WATER POLLUTION CONTROL RESEARCH SERIES • 12010 EZV 02/70
Treatment of Waste Water-
Waste Oil Mixtures
U.a DEPARTMENT OF THE INTERIOR • FEDERAL WAflW POLLUTION CONTROL ADMINISTRATION
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WATER POLLUTION CONTROL RESEARCH
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TREATMENT OF
WASTE WATER - WASTE OIL MIXTURES
FEDERAL WATER POLLUTION CONTROL ADMINISTRATION
DEPARTMENT OF THE INTERIOR
By
ARMCO STEEL CORPORATION
703 CURTIS STREET
MIDDLETOWN, OHIO 45042
PROGRAM NO. 12010 EZV
GRANT NO. WPRD-169-01-68
MAY, 1970
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FWPCA Review Notice
This report has been reviewed by the Federal
Water Pollution Control Administration and
approved for publication. Approval does not
signify that the contents necessarily reflect
the views and policies of the Federal Water
Pollution Control Administration.
For rale by tbt Superintendent of Document*, U.S. Government Printing Office
Wtthlngton, D.C. 20*02 - Price 12.60
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ABSTRACT
Cold reduction of steel strip results in the nroduction of large
quantities of waste water containing variable amounts of oil. A
five stand tandem cold mill located at Armco Steel Corporation's
Ashland, Kentucky T-forks produces 200 to 500 gpm of waste water
containing UOO to U,000 ppm of oil. The COD of the waste varies
from UOO to 20,000 npm.
A treatment process and facility was developed, constructed, and
demonstrated, on full scale, for the treatment of cold mill wastes.
The treatment process utilized chemical coagulation to break the
emulsions. The chemicals employed included alum, lime, clay and
organic polyelectrolyte. The process consisted of the following
treatment steps; equalization, chemical addition and rapid mixing,
flocculation, and dissolved air flotation. A number of treatment
variables were studied in the laboratory and in the field in order
to establish process kinetics and optimum treatment efficiency.
Zeta potential, streaming current, and particle size distribution
were used in laboratory studies to describe the effect of the
following variables on process kinetics; acid number, initial oil
concentration, type of emulsifier, chemical dosage, order of
chemical addition, reaction time, and final pH. Based on these
studies, an hypothesis of the emulsion breaking mechanism was
proposed.
Oil, COD, and turbidity were used in field studies to establish
the effect of the following variables on treatment efficiency;
chemical concentration, order of chemical addition, chemical
mixing time, flocculation mixing time and speed, and air flotation
time and recirculation rate. Based on these studies, design
criteria and operating costs for this process were presented.
This report was submitted in fulfillment of Program No. 12010 EZV
between the Federal Water Pollution Control Administration and the
Armco Steel Corporation.
KEY WORDS
Waste Water Treatment / Emulsions1/ ^ Flotationx/
Industrial Wastes Coagulation / ' Kinetics y
Steel Wastes Flocculation^ Zeta Potential
Oily Wastes
iii
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CONTENTS
ABSTRACT
SECTION I
SECTION H
SECTION III
SECTION IV
SECTION V
SECTION VI
SECTION VII
SECTION VIII
SECTION IX
- CONCLUSIONS
- RECOMMENDATIONS
- INTRODUCTION
- BASIC RESEARCH
Present Theory and Practice
Experimental
Discussion
- FIELD STUDIES
Batch Treatment of Concentrated
Coolant
Continuous Treatment of Rinse
Waters
Comparative Economics
- ACKNOWLEDGMENTS
REFERENCES
GLOSSARY
APPENDICES
APPENDIX A
APPENDIX B
APPENDIX C
APPENDIX D
APPENDIX E
APPENDIX F
APPENDIX G
Research & Development
Work
Detailed Engineering
Report and Drawings
Kinetic Relationships
Employed
Kinetic Relationship
for Sequential Reactors
Use of Single Channel
Coulter Counter for
Flocculation Kinetics
Samples and Reagents
Experimental Equipment
Page No.
xii
1
5
9
11
11
19
hh
57
58
65
79
85
87
89
93
93
103
127
129
131
135
137
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LIST OF FIGURES
Figure
No. Title
1. VARIATION IN ZETA POTENTIAL AND pH OF SPENT
MILL EMULSION WITH DILUTION
2. VARIATION IN ZETA POTENTIAL WITH ALUM ADDITION
3. VARIATION IN ZETA POTENTIAL WITH ADDED EMULSIFIER
(TRITON X-100)
U. VARIATION IN TURBIDITY OF "TYPICAL" ROLLING OIL
EMULSION WITH TRITON X-100 EMULSIFIER
5. FRACTION CHANGE IN TURBIDITY WITH DILUTION
6. JAR TESTING OF ALUM, LIME, AND CLAY
7. JAR TESTING OF "TYPICAL" NONIONIC EMULSION
8. JAR TESTING OF "TYPICAL" ANIONIC EMULSION
9. KINETICS OF FLOCCULATION: FRACTIONAL RESIDUAL
TURBIDITY VS. FLOCCULATION TIME AT VARIOUS
VELOCITY GRADIENTS
10. FRACTIONAL RESIDUAL TURBIDITY AS A FUNCTION OF
VELOCITY GRADIENT AT CONSTANT FLOCCULATION TIMES
11. SIMPLIFIED FLOW DIAGRAM BATCH TREATMENT SYSTEM
12. SIMPLIFIED FLOW DIAGRAM CONTINUOUS TREATMENT
SYSTEM
B-l thru B-10 DETAILED ENGINEERING DRAWINGS
E-l FLOCCULATION OF A 200 ppm COMMERCIAL ROLLING
OIL WITH ALUM VIA COULTER COUNTER TECHNIQUES
E-2 FLOCCULATION OF A UOO ppm COMMERCIAL ROLLING
OIL EMULSION WITH ALUM VIA COULTER COUNTER
TECHNIQUES
E-3 FLOCCULATION OF A 1,000 ppm COMMERCIAL ROLLING
OIL EMULSION WITH ALUM VIA COULTER COUNTER
TECHNIQUES
Page No.
23
25
28
31
36
39
10
la
15
53
61
66
APPENDIX B
APPENDIX E
APPENDIX E
APPENDIX E
vii
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LIST OF TABLES
Table
No. Title Page No.
I. TYPICAL ROLLING OIL COMPOSITION 13
II. GENERAL STABILITY CHARACTERISTICS RELATIVE 16
TO ZETA POTENTIAL
III. STOCK OIL EMULSION COMPOSITIONS 20
IV. VARIATION IN ZETA POTENTIAL OF 6,000 ppm 22
SPENT MILL EMULSION WITH MILL MIX AND/OR
C02-FREE DEIONIZED WATER
V. JAR TESTING COMMERCIAL ROLLING OIL EMULSIONS 2?
WITH ALUM, LIME, CLAY AND POLYMER
VI. COMPARISON OF ROLLING OIL EMULSIONS AT DIFFERENT 29
ACID NUMBERS
VII. INITIAL COMPARATIVE JAR TEST RESULTS 32
VIII. JAR TEST RESULTS FOR CONCENTRATION VARIATION 33
IX. COMPARISON OF MEASURED AND CALCULATED 35
TURBIDITIES
X. pH OPTIMIZATION JAR TEST RESULTS 38
XI. THE EFFECT OF ORDER OF ADDITION OF ALUM AND U2
LIME DURING JAR TESTING
XII. RESIDUAL TURBIDITY VS. VELOCITY GRADIENT AND hi
ACID NUMBER
XIII. SUMMARY OF INFLUENT ANALYTICAL DATA FOR 59
CONCENTRATED COOLANT
XIV. SUMMARY OF BATCH TREATMENT RESULTS 62
XV. SUMMARY OF BATCH TREATMENT PERFORMANCE FOR 63
VARIABLES EVALUATED
XVI. INFLUENT ANALYTICAL DATA - RINSE WATER 70
XVII. CHEMICAL ADDITION SCHEMES 72
XVIII. DEMONSTRATION PROGRAM FIELD DATA 77
ix
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Table
No. Title Page No>
XIX. TREATMENT COST PER TON OF STEEL ROLLED ($) 81
XX. TREATMENT PLANT CAPITAL COSTS 82
AI and All. EFFECTS OF VARIOUS PARAMETERS ON MINIMUM APPENDIX A
CHEMICAL DOSAGES REQUIRED TO TREAT WASTES
CONTAINING 1,000 ppro OIL
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SECTION I
CONCLUSIONS
An evaluation of the results of the studies conducted as part of this
grant has resulted in a number of conclusions. Since the basic research
studies and the field studies were to a great extent independent efforts,
the conclusions drawn from the results of these two study areas will be
presented separately.
