Corporation, Mariana Mine No. 58, Pennsylvania Coal Research
Board, Department of Mines and Mineral Industries, Harrisburg, Pennsyl-
vania, January 1966.

8. Conrad, J.W. Proceedings of the Illinois Mining Institute Annual Meet-
ing. Springfield, Illinois, 1966.

9. Steinman, A.E. Coal Mine Drainage Treatment. Fortieth Annual Confer-
ence of the Water Pollution Control Federation of Pennsylvania, Univer-
sity Park, Pennsylvania, August 1968.

10. Pennsylvania Department of Environmental Resources. Mine Drainage
Manual. 2nd ed. Harrisburg, Pennsylvania, September 1973.

11. Glover, H.G. The Control of AMD Pollution by Biochemical Oxidation and
Limestone Neutralization Treatment. Prop. 22 Industrial Waste Confer-
ence, Purdue University, Part 2, 823-847, May 1967.

12. Rummel, W. Production of Iron Oxide Hydrate from Mine Waters in the
Lausitz Region. Institute Fuer Wasserwirtschaft, E. Germany, Wasser-
wirtsch - Wasser tech., 7: 344-348, 1957.

13. Akers, D.J., Jr., and E.A. Moss. Dewatering of Mine Drainage Sludge,
Phase II. EPA-R2-73-169, Research Series, February 1973.

14. Grube, W.E., and R.C. Wilmoth. Disposal of Sludge From Acid Mine Drain-
age Neutralization. National Coal Association, Bituminous Coal
Research, Inc., Sixth Symposium, Coal Mine Drainage Research, Louis-
ville, Kentucky, 1976.

15. Kostenbader, P.O., and G.F. Haines. High-Density Sludge Process for
Treating Acid Mine Drainage. Preprints of Papers Presented before Third
Symposium on Coal Mine Drainage Research, Coal Industry Advisory Commit-
tee to the Ohio River Valley Water Sanitation Commission, Pittsburgh,
Pennsylvania, 1970.

16. Temmel, F.M. Treatment of Acid and Metal-Bearing Wastewaters by the
High-Density Sludge Process. Prepared for Presentation at San Francisco
Regional Technical Meeting of American Iron and Steel Institute, Novem-
ber 1971.

150
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17. . Case History on Acid Mine Drainage Control. Presented at the
Mining Convention/Environmental Show of the American Mining
Congress, Denver, Colorado, September 1973.
151
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CHAPTER 8

ELECTRICAL REQUIREMENTS AND INSTRUMENTATION

8.1 Introduction

The purpose of this chapter is to inform the designer about electrical power,
motors, and various instrumentation. Explained herein are some common sense
rules in the design and arrangement of equipment, along with suggestions for
a clean, efficient operation.

8.2 Electrical Power

The design of equipment for a mine drainage treatment plant depends upon
quantity and quality of raw water to be treated. Consequently, process
units, mixers, aerators, sludge pumps, and other related equipment requiring
electricity must be sized accordingly.

The question of electric power supply is a very individual situation, depend-
ing upon the size of the plant and the form in which local electrical power
is available.

Three-phase power is almost always available and should always be used to (1)
keep conductor size down; (2) allow use of standard motors; (3) keep motor
maintenance at a minimum; and (4) provide the most reliable plant operation.

The voltage to be used (230 or 460) is somewhat dependent on what voltage is
locally available. In remote areas, high-tension power is the most likely
possibility; the local power company will provide the necessary step-down
transformers. In populated areas, even though 230- or 460-V (volt) power may
be available, the plant load requirements may still make it necessary for the
power company to provide a separate step-down transformer off the high-ten-
sion feeder. Simply stated, either 230- or 460-V 3-phase power can usually
be provided by the local power company. Actually, the problem of fulfilling
the electric power requirements for a plant can readily be solved as follows:

1. Determine from the equipment manufacturers the load requirements
(hp) of each motor (including the largest to be used in the plant).

2. Use 460-V supply if there are motors larger than 10 hp, since 460-V
supply is more economical.

3. Use motors of the same voltage throughout the plant.

152
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4. Always use 3-phase power.

5. After the above requirements are determined, contact the local
electric power company; they will decide how they can best provide
the service required.

8.3 Motors and Electrical Controls

Every motor requires a degree of control. The control is dependent upon the
sophistication required. As a minimum and in accordance with the usual
application, each motor is controlled from its own motor control center (MCC)
compartment, which contains the motor starter contactor, motor overload pro-
tection, and usually a "start-stop" push button and "run" light on the front
cover of the individual compartment. It is wise to provide a "lock-out"
feature in the MCC to protect maintenance personnel when a motor is out of
service.

More sophistication and larger plant controls may require a control panel on
which motor controls (usually start and stop) are located remotely from the
MCC, as well as fault annunciators. Control cabinets may also be required to
house control devices if automatic sequencing is involved.

No matter what degree of motor control is used, the control equipment should
be located in the cleanest area possible away from dust, fumes, and moisture.
A separate control room is usually the best solution.

The equipment manufacturers will recommend the appropriate housing for the
motors. This is directly related to the type of service and the environment
the equipment will experience.

8.4 Instrumentation

8.4.1 pH Control Systems

The heart of most mine drainage treatment plants is the pH control system.
Generally, it is wired to an interlock system that terminates operation if
the pH wanders outside preset limits (these limits are usually pH 6.0 and
9.0).

There are many types of pH electrodes on the market, each with advantages and
disadvantages. The designer should look for probes that can withstand daily
handling and a certain amount of abuse, and are easy to standardize. The
extrasensitive laboratory-type probes should be avoided, along with those
directly mounted in pipeline assemblies.

The typical immersion assembly, where the electrodes contact only the neu-
tralized stream and the cable assembly remains emerged, are most ideal for
mine drainage treatment processes. This assembly allows free access and easy
daily observation of electrode condition.

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Electrode fouling has long been a problem with mine drainage pH control
systems. Self-cleaning accessories for electrodes, however, are available
from many vendors. These cleaning mechanisms can be a membrane or nylon
brush or an ultrasonic transducer.

The membrane wiper is employed where electrode filming occurs, perhaps from
ferrous or ferric iron. The membrane has a reciprocating action (up-and-down
motion) across the measuring electrode. This will remove buildup of material
that can cause an error in measurement.

Another electrode cleaner is a nylon brush, which functions basically like
the membrane wiper, except that it is used where crusting solids (calcium
carbonate or calcium sulfate) deposit on the electrode.

These devices should be considered an auxiliary method for a more reliable pH
control system and should reduce the amount of maintenance required for clean
probes. Nevertheless, pH electrode cleaning remains a daily maintenance
duty. Therefore, provisions should be made by the designer for easy access
to the probes and with certain consideration for operation personnel. Probes
should not be positioned in confined areas or places that will expose person-
nel to unnecessary dangers.

The immersion pH assembly is found more often in mine drainage plants than
any other type, and is positioned in an open transfer trough where the mea-
sured liquid maintains scour velocity. If this type of arrangement is impos-
sible, a small auxiliary control loop in conjunction with a pump can be used.
More simply, a recirculatipn loop with flow-through pH electrodes has pro-
duced a good process loop with reliable performance.

The placement of pH electrodes inside the neutralization or flash mix tank
should be avoided. If unavoidable, certain provisions such as a still well
for the probes should be provided. An example of an arrangement is shown in
Figure 8-1.

When a pH probe is used as the primary control device, it is highly recom-
mended that it be positioned between the flash mix tank and the subsequent
treatment processes. Excellent control was observed by Wilmoth by placing
the pH probe in the inlet of the clarifier, thus minimizing problems of probe
fouling.

Many engineers/designers believe that a variable-speed lime feeder integrated
with a pH controller can be used to pace the addition of neutralizing agent.
Although this has been done successfully, it is subject to failure in highly
variable situations (extreme fluctuations in acidity), where the pH control-
ler has insufficient reaction time to changes in acidity. This "lag time"
causes the feeder to fall behind—sometimes to a point where it is impossible
to correct itself.

More successful systems have utilized on-off lime feeder arrangements with
enlarged flash mixing tank to provide a better buffering zone. Many existing
mine drainage plants utilize 1- to 3-minute detention volumes. However,
larger mixing tanks with 30-minute to 1-hour detention volumes have been

154
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UME SUURRV
TO RCCORDC*

rj*—pH PROBE
BAFFLE
— rtom
HALF-ROUND PIM
All 8TILL WIU
Figure 8-1. Flash mix tank with pH probe installed.
employed. This reactor allows more pH control reaction time and a steadier
state operation.
8.5 Level Controls

Level controls are incorporated into many of the processes used in mine
drainage treatment. Mercury float controls are used with the raw water
pumps, solid level indicators in the silo, and sensing probes or sonic level
controls in the slurry or stabilization tank.

Most raw water pumps, generally in the mine, are activated by hand and run
constantly. Sometimes, the equalization basin has transfer pumps that uti-
lize mercury controls or the conventional float with rod.

When using a lime silo, at least a high-level and a low-level indicator
should be installed. The exact position of these probes is up to the design-
er's discretion. The low-level indicator, however, should allow at least 1
or 2 days supply of lime when activated. This precaution will allow continu-
ous operation between deliveries.

Perhaps the best noncontact silo level indicator is the strain gauge. These
microcells are placed on the legs or skirt of the silo and measure the

155
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deflection of the steel. Equipped with a recorder, a display in the percent
full gauge can be read with remarkable accuracy. These strain gauges offer
an excellent choice as a solids level indicator.

The slurry or stabilization tank will probably contain the principal level
controls. These high and low levels, along with alarms, usually control
other equipment such as slurry pumps and lime feeder. These level controls
perhaps represent the most difficult application. In the past, a set of
physical probes (rods) have performed well, but required frequent cleaning.

Today's market offers the sonic level indicator that can be programmed to
control system equipment. These noncontact devices are an excellent alterna-
tive for nonconfined (i.e., trough, large mixing tank) liquid level measure-
ment.

These devices have performed poorly in lime silos, where "echoing" produces
erroneous readings.

Many types of controls can be employed. The decision becomes a designer
preference based upon experience and past success.

In summary, an initial awareness of electrical requirements with respect to
voltage, phase, transformers, and other hardware should alleviate later com-
plications and save costly construction time. Also, the instrumentation and
process control panel should be isolated from the active operation; ideally,
it should be placed in a separate building. The small additional capital
investment for this system will have a hundredfold return through a cleaner,
healthier environment for maintenance personnel, which will result in better
operation and longevity of equipment.
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CHAPTER 9

REVERSE OSMOSIS
9.1 Introduction
The application of reverse osmosis to the treatment of AMD has been exten-
sively studied by the U.S. Environmental Protection Agency over the last
decade. These studies have demonstrated that reverse osmosis (RO) can be
highly effective in removing most of the dissolved solids in acid mine drain-
age. The current purpose of applying reverse osmosis to AMD treatment is to
produce a potable water while achieving maximum recovery. The product water
will be low in dissolved solids, usually less than 100 mg/1, but may contain
chemical or bacterial constituents that exceed drinking water standards.
Reverse osmosis product water is not initially acceptable as potable water,
but it is of excellent quality and should be considered for other uses,
including boiler feedwater, cooling water, bathhouse shower water, or for a
variety of other industrial purposes. Although feedwater recoveries of 60%
or more can be obtained, design recoveries of 50% are more practical since
the emphasis is on producing specific quantities of high-quality effluents
and not on treating the entire volume of AMD by reverse osmosis.

This chapter is intended to provide the basic design criteria and operating
parameters applicable for using the reverse osmosis process. Using these
criteria and operating parameters, the designer can evaluate this process for
a particular application. It must be emphasized that reverse osmosis is a
complicated process as compared to other methods of treatment; therefore,
pilot testing must be performed to establish the design parameters as well as
pre- or post-treatment needs.

9.2 Operational Considerations

Osmosis occurs if two solutions of different concentrations-in the same sol-
vent are separated from one another by a membrane. If the membrane is semi-
permeable (i.e., permeable to the solvent and not to the solute), then the
solvent will flow from the more dilute solution to the more concentrated
solution until an equal concentration results. In reverse osmosis, the
direction of solvent flow is reversed by the application of pressure to the
more concentrated solution. As a result, the concentrated solution, termed
the solute or brine, becomes more concentrated. The solvent, termed the
permeate, is the product from the process.
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9.2.1 Membrane Type and Configuration

Tubular, hollow-fiber, and spiral-wound membrane types have been tested for
use in treating acid mine drainage. Studies performed by Wilmoth (1) in 1972
indicated that the spiral-wound configuration with a formamid-modified cellu-
lose acetate membrane was slightly superior to others with respect to the
average flux (permeate flow rate), long-term flux stability, and dissolved
solids rejection. Since then new membrane materials have become commercially
available. It is recommended that these be treated in the spiral-wound
configuration for any RO application.

9.2.2 Pretreatment

Problems with membrane fouling can occur as the concentrations of various
compounds increase during the process. Most important is the potential for
iron foulng and calcium sulfate (gypsum) formation. Iron fouling has been
minimized by lowering the feedwater pH, or flushing the RO membrane by oper-
ating at lower pressures for short periods. When the raw AMD contains high
concentrations of sulfates, gypsum can form if its solubility is exceeded.
In this case, this process may not be applicable.

9.2.3 Prefiltration

Cartridge or bag filters should be used to minimize fouling by suspended
solids in the feedwater. This can increase membrane life and improve rejec-
tion levels. The filters should be capable of removing particles larger than
20 urn. The filters are placed at the suction side of the RO feed pumps.
Duplicate units should be provided in parallel to eliminate the need to shut
down in the RO system when cleaning or filter replacement is necessary.

9.2.4 pH Control

In treating AMD, the pH of the feed should be maintained between 2.8 and 3.0.
Adjustment of the feed pH is required to prevent the precipitation of
slightly soluble inorganic salts such as calcium and iron. At a pH less than
3.0, ferric iron (Fe3*) remains dissolved. When pH values exceed 3.0, ferric
hydroxide may begin to precipitate on the membrane surface.

Although a low pH is necessary to improve operating conditions when treating
AMD, it is lower than the optimum range of 5.0-6.5 for the cellulose acetate
membranes. The life of this common membrane formation will be decreased, but
newer, pH-resistant membranes are now available.

9.2.5 Disinfection

Another method used to reduce iron fouling problems is to provide disinfec-
tion to inhibit microbial activity in the raw mine drainage feed. Ultravio-
let light, proven to be an effective bactericide, is recommended to prevent

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an accumulation of iron-oxidizing bacteria on the membrane surface.

9.2.6 Space Requirements

In the design of any system, sufficient space must be available to accommo-
date the entire system, including all major mechanical and electrical equip-
ment, auxiliary equipment, and storage facilities. The initial design should
consider the ease of installation and modularity, or the ability to add
modules, stages, or racks to the system as needed.

9.2.7 Design Capacity

A loss in flux or permeate flow rate, due to compaction or chemical fouling
of the membrane surface, will gradually reduce the production capacity of the
system. This should be offset by appropriate sizing during the initial
design. The RO system should be sized to process the daily AMD flow in 20
hours at the average flux rate. This provides an adequate safety margin for
processing the daily flow and sufficient time for daily maintenance and
membrane cleaning. The designer should also consider providing storage
capacity in the event the system is out of service for a prolonged period.
This can be accomplished during the initial design phase by oversizing feed-
water holding tanks or by including larger storage ponds. There is no spe-
cific storage capacity design value, but a volume of several days flow should
be considered.
9.2.8 Flux Rate—Feed, Product, and Concentrate

One of the design factors critical to a successful RO operation is an accu-
rate permeation rate (flux rate) over the life of the membranes. This is
essential to estimate the quantity of installed surface area, cleaning cyles,
and membrane replacement. Initially, with the feed rate constant, a decline
of permeate flow will occur due to membrane fouling. Even without fouling, a
slight flux decline will be observed because of membrane compaction.

The system should be designed to produce a constant permeate output based on
the daily design flow. This is normally accomplished by using pressure-
compensating flow controls that automatically adjust for flow variations.
Once preset, the control valve automatically adjusts the operating pressure
to maintain the permeation rate at its predesignated flow rate.

Tests were performed on a 15.14-m3/d (4,000 gal/d), once-through, continuous
operation; an average flux rate of 500 l/m2/d (12.3 gal/d/ft2) operating at
28.2 atm (400 Ib/in2g) and 75% recovery was realized.

It is difficult to state a specific flux rate for design purposes, since flux
varies with operating pressure, concentration of the feed stream, and overall
recovery. Perhaps one of the best methods for determining a flux rate is to
extrapolate annual flux values,from pilot tests and apply these to the design
of a full-scale system.

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9.2.9 Operating Pressure

The system should be designed to operate at 281 kg/cm2 (400 lb/in2). At this
operating pressure, minimum membrane compaction will be experienced while
maintaining adequate flux ratios to assure high effluent quality. (Effluent
quality will decrease as operating pressure decreases.)

As the flux declines, the pressure control system will compensate for the
decreased flow by increasing the operating pressure, thus maintaining a
constant product flow. A loss in flux, due to compaction, is typically off-
set by appropriate sizing and startup at a reduced pressure with gradual
increase in operating pressure over the life of the modules. To minimize
fouling, each module should operate with a 10:1 brine-to-produce flow ratio.
This brine velocity presents "boundary layer" development, which is a layer
of stagnant water against the membrane surface.

9.2.10 Module Configuration

The type of pressure vessel manifolding arrangement (series or parallel) is
dictated by the desired recovery level and the need to maintain an adequate
brine-to-product flow ratio. Pressure vessel arrangements and modules are
designed so the raw feed enters a parallel bank of pressure vessels, and the
concentrate from this bank is used as the feed for the next parallel arrange-
ment of vessels.

In high-recovery continuous flow systems, it is advantageous to design the
system so only a small number of modules have to process the most concen-
trated portion of feed stream. Then if fouling due to chemical precipitation
occurs, it is confined to a minimum number of modules.

In research and testing done at the EPA Crown Mine Drainage Control Field
Site, the vessel array configurations and corresponding recovery levels
were as follows (1):

Arrangement Recovery Level
%

8-6 40-60

7-4-3 60-75

5-4-3-2 75-80

5-4-2-2-1 85-90

At high recovery levels of 85%-90%, only one pressure vessel (6 modules) would
be subjected to the most concentrated material. Should precipitation occur,
only these modules would be severely fouled.
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9.2.11 Safe Recovery Levels

There are no established design parameters for determination of ultimate
recovery levels. Rather, two controlling factors limit the overall recovery
of water from the treatment process.

The first is the precipitation of calcium sulfate (CaSOit). Acid mine waters
typically contain high concentrations of calcium and sulfate ions. As the RO
process progresses, calcium and sulfate concentrations increase. Once the
solubility of calcium sulfate is exceeded, precipitation occurs. Wilmoth
determined the empirical limit for RO recovery levels with AMD to be R =
100 - 0.55 (Ca x SOO"^ where Ca and S0*f concentrations (mg/1) are those used
in the raw feed. It is well known that gypsum (calcium sulfate) is only
slightly soluble. When concentrated, it will precipitate and form a hard,
tenacious scale on tanks, piping, and, more importantly, the membranes.

The second factor influencing recovery rates is the desired quality of the
permeate. As the drainage is processed, the concentration of total dissolved
solids (TDS) in the permeate increases almost linearly with the TDS in the
concentrate. Thus, increasing recovery increases the concentration of pollu-
tants in the waste (reject) stream and in the product water. The final use
of the permeate determines the maximum recovery of the process.

9.2.12 Dissolved Solids Rejection

The percentage rejection of waste stream contamination is greater than 90%;
in most cases, the spiral-wound cellulose acetate membranes will reject 99%
or more of the dissolved salts in the raw AMD feed. Table 9-1 shows antici-
pated permeate water quality.

As the pollutants in the reject are concentrated, more dissolved solids con-
tact the membrane; thus, more will pass through the membrane and deteriorate
the permeate quality.

9.2.13 Membrane Life Expectancy

Membrane life is affected by pH, temperature, and operating pressure. One
manufacturer defines useful membrane life as the time taken for the membrane
to lose 40% of its initial flux. Any attempt to accurately determine how
long a membrane will last under normal operating conditions-must be accom-
plished through long-term studies. Under applications of this nature, manu-
facturers project a cellulose membrane life of 3 years. For operating cost
projections, it is suggested that the designer assume that 50% of the mem-
brane area will be replaced each year.

9.2.14 Concentrate Treatment and Disposal

As the RO system removes dissolved solids, the process generates a highly
concentrated waste stream that requires treatment before disposal. The exact

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TABLE 9-1

ANTICIPATED PERMEATE WATER QUALITY
Parameter3
pH (units)
Specific conductance ( mhos)
Acidity
Calcium
Magnesium
Iron, total
Iron, ferrous
Aluminum
Manganese
Raw Water
Quality
3.4
1,020
210
150
115
110
71
15
43
Product Water
Quality
4.3
32
32
1.2
1.4
1.2
0.8
0.8
0.4
aAll values expressed as mg/1 unless otherwise noted.
volume and salt content of the concentrate stream depends on the influent
quality as well as the recovery rate. For example, an RO system with a 90%
recovery rate creates a waste stream with a pollutant concentration 10 times
that of the feedwater, but with a volume of only 10% of feed. This concen-
trate must be treated and/or disposed of in some environmentally acceptable
manner. The possible treatment and disposal methods include the following:

1. lime neutralization;

2. evaporation - mechanical and/or atmospheric;

3. contract disposal.

Lime neutralization of the waste stream from the RO process is a practical
disposal and treatment method. This method is adequately described in this
manual.

Possible evaporation techniques include the mechanical, wiped-film unit,
which is capable of reducing the volume by 75% or more. In drier parts of

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this country, evaporation ponds can be considered. These provide for atmos-
pheric evaporation to reduce the volume, possibly to dryness.

Contract hauling and disposal by an approved waste hauling firm is another
alternative.

9.2.15 Capital and Operating Costs

Accurate cost data relative to RO treatment of acid mine drainage are virtu-
ally nonexistent. Extensive studies of actual commercial units necessary to
develop these costs have never been performed. Capital and operating costs
presented in the literature are only estimates based on system size and
degree of sophistication. Operating costs are also estimates, for the most
part based upon assumed values.

9.3 References

1. Wilmoth, R.C. Applications of Reverse Osmosis to Acid Mine Drainage
Treatment. EPA-670/2-73-100, Environmental Protection Technology
Series, Cincinnati, Ohio, December 1973.

9.4 Other Selected Readings

Rex Chainbelt, Inc. Reverse Osmosis Demineralization of Acid Mine Drainage.
Program No. 14010 FQR, U.S. Environmental Protection Agency, Water Pollution
Control Research Series, March 1972.

Gulf Environmental Systems Company. Acid Mine Waste Treatment Using Reverse
Osmosis. Program No. 14010 DYG, U.S. Environmental Protection Agency, Water
Pollution Control Research Series, August 1971.
163
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CHAPTER 10

ION EXCHANGE
10.1 Introduction
Ion exchange, like reverse osmosis, can be utilized to treat acid mine drain-
age for the removal of unwanted dissolved ions to produce a water of excel-
lent quality for many industrial uses. Ion exchange can also produce a
potable grade of water; but such a system would have to be followed by fil-
tration and disinfection to comply with public health regulations.

Ion exchange in water treatment is defined as the reversible interchange of
ions between a solid medium and the aqueous solution (1). To be effective,
the solid ion-exchange medium must contain ions of its own, be insoluble in
water, and have a porous structure for the free passage of the water mole-
cules. Within the solution and the ion-exchange medium, a charge balance or
electroneutrality must be maintained; i.e., the number of charges, not the
number of ions, must stay constant. Ion exchange materials usually have a
preference for multivalent ions; therefore, they tend to exchange their
monovalent ions. This reaction can be reversed by increasing the concentra-
tion of monovalent ions. Thus, a means exists to regenerate the ion exchange
material once its capacity to exchange ions has been depleted (2, 3).

The most common ion exchange use is the softening of "hard" or mineral-bear-
ing water for domestic or commercial purposes. The hardness in water is
attributed to its calcium and magnesium content. Initially, the ion exchange
material is charged with monovalent cations, usually sodium (sodium chlo-
ride). The hard water is passed through a bed of ion exchange material, and
the divalent calcium and magnesium cations are exchanged for sodium ions as
fol1ows:

Ca++ + 2Na+ (resin) <*• Ca++ (resin) + 2Na+ (30)

Ion exchange materials tend to form stable compounds through this exchange
principle. When more than one type of cation is available, the material will
have an affinity for certain ones more than others. In commercial or indus-
trial applications, the ion exchange resin is usually operated in the proton
(H+) or acid cycle. Here, sodium is replaced with a proton (H+) and the
exhausted resin is regenerated with sulfuric or hydrochloric acid.

The earliest ion exchange materials were either natural or synthetic zeo-
lites--a mineral produced from mixtures of aluminum salts and silicates. In

164
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the 1930's, plastic materials called resins were developed that expanded the
applications of ion (cation) exchange. An anion exchange resin was developed
in 1949 that enabled the process to be used for total demineralization of
water. In the present-day technology of ion exchange, the resins available
can be classified as strong-acid cation, weak-acid cation, strong-base anion,
and weak-base anion types. Combinations of the available resins have been
used in systems for treatment of different waters for specific purposes (4,
5).

The application of these ion exchange systems for the treatment of mine
drainage has been studied mainly to produce potable water where a reduction
in the total dissolved solids is required. Processes developed include the
Sul-biSul Process, the Modified Desal Process, and the Two Resin Process.
The operation and performance of the first two of these processes have been
demonstrated in full-size facilities; the latter process, in pilot units. It
has been concluded that ion exchange can be used to demineralize mine drain-
age and produce water with a quality acceptable for potable or industrial
use, but the costs of operation do not appear competitive with other methods.

10.2 Sul-biSul Process

The Sul-biSul Process was developed by the Nalco Chemical Company but is now
assigned to the Dow Chemical Company (6). The process employs a two- or
three-bed system, depending upon the mine drainage quality. Cations are
removed by a strong-acid resin in the hydrogen form, or by a combination of
weak-acid and strong-acid resins (7, 8). The AMD feed is first passed
through the cation exchanger, which removes the metal cations and exchanges
these for hydrogen protons, or (H+) ions. This reaction is expressed by
Equation 31.

Fe+2SOlf + 2HR-^ MR2 + H2S04 (31)

where R represents the strong-acid exchange groups on the resin, and M repre-
sents a divalent metal cation, such as iron (ferric), calcium, or manganese.

The product water from this first exchange contains additional sulfuric acid
from the displaced proton (H+). Following this, the water is decarbonated to
remove carbon dioxide formed during the cation exchange process. Then a
strong-base anion resin (R1) operating in the sulfate-to-bisulfate cycle
removes both the sulfate and hydrogen ions during this exchange reaction
(Equation 32):

R'SO^ + H2S(\+ R'(HS002 (32)

because of the high acidity of the feed, sulfate ions in solution and on the
resin are converted to the bisulfate form. This conversion of bivalent sul-
fate to monovalent bisulfate provides for twice the amount of sulfate to be

165
-------
stored on the resin. Removal of the sulfate results in good-quality water.
Regeneration of the cation exchange bed is accomplished with either hydro-
chloric or sulfuric acid. In the regeneration of the anion bed, bisulfate
ions are converted back to the sulfate form by the feedwater in a reversal of
Equation 31. The addition of lime slurry to the regenerant will speed this
reaction. The unusual feature of this process is the removal of sulfate from
feedwater by anion exchange using only water or water with a little added
alkali as the regenerant (9). The product water must be filtered and chlori-
nated according to public health regulations before use as a potable water.
Wastes from the regeneration process would have to be treated before dis-
charge.

The Sul-biSul Process can be used to demineralize brackish water containing
predominantly sulfate anions with a dissolved solids content up to 3,000
mg/1. The raw water should have an alkalinity content about 10% of the total
anion concentration and a sulfate-to-chloride ion ratio of at least 10:1.
The process is especially suited to alkaline waters containing calcium sul-
fate, such as those contaminated by mine drainage (9).

Limitations of the process center on the low exchange capacity of the anion
exchange resin and its inefficient method of regeneration. The exhausted
anion resin can be regenerated by the raw water itself; however, this re-
quires a considerable volume of water and takes a significant length of time
if the sulfate content is low. The addition of a cheap alkali such as lime
is reported to improve the regeneration; however, one study showed poor
results (10).

One problem is the requirement for disposal of this large volume of regener-
ants. This water must be sufficiently alkaline and abundant so that it can
be used as the regenerant and then discharged to the stream. If the raw
water cannot be used as the anion bed regenerant, other alkalis must be
employed. When this is necessary, tests have indicated that there may be a
negative net production of water and the process may not be economically
competitive (9).

A water treatment plant using this process began operation in 1971 at Smith
Township, Pa. The plant was designed for the production of 1,900 m3/d (0.5
Mgal/d) of potable water. The raw water supply at Smith Township is a small
stream that is affected by mine drainage. Studies indicated that it con-
tained sulfates, iron, and manganese, yet remained alkaline. Pilot tests
confirmed the applicability of the Sul-biSul Process (11). Projected raw and
finished water quality for this plant are shown in Table 10-1, and a flow
schematic of the system in Figure 10-1 (12).

Cost data for the Sul-biSul Process are limited to the projections from a few
studies and the Smith Township plant. At Smith Township, a continuous ion
exchange-regeneration system was installed at a capital cost of $898,000.
Operating costs are not available because the plant did not meet design
capacity specifications and is not being operated pending litigation.

Projected operating costs were estimated to be in the range of $0.10-$0.13/m
($0.40-$0.50/1,000 gal). These costs have more than doubled if extrapolated

166
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TABLE 10-1

PROJECTED RAW AND FINISHED WATER QUALITY
SUL-BISUL PROCESS AT SMITH TOWNSHIP, PA.
Typical Quality
Parameter Raw Water Finished Water

pH 6.5 - 8.4 8.0

Alkalinity, mg/1 76 10 - 30

Dissolved solids, mg/1 1,500 - 2,000 300

Sulfates, mg/1 400 - 1,300 50 - 100

Hardness, mg/1 1,600 150

Chlorides, mg/1 16 2
to a 1980 cost basis. Unless a sufficient supply of excessively alkaline
water is available for regeneration of the anion resin, the Sul-biSul Process
cannot economically produce potable water from acid mine drainage (9).

10.3 Modified Desal Process

The Modified Desal Process is another ion exchange process that has been
investigated for treatment of AMD to recover potable water (8, 9, 10, 13).
This process uses a Weak-base anion resin in the free-base form, which is
converted to the bicarbonate form to treat the raw AMD. The weak-base resin
exchanges sulfates (or other anions) for bicarbonate, allowing the cations to
pass through the bed according to the following reaction:

VSOn + 2R"HC03 + R"2SQk + M(HC03)2 (33)

where R" = is the weak-base exchange group on the resin matrix

M = a divalent metal ion

The solution of metal bicarbonates is aerated to oxidize ferrous iron to
ferric iron and to purge the carbon dioxide gas. The effluent is then
treated with lime to precipitate metal hydroxides, settled to remove sus-
pended solids, and then filtered and chlorinated if it is to be used as a
potable water (7, 9, 13).