With regard to the field studies, it should be noted that in general
no two cold mills are operated identically. On this basis, the con-
clusions made with regard to the Ashland Works cold mill and the
success of the related treatment plant cannot necessarily be extended
to treatment of all cold mill wastes.
BASIC RESEARCH STUDIES
1. The stability of concentrated wastes as measured by zeta potential
is less than the same wastes diluted to oil concentrations equiv-
alent to rinse water. This was indicated by a 2l\. percent reduction
in zeta potential. This is due in part to the higher specific
conductivity of the water phase in well used emulsion systems.
(Pages 22-23, U9)
2. The particle size distribution of rolling oils is measurably
shifted toward smaller mean particle sizes during the first few
days of use, apparently as a result of the temperature, pressure,
and shear to which the emulsion is subjected during use.
(Page 22)
3. Colloid titration techniques were too insensitive for estimation
of intrinsic emulsion stability. (Pages 20-21)
U. Relative streaming current techniques for measuring emulsion
stability appear both feasible and amenable to process control.
(Pages 21-22)
5. Maximum flocculation efficiency occurred:
a. Mhere anionic emulsifier was substituted for nonionic
emulsifier at relatively low velocity gradients.
However, the converse appears possible at velocity
gradients higher than lUi sec . (Pages 27,
b. At a pH value of 6.0 - O.U. (Pages 39-l£,
c. When alum and lime flocculants are added simultaneously.
(Page
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6. The effect of acid number on overall process efficiency was
insignificant to at least acid number 16 for "typical" rolling
oil formulations containing added emulsifier. (Pages U3, H8)
7. The following generalized mechanism for coagulation and floc-
culation is proposed. (Pages
a. Coagulation occurs via reaction of positively charged
hydroxo complexes of aluminum with oil droplets to form
positive, negative, or neutral particles depending on
emulsifier type and pH;
b. These particles subsequently react through coulombic and
van der Waals forces with growing positively charged alum
floe ; and
c. The resulting floe grows with concomitant reduction of floe
charge .
8. The increase of optimum alum dose is small over the oil con-
centration range 500 to 1,000 ppm. (Page 50)
9. Ideal flocculation kinetics were found to dominate commencing
at mean velocity gradient values of between 28 to 60 sec"1
for AN16 nonionically emulsified systems. Redispersion of
floe does not occur at velocity gradients below the highest
value tested (lUU sec'-1-). The composite flocculation rate
constant (kD) was found to be about 10"^ (mg/l)~l. Bench
• level testing proceeded under the influence of this rate
constant for three minutes at which point the overall rate
of flocculation decreased abruptly. (Pages 52-55)
10. Turbidity measurements in this work were found to be adequate
for kinetics studies and the feasibility of using a single
channel Model B Coulter Counter for more refined kinetics work
was demonstrated. (Pages 127-128, 131-13U)
11. The kinetics relationships presented indicate that the mean
velocity gradient via shaft torque or shaft horsepower is
preferred over tip speed measurements for jar testing.
(Pages 127-128)
FIELD STUDIES
Batch Treatment of Concentrated Coolant
1. The particular operating practice of the Ashland '-forks Cold Mill
and the subsequent destination of the cold rolled product
permitted operation of the mill on a recirculated coolant system
which reduced waste water flows from the 1,500 gpm plant design
flow to approximately 100,000 gallons per week. (Pages 57-58)
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2. The type of rollinc; oil used and the type of emulsion breaker
employed, determine the optimum treatment requirements.
(Pages 60, 67, 32)
3. Treatment efficiency is reduced by either underdosing or over-
dosing with emulsion breaker. (Page 59)
Lu The results of jar testing conducted during this study were more
sensitive to changes in operating parameters and chemical con-
centrations than the results from the full-scale facility.
(Page 65)
5. Under the conditions evaluated, heating the waste up to 150 F
offered no advantage in treatment efficiency. (Pages 60-6U)
6. Although some physical separation of oil (tramp) from concentrated
coolant can be achieved, total oil removal is significantly
greater with the use of a cationic polymer as an emulsion breaker.
(Pages 62-63)
7. The treatment process and facility as operated is capable of
greater than 90 percent removal of total oil and COD. (Pages 62-63)
8. The recovered oil is of fuel oil quality. (Page 6U)
Continuous Treatment of Rinse Water
9. Effective treatment of combinations of oily rinse water and con-
centrated coolant can be obtained using alum, lime, clay, and
organic polymer. Typical chemical dosages for the waste water
received during this study were 175 ppw alum, 36 ppm lime,
15 ppm clay, and 0.5 ppm polymer. This represents a chemical
cost of approximately $0.05 per 1,000 gallons. (Pages 70-71)
10. Simultaneous addition of alum, lime, and clay to the same mix
tank at the detention times studied did not significantly change
treatment efficiency. (Pages 71-73)
11. A mixing period between the addition of other chemicals and
polymer is necessary for efficient utilization of polymer.
(Pages 71-73)
12. Equalization is an essential requirement for continuous effective
treatment. (Pages 68-69)
13. Control of pH is a significant factor in maintaining effective
treatment of oily rinse waters by coagulation with inorganic
salts. (Page 69)
Hi. Air flotation of the oily floe produced by this process is required
to achieve satisfactory separation with a separation detention
time of 57 minutes or below. (Page 73)
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15. For the system studied, flocculation equipment was unnecessary
to achieve effective treatment of oily rinse waters. (Pages 73-7U)
16. An air flotation detention time of 2£ minutes including a recycle
rate of 20 percent of the influent flow is adequate for floe
separation. (Page 73)
17. The treatment process and facility as operated is capable of
90 percent removal of total oil, COD, and BOD. (Pages 75-79)
18. The recovered oil is adequate for use as road oil but due to its
high water and solids content, it is unacceptable as a boiler
fuel. (Pages 76-79)
19. Treatment plant operating costs range from $O.OU to $0.07 per ton
of steel rolled depending upon the type of rolling oil and the
waste volumes. (Page 81)
20. Based upon treatment considerations for concentrated coolant and
rinse water, operation of the Cold Mill on a recirculated coolant
system is preferred over the rinse water system. Batch treatment
of concentrated coolant is favored on the basis of less discharged
pollutants and more economical treatment. (Pages 79-83)
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SECTION II
RECOMMENDATIONS
On the basis of the conclusions drawn, a number of recommendations are
made below. Many of these recommendations are presented relative to
the logical extension of this work. It is hoped that those recommenda-
tions based on experimental conclusions will assist the technologist
in the field and that those pertinent to the extension of this work
will provide a measure of guidance to those who wish to expand present
knowledge in this area of pollution abatement.
BASIC RESEARCH STUDIES
1. Research should be expanded relative to the efficiency of emulsion
breaking of "typical", virgin, used, and spent commercial rolling
oils as a function of emulsifier content and type. The analytical
procedure recently reported by Ludwig (7) may be of value in this
regard.
2. Research should also be done relative to the lard oil to mineral
oil ratio and its effect on flocculation efficiency.
3. A study of optimum alum-lime dose versus emulsion concentration
should be conducted if quantitative relationships for emulsion
flocculation are to be ultimately achieved.
li. Refined kinetics studies should also be performed, especially via
the Coulter Counter in order to define effective rate equations
for this process.
5. A study of velocity gradient should be extended to include the
advantages of employing initially high velocity gradients followed
by the lower values necessary to avoid redispersion in the latter
stages of the process.
6. The velocity gradient effect should be investigated at the pilot
plant level via the use of shaft torqueometers or other such
means to estimate power dissipation versus process efficiency
and engineering kinetics.
7. Pilot plant research of concentration and compositional effects
on process efficiency should be performed.
8. Compositional effects on rolling characteristics should be studied
in order to ultimately optimize emulsion breaking efficiency with
rolling performance.
9. The use of relative streaming current devices as process control
instruments should be investigated at both the pilot plant and
full-scale plant level.
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10. Some attempt should be made to monitor and control pH during
waste emulsion flocculation at the full-scale plant level.
FIELD STUDIES
1. An extensive evaluation of the roll coolant and rinse water
requirements of the production facility should be made prior
to design of the associated treatment facility. Where possible,
the economic trade-off between production flexibility and
capital expense of the corresponding treatment facility should
be fully evaluated in the interest of economic feasibility.
2. The use of a rinse water system on the tandem cold mill if not
entirely necessary should be discouraged in the interest of
reduced water pollution, and capital and operating expense.