167
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RAW WATER
REGENERANT^
SULFURIC OR
HYDROCHLORIC
ACID
LIME
SLURRY
STRONG - ACID
WEAK - ACID
CATION EXCHANGER
DECARBONATOR
1
STRONG- BASE
ANION EXCHANGER
I
LIME
SAND FILTRATION
NEUTRALIZATION
TANK
DISINFECTION AND
pH ADJUSTMENT
1
1
MECHANICAL
CLARIFIER
POTABLE WATER SUPPLY
TREATED WATER
TO STREAM
Figure 10-1. Sul-biSul Process continuous ion exchange flow sheet.

168
-------
Ammonia is used as the alkaline regenerant to displace sulfate from the
exhausted resin. Lime is used to precipitate the ammonia regenerant for re-
use. In this way, ammonia is recycled in the process. It is possible to
recover the carbon dioxide and lime used in this process by roasting lime
sludge wastes in a kiln. If this were done, all of the principal chemicals
used in the process would be recycled. The net result would approach a zero
discharge process, with fuel, power, and makeup quantities of chemicals fed
into the plant, and only potable water, iron hydroxide, and calcium sulfate
sludge discharged from it (9).

The Modified Desal Process is not limited by total dissolved solids or pH
levels; however, large quantities of carbon dioxide are required to achieve
good resin utilization for high total dissolved solids or alkaline feed
waters. The process is limited in application to waters containing less than
2,200 mg/1 of sulfate. Another limitation is that mine waters containing
iron in the ferric form may cause fouling of the anion bed because of precip-
itation of ferric hydroxide (9, 13).

A demonstration plant for treatment of AMD by the Modified Desal Process was
constructed in 1972-73 by the Pennsylvania Department of Environmental
Resources at Hawk Run near Philipsburg, Pa., at a capital cost of $2,335,000.
The purpose of this plant was to demonstrate the applicability of the Modi-
fied Desal Ion Exchange Process for treating acid mine drainage. Its second-
ary purpose was to provide a drinking water supply with a capacity of 1,892.5
m3/d (0.5 Mgal/d) for the nearby Philipsburg area. A typical water analysis
is shown in Table 10-2.

Operating costs in 1975 were $118,925 or $0.14/m3 ($0.54/1,000 gal) of plant
capacity based on an output of 2,271 m3/d (0.6 Mgal/d). This operating cost
does not include depreciation of the capital cost,.which was estimated to be
$214,000 annually. It should be pointed out that this plant was designed and
operated as a research facility, so all costs are higher than would be
expected in a normal production facility (13). Estimated operating costs for
a 750 mg/1 sulfate feed were $0.49/m3 ($1,85/1,000 gal) including deprecia-
tion (9). The 1975 actual costs of water produced by the Modified Desal
Process were about $0.48/m3 ($1.82/1,000 gal) when depreciation was included.

Several optimizing modifications have been made at the Hawk Run Plant to
increase its efficiency. An important one involves precarbonating the acid
mine drainage, which has resulted in a significant increase in capacity from
1,893 to 3,028 m3/d (0.5 to 0.8 Mgal/d). A schematic of this process at Hawk
Run is shown in Figure 10-2.

The waste regenerant is composed of an ammonium sulfate solution. This is
1ime-treated to form calcium sulfate, which is then removed by filtration.
The filter effluent is sent to a distillation process where 92%-95% of the
ammonia is recovered for reuse as the first-stage regenerant.

The Hawk Run Plant was constructed to demonstrate the process for augmenta-
tion of a degrading water supply. Water quality has now improved to the
ipoint that the plant is currently not needed, and it has been placed in the
standby mdde pending its future use. The plant has also been offered for

169
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TABLE 10-2

TYPICAL WATER ANALYSIS
HAWK RUN AMD TREATMENT PLANT
Raw Product
Water Water
Hot free mineral acidity
(mg/1 CaC03) 384

Free mineral acidity
(mg/1 CaC03) 362

Iron (mg/1) 101 0.2

Calcium (mg/1 CaC03) 295 85

Magnesium (mg/1 CaC03) 100 99

Total hardness (mg/1 CaC03) 395 184

Sulfate (mg/1) 684 192

Total dissolved solids (mg/1) 1,084 284

pH (Standard Units) 3.7 9.5
sale to the nearby community for $1, but it was not accepted because of its
high operating cost compared to other water treatment processes (13). While
it was operated, it performed extremely well (14).

10.4 Two Resin Process

In 1972, the Culligan International Company (10) investigated a standard
two-resin system. This process has been further investigated by the U.S.
Environmental Protection Agency at their Crown, W. Va., Research Facility
(14).

The Two Resin Process involves the use of a strong-acid cation exchanger in
the acid (H+) form followed by a weak-base anion exchanger in the free-base
(OH~) form as shown in Figure 10-3. In the cation column, protons (H+ions)
are exchanged for the metal ions in the acid mine drainage. Following cation
exchange, the anion column feed is rich in sulfuric acid.

Total metal cation removal greatly increases the regenerant dosage and the
operating cost of the system. It has been demonstrated that significant cost

170
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MINE DRAINAGE
LIME
WEAK BASE

CATION

EXCHANGER
DECARBONATOR
I AERATOR
\
REGENERANT
DISPOSAL
OR
RECOVERY
/

/ IRON
SETTUNG
\
SLUDGE
/ LIME
\ SLUDGE
LIME SOFTENING
LAGOON
PRODUCT WATER
Figure 10-2. Modified Desal Process flow diagram.
reductions can be realized by operating the system consistent with the needs
for product water end-use. The concentration of residual metals in the
cation exchanger effluent can be optimized by controlling the dosage of
regenerant.

Feed to the anion exchanger is predominantly sulfuric acid, which is totally
absorbed by the resin. A weak-base anion exchange resin only absorbs acids;
it cannot split neutral salts. The anion exchange effluent is alkaline, and
some precipitation of residual iron and aluminum ions can be expected. The
effect of this accumulation on the anion resin efficiency and capacity must
be monitored, but was observed by Wilmoth to have no significant adverse
effects in 900 regeneration cycles.

Either sulfuric acid or hydrochloric acid may be used for regenerating the
cation exchanger. Sulfuric acid is usually preferred because of its lower
cost; however, gypsum may form if sulfuric acid is used. If so, ti2SO^ regen-
erate concentration must be limited to 2% by weight. Treatment for disposal
of both regenerant streams is necessary.

Water quality results from the EPA Crown studies are summarized in Table
10-3. As could be expected. *n increase in the cation regenerant dosage
171
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SULFURIC
ACID
SODIUM
HYDROXIDE
ACID MINE
DRAINAGE
TO WASTE

N
TO WASTE
X
S

BACKWASH )
^
t

\
/
CATION
EXCHANGER

s
BACKWASH

'x
N

/
\
\

\
/
BACKWASH)/
ANION
EXCHANGER

"s
/BACKWASH v
/

TO WASTE
(REGENERATION
AND RINSES)
TO WASTE
(REGENERATION
AND RINSES)
PRODUCT
(TO pH ADJUST-
MENT, FILTRATION,
8 CHLORINATION)
Figure 10-3. Two-resin ion exchange system.
172
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TABLE 10-3

SUMMARY OF ION EXCHANGE SYSTEM CHEMICAL ANALYSES*
GO
Sampl e
Acid
Cond
Total
Ca Iron Fe2+
Na
Al
Mn
j*All units are expressed as mg/1 except for conductivity (^mhos/cm) and pH (standard units).
Acidity, alkalinity, and exchangeable cations are expressed as
TDS

Raw feed
Cation effluent
An ion effluent

Raw feed
Cation ef f Tuent
An ion effluent

Raw feed
Cation effluent
An ion effluent

2,870
8,890
1,230

2,870
8,910
1,370

2,790
9,700
1,430
OPERATION AT
410
1,780.
280°
OPERATION AT
410
1,730.
340b
OPERATION AT
430
2,000.
290°
48-g/l
4.9
1.54
9.4
96-g/l
5.1
1.55
9.3
144-g/l
5.0
1.58
9.5
(3-lb/ft3) DOSAGE OF
400 210
42 23
38 3.1
190
22
0.3
(6-lb/ft3) DOSAGE OF
330 200
29 14
21 2.4
190
13
0
(9-lb/ft3) DOSAGE OF
340 210
24 16
19 1.4
LONG-TERM OPERATION AT 48-g/l (3-lb/ft3
Raw feed
Cation effluent
An ion effluent
2,770
8,220
1,310
500
1,930.
290b
4.5
1.61
9.3
350 180
45 22
38 6.7
200
13
0
SULFURIC ACID
360
260
320
9.3
0.6
0.1
5.0
0.6
0.4
2
2

,600
,460
610
2,690
2,790
980
SULFURIC ACID
340
240
380
8.7
0.6
0.3
5.0
0.3
0.2
2
2

,430
,380
570
3,410
2,670
970
SULFURIC ACID
350
190
400
) DOSAGE OF
170
19
1.1
330
250
330
7.8
0.7
0.5
SULFURIC
8.5
0.74
0.21
5.2
0.8
0.2
ACID
5.3
0.62
0.44
2
2

2
2

,470
,280
730

,440
,370
660
3,500
2,550
1,150

3,420
2,700
1,050
-------
resulted in a decrease of the divalent cations in the exchanger effluent.
The raw AMD was found to be very high in sodium, but very little removal of
this monovalent ion was achieved. The anion exchange column effectively re-
moved all acidity and imparted alkalinity. Precipitation of iron within the
column was also observed, but no deleterious effects could be documented.

Long-term testing was performed at minimal cation regenerant dosage. A
deterioration in the performance of the cation column was observed. Although
the utilization efficiency of the regenerant decreased, the product water
quality remained acceptable. Unexpectedly, there was no apparent reduction
in the anion exchanger efficiency, even though iron precipitation did occur.
Using the low-pH cation column effluent for backwashing the anion column
effectively prevented the iron hydroxide precipitate from inhibiting anion
column performance.

Except for the presence of sodium ions, the water produced at Crown would
meet potable requirements following filtration and chlorination. Estimated
capital costs for plants utilizing the Two Resin Process on a 1978 cost basis
range from $1,000,000 for a l,893-m3/d (0.5 Mgal/d) plant to $1,700,000 for a
facility capable of producing 3,785 m3/d (1.0 Mgal/d) (10). For the Crown
raw water, which is more polluted than most AMD, operating costs appear to be
about 50% greater than those for the Modified Desal Process.

10.5 References

1. The Dow Chemical Company. Dowex: Ion Exchange. The Dow Chemical
Company, Midland, Michigan, 1964.

2. Calmon, C. Modern Ion Exchange Technology. Industrial Water Engineer-
ing, April/May 1972.

3. Fair, G.M., J.C. Geyer, and D.A. Okun. Water and Wastewater Engineer-
ing. Vol. 2. John Wiley & Sons, New York, 1968.

4. Lynch, M.A., Jr., and M.S. Mintz. Membrane and Ion-Exchange Processes
— A Review. Journal American Water Works Association 64(11), 711-719,
1972.

5. The Oow Chemical Company. Fundamentals of Ion Exchange. Idea ± Ex-
change 1 (1), January 1971.

6. U.S. Department of the Interior. Sul-biSul Ion Exchange Process —
Field Evaluation on Brackish Waters. Progress Report No. 446, Office of
Saline Water, May 1969.

7. Burns and Roe, Inc. Preliminary Design Report — Acid Mine Drainage
Bemonstration Project, Philipsburg, Pennsylvania. Report to the Penn-
sylvania Department of Mines and Mineral Industries, 1969.
174
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8. Pollio, F., and R. Kunin. Ion Exchange Processes for the Reclamation of
Acid Mine Drainage Waters. Environmental Science & Technology 1 (3),
March 1967.

9. Burns and Roe, Inc. Evaluation of Ion Exchange Processes for Treatment
of Mine Drainage Waters. Report to the Commonwealth of Pennsylvania
Department of Environmental Resources and the U.S. Department of the
Interior, Office of Saline Water, December 1973.

10. Holmes, J., and E. Kreusch. Acid Mine Drainage Treatment by Ion Ex-
change. EPA-R2-72-056, Environmental Protection Technology Series,
Washington, D.C., November 1972.

11. Zabban, W., T. Fithian, and D.R. Maneval. Conversion of Coal-Mine
Drainage to Potable Water by Ion Exchange. Journal American Water Works
Association 64 (11), November 1972.

12. Skelly and toy and Penn Environmental Consultants, Inc. Processes, Pro-
cedures, and Methods to Control Pollution from Mining Activities.
EPA-430/9-73-011, Washington, D.C.

13. Kunin, R., and J.J. Demchalk. The Use of Amberlite Ion Exchange Resins
in Treating Acid Mine Waters at Philipsburg, Pennsylvania. Rohm and
Haas Company, Philadelphia, Pennsylvania.

14. Wilmoth, R.C., R.B. Scott, and E.F. Harris. Application of Ion Exchange
to Acid Mine Drainage Treatment. 32nd Annual Purdue Industrial Waste
Conference, May 1977.

10.6 Other Selected Readings

Rose, J.L. Treatment of Acid Mine Drainage by Ion Exchange Process. Pre-
prints, Third Symposium on Coal Mine Drainage Research, Mellon Institute,
Pittsburgh, Pennsylvania, May 1970.

The Dow Chemical Company. Basic Demineralization. Idea ± Exchange 2 (1),
January 1972.

The Dow Chemical Company. Cation Resin ± Hydrogen Cycle. Idea ± Exchange 2
(2), April 1972.

The Dow Chemical Company. Weak Acid Cation Resins. Idea ± Exchange 2 (3),
July 1972.

The Dow Chemical Company. Anion Resin - Hydrogen Cycle. Idea ± Exchange 2
(4), October 1972.
175
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CHAPTER 11

CHEMICAL SOFTENING
11.1 Introduction
Chemical softening is employed as a treatment process to remove dissolved
ions from AMD only when considering the effluent for industrial use or pos-
sibly as potable water. Softening processes have been used extensively to
remove hardness (calcium, magnesium, iron, manganese, and aluminum) from
municipal water supplies, and can be adapted to treat acid mine drainage.

Two processes that merit consideration for possible potable water production
are lime-soda and alumina-lime-soda. The former has been the most common
process used by water plants" to treat hard water. The latter was developed
during 1971 as a brackish water desalination process (1). The advantages
chemical softening has over other water recovery methods—e.g., ion exchange
or reverse osmosis—are that only conventional treatment equipment is re-
quired and no waste streams other than sludge are produced.

11.2 Lime-Soda Softening Process

The lime-soda softening process has been widely used to remove hardness,
iron, and manganese from municipal water supplies (2, 3, 4, 5). This process
takes advantage of the low solubilities of calcium and magnesium compounds
and removes these precipitated cations by sedimentation. , Calcium is precipi-
tated as calcium carbonate by increasing the carbonate concentration in the
water. Magnesium is precipitated as magnesium hydroxide by increasing the
hydroxide concentration. Lime and soda ash are the chemicals most often used
to bring about these chemical reactions. Precipitation is responsible for
iron and manganese removal to levels within drinking water standards or
industrial quality.

The total dissolved solids in the water, however, are not greatly affected by
this process. Calcium and magnesium ions are replaced by sodium ions while
the sulfate concentration remains constant. The choice of this process
depends on the divalent cation concentrations in the raw AMD and the concen-
trations desired in the effluent water.

For the application of lime-soda softening to acid mine drainage, the first
four unit processes are the same as for conventional lime neutralization (6).
These are equalization, neutralization, iron oxidation, and solids removal,
as illustrated in Figure 11-1. Then, the effluent from the solids removal
unit enters a flash mix tank for chemical addition, the first softening

176
-------
RAW WATER

O
0

\
3
U
w

FILTRATE
OT
BACKWASH LAGOON
Figure 11-1. Unit processes of lime-soda softening.