3. Close communication between production and treatment facilities
should be fostered and maintained.
U. Pilot plant studies should be encouraged prior to installation
of the treatment facility, especially when functioning on a
continuous process, to attempt to bridge the gap between batch-
wise bench studies and continuous full-scale.
5. Jar testing should be used as a basis for chemical requirements,
but continued reduction of chemicals at the full-scale facility
should be attempted as long as effective treatment is maintained.
6. Changes in rolling oils should prompt immediate reevaluation
of the chemical requirements.
7. Equalization capacity should be as large as is feasible. A
detention time of at least one hour is recommended.
8. pH should be continuously monitored and used to control lime
additions.
9. Centrifugal pumping of lime slurry is recommended over positive
displacement pumping.
10. The design and use of flocculation equipment should be given
further study to fully optimize the flocculation step of the
process. Pilot plant studies should be undertaken to provide
better direction on this process parameter.
11. The operation of the skimming mechanism should be performed at
a frequency and duration so as not only to minimize the removal
of unnecessary water in the scum but also to prevent the scum
from dewatering to the point of being unmanageable in gravity
flow lines.
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12. Some type of shelter should be provided for the flotation unit
to prevent destruction of the floated floe and reentraintnent
due to rainfall.
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SECTION III
INTRODUCTION
The process of cold rolling steel strip to light gauges requires
lubrication and cooling of the rolls. Oil alone is inadequate as
a coolant for a high-speed cold mill, so an oil-in-water (o/w) emul-
sion is used as a combination lubricant-coolant. Modern tandem cold
mills have recirculating coolant systems capable of flooding each
reduction stand with a large quantity of emulsion; this emulsion can
be reused for weeks or months before it deteriorates to the point
that it can no longer be used. In cases where strip cleanliness is
of major concern, once-through rinse water is used in place of emul-
sion on the first and/or the last stands of the mill. This type of
mill, then, produces two types of oily wastes: periodic dumps of
spent emulsion from the coolant system; and, once-through rinse water
from the first and/or last stands.
Commercial rolling oils are complex mixtures containing principally
fats and mineral oil, with lesser amounts of free fatty acids, emu]-
sifiers, bactericides, and sometimes rust inhibitors. Most rolling
oil formulations are trade secrets, and historically the rolling oil
user has not been successful in determining which parameters are most
critical for rolling. Indeed, even seemingly similar rolling mills
may require different oil formulations for optimum rolling efficiency.
The chemical composition of a cold mill waste is, then, not only
largely unknown; it also varies from mill to mill and as oil suppliers
are changed on the same mill.
The more concentrated oily wastes, such as from batch dumps, can be
disposed of relatively economically (contract hauling, chemical emul-
sion breaking, etc.); these are usually on the order of a few
thousands of gallons per day or less. Much more difficult is the
problem of efficient and economical treatment of rinse waters, which
are on the order of a thousand gallons per minute or more. Past
practice throughout the industry has been to send these wastes
directly to the sewer, to discharge them to public waterways through
lagoons in which oil can separate from unstable emulsions, or to
apply some sort of chemical coagulation process which usually results
in the formation of large quantities of an almost unmanageable sludge
which is, in turn, a disposal problem. The coagulation of dilute
o/w emulsions has been neither a science nor an art; little is known
of the physico-chemical aspects of efficient treatment or of the
mechanisms and kinetics involved.
An improved chemical coagulation process for treating cold mill rinse
waters was developed by the Armco research laboratories. (Appendix
A) . Coagulation is initiated with relatively small amounts of alum
and/or a cationic organic polyelectrolyte. Lime is used for pH
control. Clay and an anionic or nonionic polymer are used to promote
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growth of the oily floe. This floe is then separated from the water
by dissolved air flotation; the water is sent to the sewer and the
oil may then be destroyed by incineration or processed further to
recover the oil.
A facility was designed and constructed by Armco (Appendix B) to
apply this coagulation process to the wastes from a new five-stand
tandem cold mill at the Ashland, Kentucky Works. A project was under-
taken to optimize and demonstrate the full-scale facility and simul-
taneously to perform basic research concerning the conceptual mecha-
nisms and apparent kinetics of the coagulation and flocculation of
dilute o/w emulsions. The results of both the basic research and the
operation of the facility are described in this report.
10
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SECTION IV
BASIC RESEARCH
PRESENT THEORIES AND PRACTICES
The clarification of water containing spent rolling oil emulsion
depends on two general factors: (a) the inherent stability of the
system towards coagulation and the means of producing emulsion
destabilization, and (b) the rate of flocculation of a destabilized
system. The complete discussion of spent rolling oil emulsion waste
treatment requires, therefore, an insight into both the various
possible thermodynamic states of the system and the kinetics of
change from one such state to another. Because of the present level
of knowledge in this area of research the former was treated in a
qualitative manner whereas the latter was studied from at least a
semi-quantitative point of view.
Emulsion Stability
The "stability" of a system may be regarded as the relative tendency
of the system to remain in the dispersed state. Most rolling oil
emulsions are true colloids, exhibiting a measureable electrical
potential at the "plane of shear"! between the oil droplet and the
continuous phase. All other factors being equal, the larger the
absolute magnitude of this "zeta" potential the more stable the
system, the lower the magnitude the less stable. At very low zeta
potentials the kinetic energy of particle motion (whether solely
Brownian or under an induced velocity gradient) may exceed the
repulsive energy between particles in which case coagulation and
ultimately flocculation will occur.2
electrical potential at the "plane of shear" is referred to as the
zeta potential (Zp) and may be defined as the potential between the
continuous or solvent phase and the boundary plane which separates the
solvent adhering to a colloid in its motion and that which can move with
respect to it. Zeta potential can be determined through the measure-
ment of electrophoretic mobility. Relative surface charge has also
been estimated through relative streaming current or colloid titration
techniques(1).
^Coagulation is generally considered to mean electrical destabilization
of particles or removal of forces of repulsion without proceeding to
agglomeration or flocculation.
11
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The zeta potential depends on the state of the oil droplet-continuous
phase interface. The chemical composition of the rolling oil and of
the continuous phase ultimately determines the state of the oil-water
interface. Since these factors are the more easily determined, the
relationship of zeta potential to chemical composition is frequently
studied.
The composition of commercial rolling oils are invariably trade
secrets. It is known however that they usually contain lard oil,
mineral oil, emulsifiers and possibly such additives as bactericide,
extreme pressure lubricants, etc. A typical rolling oil composition
is given in Table I in terms of the major components(2).
The individual rolling oil components may be classed as active or
inert towards emulsification. The mineral oil portion may be
considered inactive, for example. Obviously, added emulsifiers are
active components and may run as high as k percent in some commercial
formulations. The primary active components in the lard oil portion
are the free fatty acids. Most of the lard oil portion is composed
of glyceryl esters of fatty acids. The unesterified free fatty acid
content is variable among commercially available oils(3) from as
low as about 3 to as much as 13 percent by weight of total formulation,
The added emulsifiers and the free fatty acids present at the oil
droplet surface act as surfactants, rendering the oil droplet stable
against coagulation and stabilizing the small colloid size of the
droplet through surface tension effects. These effects will vary as
the relative levels of added emulsifier and percentage of free fatty
acid present, being generally more stable as these levels increase.3
Any changes in intrinsic stability of an emulsion may be regarded
as arising from interactions and reactions at the continuous phase -
oil droplet interface. These changes may arise from numerous factors
among which are: pH, ionic strength, specific ion interaction and
temperature.
The pH of the system exerts a significant influence on oil emulsion
stability. Lowering the pH should lower the zeta potential by
suppressing ionization of the free fatty acid. This is especially
likely as the solvation radius of the proton is sufficiently small
to cause it to adsorb beneath the plane of shear of the emulsified
particle
are limits to this general rule however, as will be noted
during the experimental and discussion portions of this work.
12
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TABLE I
TYPICAL ROLLING OIL COMPOSITIONa
Non-Ionic
Lard Oil U8.10 to 52.30$
Mineral Oil -200 sees @ 100F U6.60 to 1*7.60$
Triton X-100 (Rohm & Haas) 2.00 to 3.00$
Bactericide 0.10$
Acid No. 15-19
Anionic
Lard Oil W.UO to 52.UO$
Mineral Oil -200 sees @ 100F U6.50 to U7.50$
Atlas 0-3300 (Sulfonate) 2.00 to 3.00$
Bactericide 0.10$
Acid No. 15-19
a Compliments of Vendor "A"
13
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The ionic strength of the continuous phase will also effect emulsion
stability. However, there exists little or no evidence to indicate
a direct relationship betwen electrical conductivity, per se, and
colloid stability. This situation probably arises from two con-
siderations. First, monovalent ions (other than hydrogen or lithium)
exert a real but small effect on zeta potential because of their
small charge/radius ratio. In general the effect of ionic strength
arising from monovalent ions will be primarily that of zeta potential
attenuation through reduction of the electrical capacitance in the
aqueous phase. Bivalent ions and trivalent ions exhibit a more
pronounced effect than monovalent ions on colloid stability and have
led to such semi-quantitative comparisons as the Hofmeister Series
and the Schulze-Hardy Rule(5). The hydrolysis products of multivalent
ions may interact with an emulsion droplet in other ways than through
reduction of field strength in the aqueous phase. These effects on
emulsion stability may overshadow those of ionic strength.
The nature of a multivalent cationic species may vary depending upon
the pH. At very low values it may exist as a simple hydrated ion.
At higher values it may form positively charged hydroxo complexes
and finally insoluble hydroxides. At sufficiently high pH values
amphoteric substances such as aluminum may even form negatively
charged hydroxo species in the form of anionic coordination complexes.
The interaction of these different species with colloids can vary
considerably. Thus, the pH range of flocculation is frequently critical.
It has been reported(6). for example, that the positive hydroxo com-
plexes of aluminum (pH u-5) most effectively destabilize colloids.
Hydroxo complexes of multivalent ions should also be adsorbable
through van der Waals forces as well as ionic forces, although this
has been rarely discussed. Amphoteric metals in solution also
exhibit a zero point of charge (ZPC) at which they demonstrate a
low solubility and therefore promote rapid phase separation. The
combination of these many effects has led to the rather wide usage
of iron and aluminum salts as coagulant-flocculants,
Adsorption-charge reduction effects will vary with temperature as
well as type. As would be expected the stability of an emulsion
diminishes with increasing temperature due to numerous effects such
as increased Brownian motion, increased reaction rates and decreased
surface tension.
Spent rolling oil emulsions are less stable than fresh emulsions.
Destabilizing factors include: (a) a build up in hardness in the
aqueous phase due to solvent evaporation; (b) reduction of free fatty
acid content through high pressure pyrolysis of unsaturated bond
linkages(7) and reaction with iron fines(7,8)j and (c) bacterial
degradation. Partially off-setting these influences is the dispersive
action of the mill. This latter effect will depend significantly on
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such factors as roll pressure, volume of the emulsion system, and flow
rate of the emulsion onto the mill stands. A significant portion of
spent rolling oil emulsion will exhibit "break-out" of "tramp" oil
on standing. The remaining emulsified oil may be rather poorly stable
(Of. Table II, and ref. 9).
The above mentioned factors will influence the stability of a system
towards coagulation. These same factors plus some additional ones
play the more pragmatic role of influencing the rate of flocculation
of a destabilized rolling oil emulsion system.
Flocculation Rates
Flocculation rates are influenced by both chemical and physical factors,
Chemical factors include interaction among destabilized emulsion drop-
lets, between such droplets and aggregates thereof with primary
flocculants such as insoluble iron and aluminum hydroxo species,
and so-called coagulant aids such as clay and polymer.
Insoluble positively charged iron and aluminum hydroxo species may
react ionically with negatively charged emulsion droplets. However,
van der Waals forces are also operative and independent of charge.
The magnitude of these effects depends upon pH and concentration.
Entrapment of oil by precipitating alum or iron floe may also occur.
Kinetic effects related to insoluble metal floe interaction are
largely unknown although previous work on the flocculation of kaolin
suggested first-order kinetics for removal of the primary clay
particles(10).
The role of clay as a floccuLant aid can be attributed to its ability
to adsorb charged particles. This results from the fact that clay
platelets possess positively charged edges and negatively charged
faces. Thus, clay particles can function as bridges between charged
particles. The density of clay is such that when incorporated into
the floe it may speed the sedimentation of the flocculated phase.
Particulate matter such as clay also appears to improve floe
strength and therefore retard "redispersion"(ll).
Long chain polymers are also used as flocculant aids. These polymers
may adsorb ionically and/or through van der Waals forces depending
on the nature of the particular polymer being used. Under optimum
conditions the polymer can adsorb on more than one colloid particle
such that a bridging effect is achieved. Under such conditions a
large interconnected network of microflocs can form leading to rapid
phase separation(6).
The kinetics of rolling oil emulsion flocculation has remained
largely unstudied. Slezak(9) has recently reported some significant
work along these lines for spent rolling oil emulsions. However,
-------
other detailed studies on the flocculation of oil-in-water (o/w)
emulsions have been performed on atypical raonodisperse systems at
unrealistically slow rates of flocculation(12,13).
TABLE II
GENERAL STABILITY CHARACTERISTICS
RELATIVE TO ZETA POTENTIAL3
Avg. Zeta Potential (mv.)
Maximum Agglomeration and Precipitation 0 to +3
Range of Strong Agglomeration and Precipitation +5> to -5
Threshold of Agglomeration -10 to -15
Threshold of Delicate Dispersion -16 to -30
Moderate Stability -31 to -UO
Fairly Good Stability -Ul to -60
Very Good Stability -6l to -80
Extremely Good Stability -81 to -100
a From Riddick, T. M., "Control of Colloid Stability Through Zeta
Potential" Vol I. p. 2, Livingston Publ. Co., Wynnewood, Pa. (1968)
16
-------
Comparisons with real systems should include not only the proper
pH and dosage control but also the appropriate energy input to the
system necessary to produce practical rates of flocculation. The
intrinsic rate constant for flocculation (k) is related to the initial
and instantaneous number of primary particles through the relationships
(Appendices C, D):
N = exp (-KDGNot) (batch process)
and
5 = (l+kDGNoi)m (series of continuous stirred-tank reactors)
NO m
where 0 is the velocity gradient (sec~l), D the particle size dis-
tribution coefficient, t the time of flocculation (sees), t/m the
retention time per reactor and, m the number of reactors. The
velocity gradient Q, can be estimated from a number of physical
parameters (lit) including the experimentally measured torque on a
stirrer being rotated in a reactor vessel(15>) or through consideration
of the shaft speed and paddle dimensions of an "infinitely thin"
vertical paddle-on-shaft stirrer as originally derived by Camp(l6).
These various approximations yield an estimate of 0^ the mean velocity
gradient, where
and Wra is the mean power dissipated per unit volume per unit time in
the system and u is the absolute viscosity.
Thus, approximation of Wro is necessary to estimate G^. For the
typical jar test, Wm may be estimated by the method of Camp(l6)
through the relationship:
Wra - 239 Cd (1-K)3 (n)3 Sum(A(rb)3)
where Cd and Kare constants, A the area of a paddle, i*b the distance
of the center of the paddle to the center of the shaft, n the
revolutions per second of the shaft, and V the volume of the system
being flocculated.
Consequently, flocculation kinetics can be determined experimentally
through measureable parameters. The value of N and N0 can be measured
directly by means of a Coulter Counterk(ll), estimated via number or
weight concentrations or possibly through turbidity measurements under
appropriate circumstances (as shown in this work). The value of D
is very difficult to estimate under kinetic conditions although
Harris(10), et al have done so through use of a computer and size
distribution data.
^Trademark Coulter Electronics Div.
17
-------
The flocculation rate increases with increasing work input to the
system up to such work input levels where degradation of the floe
becomes serious. Additionally, the rate constant, k, will not
necessarily remain constant beyond some point at which the primary
flocculant may change in character. Thus, optimization of the
flocculation rates in real systems will require optimization of
mean velocity gradient and reaction time as well as flocculant dosage
and pH.
Typical Commercial Processes
Historically a few general processes for emulsion breaking of o/w
emulsions have developed. These approaches provide for the destabi-
lization of the emulsion by acid, polyvalent ions or both. If
polyvalent ions are employed, the pH is generally regulated by lime
and controlled at a pH corresponding to the ZPC of the hydrous
polyvalent metal floe. The pH for the zero point of charge of '
A1C13 may be as high as 8-9 whereas for Al2(SO^)3«l8H20 it would
be significantly less (pH 6-7) because the sulfate ion is more
readily incorporated within the coordination sphere than are halide
ions. Primary flocculation with such flocculants as alum and lime
is frequently followed by clay addition in the same general pH
range. Further treatment with polymer is also often employed.