177
-------
process. This step is followed by the softening reaction (flocculation)
tank, settling basins, a recarbonation chamber, filters, and chlori nation.
Provisions also must be made for sludge recirculation and sludge handling, as
well as for filter backwash equipment.

The functions of the neutralization and iron oxidation stages are the same as
a typical AMD plant: pH adjustment and iron and manganese removal. Lime is
added at this stage for neutralization of the mine drainage acidity and the
precipitation of iron, manganese, and aluminum as hydroxides.

Lime is again added to the sedimentation basin effluent as the first step in
the softening process. Lime is required for further manganese and magnesium
removal, both of which are precipitated as their respective hydroxides.
These reactions are expressed by Equations 34 and 35 (6):
MgS04 + Ca(OH)2 -»• Mg(OH)2 + CaS04 (34)

+ Ca(OH}2 •*• Mn(OH)2 + CaSO^ (35)
Free carbon dioxide and carbonate hardness also exert a lime demand in this
stage. Soda ash (Na2COs) is then added to remove noncarbonate calcium hard-
ness or calcium sulfate (CaSCK), as illustrated in Equations 34, 35, and 36.
The calcium ion is precipitated as calcium carbonate while the sulfate
remains in solution.

CaSO^. + Na2C03 •»• CaC03 + Na^Oi* (36)

The pH must be maintained at 9.5 or higher for this precipitation to occur
(6).

A lime-soda softening plant in Altoona, Pa., was built by the Pennsylvania
Department of Environmental Resources for the purpose of treating streams
affected by mine drainage to augment the supply of drinking water for that
city (see Table 11-1). It is important to note that the sulfate and specific
conductance concentrations in the raw water are low, thus qualifying it as a
potential source of potable water.

A schematic of the treatment plant process was shown previously in Figure
11-1. The finished water quality of this plant generally met the EPA Drink-
ing Water Standards (6).

During a 5-month study, several important observations concerning this pro-
cess were noted (6). The most important was that the softening system worked
well only when the reaction zone solids were in the range of 10%-15% settle-
able solids by volume. A failure to build up the solids in the reactor-
clarifier unit resulted in an increased effluent hardness. Also, a minimum
pH of 11.0 and minimum temperature of 12°C (54°F) were required for favorable

178
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TABLE 11-1

TYPICAL BLENDED RAW WATER CHARACTERISTICS,
LIME-SODA SOFTENING PLANT, ALTOON, PA.
KITTANNING RUN:IMPOUNDING DAM = 1:1.33
Parameter Value

pH 3.0

Acidity, mg/1 as CaC03 170

Calcium, mg/1 28

Magnesium, mg/1 18

Iron, mg/1 17

Manganese, mg/1 4.5

Aluminum, mg/1 13

Sodium, mg/1 1.8

Hardness, mg/1 as CaC03 260

Sulfates, mg/1 270

Specific conductance, ymhos/cm 820

softening reactions.

The softener unit (solids contact clarifier) required sludge recirculation to
the reaction zone where raw water and chemicals were added. A buildup of
reaction zone solids (10%-15% settlcable solids by volume in 15 minutes) was
a prerequisite for proper operation according to the manufacturer's recommen-
dation. Only during 3 weeks of the study was it possible to obtain this type
of operation. The reason for failure of the unit to build up solids was not
understood and resulted in disenchantment among the experimenters. The
minimum attainable hardness during the study was 120 mg/1 CaC03, regardless
of the soda ash dosage. The authors of the study recommended against the use
of the particular solids contact clarifier installed in the Altoona plant.

The most effect removal of both manganese and magnesium in the neutralization
stage occurred at a pH of 11.0. Provisions for the addition of potassium
permanganate (KMn04) prior to neutralization were made to insure complete
manganese oxidation when the pH was less than 11.0. The filters removed
significant amounts of iron and manganese, enabling the effluent to meet the

179
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drinking water standards of 0.3 mg/1 iron and 0.05 mg/1 manganese.

The reported costs in 1975 for this treatment process were $0.030/m3 ($0.114/
1,000 gal) for the neutralization in the first stage of the system. The raw
water averaged 170 mg/1 of acidity and 17 mg/1 of total iron. For the over-
all system operation, costs ranged from $0.098/m3($0.371/1,000 gal) for water
with 120 mg/1 of hardness to $0.088/m3 ($0.333/1,000 gal) for 200 mg/1 hard-
ness. These operating costs include power, chemicals, personnel, maintenance
supplies, and depreciation.

11.3 Alumina-Lime-Soda Process

The alumina-lime-soda process was originally developed as a method of desal-
inating brackish water to produce high-quality potable water. It is specif-
ically suited for waters in which the principal source of salinity is sul-
fate. It is capable of removing heavy metals and hardness as well. Economic
considerations indicate the process is most useful for treating AMD with
sulfate concentrations between 400 and 1,200 mg/1. The practical lower limit
for this process is 100 mg/1, where sulfate removal is not economically
feasible. The drinking water standards limit sulfate concentration to 250
mg/1.

The alumina-lime-soda process is divided into two stages, as illustrated in
Figure 11-2. The raw AMD is split into two streams, the larger of which is
treated with lime and sodium aluminate (NaA102) (Stage I). The ratio of AMD
treated in Stage I to that bypassed to Stage II depends on the sulfate con-
centration desired in the total plant effluent. The effluent from Stage I
will contain approximately 100 mg/1 sulfate, while the sulfate concentration
bypassed to Stage II remains constant. The sulfate concentration in the
final blended flow can be calculated by a simple mass balance. Stage I
effluent is mixed with the smaller AMD stream while carbon dioxide is added
for pH control (Stage II). Both stages produce solids, which are removed by
filtration.

The key process reactions occur in Stage I. The sodium aluminate and lime
neutralize the raw acidity, precipitate the heavy metals and magnesium, and
remove calcium sulfate. Pilot studies conducted at the Commonwealth of
Pennsylvania, Acid Mine Drainage Research Facility at Hollywood, Pa., indi-
cate that maximum sulfate removal occurs when the Stage I pH is held at 12.0
and the alkalinity level is 600 mg/1 as CaC03. Values less than these result
in ineffective sulfate removal. Values greater than these will not effec-
tively remove more sulfate, resulting in wasted lime and increased operating
costs.

Sulfate is removed by the sodium aluminate, and to a lesser extent, by the
iron and aluminum present in the raw acid mine drainage. The reaction be-
tween the sodium aluminate, lime, and AMD produces insoluble calcium sulfoa-
luminates (SCaSO^ • A1203 • 3CaO • x H20 and CaS04 • A1203 • 3CaO • x H20).
The latter form seems to dominate the resulting sludge. Each mole of sodium
aluminate removes 1 mol of sulfate. The iron present in the AMD forms insol-
uble calcium sulfoferrites. The lime stabilizes the precipitates as well as

180
-------
RAW AMD
NaAlO,
Ca(OH)2
STAGE I
ALUMINA - LIME -

SODA TREATMENT
SETTLING

TANK
FILTRATION
NIX
SOLIDS DISPOSAL
\

\
/
N
COL
\
7

\
7
STAGE II
RATIO MIXING
OF RAW AMD/
STAGE I EFFLUENT
AND CARBONATION
NX
FILTRATION
\

7
PH

ADJUSTMENT
PRODUCT WATER
Figure 11-2. Stages of the alumina-lime-soda process.

181
-------
adding low-cost alkalinity to maintain the reaction pH of 12.0.

The Hollywood experiments found that a minimum Stage I retention time of
80-100 minutes was necessary to provide maximum sulfate removal. The com-
bined factors of high retention time, high pH, and a reaction vessel open to
the air resulted in complete oxidation of the ferrous iron and manganese,
which eliminated the need for a separate oxidation unit.

Following the Stage I reaction, precipitating solids are removed in a mechan-
ical settling unit. The resulting sludge is usually less than 2% solids by
weight, and should be further concentrated by pressure or vacuum filtration
to recover the softened filtrate. Dewatered sludge with 10%-12% solids by
weight was demonstrated.

The Stage I water contains excess lime, which is removed in Stage II by mix-
ing with the smaller stream of raw acid mine drainage. This excess lime will
neutralize the acid in the raw mine drainage stream. Carbon dioxide is also
added to lower the pH to 10.3, the minimum solubility of calcium carbonate.
Calcium carbonate is formed by the reaction

20H- + Ca+2 + C02 ->CaC03 + H20 (37)

Excess carbon dioxide will redissolve calcium to form calcium bicarbonate.
Too little carbon dioxide will leave free hydroxide alkalinity in the water.

The calcium carbonate and metal hydroxide precipitates are removed by sand
filtration. The resulting filtrate will have a pH of 10.3 and will contain
about 35 mg/1 of dissolved calcium carbonate, its minimum solubility. Addi-
tional carbonation will drop the pH to a value acceptable for potable water.

Costs for the alumina-lime-soda process depend upon the cost of sodium alumi-
nate. This chemical is available in two forms: the commercially available
"dry" form, and the "calcined" form produced by heating a mixture of soda ash
and bauxite. The latter form is less expensive, but requires the installa-
tion of an aluminate slaker. The quantity of sodium aluminate needed depends
on the sulfate, iron, and aluminum concentrations in the acid mine drainage.
Increased concentrations of the latter two tend to lower aluminate require-
ments because of their ability to precipitate sulfate.

Construction and operating costs were estimated for AMD treatment facilities
of three sizes in 1975. The costs are summarized in Table 11-2.

The chemical costs include sodium aluminate made by a bauxite-soda ash
slaking plant. These cost estimates should be doubled if a "dry" aluminate
plant is to be built. Depending on the quantity delivered, the March 1978
cost of dry sodium aluminate is $545-$875/kkg ($494-$794/ton). The sodium
aluminate cost at the time of this study was $303/kkg ($275/dry ton), so the
total operating costs have increased by approximately 50%.
182
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TABLE 11-2

ESTIMATED COSTS OF THE ALUMINA-LIME-SODA PROCESS (1)

1,893 m3/d 3,785 m3/d 18,925 m3/d
(0.5 Mgal/d) (1.0 Mgal/d) (5.0 Mgal/d

Total construction cost $352,000 $516,000 $1,382,000
(excluding engineering
and legal costs)

Operation and maintenance
costs (excluding depreciation)
($/m3) 0.27 0.24 0.21
($1,000 gal) 1.04 0.92 0.79
11.4 References

1. Nebgen, J.W., D.F. Weatherman, M. Valentine, and E.P. Shea. Treatment
of Acid Mine Drainage by the Alumina-Lime-Soda Process. EPA-600/2-76-
206, Technology Series Report, Cincinnati, Ohio, September 1976.

2. The American Water Works Association, Inc. Water Quality and Treatment:
A Handbook of Public Water Supplies. 3rd ed. McGraw-Hill, New York,
1971.

3. Riehl, M.L. Water Supply and Treatment, llth ed. National Lime Asso-
ciation, Bulletin 211, Washington, D.C., 1976.

4. Sawyer, C.N., and P.L. McCarty. Chemistry for Sanitary Engineers. 2nd
ed. McGraw-Hill, New York, 1967.

5. Clark, J.W., W. Viessman, Jr., and M.J. Hammer. Water Supply and Pollu-
tion Control. 2nd ed. International Textbook Company, Scranton, Penn-
sylvania, 1971.

6. Long, D.A., J.L. Butler, and M.J. Lenkevich. Soda Ash Treatment of
Neutralized Mine Drainage. EPA-600/2-77-090, Cincinnati, Ohio, May
1977.
183
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CHAPTER 12

CALCULATIONS AND PROCEDURES FOR DESIGN OF A MINE DRAINAGE
TREATMENT PLANT
12.1 Introduction

This chapter will outline a procedure for the design of a treatment process
and related equipment for a mine drainage treatment plant. The purpose of
this outline is to give the designer insight into the evaluation of possible
alternatives and justification for choosing a certain method or process.

12.2 Design Example I

This particular mine drainage plant will have a daily average flow of 2,880
m3/d (0.76 Mgal/d), and a raw water quality as shown in Table 12-1.

Several assumptions concerning the operation of this treatment plant have
been made to aid in its design. The plant is designed for continuous opera-
tion with as little operator supervision as possible. Therefore, the design
will be automated as much as possible, and have gravity flow between pro-
cesses. The processes requiring design are flow equalization, neutraliza-
tion, aeration, solids separation, and sludge disposal.

12.2.1 Equalization Basin Design (see Chapter 2)

For this design situation, it is assumed that underground storage is avail-
able, but it might not be adequate during high flow periods in the spring.
Therefore, an equalization basin of 2 days storage volume will be sized. The
basin will operate at 25% capacity.

12.2.1.1 Sizing

The equalization volume is

Volume = 2?8{*° m x 2 d = 5,760 m3 (203,400 ft3)
d
184
-------
TABLE 12-1
RAW WATER QUALITY AND EFFLUENT LIMITATIONS
Raw Water
Effluent Limitations
30 Consecutive Maximum
Day Average Daily
Parameter

PH
Net alkalinity, mg/1 as CaC03
Sulfate, mg/1
Suspended solids, mg/1
Iron (total), mg/1
Iron (ferrous), mg/1
Manganese, mg/1
Nickel, mg/1
Zinc, mg/1
Aluminum, mg/1
Assuming a 3-m (10-ft) water depth with a 1-m (3-ft) freeboard and choosing
to make a square equalization basin, the approximate water surface area is

Approximate surface area = 5>7^°rnm = 1,920 m2 (20,670 ft2)
O HI
3.1
-600
1,500
55
100
95
3
1
1
30
6-9
>acidity
—
35
3.5
—
2.0
—
—
——
6-9
>acidity
—
70
7.0
—
4.0
—
—
—-.
The equalization basin will be a square earthen basin with 2%:1 inside slopes
and 3:1 outside slopes. Top berm width will be 4.6 m (15 ft) and the inside
berm-to-berm length 36.6 m (120 ft). These dimensions provide the required
volume, as shown by the average end area calculation.
2 2
Water Volume = (AI* A^) H = (36.6) + (51.6) (3) = M04 m3 (211,941 ft3)
where Ax and A2 are the bottom and top water surface areas, respectively, and
H is the water depth
185
-------
The following calculations show how the cut and fill volumes are balanced to
determine the height of the dikes above ground level (h). This assumes the
proposed site is relatively flat (see Figure 12-1).

Cut Volume = Fill Volume

(36.6)2 + (56.4 - 5h)2 (4 . h) = 4.6 + (4.6 + 5.5h)(h)(4(36>6) +

2 2

4(14.5 + 3h))

(4,521 - 564h + 25h2)(4 - h) (%) = (%) (h) (9.1 + 5.5h) (204 + 12h)

17,903 - 2,233h + 99h2 - 4,521h + 564h2 - 25h3 = l,865h + l,232h2 + 66h3

8,619h + 569h2 + 91h3 = 17,903

This equation can be solved for the height of the dikes above ground level
(h) by trial and error. Values of h are substituted into the left side of
the equation until the solution is reasonably close to the desired value.

h f (h)
1.52 m (5 ft) 14,779

1.83 m (6 ft) 18,236

Cut Volume = 3,825 m3 (5,000 yd3)

Fill Volume = 3,980 m3 (5,200 yd3)

If the proposed site is level, 2.1 m (7 ft) of excavation will provide the
fill necessary to build the dikes 1.8 m (6 ft) above ground level, giving a
total height of 4 m (13 ft).

The total pond area is

Pond area = (76.5 m)2 = 5,850 m2 (63,000 ft2)
= 0.6 ha (1.5 ac)

12.2.2 Theoretical Lime Requirement (see Chapter 3)

Assume a 70* hydrated lime efficiency. The net alkalinity, determined by
analysis, is -600 mg/1 as CaCO . The theoretical lime requirement from Table
1-2 is as follows:
186
-------
65.5m (215ft)
56.4m (185ft)
'•• —
\ 3.96m-h
\ (13ft-h!