Scope of Unresolved Questions
A great deal of fundamental (and therefore practical) information is
still lacking. The areas of deficiency include:
1. The effect of rolling oil composition on emulsion stability
2. The effect of history on emulsion stability
3. The mechanism of emulsion destabilization and flocculation
under optimum dosage and pH conditions
h. The rate of flocculation
These questions have received the attention of this work, particularly
at the 1,000 ppra emulsion concentration level. Dilute spent rolling
oil wastes, normally expected to be in this range, constitute a more
serious process problem than more concentrated wastes.
18
-------
EXPERIMENTAL
The experimentation performed in this work involved the following:
1. Emulsion stability was studied in terms of zeta potential and
in a few cases via relative streaming current or colloid
titration techniques. In a number of instances particle size
distributions were also made.
Emulsion stability was examined in the following areas:
a. As a function of composition of oil and/or aqueous phase
for commercial and "typical" rolling oil emulsions;
b. As a function of mill time, ionic conductivity, pH and
dilution for mill spent commercial rolling oil emulsions.
2. Turbidities of emulsions were examined as a function of con-
centration and composition for laboratory prepared commercial
and "typical" rolling oil emulsions.
3. Jar testing was extensively performed in the following areas:
a. On a comparative basis among several laboratory prepared
commerical rolling oil emulsionsj
b. On a laboratory prepared commercial rolling oil emulsion
in terms of alum-lime addition and relative streaming
current;
c. On "typical" rolling oil emulsions as a function of pH,
chemical additions and order of addition thereof, as well
as a function of oil composition, velocity gradient and
time. (For stock solution compositions, see Table III.)
h. Jar test kinetics studies were conducted on a "typical" rolling
oil emulsion. An approximate composite rate constant was
determined. An evaluation was also made of a direct particle
counting technique for determining rate constants for flocculation
of rolling oil emulsions at practical concentrations.
Stability and Properties of Commercial Rolling Oils
The relative stability of unused and mill-spent waste commercial
rolling oils was investigated from several points of view. These
included particle size distribution analysis and zeta potential.
19
-------
The particle size distribution of several samples were ran. These
included brand "C" emulsion after 2 hours and after 2 days use on
the Ashland mill and a spent brand "A" emulsion. The "C" samples
were run within 72 hours after the last sample was taken whereas
the spent brand "B" emulsion was run about one month after sample
taking.
The two hour "C" emulsion exhibited a slight maximum in weight
percent at about 3-5 micrometers particle diameter as did the spent
month old "A" waste emulsion. The two day "C" emulsion exhibited a
shift to smaller particle sizes and a disappearance of the 3-5
micrometer maximum. As near as could be determined within the
limitations of the instrument, dilution of the 2 day used emulsion
did not measurably alter the particle size distribution. Unfor-
tunately, comparison could only be made of this effect at relatively
large particle diameters (greater than 12-15 micrometers) at which
the total weight percent of oil is very small compared to that below
about 12 micrometers particle diameter.
Stability estimates on laboratory prepared commercial rolling oil
emulsions were made primarily via zeta potential measurements. Before
receipt of the Zeta-Meter however, an evaluation of colloid titration
techniques was conducted and use was made of the Relative Streaming
Current Detector.
TABLE III
STOCK OIL EMULSION COMPOSITIONS
Stock Emulsion Cone, (ppm) A.N.a (±) Mineral Oil Fraction13
1 8000 6 0.500
2 8096 8 O.U9U
3 821*0 11 O.U85
k 8U80 16 0.1*72
5 8720 21 0.1*59
6 8960 26
a Acid number of oil
b Qnulsions contain only mineral oil and lard oil
20
-------
The colloid titration technique proposed by Kawumura and Tanaka(l)
was examined. This technique involves treatment of a colloid sample
with a positive polymer and back-titration with a negative polymer
in the presence of an indicator dye sensitive to positive and
negative colloids. The number of milliequivalents of positive
polymer necessary to neutralize a milligram of negative colloid
is referred to as the "colloid charge". To test this method a
10.0 weight percent "B" oil emulsion was prepared, and by successive
binary dilutions, emulsions of 5.00, 2.50 0.00988 weight percent
were also prepared. A 10.0 ml aliquot of 5.00 percent emulsion was
found to be neutralized by about 8.7 x 10-3 meqs of "Cat-floe" when
back-titrated with 9.U11 x 10-% PVSK and o-toluidine blue indicator.
The end point was somewhat sluggish. At concentrations lower than
about 1.00 percent oil, the results were poor even under corrected
blank conditions. Dilution of the indicator to a point where its
blank value was acceptable resulted in indiscernable end points
for emulsions less concentrated than about 0.2 weight percent.
The Relative Streaming Current Detector was also utilized for
stability estimates. Brand "B" oil emulsions in deionized water
gave relative streaming current (RSC) values of -128 for 1000 ppm,
-lOlj. for 500 ppm and -80 for 250 ppm oil concentrations. By com-
parison an "A" emulsion yielded a -3U to -1*0 value for a 500 ppm
oil concentration. Treatment with 22 ppm alum resulted in a final
RSC value of +7.5 in the latter case whereas all the "B" emulsions
remained negative even at 1050 ppm in alum. Above about 20-25 ppm
in alum the 1000, 500 and 250 ppm "B" emulsion exhibit RSC's of
about -30 to -hO which diminished only very slowly upon further
alum addition.
Various rolling oil emulsions were titrated with alum and lime with
simultaneous measurement of the RSC. Most titrations were performed
on 900 ml aliquots of 500 ppm "A" emulsions under gentle stirring
conditions. The original "A" emulsion exhibited a pH of 5.9-6.1 and
an RSC of -3U to -iiO. Upon alum and/or lime addition the RSC was
always negative for pH values greater than 6.5-7.0, positive for pH
values less than 6.5-7.0 and zero for pH values in the range 6.5-7.0.
When alum and lime were added incrementally and alternately to the
emulsion so as to avoid large variations in pH, pinpoint flocculation
would occur in the range 6.5-7.0 at a definite minimum amount of
added alum and lime. This minimum amount also corresponded closely
to that necessary to produce a visible alum floe in distilled water
at pH 6.5-7.0. In one experiment an excess of alum was added
(0.1 g/O.U5 g oil) and back-titrated with lime to produce a floe on
the acid side. In this instance again, however, the RSC was zero
at the flocculation point.
21
-------
Zeta potential values were employed throughout the bulk of this
study for comparative stability estimates. Because of the fact
that commercial rolling oils were of unknown composition, such
comparisons were made in reference to the effects of changes in
the continuous water phase.
The effects of pH, dilution and specific conductivity were examined
via zeta potential on the spent "B" rolling oil emulsion. The
emulsified oil content of the sample was 6,OCX) ppm whereas the tramp
oil content was 7,500 ppm. Samples more concentrated than about
300-500 ppm in emulsified oil were ultracentrifugated at lli,ltOO G's
for 1.5 hours or more in order to obtain samples translucent enough
for zeta potential measurement(18). Dilutions were made with
C02-free deionized water and as-received Ashland mill mix water
in order to bring out the effects of pH and ionic conductivity.
The results are plotted in Figure 1.
This spent "B" emulsion was also diluted to a given concentration
(1200 ppm or a k to 1 dilution) with C02-free deionized and mill
mix water in different proportions with subsequent zeta potential
and pH measurement. The results appear in Table IV. Pickup of
CO2 by the C02-free deionized water during the course of these
measurements affected the measured pH values slightly.
The "C" used emulsion was also investigated and found to have a
zeta potential of -U0.7 mv at "infinite dilution" (dilution to
30 ppm or less) in COg-free deionized water after two hours mill
time. The value after two days mill time was -1*5.8 mv. The specific
conductivities in these measurements were 8 and 9 micromhos, respec-
tively. The difference in the two values was due to the fact that
in the two day mill sample virtually all the emulsion droplets
exhibited the same electrophoretic mobility whereas the two hour
sample was composed of particles 70-90 percent of which exhibited
this same electrophoretic mobility while 10-30 percent of which
moved more slowly under the 200 V. applied potential.
TABLE IV
VARIATION IN ZETA POTENTIAL OF 6,000 PPM SPENT MILL
EMULSION DILUTED WITH MILL MIX AND/OR 002-FREE DEIONIZED WATER
Sample Emulsion Mill Water/ Zp
No. Cone, (ppm) Deionized Water S.C, p_H (mv.)
1 1200 Inf. 0.590K 6.UO -2U.1
2 1200 1 O.U70K 6.25 -25.0
3 1200 0.33 0.295K 5.90 -27.7
h 1200 0 0.205K 6.10 -29.2
22
-------
FIG. 1
VARIATION IN ZETA POTENTIAL AND pH
OF SPENT MILL EMULSION WITH DILUTION
O Zp after dilution with mill mix water (rav.)