L 36.6m (120ft) J
p* •*!
^ 56.4m-5h (185ft-5h) r
65.5rTM-6h (215ft*6h)
4.57m-*-5.5h
^ (15ft-*-5.5h)
»^_
5.5m
(18ft)
\
/

\
/

^4.6m
( 15ft)
76.5m
/
\

/
\
^ 20.0m „
(65.5ft)
(251ft)
Figure 12-1. Equalization basin, Design Example I.
187
-------
600 mg/1 CaC03 x jj^fi^x _1_ = 635 mg/1 Ca(OH)
The theoretical daily hydrated lime requirement is

3
635 mg Ca(OH)2 x 2>88° m x l'OQ° x -& = i,830 kg/d (4,025 Ib/d)
An operating pH of 8.5 is desired. The theoretical lime requirement, based
on the acidity analysis, will raise the pH to 8.3. It is estimated that a
small increase in lime must be added to maintain the desired operating pH.
Any significantly higher pH would require that a titration curve be estab-
lished and the lime requirements established on that basis.

12.2.3 Lime Requirement from Treatability Test

A treatability test was conducted according to the procedure outlined in
Chapter 6, Section 6.3. This is the better method of determining actual lime
requirements. The test indicated that 627 mg/1 of lime would be needed to
achieve the desired neutralization pH of 8.5.

Therefore, the actual daily lime requirement equals

L . 1>800 kg/d (4,000 Ib/d)
Since the daily lime usage of 1.8 kkg/d (2 tons/d) is low, the designer's
best choice would be to employ a hydrated lime system. A quicklime operation
with slaker requires a capital investment and maintenance cost that would be
difficult to justify.

12.2.4 Lime Silo Sizing (see Chapter 3)

Considering that most mine drainage plants are in rural areas, the designer
should provide at least seven days lime storage capacity in a silo. In this
case, the minimum silo capacity should be 12.7 kkg (14 tons). To take advan-
tage of pneumatic bulk delivery (minimum 18.1 kkg (20 tons) per delivery), a
silo larger than 18.1 kkg (20 tons) is necessary. The cost per unit volume
of silos is perhaps the least expensive item in the entire plant. Thus, the
designer can provide excess capacity at relatively low costs.

A 27.2-kkg (30-ton) silo is chosen. This will permit bulk delivery, allow
a 9.1-kkg (10-ton) floating freeboard for operation, and give the operator a
5-day leeway for delivery.

188
-------
Assuming a hydrated lime density of 481 kg/m3 (30 lb/ft3), the silo volume
required to store 27.2 kkg (30 tons) is
Silo volume = ? = 56'5 m3 (2>°°° ft3)
The silo specifications are
diameter = 3.05 m (10 ft)
sidewall height = 7.3 m (24 ft)
60° hopper bottom

12.2.4.1 Bin Activator (Vibratory)
A vibratory bin activator is highly recommended for hydrated lime silos. Bin
activators are sized as one-half the silo diameter for silos up to 6.1 m (20
ft) in diameter, and one-third the diameter for larger silos. Therefore, a
1.52-m (5-ft) vibratory bin activator should be used.
12.2.4.2 Other Silo Equipment Required
1. Dust collector, whose size varies with model;
2. 10.16-cm (4-in) fill line with long radius turn and quick-connect
coupling;
3. bin level indicators (side ports);
4. OSHA side-mounted ladder and a hand railing atop the silo;
5. foundation.
12.2.4.3 Silo Slide Valve
It is necessary to have a silo slide valve on the bin activator bottom. This
enables the operator to close the silo at any time, and is extremely impor-
tant when the feeder malfunctions or when an empty silo is being filled.
When filling an empty silo, this valve must be closed. Otherwise, lime will
be blown throughout the feeder area.

12.2.5 Lime Feeder (see Chapter 3)
The average hydrated lime feed rate is as follows:
189
-------
75 kg/hr (167 ib/hr)

The volumetric feed rate is

3" 0-16 m3/hr (5.52 ft3/hr)
The designer has a number of feeder types from which to choose; the choice
can be a matter of personal preference. A variable screw feeder with a
delivery range of 0.065-0.65 m3/hr (2.3-23 ft3/hr) is chosen.

12.2.6 Optional Designs for Lime Feeding

Since the lime requirement for this particular plant is low, the designer has
methods available other than a conventional lime slurry feed system. A dry
feed system or a volume slurrying system are viable alternatives.

12.2.6.1 Dry Feed System

There are various opinions on the feeding of dry lime directly into the raw
water. There exists a cutoff point, yet to be determined, where dry lime
feed has prominent disadvantages. This stems from the low solubility of
lime. In this case, however, where the lime requirement is only 0.625
kg/1,000 1 (5.26 lb/1,000 gal), a dry lime system is entirely feasible.

The designer, as a modification to the previous design, can eliminate the
slurry feed system, but should provide a longer reaction time in the flash
mixer.

12.2.6.2 Volume Slurrying

Volume slurrying involves blowing the hydrated lime into a large closed-top
tank equipped with a mixer and large dust collector. The tank contains a
specific amount of water that depends upon the slurry concentration desired.
The lime slurries as it enters the tank and is continuously agitated. The
slurry tank must have the capacity to handle 18.1 kkg (20 tons) of dry lime
put in solution. If a 15% slurry is desired, a minimum tank capacity of 122
m3 (32,260 gal) is required to accept a full pneumatic truck load. To main-
tain continuous treatment, at least 4.54 kkg (5 tons) or 30.3 m3 (8,000 gal)
capacity would have to be added, increasing the minimum slurry tank volume to
151.4 m3 (40,000 gal).
190
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12.2.6.3 Lime Slurry System

A 10% (by weight) lime slurry concentration is initially chosen as the feed
(see Table 3-1). The slurry weight is 1.1 kg/1 (8.8 Ib/gal). The lime
(Ca(OH)2) concentration is 0.10 kg/1 (0.83 Ib/gal). The required slurry
makeup rate is

' 12'5 1/min (3-3 9al/min)
However, the slurry system makeup rate should be at least two times that
required. Therefore, a 25 1/min (6.6 gal/min) makeup rate should be used,
with a water-to-lime weight ratio of 9.7:1. The lime feed should be

25 1/min x °'^ k9ca(OH)2 x 60J1n = 150 kg Ca(OH)2/hr (330 Ib/hr)

= 0.3 m3/hr (n ft3/hr)

=2.5 kg/min (5.5 Ib/min)

The water feed should be

2.5 kg/min x 9.7 = 24.2 kg H20/min (53.4 Ib/min)

= 24.2 1 H20/min (6.4 gal/min)

12.2.6.4 Slurry Tank

The slurry tank will be designed to provide 20 minutes detention time when
half full. The tank will provide slurry preparation, storage, and stabiliza-
tion.

The required tank volume is

i 3
Volume = 40 min x 12.5 1/min x 1 ^ ] = 0.5 m3 (17.7 ft3)
Design a circular tank with the diameter equal to the height.

Volume = 0.785 (D2) (H) Set D = H

191
-------
0.5 = 0.785 (D3)

D = 0.86 m (2.8 ft)

Make the tank diameter 0.9 m (3 ft) and the overall height 1.22 m (4 ft),
which includes 0.3 m (1 ft) freeboard. The total working volume is 592 1
(160 gal). The lime feeder and water will be turned on when the low level is
reached, and turned off at high level.

The control level positions, with respect to tank bottom, will be as follows:

Low level (turns feeder and water on) at 0.46 m (1.5 ft)
High level (turns feeder and water off) at 0.91 m (3 ft)
Emergency high level (shutdown of system) at 1.07 m (3.5 ft)

12.2.6.5 Slurry Mixer

A standard rule of thumb to size the mixer is 0.2 kW/m3 (1 hp/1,000 gal).
This will give the designer a good approximation of the mixer size until the
final sizing can be determined with the help of the vendor.

Since the slurry tank volume is only 0.592 m3 (160 gal), the required horse-
power equals

0.592 m3 x (0.2 kW/m3) = 0.11 kW (0.15 hp)

Mixer motors this size operate at 30%-50% efficiency. Therefore, a 0.37 kW
(0.5 hp), angular, side-mounted propeller mixer should be provided. The
propeller shaft should extend below the low-level control to insure that the
contents will be mixed at all times.

12.2.6.6 Slurry Feed System

One variable-speed, volumetric slurry feeder should be provided, with an
operating range of 0.314-113.6 1/min (0.083-30 gal/min) and a 0.19-kW (0.25-
hp) feeder motor. This variable speed feeder is controlled by a pH monitor
placed immediately after the flash mixing tank. The pH probes should not be
placed inside the flash mix tank, but as near to the inlet of the next Dera-
tion or settling) unit as possible.

12.2.6.7 Piping and Feed Pump

Ideally, the slurry line velocity within the slurry feed loop should be in
the range of 1.0-1.2 m/s (3-4 ft/s). The pipe diameter required to maintain
this velocity at a flow rate of 12.5 1/min (3.3 gal/min) is calculated as
follows:

192
-------
Pipe cross-sectional area = -^__ = ^.5 1/min jin. 1,000 cm3x m^
Velocity 1.2 m/s 60 s 1 100 cm

= 1.71 cm2 (0.266 in2)

This cross-sectional area would indicate a commercial pipe of 1.27-cm (0.5-
in) diameter; however, it is also recommended that any slurry line not be
less than 2.5 cm (1 in) in diameter. Assuming a 2.5-cm (1-in) minimum pipe
diameter for this slurry line, and use of head loss tables for standard water
pipe, a flow of 37.9 1/min (10 gal/min) will produce a velocity of about 1.1
m/s (3.7 ft/s), with a head loss of 12 m/100 m (12 ft/100 ft). These are
acceptable design values.

In some cases, revisions to the sizing of the slurry tank, its mixer, or
other related equipment may be necessary. In this case, the slurry is to be
recirculated from the slurry makeup tank to the volumetric feeder. Excess
flow is returned to the slurry tank and a pH system controls the feed from
the volumetric feeder to the flash mix tank.

The head loss in a 2.54-cm (1-in) diameter pipe is approximately 12 m/100 m
(12 ft/100 ft). If it is assumed that the length of the slurry loop is 7.62
m (25 ft) and the static head is 3.0 m (10 ft), the total pump head is

Total head - 3.0 + 7.62 x - = 3.91 m (12.8 ft)
The slurry pump power needed is

kW =

where r = slurry density (kg/m3)

Q = slurry flow (m3/s)

H = head (m)

kW
101.97
1.060 kg/m3 x 0.0024 m3/s x 3.91 m
101.97

kW = 0.098 kW (0.13 hp)

According to vendor catalogs, pumps in this range operate at 35% efficiency,
which is very low. Thus, the required pump size would be as follows:
193
-------
0.28 kW (0.037 hp)
A slurry pump with a 0.37-kW (0.50-hp) motor with a 2.54-cm (1-in) discharge
and 3.18-cm (1%-in) suction should be selected.
12.2.7 Flash Mix Tank (see Chapter 4)
Assume an effective detention time of 5 minutes. The volume required must
include both the AMD flow and the slurry flow.
Vol = 5 min x ((2,880 m3/d x M40dm1n x >) + 12.5 1/min)

= 10,063 1 (2,660 gal)
Design a circular tank with H:D approximately equal to 1.
Volume = 0.785 (D2) (H)
10.06 m3 = 0.0785 D3
D = 2.34 m (7.68 ft)
The final flash mix tank dimensions are
Diameter = 2.44 m (8 ft)
Height = 2.44 m (8 ft) •*• 0.61 m (2 ft) freeboard = 3.05 m (10 ft)
The inlet and outlet positions are 180° apart.
Inlet - top entry of raw water and slurry pipes with air break.
Outlet - pipe insuring 1.0-1.2 m/s (3-4 ft/s) exit velocity.
The outlet pipe diameter calculations are

Flow
Area =
Velocity
194
-------
(2'880
A = 0.034 m2
D = 0.2 m

Use a 20-cm (8-in) pipe. Use a 20-cm (8-in) inverted elbow with a 0.61-m
(2-ft) nipple entering at the 2.44-m (8-ft) height.
The tank will require baffling. Place four baffles 90° apart. Design the
baffles according to the following specifications:
Baffle width 1/18 x 244 cm = 13.6 cm (5.3 in)
Make the baffle width 15.24 cm (6 in).
Wall clearance = 0.10 x 15.24 cm = 1.42 cm (0.6 in)
Make the minimum clearance 2.54 cm (1 in).
The baffles should end a minimum of 6.35 cm (2.5 in) above the tank bottom.
They should extend at least 15.24 cm (6 in) above the static water level.

12.2.8 Aeration Tank (see Chapter 5)
The average ferrous iron concentration is 95 mg/1. The daily ferrous iron
loading is
1.8801 „ __ , 274 kg Fe+2/d (603 ,„

The theoretical oxygen requirement for iron oxidation is determined by the
chemical relationship that
7 kg (15.4 Ib) of iron are oxidized by 1 kg (2.2 Ib) of oxygen
3.2 kg (7 Ib) of iron are oxidized by 0.454 kg (1 Ib) of oxygen

= 39 kg °2/d (86 lb °2/d) = 1>63 kg/hr (3'59 lb/hr)
195
-------
The approximate kW (hp) requirements, using 2.13 kg (h/kW-hr (3.5 Ib 02/hp-
hr) oxygen transfer rate, can be determined. A 0.75-kW (1-hp) aerator is
initially sized, although mixing requirements must still be considered and
are of the utmost importance.

Detention time required in an aeration basin can best be estimated from the
iron oxidation-vs.-time curve produced during the treatability study. The
slope of this curve given in milligrams per liter per minute can be used to
determine the minimum detention time necessary for complete ferrous iron
oxidation at the operating pH chosen.

From the treatability test, an 11.25 mg/l/min iron oxidation rate was
derived. Assuming complete oxidation in the basin, the detention time
required is
Detention tta. = - • 8.44 min
A high degree of short-circuiting can occur in an aeration basin. Therefore,
a scale-up or safety factor must be applied. Since this calculated detention
period is short, a factor of at least 4.0 (Table 5-2) must be applied. Thus,
a minimum 34.0-minute detention time should be provided.

Aeration tank size:

Volume = Q x t

= ((2,880 m3/d x 1>440dm1n ) +(12.5 1/min x 1>QQ0 1 )) x 34 min

= 68.4 m3 (2,400 ft3)

where Q = total flow (raw water and slurry)

t = 34 min

Assume a 1.83-m (6-ft) water depth and circular basin.

Surface Area . «$£

= 37.4 m2 (400 ft2)

Basin Diameter - 6.9 m (20.7 ft)

The designer now selects an aerator that satisfies both oxygen transfer and
mixing requirements.