OO Zp after dilution with C02 free deionized water (nv.)
A pH at given dilution with mill mix water
-20
6000
Oil Concentration, ppm
23
-------
A spent "A" rolling oil emulsion was also examined. After filtra-
tion over a 0.8 micrometer Millipore filter pad a 3,100 ppm emulsion
was obtained. The zeta potential of the system was approximately
-2h mv. Dilution with mill mix water in binary steps to 39 ppm
failed to produce any significant change in zeta potential.
A sample of this same system diluted to 300-UOO ppm with mill mix
water was treated with dilute alum and lime. Upon the addition of
alum alone the zeta potential of the emulsion increased from -2k mv.
at pH 6.96 to +8.3 mv. at pH U.90. Back treatment with lime to
pH 10.3 yielded a zeta potential of -18.5 mv. Upon adding more
alum and lime so as to form an alum floe and subsequently filtering
over a 0.8 micrometer Millipore filter, a zeta potential of -21 mv.
was observed on the few emulsion particles remaining in the system
at pH 7.03. More alum was added to this treated and filtered
system so as to reduce the pH to U.$7. The remaining particles
were observed to be mostly immobile with a relative few having a
zeta potential of -3.8 mv. with fewer still having a very small
positive zeta potential.
The effect of polyvalent ion concentration on zeta potential was
also investigated by treating another filtered and diluted "A" spent
emulsion with various amounts of alum. The concentration of the
filtered spent emulsion was 7,100 ppm and the sample tested was
6h5> ppm after dilution with deionized water. In a first series of
tests the alum to oil ratio was varied by cumulative addition of
alum to the original emulsion sample. A second series was also run
in which the oil content was 260 ppm and each sample of emulsion
plus alum was prepared separately and subjected to zeta potential
measurement immediately after preparation. These results appear in
Figure 2 wherein the ratio of the fractional change in zeta potential
after alum addition is plotted versus the alum to oil ratio.
A "B" emulsion was also treated with alum. This emulsion did not
exhibit a positive zeta potential in the presence of even large
excesses of alum.
The turbidity of another laboratory prepared "B" rolling oil emulsion
was determined. The as-prepared 9k2 ppm emulsion was then diluted
with various amounts of deionized water and the resultant turbidities
measured. The relationship between concentration and turbidity was
found to be:
0.862
InCd
-------
FIG. 2 VARIATION IN ZETA POTENTIAL
WITH ALUM ADDITION
1.0
0.5
•rl
0.1
O 011:645 ppm Spent Emulsion
O 011:260 ppm Spent Emulsion
I » I I I 1 I
J_
1.0 0.5 0.1 n 0.05
Millequivalents ALUM/G. OIL
25
-------
where:
JTUd = turbidity of diluted emulsion in JTU
Cd = concentration of diluted emulsion in ppm
JTU0 - turbidity as-prepared emulsion
GI = concentration of as-prepared emulsion
Jar testing was performed on laboratory prepared commercial rolling
oils early in the grant period before the receipt of the Zeta-Meter
or the turbidimeter. In these cases "relative clarity" was used as
a criteria rather than turbidity. A pH of flocculation (6.5-7.0) was
selected on the basis of current practice. The results of comparative
jar testing results among the various commercial rolling oils is
given in Table V for alum, lime, clay and polymer additions. In
each instance alum was added first followed by lime addition.
Stirring times of 5.0 minutes were employed throughout. The zeta
potentials at "infinite dilution" given in Table V were determined
considerably later from freshly prepared emulsions.
Qnulsion Stability and Properties of "Typical" Rolling Oils
The effect of chemical composition on resultant zeta potential was
studied for synthetic "typical" rolling oils having the mineral oil
fractions given in Table III.
The effect of emulsifier content on synthetic "typical'1 rolling oils
was studied on AN11 and AN16 "typical" oils. The oils in question
contained only mineral oil and lard oil in the ratios 0.970 and
O.ylik for the AN11 and AN16 cases, respectively. The emulsions
were prepared from chloroform standard solutions at 850 ppm in
deionized water with emulsification via the Waring Blendor for
60 i 1 sees. Each of these emulsions was then treated with various
amounts of Rohm and Haas Triton X-100 (a nonionic polyethoxylated
alkyl phenol emulsifier) and the zeta potential of each resultant
system measured. In the AN16 case the oil concentration diminished
(as a result of dilution by added emulsifier) from 850 ppm to 607 ppm.
In the AN11 case the concentration was maintained at 699 ppm except
at the 0.613 and 0.805 emulsifier to emulsion ratio levels by means
of deionized water adduncts. In each case the systems were stirred
gently after each addition. These data are presented in Figure 3.
The zeta potential of "typical" oil emulsions of different oil acid
numbers were correspondingly investigated. These included emulsions
with and without added emulsifier (Triton X-100 at emulsifier to oil
ratio 0.021*2) at similar oil in water concentrations. These results
appear in Table VI.
26
-------
TABLE V
ro
(S
•H
O
"B"
"D"
"C"
"A"
"B"
"D"
"C"
"A"
"B"
"D"
"C"
"A"
JAR TESTING
16
3>5
"""* d •
• (DO
a> C C
E HO
N « — • fa O
-35.6 138
-37.7 138
-U3.7 138
-32.5 138
13U
13U
13U
13U
138
138
138
138
COMMERCIAL ROLLING OIL EMULSIONS WITH ALUM, LIME,
Alum
51
'^^
a o
38
fc 0
56
56
56
56
56
56
56
56
56
56
56
56
, Lime Addition
f
o
-------
FIG. 3. VARIATION IN ZETA POTENTIAL WITH
ADDED EMULSIFIER (TRITON X-100)
10°
10-1
lo-2
Original Oil Cone:850 ppm
10-3
-I
.
-30
Zp (mv.)
-70
-80
28
-------
TABLE VI
JVJ
COMPARISON OF ROLLING OIL
EMULSION OF DIFFERENT ACID NUMBERS
Etnulsifer Free
Cone, (ppra)
800
712
825
850
872
Zp (rav. )
-U5.7
-U9.U
-55.8
-83.9
-58.6
Turbidity
585
565
285
285
325
Containing BnuLsifier3
(JTU) Cone, (ppm)
800
800
800
800
800
Zp (mv. )
-37.6
-U9.5
-55.7
-U6.6
-1*3. U
Turbidity (JTU)
270
2UO
210
2UO
250
AN(±1)
6
8
11
16
21
26
a Triton X-100 emxasifier/oil = 0.02U2 in all cases
-------
The effect of emulsifier content on particle size distribution was
also studied. Particle size distributions were made on AN16, 81;8 ppm
"typical" rolling oil emulsions treated with various amounts of
Triton X-100 and dispersed for 120 i 1 seconds via a Waring Blendor.
These particle size distribution measurements revealed that the
higher the emulsifier to oil ratio the smaller the mean particle
diameter up to a ratio of 0.121. At a ratio of 0.6?7, however,
the system exhibited a particle size distribution shifted toward
larger diameters. It was also observed that creaming did not occur
in the 0.121 ratio case even after two weeks shelf aging whereas in
each of the other cases creaming was pronounced within less than
2k hours.
A comparison of particle size distributions resulting from different
emulsifier types was also made. An acid number 16 oil emulsion was
prepared with Atlas G-3300 (an anionic sulfonate emulsifier) for
example, instead of Triton X-100 at a comparative emulsifier to oil
ratio of 0.02lil and the particle size distribution determined. No
difference in particle size distribution could be detected between
this emulsion and an AN16 nonionically emulsified oil. Careful
analysis revealed, however, that about 75 percent by weight of
both emulsions was composed of particle sizes below the lower
limit of instrumental detection.
Another "typical" emulsion, identical save the fact that it con-
tained no emulsifier, yielded a particle size distribution measureably
shifted to larger particle diameters.
The turbidity of "typical" rolling oil emulsions were investigated
in reference to emulsifier content. In these experiments the
concentration of an ANl6 rolling oil was maintained constant at
81i8 ppm and the emulsifier content varied. This was achieved by
dispersing a 50.0 ml aliquot of 8,U80 ppm oil emulsion containing
no added emulsion in Ui>0 ml deionized water containing various
amounts of Triton X-100 via a Waring Blendor for 120 i 1 seconds.
The results appear in Figure k. Dilution of the resultant 81i8 ppm
emulsions to U2lj and 212 ppm yielded the same relative turbidity
distributions.