196
-------
The type of aerator chosen here is a floating, 0.89-kW (1.2-hp) aerator with
a working volume of 8.23 m (27 ft) in diameter and 1.83 m (6 ft) in depth,
and an oxygen transfer rate of 2.3 kg 02/kW-hr (3.8 Ib 02/hp-hr).
It is the designer's option either to build the aeration basin from concrete
and steel or to use an earthen basin with an erosionproof bottom.
12.2.9 Estimated Sludge Production (see Chapter 7)
The sludge produced by the treatment of this mine water can be estimated as
the sum of the metal hydroxides removed in significant concentrations (iron
and aluminum), the suspended solids in the mine drainage, and unused lime.
This estimation assumes 100% solids removal.
Hydrolysis of ferric iron, assuming complete oxidation of ferrous iron:
Fe+3 + 3H20 Fe(OH)3 + 3H +
56g/g-mol 107 g/g-mol
100 mg/1 - (10Q mg/1) = 191 mg/1
Aluminum hydrolysis:
Al+3 + 3H20 A1(OH)3 + 3H +
27 g/g-mol 78 g/g-mol
30 mg/1 (30 mg/1) - 87 mg/1
Theoretical daily solids production:
Ferric hydroxide
W1 mg/1 x l^JSi x i«g°l x -j^ - 550 kg/d (1,212 lb/d)

Aluminum hydroxide
87 mg/1 x l^SOjnl x i^J_ x J kg = 251 kg/d (552 lb/d)

Suspended solids
O QQn m 10001 1Un
55 mg/i x ^88Ujn_ x l^LL x ^6k9g = 158 kg/d (349 lb/d)
197
-------
Unused lime (assume 15% lime wastage, which is not unusually high and
is due primarily to the insolubility of lime)

1,806 kg/d x 0.15 = 271 kg/d (596 Ib/d)

Assuming the sludge withdrawn from the settling basin will be at only 1%
solids, the liquid sludge weight is
1>23° k = 123,000 kg sludge/d (270,600 Ib sludge/d)
For simplicity, assume the sludge weighs the same as water.

= 123'000 ] sludge/d (32,497 gal sludge/d)

= 123 m3 sludge/d (4,343 ft3 sludge/d)

The designer can now cross-check this theoretical sludge volume estimation
against results found in the treatability study. In this case the settling
test gave a final settled volume of 35 ml. The sludge volume then can be
calculated as a simple percent.

= o.035 or 3.5% of the flow

= 0.035 x 2,880 m3/d = 101 fli3/d (3,560 ft3/d)

Therefore, the- designer should consider the higher sludge volume and design
accordingly.

12.2.10 Settling Basin Design

In this case, land is available, so an earthen settling basin is chosen. The
design parameters are as follows:

1. detention time equals a minimum of 12 hours clear water storage and
a minimum of 3 days sludge storage capacity;

2. equipped with automatic sludge removal device;

3. influent flows equal 2,880 m3/d (0.76 Mgal/d);

4. 0.91 m (3 ft) freeboard;
198
-------
5. exterior slopes 3:1;

6. interior slopes 2^:1.

Calculations of basin volume follow:

Three days sludge storage = 123 m3/d x 3 d = 369 m3 (13,030 ft3)

12 hours clear water storage = 2,880 m3/d x ?I hr/d

= 1,440 m3 (50,900 ft )

Total Volume Lime = 1,809 m3 (63,785 ft3)

Assume a 3.05-m (10-ft) depth using a length-to-width ratio of 3:1 for
bottom.

12.2.11 Settling Basin with Hydraulic Sludge Removal System

The designer has decided to use an hydraulic sludge removal system to mini-
mize labor costs associated with manual sludge removal. The manufacturer of
these systems has recommended an area of 9.1 m x 18.3 m (30 ft x 60 ft) to be
covered by the system. Thus, the pond bottom should have a width of 9.1 m
(30 ft) and length of 24.4 m (80 ft) to obtain the necessary volume (see
Figure 12-2). The length-to-width ratio is 2.67 and the volume provided is

Volume - (9.lx 24.4) ^(24.4x29.6) (38n m3 (63>952 ft3)

which is about equal to the required volume of 1,809 m3 (63,875 ft3).

Calculations to determine cut and fill volumes, assuming the pond is to be
built on level ground, are

Cut Volume = Fill Volume

(24.4) (9.1) + (44.2 - 5h) (29.0 - 5h) /3 96 _ h) =

(A C7 + (A C7 + C Cfr\ ...... . .
V i • *"» / * V ~ • *J / ' l C i O U. \ i fO\ /*"! A A i 1 H I" i *M_\\
-a — i . -. - .„. . i— — _* (n) I I ^ ) V y 1 ~r 14 j T on ) ~F { X J \ r_ 44 i 14 D T OM ) )

(222 + 1,282 - 145h - 221h - 25h2) (%) (3.96 - h) =
(h) (9.14 + 5.5h) (125 + 12h)
(1,504 - 366h + 25h2) (3.96 - h) = (h) (9.14 + 5.5h) (125 + 12h)
199
-------
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(5,956 - l,449h + 99h2 - l,504h + 366h2 - 25h3 =
l,143h + 688hz + 110hz + 66h3)
4,096h + 333h2 + 91h3 = 5,956
This equation can be solved by trial and error.

h f(h)
1.22 m (4 ft) 5,658
1.37 m (4.5 ft) 6,470
1.30 m (4.25 ft) 6,087
Choose h - 1.30 m (4.25 ft)
Cut Volume = 1,430 m3 (1,870 yd3)
Fill Volume = 1,480 m3 (1,940 yd3)
Overall Dimensions =9.1+2 (14.5 + (3) (1.30)) = 45.9 m (150.5 ft)
= 24.4 + 2 (14.5 + (3) (1.30)) = 61.2 m (200.5 ft)
Total Basin Area = (45.9 m) (61.2 m) = 2,809 m2 (3,240 ft2)
= 0.28 ha (0.69 ac)
Total Water Area = (24.4) (39.6) = 966 m2 (10,400 ft2)
= 0.10 ha (0.24 ac)

12.2.12 Sludge Disposal Pond
Based on past experience, the following assumptions have been made for oper-
ating a lagoon disposal: a 20-year mine life, and sludge withdrawn at 1%
solids for the primary settling basin.
Air-dried sludge can achieve 16%-18% solids in a disposal lagoon, but 12%
will be used on a conservative basis. The decanted water is returned to the
primary settling pond.
The sludge produced (from treatability test) is as follows:
Sludge produced per day = 123 m3 (4,343 ft3) at 1% solids
Final sludge volume at 12% = 10.3 m3 (362 ft3)
201
-------
12.2.12.1 Required Sludge Disposal Pond Volume

- x 20 yr = 75,190 m3 (2,655,000 ft3)
" j •

Provide a 1.52-m (5-ft) clear water depth plus a 0.91-m (3-ft) freeboard at
capacity. Assume a pond depth of 10.67 m (35 ft), thus providing an 8.23-m
(27-ft) sludge depth.

Area required = 75^^ ™3 = 9,136 m2 (98,303 ft2)

= 0.91 ha (2.26 ac)
Preliminary pond volume = 95.6 m x 95.6 m x 10.67 m (315 ft x 315 ft x 35 ft)
12.2.12.2 Sludge Pond Layout
Assuming the following parameters, the sludge pond was sized.
Pond inside slope 2%:1 Outside slope 3:1 Berm width = 4.51 m (15 ft)

a. First Trial
Estimate square pond bottom dimensions at 76.2 m x 76.2 m (250 ft x 250 ft):
Effective Volume = (76-2) + ^(117.3) (8>23) = 80j513 m3 (2,843,274 ft3)

which is slightly greater than the required volume, 25,190 m3 (2,655,000
ft3).
Cut Volume =fill Volume
(76.2)2 + (129 - 5h)2 /,n -, UN _ 4.6 + (4.6 + 5.5h) ,UA
o \iu«/*"nj"" « \ •• /
(4(129.7 + 5h) + 4(4.6 + 5.5h))
(22,629 - l,297h + 25I12) (10.7 - h) (%) = (9.2 + 5.5h) (h) (537 + 2h)
242,130 - 13,878h + 268h2 - 22,629h + 1.297I12- 25h3 = 4,940h + 18.4h2 +
2,954h2+ llh3
242,130 - 41,447h - 1,4071^ - 36h3 = 0
36h3 + l,407h2 + 41,447h = 242,130

202
-------
This equation can be solved by trial and error.

h f(h)

4.5 m (14.7 ft) 218,284

5.0 m (16.4 ft) 246,910

4.9 m (16.1 ft) 241,108

h = 4.9 m (16.1 ft)

Cut Volume = 48,933 m3 (64,726 yd3)
Fill Volume = 48,446 m3 (64,082 yd3)

The volume to be excavated is equal to the fill needed for the dikes when the
depth of excavation is 5.8 m (19.0 ft) and the dikes are raised 4.9 m (16.1
ft) above the existing ground level, making a total depth of 10.7 m (35.1
ft). This assumes level ground at the excavation site.

Total area required = IO^OQ $/ha = 2.6 ha (6.4 ac)
12.3 Design Example II

This treatment plant is to be designed to treat a flow of 11,520 m3/d (3.04
Mgal/d) with a raw water quality as shown in Table 12-2.

Several assumptions concerning the operation of this plant have been made to
aid in its design. The plant is designed to operate continuously with as
much automation as possible, reducing the need for operator supervision to
one visit per day. Gravity flow between the unit processes should be pro-
vided. These processes should include flow equalization, pH adjustment,
aeration, solids separation, and sludge disposal. It is also assumed that
limited land area is available, and that final sludge disposal is to a deep
mine through a borehole.

12.3.1 Equalization Basin

Because large underground storage is available, only a small equalization
basin need be constructed (1 day storage). This basin will be at least
11,520 ms (3.04 Mgal) in size.

Assume a 3.0-m (10-ft) water depth and a 0.9-m (3-ft) freeboard above the
maximum water level.
203
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TABLE 12-2
RAW WATER AND EFFLUENT QUALITY LIMITATIONS
Effluent Limitations
Parameter

PH
Acidity
Sulfate, mg/1
Suspended solids, mg/1
Iron (total), mg/1
Iron (ferrous), mg/1
Manganese, mg/1
Nickel, mg/1
Zinc, mg/1
Aluminum, mg/1
Raw Water
2.9
2,200
2,000
165
480
440
11
1.5
3
150
30 Consecutive
Day Average
6-9
net acidity
—
35
3.5
—
2.0
—
—
— _
Maximum
Daily
6-9
net acidity
—
70
7.0
—
4.0
—
—
— mm
Approximate Water Surface Area Requirement =
11
= *
m

= 3,840 m2 (41,330 ft2)
Design the equalization basin to be a square earthen basin with 2%:1 inside
slopes and 3:1 outside slopes. Make the top berm width 4.6 m (15 ft) and the
inside berm-to-berm length 54.9 m (180 ft). These dimensions will give the
volume required, as shown by the following calculation:
Water Volume =
H - (54.9)2 + (70.1)2 (3) = n>900 ^ (420j000 ft3}
where Ax and A2 are the bottom and top water surface areas respectively, and
H is the water depth.
The following calculations show how the cut and fill volumes are balanced to
determine the proper excavation depth. It is assumed that the land area
204
-------
before excavation is flat (see Figure 12-3).
Cut Volume = Fill Volume
(54.9)2 + (74.7 - 5h)2 (3_g6 _ h) = 9M^ 5.5h
(3,014 + 5,580 - 747h + 25h2) (3.96 - h) (%) = (%) (9.1 + 5.5h) (h) (277.4
+ 12h)
34,032 - 2,958h + 99h2 - 8,594h + 747h2 - 25h3 = 2,524h + l,526h2 + 109h2 +
66h3
34,032 - ll,552h + 846h2 - 25h3 = 2,524h + l,635h2 + 66h3
14,076h + 789h3 + 91h 3 = 34,032

This equation can be solved for the height of the dikes above ground level
(h) by trial and error, choosing a value of h and solving for the value to
the right of the equal sign.

h f(h)
1.5 m (5 ft) 23,196
2.4 m (8 ft) 39,585
2.1 m (7 ft) 33,882
Cut Volume = 6,500 m3 (8,500 yd3)
Fill Volume - 6,340 m3 (8,300 yd3)
If the equalization pond site is level, 1.8 m (6 ft) of pond excavation will
provide the fill necessary to build the dikes 2.1 m (7 ft) above ground
level, giving a total height of 3.9 m (13 ft).
Total Water Surface Area = (70.1 m)2 = 4,915 m2 (52,900 ft2)
= 0.49 ha (1.21 ac)
Total Pond Area = (96.6 m)2 = 9,340 m2 (100,500 ft2)
= 0.94 ha (2.31 ac)
205
-------
83.8m (275ft)
74.7m-5h <245ft-5h>
83.8m+6h (275ft+6h>
6.4m
(21ft»
\
/

\
/

4.6m
(15ft)
96.7m (317f«
/
\

X
\
20.9m ^
(68.5ft)

Figure 12-3. Equalization basin, Design Example II

206
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12.3.2 Theoretical Lime Requirement

Assume a 70% lime efficiency. The net alkalinity, as determined from analy-
ses, is -2,200 mg/1 as CaCO,. The theoretical lime requirement from Table
1-2 is thus
2'200 m9 x X = 2>325 m^ Ca(OH)2
The theoretical daily lime (Ca(OH)2) requirement is

2.32JLJK1 x ll^Ojnl x 1^1 x _m . 2

= 26.8 kkg/d (29.5 tons/d)

From a titration curve, it was found that 12% excess lime is needed to main-
tain a pH of 9.0 for manganese removal, but 15% will be used in this calcu-
lation. The theoretical daily lime (Ca(OH)2) requirement is then

26,784 kg/d x 1.15 = 30,802 kg/d (67,764 Ib/d), or

= 30,8 kkg/d (33.9 tons/d)

The above procedure can be used as a preliminary estimation of the amount of
lime that must be handled. It is a fairly accurate method, assuming the raw
water quality will not change significantly. This method of approximating
lime usage is the best alternative when a treatability test cannot be per-
formed. When designing a plant of this size, however, a treatability test
should be performed.

12.3.3 Actual Lime Requirement from Treatability Test

A treatability test was conducted according to the procedures outlined in
Chapter 6, Section 6.3. This test indicated that 2,010 mg/1 of lime are
required to neutralize to pH 9.0. Therefore, the actual lime requirement,
assuming constant water quality, is
Xfl 23j200 kg/d {51j000 lb/d)

= 16.1 kg/min (35.5 Ib/min)

= 23.2 kkg/d (25.6 tons/d)

207
-------
At this point, the designer must consider the type of lime to be used for
neutralization. In the past, the obvious choice would have been to use
quicklime, because a significant price difference existed between quicklime
and hydrated lime. The savings that resulted from using quicklime eventually
offset the initial capital investment and maintenance costs associated with a
slaker. The difference in cost between hydrated lime and quicklime has
recently narrowed to within $2.20/kkg ($2.00/ton), making the choice less
clear.

In this case, since the hydrated lime requirement is over 23 kkg/d (25 tons/
d), a $50/d savings can be realized by installing a slaker and using quick-
lime. This is a significant savings that will offset the slaker capital cost
and any anticipated maintenance costs. Therefore, the quicklime system is
chosen.

12.3.4 Quicklime Requirements

It is important to realize that a slaker has an average efficiency near 90%.
The remaining 10% is lost to grit, and lime leaving with the grit. Many
vendors claim higher efficiencies; these claims may be true, making our value
conservative. Slaker efficiency also depends on the quality of lime used.

The quicklime (CaO) requirement is calculated as follows:

CaO Equivalent = 23.2 kkg Ca(OH)2/d x 745g6 ga(oS)%l

= 17.6 kkg CaO/d (19.4 tons/d)

Actual quicklime requirement = 17.6 kkg CaO/d = ig>6 kkg Ca0/d (2Lg tons/d)
12.3.5 Lime Silo

The lime silo should be sized to provide at least 7 days storage.

Silo Capacity = 7 d x 19.6 kkg/d = 137.2 kkg (151 tons)

Assuming a quicklime density of 882 kg/m3 (55 lb/ft3), the silo volume
required for 7 days quicklime storage is
Silo Volume = kg/rn = 155'6 m (5'500

Silo Specifications:

diameter = 3.65 m (12 ft)

208
-------
sidewall height = 13.7 m (45 ft)

60° hopper bottom

12.3.5.1 Bin Activator

A bin activator is highly recommended for silos storing quicklime. Many
types of hoppers or bin activators are on the market. This designer chooses
a 1.83-m (6-ft) vibratory bin activator, one-half the silo diameter.