Because of the peculiarity in the turbidity distribution with
emulsifier content, a comparison was made of the ultraviolet
adsorption characteristics of the pure oil, the deionized water,
the emulsifier in water and an emulsified oil. It was found that
the emulsifier absorbed increasingly strongly as the wavelength
was decreased below UOO millimicrometers whereas the oil exhibited
essentially constant absorption characteristics in this same range.
The acid number and concentration effects on the turbidity of as-
prepared emulsifier-free "typical" rolling oil emulsions are presented
in Tables VII and VIII along with the values for turbidity after
breaking the emulsions with the indicated dosages of alum and lime.
30
-------
FIG. 4 VARIATION IN TURBIDITY OF "TYPICAL" ROLLING
OIL EMULSION WITH TRITON X-100 EMULSIFIES
Oil: AN16, 848 ppm
600
500
400
300
200
10-3
10'
,-2
10'1
10C
EMULSIFIER/OIL
31
-------
TABLE VII
INITIAL COMPARATIVE JAR
sec -
RESULTS
ro
Sample
No.
Blank
1
2
1*
5
6
7
8
9
10
11
12
13
11*
AN(a)
6
8
11
16
21
26
6
8
11
16
21
26
6
16
Oil Cone.
(ppm)
.
5oo
506
515
530
51*5
560
1000
1012
1030
1060
1090
1120
2000
2120
As-prepared
Turbidity
(JTU)
0.2
180
200
225
2UO
285
310
310
31*0
390
1*70
510
51*5
525*
680*
250 ppm
alum
pHfD
5.1*
5.1*
5.1*
5.3
5.5
5.2
5.3
5.3
5.2
5.1*
5.2
5.2
5.2
d
d
+ lime
JTUf D, c
.
—
-.
226
28
1*3
1*1*
1*
8
9
27
1*2
53
21
. 53
250
alum -i
pHf
7.1*
_
_
_,
_
-
_
7.6
«
_
7.5
_
7.5
•
ppm
•• lime
JTUf
_
_
_
_
_
M
13
_
13
22
IK
-
250 ppm
alum + lime
8.2
8.1
8.1
7.9
8.3
8.3
11*
18
12
13
23
a At 2000 ppm and above the measured turbidity values are imprecise and may even decrease with
increasing concentration as a nephelometric method was employed in these studies.
b pHf, JTUf are pH and turbidity values after flocculation.
c All turbidity values taken one hour after flocculation.
d pH meter became defective.
e Turbidity not attainable due to dispersion of large floe particles.
-------
TABLE VIII
JAR TEST RESULTS FOR CONCENTRATION VARIATION
Dosage: Alum, 125 ppm; Lime, 25 ppm
As-prepared Nature of
Cone., ppm AN(+1) pH orig.a Turbidity (JTUj) pHf& JTUfb Floe
Uo
103
206
515
1030
2060
11
11
11
11
11
11
6.2
6.2
6.2
6.2
6.2
6.2
25
5U
90
230
-
_
6.0
6.0
5.7
5.8
6.0
5.9
M^H^H^^^^
2
6
15
30
100
250
Settled
Settled
Settled
Suspended
Suspended-
Floated
Floated
a The pH values reported are questionable as the pH meter was found to
be functioning erroneously shortly after these tests. A probable
final pH of 5.0 to 5.5 is assumed on the basis of the alum and lime
dosages.
b Measured post-flocculation turbidity in JTU's.
33
-------
An empirical relationship among acid number, concentration and
turbidity of these as-prepared emulsifier-free emulsions was
established as:
JTUi* - (Ci)0.7h7 (AN) 0-311
where:
JTUi* " calculated turbidity value expressed in Jackson Turbidity
Units
GI = as-prepared oil concentration of the emulsion in ppra
AN * acid number of the oil
These calculated and measured turbidity values are presented in
Table IX. Also included in Table IX are turbidity values of some
of these emulsions after being flocculated with 250 ppm added alum
and enough lime to produce a final pH of about 5.2. Flocculation
was performed for five minutes under a mean velocity gradient of
l!4i sec.-l. An empirical relationship was established among acid
number, concentration, as-prepared turbidity and post-flocculation
turbidity (under these specific conditions) as:
JTUf* - 1.1 x 10-3 JTUi* (AN)1'1*0
JTUf* - 1.1 x 10-3 (Ci)°«7k7 (AN)1.72
where:
JTUf* - calculated turbidity of the emulsion after flocculation
In Figure 5 is presented a plot of turbidity ratio versus concen-
tration ratio for a "typical" rolling oil emulsion system. The
various concentrations were produced by dilution of an as-prepared
1,060 ppm AN16 rolling oil emulsion having an emulsifier (Triton X-100)
to oil ratio of 0.021*2. The slope of the resulting curve was found
to be such that:
0.850 lr£l «
A good deal of experimentation was conducted in reference to the
jar testing of "typical" rolling oil emulsions of various compositions.
These emulsions were prepared in deionized water by dilution of stock
emulsions whose compositions arc given in Table III for emulsifier
free systems. Systems containing nonionic or anionic emulsifiers
were prepared from stock emulsions having the same mineral oil
fractions as given in Table III but additionally containing added
emulsifier.
or
-------
TABLE IX
COMPAEISON OF MEASURED AND CALCULATED TURBIDITIES*
Oil Cone.
(ppm)
500
506
515
530
5U5
560
1000
1012
1030
1060
1090
1120
2000
2120
22UO
Uo
103
206
515
AN(±1)
6
8
11
16
21
26
6
8
11
16
21
26
6
16
26
11
11
11
11
JTUi (meas.)
180
200
225
2liO
285
310
310
3UO
390
U70
510
5U5
525
680
_
25
5U
90
230
JTUi*
(calc.)
181
200
22k
257
285
311
30U
335
375
U31
U79
522
. 510
723
876
33
67
113
22k
JTUf
(me as. )
_
_
28
U3
Ui
I*. 2
8.2
9
27
U2
53
21
53
1145
«•
—
.
mm
JTUf*
(calc. )
.
_
_
Hi
23
3U
U
7
12
2U
38
57
8
Uo
95
_
*»
-
-
a For as-prepared emulsions and emulsion flocculated at G^ =
sec-1 for 5.0 i 0.02 minutes via dosage with 250 ppm alum, 1*9 ppn
lime.
35
-------
u*
i r
I i i 111 I i i | iT
FIG. 5. FRACTIONAL CHANGE IN TURBIDITT WITH DILUTION
(Ci - 1060 ppm; JTUO - 295)
Oil: AN16; Triton X-100/oil - Q.Q2U2
i I I i i i l
Cd/Ci
I
I I I I I i ll
$ -
•-»
I I I I I 1 I
0.01
0.1
0.5
0.1
0.05
0.01
1.0
-------
The jar test results for emulsifier-free oils is given in Tables VII
and VIII for various emulsion concentrations, alum-lime dosages
and oil acid numbers.
An extensive jar testing program was also conducted on 1060 ppm,
AN16 "typical" oil emulsions containing added emulsifiers. (The
emulsifier to oil ratios were 0.02^1 for anionic Atlas G-300 and
0.02U2 for nonionic Triton X-100). In each case 125 ppm alunP was
added first, immediately followed by various amounts of lime and
5.0 minutes stirring under a mean velocity gradient of 59.6 sec-1.
Following this treatment the system was treated with 21-25 ppm
montmorillonite clay and again subjected to 5.0 minutes stirring
at a Gm of 59.6 sec-1.
In addition to the jar testing of the oil emulsions with alum, lime,
and clay, testing was also performed on the individual components of
the emulsion breaking process. That is, the alum-lime and alum-lime-
clay reactions were examined in the absence of oil. The data for all
these various tests are given in Table X in which the pH, specific
conductivity, turbidity and zeta potential are reported in reference
to alum-lime dose. A graphical summary of these data (except
specific conductivity) appears in Figures 6-8.
The optimum dose for equivalent flocculation of an AN16 nonionically
emulsified oil at 530 ppm was found to be 95 ppm in alum in comparison
with the 119 ppm required for the same emulsion at 1060 ppm in oil.
This experimentation was performed via successive approximation at
the optimum flocculation pH of 6.0 i O.U. Successive approximation
techniques also were used in establishing that only about 5 ppm of
alum was necessary to neutralize the zeta potential of a 500 ppm
system of this oil.
The effect of order of addition of alum and lime on post-flocculation
turbidity of emulsifier containing emulsions was examined over the
optimum pH range determined above. These results appear in Table XI.
These runs were also conducted at a mean velocity gradient of 59.6
sec"1 for the usual 5.0 minutes stirring time.