12.3.5.2 Silo Slide Valve

This simple valve is considered mandatory on every lime storage silo. The
slide valve fits over the hopper opening during filling to prevent lime from
being blown into the feeder room below the silo.

12.3.6 Quicklime Feeder (to Slaker)

The required feeder rate is

= 820 kg/hr (1,800 Ib/hr)

= 0.94 m3/hr (32.3 f13/hr)

A variable screw feeder is selected with an operating range of 0.56-5.60 m3/
hr (20-200 ft3/hr), having a 0.56-kW (0.75-hp) motor.

12.3.7 Slaker

An automatic, thermostatically controlled, continuous slaker is chosen to
maintain a slaking temperature of 77°C (170°F), and to provide an efficient
and safer method of slaking. A slaker with a maximum rated capacity of 1,135
kg/hr (2,500 Ib/hr) will meet the slaking needs of this plant.

12.3.8 Slurry Feed System

The operational criteria for the slurry feed system are listed below.

1. Use a loop system with a pH-controlled pinch valve. A 0.75-kW (1-
hp) compressor, providing 5-atm (60-1b/in2) pressure, is needed to
operate the pinch valve. A pH control system will regulate the
pinch valve.

2. Slurry addition should not exceed IQ% of the raw water flow.

209
-------
3. A minimum velocity of 1.22 m/s (4 ft/s) should be maintained in the
slurry loop.

4. Circulate slurry at three times the average bleed-off rate.

5. Assume the slurry loop total length is 15.25 m (50 ft), the static
head is 3 m (10 ft), and the friction head loss is 3 m (10 ft).

12.3.8.1 Evaluating Possible Slurry Concentrations

The slurry system is to deliver 16.28 kg/min (35.9 Ib/min) Ca(OH)2. Slurry
concentrations are examined in Table 3-1.

The final choice of slurry concentration will involve a tradeoff between
higher maintenance costs associated with high slurry concentrations and the
resulting lower pump and power costs. This designer will use a 15% slurry in
the loop system.

In this case the sulfate concentration (2,000 mg/1) in the raw AMD is safely
below the concentration (3,000 mg/1) where gypsum formation will occur, so
raw AMD can be used as makeup water. Where gypsum will be a problem, a sepa-
rate water source such as a well is recommended.
12.3.9 Stabilization Tank

The tank will be designed to utilize the upper one-half of the tank volume.
This will allow slurry storage for peak demand if the raw water requires more
lime. Even at the low level point, the tank should provide 20 minutes deten-
tion time for the incoming slurry.

One-half tank volume = 20 min x 1.8 1/s x 60 s/min = 2,200 1 (581 gal)

Provide a tank twice this size to provide 20 minutes detention when the tank
is half full.

Tank volume = 4,400 1 (1,162 gal)

= 4.40 m3 (155.4 ft3)

Design a circular tank with the diameter equal to the height. Assume H=D.

Vol = 0.785 (D2) (H)

4.40 = 0.785 (D)3

D = 1.78 m (5.8 ft)

The final tank dimensions are as follows:

Diameter = 1.83 m (6 ft)

210
-------
Height = 2.44 m (8 ft), which includes 0.61 m (2 ft) freeboard
Volume = 4.8 m3 (1,269 gal)
12.3.9.1 Level Controls
Level controls should be placed in the stabilization tank as follows:
1. high-level emergency shutoff at 2.13 m (7 ft);
2. high-level shutoff at 1.83 m (6 ft);
3. low-level slaker starting switch at 0.91 m (3 ft).

12.3.9.2 Miscellaneous Tank Requirements
The following are also included in the design:
1. full drain bottom nozzle;
2. elevated tank to provide flooded suction;
3. a centrifugal slurry pump and a standby slurry pump piped in paral-
lel.

12.3.9.3 Slurry Tank Mixer
The previous rule of thumb, 0.2 kW/m3 (1 hp/1,000 gal), will give a good
approximation of the required mixer size. Therefore, a 4,800 1 (1,269 gal)
tank requires a 0.95-kW (1.27-hp) mixer. Use a 1.1-kW (1.5-hp) propeller
mixer with 1.5 pitch at 350 r/min. The designer has the option to side-mount
or fix-mount the mixer on a small cross beam at a slight angle. This would
eliminate the need for a baffled slurry tank. The designer can also choose a
top-mounted, vertical, on-center mixer with a baffled tank.
12.3.10 Flash Mix Tank
The flash mix tank should provide approximately 5 minutes detention time for
neutralization. The required volume is
Volume = Q(flow) x t (time)
= (133.3 + 1.8 1/s) x (5 min) (60 s/min)
= 40,550 1 (10,713 gal)
Volume = 40.55 m3 (1,432 ft3)

211
-------
Design a circular tank with the diameter equal to the height.
Volume = 0.785 D2 (H) Set D = H
40.55 m3 = 0.785 D3
D3 = 51.7 m3
D = 3.72 m (12.22 ft)
Make the diameter 3.81 m (12.5 ft). The final tank dimensions are
Diameter = 3.81 m (12.5 ft)
Height = 4.42 m (14.5 ft), which includes 0.61 m (2 ft) freeboard
12.3.10.1 Baffles
Four baffles should be placed in the tank, 90° apart. Design specifications
are
Width = (3.81 m) x (1/12) = 0.32 m (12.5 in)
Wall clearance = 0.15 x (0.32 m) = 0.048 m (2 in)
Bottom clearance = 0.5 (0.32 m) = 0.16 m (6 in)
Extended baffles = 0.30 m (1 ft) above static water level
12.3.10.2 Mixer
Using 0.2 kW/1 m3 (1 hp/1,000 gal),
40.55 m3 x 0.2 kW/1 m3 = 8.0 kW (10.7 hp)
A 7.46-kW (10-hp) mixer will be satisfactory. It should be a top-mounted,
vertical, on-center, axial turbine mixer. The influent pipes should enter
the top with an air break, while the outlet should be an inverted elbow at
3.81 m (12.5 ft) with a 0.91-m (3-ft) submerged nipple.
The inlet and outlet must be 180° apart.
12.3.11 Aeration Tank
The theoretical iron loading is as follows:
212
-------
440 mg x 11.520m x 1.000 1 x ^kg^ = M69 kg/d (lljl65 1b/d)
The theoretical oxygen requirement per day is

k9 = 724 kg/d (1,595 Ib/d)
A mechanical or surface aerator is selected for use. Based on an oxygen
transfer rate of 2.13 kg/kW-hr (3.5 lb/hp-hr)» an aerator of approximately
14.9 kW (20 hp) is needed. The mixer must be checked after the aeration tank
is sized to insure complete mixing.

12.3.11.1 Aeration Detention Time

From the treatability study, an iron oxidation rate of 15 mg/l/min was deter-
mined. Based upon this rate, a detention time is calculated, assuming total
oxidation.
D = 440 mg/1 = 2g .
ut 15 mg/l/min ^ mn

Employing a safety factor of two, which takes into consideration lower oxida-
tion rates from weather, short-circuiting, higher flows, and an increase in
iron concentrations, a total detention time of 58 minutes should be provided.

The aeration basin volume required is

Volume = flow x time

= (133.3 + 1.8 1/s) (58 min) (60 s/min)

= 470,148 1 (124,213 gal)

= 470.1 m3 (16,606 ft3)

Because the area available is limited, the aeration basin will, be constructed
of either steel or concrete. Also, it will be circular.

Volume = 0.785 D2H Assume: H = 4.6 m (15 ft)

470.1 m3 = 0.785 D2 (4.6 m)

470.1 = 3.61 D2

D2 = 130 m2

D = 11.4 m (37.4 ft)

213
-------
The aeration tank dimensions are
Diameter = 12.2 m (40 ft)
Depth = 5.2 m (17 ft), which includes 0.61 m (2 ft) freeboard
It is imperative to insure that the aerator has enough power to turn over the
basin contents. The geometry of the basin must also be within the aerator
limits.
For the aerator selected, the specifications indicate it will provide a zone
of complete mixing 17.6 m (58 ft) in diameter with a 4.5-m (15-ft) depth.
This is adequate for this application and should provide a margin of safety.

12.3.12 Estimated Sludge Production
The sludge produced by the treatment of this water can be estimated as the
sum of the metal hydroxides removed in significant concentrations, iron and
aluminum, suspended solids in the mine drainage, and unused lime. This
estimation assumes 100% removal.
Hydrolysis of ferric iron, assuming complete oxidation of ferrous iron:
Fe+3+ 3H20 + Fe(OH)3 + 3H +
56 g/g-mol 107 g/g-mol
480 mg/1 -^ (480 mg/1) = 917 mg/1

Aluminum hydrolysis:
Al+3 + 3H20 * A1(OH)3 + 3H +
27 g/g-mol 78 g/g-mol
150 mg/1 || (150 mg/1) = 433 mg/1

Theoretical daily solids production:
Ferric hydroxide
917 mg/1 x ii.520j.il x M^i x __^g__ . 10>560 kg/d {23j270 lb/d)

433 mg/1 x 11 .520 "3 x JM x = 4,990 kg/d (10,990 Ib/d)
165 mg/1 x . m x . Xl^r = 1,900 kg/d (4,190 Ib/d)

214
-------
Unused lime (assume 10%)

23,200 kg/d x 0.10 = 2,320 kg/d (5,100 Ib/d)

Total solids, dry weight = 19,770 kg/d (43,550 Ib/d)

Assuming the sludge solids content is 1%, the daily wet sludge production is

siud,. - '
= 1,977,000 kg/d (4,355,000 Ib/d)

Assuming the sludge weighs the same as water, the volume of sludge produced
daily is

Sludge Volume = 1>97^>{^1kg/d = 1,977,000 1/d (522,300 gal/d)

= 1,977 m3/d (69,800 ft3/d)

12.3.13 Actual Sludge Production from Treatability Test

The above calculations could be used when a treatability test is not per-
formed. The better way to design a plant of this size, however, is with data
obtained from a treatability test.

The sludge volume based upon the treatability test amounted to 11% of the
plant flow, 1,267 m3/d (334,400 gal/d), which is lower than the theoretical
value. This could be caused by several factors: (1) high suspended solids
in the supernatant, whereas the theoretical value assumed removal to 35 mg/1 ;
(2) better lime efficiency; and (3) ideal settling conditions producing a
sludge density greater than 1%. Therefore, it is the designer's option as to
which volume to use. This designer chooses to be conservative and use the
larger volume, 1,267 m3/d (334,400 gal/d).

12.3.14 Settling Unit

Because of limited land availability, a mechanical clarifier will be pro-
vided. Using a rise rate of 0.006 m/min (0.02 ft/min), the clarifier diam-
eter is sized accordingly.

Area = Flow = 11.520 m3/d m , 33Q 2 fl
Area Rise Rate 0.006 m/min x 1,440 min/d lfJJU m U'

Area = 0.785 D2

215
-------
1,330 m3 = 0.785 D2
D = 41.2 m (135 ft)

The depth of the clarifier or thickener is a function of detention time,
raking mechanism, mode of operation, and type of separator chosen (solids
contact vs. conventional). The designer chooses an upflow clarifier with 2
days sludge storage capacity, and 12 hours clear water depth.
The clarifier volume is
2 days sludge storage = 2,530 m3 ( 668,800 gal)
12 hours clear water depth = 5.760 m3 (1.522.000 gal)
Total Volume = 8,290 m3 (2,190,800 gal)
The clarifier depth (H) is
Volume = 0.785 (D)2 (H)
8,290 m3 = 0.785 (41.2 m)2 (H)
H = 6.22 m (20.4 ft)
The final clarifier dimensions will be those of the closest standard-size
clarifier.

12.3.15 Final Sludge Disposal
Final sludge disposal will be to a borehole.
216
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CHAPTER 13

COST ESTIMATING
13.1 Introduction
Determining the capital investment and construction cost of any proposed AMD
treatment plant, as well as its daily or annual cost of operation, is an
important part of the designer's role.

There are many ways to estimate the cost of construction, but this will vary
with the designer's capability. Most designers are able to perform the
engineering and construction in-house, while others might hire a consultant
to do the engineering design and then contract for the plant's construction.
All of these factors influence the total cost of a facility and lead to large
cost differences between similar plants.

Construction cost estimates can be prepared in different ways. For initial
budget purposes, determining the installed or constructed costs for the
various major units of a treatment facility by rule-of-thumb methods and
adding in the costs for associated facilities or services will enable the
designer to establish an initial budget within 25% of the final cost. This
method of estimating can also be used for comparing costs among any options
within a unit process. The cost items to be considered could include the
following:
Major Cost Items

Borehole and pumps

Equalization basin and pumps

Lime system

Aeration system

Settling unit(s)

Sludge dewatering

Sludge disposal
Associated Facilities

Land and access roads

Zoning and permits

Power supply

Interconnecting piping

Panels and wiring

Instrumentation

Special foundations or site
preparation

Fencing

217
-------
Contingency

Engineering fees

Escalation of cost by inflation
during the construction period

Contractor's overhead and profit

Once an initial budget for the project is established, the cost estimate can
be refined or finalized only by completing the engineering design necessary
for construction of the facility and installation of its equipment. The
degree of effort here will determine the accuracy of the cost estimate.
Complete design and detail drawings will enable the estimator to refine his
final budget cost to a contingency of less than 10%. In addition, this
effort will also assist the contractors bidding for the work to provide more
accurate prices.

For rapid estimating purposes, which will determine the "ballpark" cost of a
treatment facility, Figure 13-1 can be used. This figure is based on the
costs for many treatment facilities with varying levels of equipment and
automation. A range of costs based on flow through the facility is shown by
the band on the figure. This covers simple to complex levels of construc-
tion.

Typical initial budget cost estimates have been prepared for the two design
examples presented in Chapter 12. These cost estimates include the rule-of-
thumb methods discussed throughout this manual, as well as pricing values
from the 1980 Dodge Manual for Building Construction Pricing and Scheduling
and the 1980 Means Catalog. These are excellent sources for up-to-date
construction estimating values.

Annual operational and maintenance costs, which include electricity and
chemicals, are also included in the examples. They are based on 1981 product
costs.
13.2 Cost Breakdown of Design Example I

This particular mine drainage plant has a daily average flow of 2,880 m3/d
(527 gal/min). Chemical requirements are 1.8 kkg/d (2 tons/d) of hydrated
lime.

Capital Costs

The following capital costs will be incurred with this plant:

Item Cost ($)

1. Raw Water Pumps (2) 500 gal/min $ 2,000

2. Equalization Basin, 6,000 m3 (1.5 Mgal)

218
-------
COST IN DOLLARS
100,000

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Item Cost ($)

Excavation and material placement
6,000 m3 (7,800 yd3) 9 $3/yd3 $23,400

Clay liner thickness 0.91 m (3 ft)
Area equals 5,850 m2 (63,000 ft2)
5,324 m3 (7,000 yd3) 9 $3/yd3 21,000

3. Lime Storage and Feed System

Silo 27.2 kkg (30 tons) including side ladder, handrail,
cage, 60° hopper, fill pipe 10,000
Bin activator 1.52 m (5 ft) 4,500
5-cm (2-in) screw feeder 2,100
Hopper slide valve 250
Bin level indicator 250
Dust collector 1,300
Concrete foundation
3.65-m (12-ft) square pad x 1.06 m (3.5 ft)
14.5 m3 (19 yd3) 9 $250/yd3 4,750
Pad excavation 29 m3(38 yd3) @ $4/yd3 150
Delivery 600
Erection (minimal because welded silo) 560

4. Lime Fed Dry Directly from Feeder No Cost

5. Flash Mix Tank approximately 11,355 1 (3,000 gal)

Diameter = 2.44 m (8 ft)
Height = 3.05 m (10 ft)
Fiberglass tank includes baffles, nozzles, mixer
mounts 4,100

6. Aeration Tank (Circular Earthen Basin)

Diameter = 6.40 m (21 ft)
Depth = 1.83 m (6 ft)
59 m3 (77 yd3) 9 $3/yd3 230

Bottom concrete formless pour 5 cm (2 in)
37 m2 (400 ft2) x .05 m = 2.0 m3
2.0 m3 (2.6 yd3) ® $200/yd3 525

7. Aerator 0.89 kW (1.2 hp) 3,500

8. Settling Basin (Earthen) with Sludge Removal System

Volume 1,782 m3 (471,000 gal)

From cut and fill
Fill + excavation = 2,910 m3 (3,810 yd3) 9 $3/yd3 11,430

220
-------
Item Cost ($)
Clay liner 0.61 m (2 ft) x 2,809 m2 (30,240 ft2)
1,713 m3 (2,240 yd3) @ $3/yd3 6,720
Hydraulic sludge removal system 15,000
9. Sludge Disposal Pond 46,000 m3 (12.15 Mgal)

From cut and fill
97,379 m3 (124,254 yd3) @ $3/yd 3 372,764
Clay liner 0.61 m (2 ft) x 28,200 m2 (303,546 ft2)
17,-202 m3 (22,500 yd3)(P $3/yd3 67,498
10. Land Cost for Treatment Site (assume 62.5 ha)
6 ac @ $5,000/ac 30,000
11. Instrumentation for Automatic Interlock System with pH
Assemblies, Panel, Annunciator, and Recorder 5,000
12. Electrical Motor Starts, Transformer, Heater, Control
Room with Insulation, Lighting 5,000
13. Piping and Miscellaneous 5,000
14. Electric Power to the Site (assume $25,000) 25,000
15. Fencing, Complete Enclosure
$33.00/m ($10/ft) x 3,100 m 31.000
Initial Construction Cost Estimate $653,627
Engineering Fees 9 10% 65,363
Contingency @ 15% 98,050
Contractor's Overhead and
Profit @ 20% 150.335
TOTAL CAPITAL COST BUDGET $967.375

Annual Operational and Maintenance Costs
1. Electricity (11.7 KwH @ $0.05/KwH) $ 5,130
2. Chemicals (lime @ $65/ton) 47,450
221
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Item Cost ($)

3. Manpower (1 worker @ 4 hr/d @ $10/hr) 14,600

4. Sludge Disposal (assume $10,000/yr for maintenance
of hydraulic sludge removal system) 10,000

$77.180
13.3 Cost Breakdown of Design Example II

This particular mine drainage plant has a daily average flow of 11,520 m3
(2,110 gal/min). Chemical requirements are 17.6 kkg (19.4 tons) of quick-
lime per day.

Capital Costs

Item Cost ($)

1. Raw Water Pumps (2) $ 4,500

2. Equalization Basin 11,520 m3 (3.0 mg)

Cut and fill
12,840 m3 (16,795 yd3) @ $3/yd3 50,385

Clay liner 0.91 m (3 ft)
9,336 m3 (100,500 ft2) @ 0.91 m (3 ft) =
8,496 m3 (11,113 yd3) @ $3/yd3 33,340

3. Lime Storage Silo Capacity 137 kkg (151 tons)

This will be a completely welded tank including caged
side ladder, handrail, 60° hopper, fill pipe 21,000

Bin activator 1.82 m (6 ft) 5,000
10.2-cm (4-in) screw feeder 3,200
Hopper slide valve 350
Bin level indicators 250
Dust collector 2,600
Concrete foundation
4.6-m (15-ft) square pad x 1.5 m (5 ft)
32 m3 (42 yd3) @ $250/yd3 10,500
Pad excavation 64 m3 (84 yd3) @ $4/yd3 328
Delivery 1,200
Erection (welded silo) 1,120
222
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Item Cost ($)

4. Lime Feed System (Slaking Quicklime)

1,135 kg/hr (2,500 Ib/hr) slaker 15,200
Slaker installation 2,500
Slurry loop (piping and pinch valve) 1,000
Compressor 0.75 kW (1 hp) and installation 1,200
Recirculation pump 0.37 kW (0.5 hp) 750
Stabilization tank 4,400 1 (1,162 gal) fiberglass
material, mixer mounts, level controls, nozzles 2,100
Propeller mixer 1.10 kW (1.5 hp)
stainless steel shaft and propeller 1,500
Miscellaneous fabrication and piping 1,000

5. Flash Mix Tank 50.5 m3 (13,500 gal)

Coated steel tank, baffles, nozzles, mixer mounts 5,500
Mixer 7.46 kW (10 hp) 8,000

6. Circular Concrete Aeration Basin

604.5 m3 (160,000 gal) volume
Excavation 1,224 m* (1,600 yd3) @ $3/yd3 4,800
Concrete for walls 78 m3 (102 yd3) @ $250/yd3 25,500
Concrete for pad 135 m3 (177 yd3) @ $250/yd3 44,250
Aerator 14.9 kW (20 hp) fix-mounted 22,000
Aerator walk-on platform 5,000

7. Settling Unit (Mechanical Clarifier)

Concrete structure with steel mechanism
Single unit 41.2 m (135 ft) in diameter
Volume is 8,290 m (2,190,300 gal)
Clarifier will be half-buried
Excavation 6,373 m3 (8,335 yd3) @ $3/yd3 25,000

Clarifier wall concrete 463 m3 (605 yd3) @ $175/yd3 105,875

Bottom structure and pad
1,737 m3 (2,272 yd3) @ $175/yd3 397,600

Mechanism 41.2 m (135 ft) @ $l,250/ft 168,750

Piping and miscellaneous 10,000

8. Sludge Disposal to Borehole Variable

9. Land Cost for Treatment Site (assume 2 ha)

5 ac @ $5,000 25,000
223
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Item Cost ($)
10. Instrumentation for Automatic Interlock System, pH
Assembly, Pumps, Control Panel, Annunciator, Recorder,
and Housing 30,000
11. Electrical Motor Starters, Transformer Heater, Lighting,
Insulation, Wiring 30,000
12. Piping and Miscellaneous 15,000
13. Electrical Power to Site Variable
14. Fencing Minimal $ 15.000
Initial Construction Cost Estimate $1,096,298
Engineering Fees @ 10% 109,630
Contingency @ 15% 164,445
Contractor's Overhead and
Profit @ 20% 252.148
TOTAL COST BUDGET $1.622,521

Annual Operational and Maintenance Costs
1. Electricity (36.5 KwH @ $0.05/KwH) $ 16,000
2. Chemicals (19.4 tons quicklime/d @ $60/ton) 424,860
3. Manpower (1 worker @ 6 hr/d @ $10/hr) 21.900
$ 462,760
224
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APPENDIX
PHYSICAL AND CHEMICAL PROPERTIES OF LIME
A.I Specifications on Lime
Use of the following standards by the American Society for Testing and
Materials specifying the nomenclature and the chemical and physical methods
of testing chemical lime products is recommended.
C 51-71 - Terms Relating to Lime
C 50-57 - Sampling, Inspection, Packing, and Marking of Quicklime and
Lime Products
C 110-71 - Physical Testing of Quick and Hydrated Lime
C 400-64 - Testing Quicklime and Hydrated Lime for Neutralization of
Waste Acid
C 25-72 - Chemical Analysis of Limestone, Quicklime, and Hydrated Lime
C 53-63 - Quicklime and Hydrated Lime for Water Treatment
C 433-63 - Quicklime and Hydrated Lime for Hypochlorite Bleach
Manufacture
C 415-72 - Quicklime and Hydrated Lime for Calcium Silicate Products
C 258-52 - Quicklime for Calcium Carbide Manufacture
C 259-52 - Hydrated Lime for Grease Manufacture
C 49-57 - Quicklime and Hydrated Lime for Silica Brick Manufacture
C 46-62 - Quicklime and Limestone for Sulfite Pulp Manufacture
C 45-25 - Quicklime and Hydrated Lime for Cooking of Rags in Paper
Manufacture
C 593-69 - Fly Ash and Other Pozzolans for Use with Lime
Copies of these standards may be obtained by writing to the American Society
for Testing and Materials, 1916 Race Street, Philadelphia, Pa., 19103.

225
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TABLE A-l

TYPICAL ANALYSES OF COMMERCIAL QUICKLIMES

High Calcium Quicklimes Dolomitic Quicklimes
Component Range Range
CaO
MgO
Si02
Fe203
A1203
H20
C02
93.25 -
0.30 -
0.20 -
0.10 -
0.10 -
0.10 -
0.40 -
98.00
2.50
1.50
0.40
0.50
0.90
1.50
0
/
55.50
37.60
0.10
0.05
0.05
0.10
0.40
C
- 57.50
- 40.80
- 1.50
- 0.40
- 0.50
- 0.90
- 1.50
aThe values given in this range do not necessarily represent minimum and maxi-
mum percentages.

TABLE A-2

pH OF CALCIUM HYDROXIDE SOLUTIONS AT 25°C

CaO £H
9/1

0.064 11.27
0.065 11.28
0.122 11.54
0.164 11.66
0.271 11.89
0.462 12.10
0.680 12.29
0.710 12.31
0.975 12.44
1.027 12.47
1.160 12.53
Since solubility of lime decreases as the temperature increases, the pH of
lime solutions is correspondingly lower at higher temperatures.

Data from F.M. Lea and G.E. Bessey, Journal of the Chemical Society,
1,612-1,615, 1937.

226
-------
ro
ro
Properties
Chemical Name
Chemical Formula
Molecular Weight
Melting Point
Decomposition Point
Boiling Point
Refractive Index
Heat of Solution at
18°C
Crystalline Form
Density
Solubility:
In Hot and Cold
Water
TABLE A-3
PROPERTIES OF THEORETICALLY PURE LIME COMPONENTS

Pure Lime
Quicklime Components Hydrated Lime Components
Calcium Oxide Magnesium Oxide Calcium Hydroxide Magnesium Hydroxide
Ca(OH),
Mg(OH),
74.096
58.336
CaO MgO
56.08 40.32
2,570°C (4,658°F) 2,800°C (5,072°F)
580°C (1,076°F) 345°C (653°F)a
2,850°C (5,162°F) 3,600°C (6,512°F)
1.838 1.736 1.574 and 1.545 1.559 and 1.580
+18.33 kg-cal
cubic
3.40
cubic
3.65
+2.79 kg-cal
hexagonal
2.343
-0.0 kg-cal
hexagonal
2.4
See Solubility Sections A.2 and A.4.
There is not complete agreement on the exact decomposition point of Mg(OH)2; however, the value given
represents the best data available.
-------
PO
ro
00
Lime
Substances
CaO
Ca(OH)2
MgO
Mg(OH)2
CaO-MgO
Ca(OH)2 -MgO
Ca(OH)2-Mg(OH)2
TABLE A-4
GRAVIMETRIC PERCENTAGES OF CRITICAL CONSTITUENTS OF LIMES
Percents of Elements
Ca_ M£ £

71.47 — 28.53
54.09 -- 43.19
60.32 39.68
41.69 54.85
41.58 25.23 33.19
35.03 21.26 41.95
30.27 18.36 48.33
H20
3.46 30.88

1.76 15.75
3.04 27.21
Percents of Compounds
CaO
100.00
75.69
—
—
58.17
49.01
42.35
MgO
—
100.00
69.12
41.83
35.24
30.44
Alkali Oxides
(CaO +
MgO)
100.00
75.69
100.00
69.12
100.00
84.25
72.79
-------
TABLE A-5

PROPERTIES OF TYPICAL COMMERCIAL LIME PRODUCTS
QUICKLIMES
High Calcium
Dolomitic
Primary Constituents
Specific Gravity
Bulk Density (Pebble Lime), lb/ft3 ....
Specific Heat at 38°C (100°F), BTU/lb. . .
Anqle of Reoose
CaO
3.2 - 3.4
55 - 60
0.19
55°a
CaO and MgO
3.2 - 3.4
55 - 60
0.21
55°a
HYDRATES
High
Calcium
Normal
Dolomitic
Pressure
Do!omi ti c
Primary Constituents . .
Specific Gravity ....
Bulk Density, Ib/ft3 . .
Specific Heat at 38°C
(100°F) BTU/lb
Angle of Repose
Ca(OH)2
2.3 - 2.4
25 - 35b
0.29
70°a
Ca(OH)2+ MgO
2.7 - 2.9
25 - 35b
0.29
70°a
Ca(OH)2+ Mg(OH)2
2.4 - 2.6
30 - 40b
0.29
70°a
The angle of repose for both types of lime (hydrate in particular) varies
considerably with mesh, moisture content, degree of aeration, and physical
characteristics of the lime (e.g., for quicklime it generally varies from
50°-55°, and for hydrated lime it may range as much as 15°-80°.

In some instances, these values may be extended. The Scott method is used
for determining the bulk density values. In calculating bin volumes, the
lower figure should be used.
229
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A.2 Solubility of Calcium Hydroxide

TABLE A-6

SOLUBILITY OF CALCIUM HYDROXIDE IN WATER

Grams/100 q Saturated Solution
t°c
0
10
20
25
30
40
50
60
70
80
90
100
CaO
0.140
0.133
0.125
0.120
0.116
0.106
0.097
0.088
0.079
0.070
0.061
0.054
Ca(OH)2
0.185
0.176
0.165
0.159
0.153
0.140
0.128
0.116
0.104
0.092
0.081
0.071
The solubility of commercial limes in water does not vary more than 7% from
the solubility of pure calcium hydroxide. The differences are probably due
to the trace amounts of sodium and potassium hydroxide in commercial limes.
Magnesia, silica, and carbonate have no effect upon the solubility of ordi-
nary lime, but may have a marked effect upon its rate of solution.

Particle size has considerable influence upon solubility. Freshly slaked
lime, which is of small particle size, is about 10% more soluble than coarse-
particle or aged slake lime. This effect is due to the slow expansion of the
dry lime particles during storage.

These solubility data are derived from A. Seidell, Solubilities of Inorganic
and Metal Organic Compounds, 3rd ed., vol. 1, 209-210, 1940.
230
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A.3 Calculating Weights of Slurry

For calculating the weights of slurry with varying percentages of water, the
following formula may be used:

u 6.237s
w " 100 - a + sa

where W = weight in pounds of slurry per cubic foot

s = specific gravity of dry lime solids

a = percent water in slurry

A.4 Solubility of Magnesium Hydroxide

Magnesium hydroxide is virtually insoluble in water. At 18°C (64°F) and 100°
C (212°F), the solubilities are 0.0098 and 0.0042 g of Mg(OH) /I respectively,
in a saturated solution. The presence of small quantities of NaCl and Na SO
in the aqueous solution will increase the solubility of Mg(OH) slightly.

These solubility data are from A. Seidell, Solubilities of Inorganic and
Metal Organic Compounds, 3rd ed., vol. 1, 982, 1940.

A.5 Heats of Reaction at 25°C

Hydration or Slaking

CaO + H20 = Ca(OH), - heat evolved = 15,300 cal/g mol
= 27,000 BTU/lb mol

MgO + H20 = Mg(OH)2 - heat evolved = 8,800 to 10,000 cal/g mol
= 14,400 to 18,000 BTU/lb mol

Carbonation

CaO + C02 = CaC03 - heat evolved = 43,300 cal/g mol
= 78,000 BTU/lb mol

MgO + C02 = MgC03 - heat evolved = 28,900 cal/g mol
= 52,000 BTU/lb mol

Derived from Int. Crit. Tables, vol. V, 195-196.
231
*US GOVERNMENT PRINTING OFFICE 1983-659-095/0571
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