The effects of velocity gradient and acid number on post-flocculation
turbidity were also studied. In these cases only nonionic emulsifier
was employed at the usual 0.02U2 emulsifier to oil ratio. These
results appear in Table XII. Use of a Waring Blendor for producing
high velocity gradients proved beneficial if employed for about 20-30
sees before transfer to the jar test rig.
minimum alum-lime dose for acceptable post-flocculation tur-
bidities under the experimental conditions employed.
37
-------
TABLE X
pH OPTIMIZATION JAR TEST RESULTS
CD
Final Cone
(ppm)
Alum/Lime
123/0
122A1.6
120/23.0
119/3U.1
n8/ltU.9
116/55.5
115/65.8
110/105
.
pH i
It. 32
lt.62
U.80
5.58
6.55
6.75
7.35
9.62
Alum
S.C.d
195
190
190
210
255
290
355
U22
+ Idmea
Zpe
+16.6
+17.3
+23.1
+23.7
+20.9
+13.2
-29.lt
JTO
O.llt
1.1
2.8
7.1
3.7
lt.l
U.O
1.1
Alum + Lime + Clayb
pH
5.00
5.26
5.U5
6.55
6.70
_
7.U5
9.18
s.c.
175
160
175
200
270
_
365
360
Zp
+13.0
+16.9
+16.5
+10.1*
+ 8.36
_
+ 6.11
- 8.U7
JTU
0.8
2.1
2.7
3.U
U. 5
_
5.U
27
Alum + Lime
pH
lt.37
U.U5
It. 81
5.75
7.05
—
8.30
9.62
S.C.
260
210
205
220
265
-
315
U60
+ Anionic Oila
Zp
-16.1
-23.7
+16.3
+10.$
-51.2
_
-U9.6
-13.3
JTU
360
325
70
32
165
_
3U5
370
Alum + Lime +
Anionic Oil + Clayb
Alum + Lime +
Nonionic Oilc
Alum + Lime +
Nonionic Oil + Clay
123/0
122/11.6
120/23.0
119/3U.1
U8/ltit.9
116/55.5
115/65.8
110/105
it.lto
It. 72
5.05
6.12
7.38
8.U2
_
9.95
235
205
195
225
270
305
_
1*05
-12.8
-11.2
+ 6.U5
-llt.lt
-U2.5
-33.3
_
-35.1
220
1U5
Ca.1-2
Ca.6.5
35
220
_
230
It. 30
U.55
1».83
5.80
6.23
7.15
_
9.1*2
225
202
200
2liO
275
360
_
380
+ 7.78
+12.6
+18.3
+12.2
-3U.9
-3lt.2
.
-31.3
390
320
375
73
170
2UO
_
31*0
U.25
It. 72
5.02
5.98
6.65
7.30
_
9.72
200
198
198
230
270
370
_
3U5
+ 9.39 170
+ 8.18 200
+11.1 26
-16.U Ca.10-20
-19.7 Ca.29
-22.5 72
_ *.
-25.0 265
a Alum always added first followed by lime and system stirred 5.00i 0.02 min. @ On, = 59.6
sec"1
D Same as a except after first 5 min. clay added (21-25 ppm) and system stirred 5.00 i 0.02
more
c Same procedure as in a but 2.58£ Triton X-100 nonionic emulsified oil was employed in this
case whereas 2.575? 0-3300 anionic eraulsifier employed in a.
^ Specific conductivity in micromhos
e Zeta potential (millivolts)
f Post-flocculation turbidity in JTU's
-------
100
FIG. 6. JAR TESTING OF ALUM, LIME, AND CLAY
original solution :/rf)5 ml 123 ppm alum
Clay added:25 ppn
\ -
—-O Zp after line addition \
O— Zp after subsequent clay addition
/TU after liae addition
v JTU after subsequent clay adoisaon
I I I I I I I I I I I I I I I I I I I I I I I I I
-------
FIG. 7. JAR TESTINS OF "TYPICAL" MONIONIC EMULSION
+15
°xj
45
&
-15
-25
100
Oil:ANl6, Triton I-100/oil = 0.0242; oil cone.
1060 ppm initially, 1008 ppm at aluaAime = 119ppn/34.1 ppm
Alum: 123 ppm added initially
Clay: 25 ppm added initially
Zp after alum, line addition
Zp after subsequent clay addition
JTU after alua, lime addition
JTU after subsequent clay addition
—--0--—
" O-..
\
\
T
\ '
k.k
5.2
i I I I I / . f / I if i I > f f If \ \ | | M I / / I I I I I I I I /
6.0 6.8 7.6 B.k
10
9.2
-------
FIG. 8. JAR TESTING OF "TYPICAL" AHIONIC EMULSION
100
Oil:ANl6; Atlas G-3300/oil = 0.02/a;
oil cone. 1060 ppm originally,
1008 ppm at alula/lime - 119 ppm/3^.1 ppm
Alum: 123 ppm added initially
Clay: 25 ppm added initially
Zp after alum, lime addition
Zp after subsequent clay addition
JTU after alum, lime addition
JTU after subsequent clay addition
l+.U
-------
TABLE XI
THE EFFECT OF ORDER OF ADDITION OF ALUM AND LIME DURING JAR TESTING
("Typical" AN16 rolling oil 2.57-2.58$ in emulsifierj emulsion is
1060 ppm initially; vel. grad., G^ = 59.6 sec-1, reaction time,
t - 300 sees)
Final
Run AlumAiroe Dose Order ofb
No. Emul.a (ppm/ppm) Addition
1 a 122/11.6 a-1
2 a 120/23.0 a-1
3 a 119/3U.1 a-1
h a 118M.9 a-1
5 n 122 Al. 6 a-1
6 n 120/23.0 a-1
7 n 119/3U.1 a-1
8 n 122M.9 a-1
9 a 122/11.6 al
10 a 120/23.0 al
11 a 119/3U.1 al
12 a 118M.9 al
13 n 122/11.6 al
lU n 120/23.0 al
15 n 119/3li.l al
16 n 118/UU.9 al
17 a 122/11.6 1-a
18 a 120/23.0 1-a
19 a 119/3U.1 1-a
20 a 118M.9 1-a
21 n 120/23.0 1-a
22 n 119/3U.1 1-a
23 n 118/UU.9 1-a
k.\£
U.8l
5.75
7.05
1|.55
U.83
5.80
6.23
U.63
1*.91
5.78
6.U3
U.60
U.98
5.90
6.60
U.55
5.02
6.25
6. la
li.95
6.10
6.88
Final
Turb.(JTU)
325
70
32
165
320
375
73
170
280
59
50
290
180
21
120
3UO
390
310
160
1U5
105
195
* Bntasifiers: a « anionic Atlas 0-3300; n = nonionic Triton X-100
b Order of Addition: a-1 - alum followed by limej al « alum and
lime added simultaneously; 1-a » lime followed by alum
-------
TABLE XII
RESIDUAL TURBIDITY VS. VELOCITY GRADIENT AND ACID NUMBER
(Final alumtlime - 119 ppm:3h.l ppmj reaction time 300 sees; oil
composition; mineral oil:lard oil - O.li5 to 0.50j eraulsifier:
oil - 0.02U2)
AN(il)
6
11
16
26
Final Oil
Cone.
(ppm)
952
980
1008
1066
*0tn « 3
95
75
220
Residual
aOm » 27.7
67.5
62.5
80
220
Turbidity
Qm " 59.6
3^a J^oD
U?a U6b
195a 202°
aOm - lUi
21
32.5
25.5
155
a Alum added first followed by lime
0 Alton and lime added simultaneously
-------
Flocculation Kinetics
Flocculation kinetics were conducted on a "typical" rolling oil
emulsion (AN16, l,060ppm, Triton X-100 to oil ratio 0. 02^2). The
oil: alum: lime ratio was 1,008 ppm: 119 ppra:3U.l ppm. Alum and lime
were added simultaneously at pH 6.0 1 0.2 in all cases. The mean
velocity gradients for the three studies were 27.7, 59.6 and ll^
sec'1. Samples were taken at various times and the turbidities
measured (Cf. Figure 9). The data were analyzed by means of the
relationship
1.18 ine = -kDNoGt (Appendix C)
from which a value of k* = kD = 1.3 x 10-7 (mg/l)"1 was obtained.
Additional Studies
A number of additional studies or calculations were made which were
of such a nature as to warrant presentation in the Appendices. These
areas include:
1. Derivation of the relationship between mean velocity gradient
and paddle tip speed, and presentation of a working relationship
for determination of the mean velocity gradient employed in this
work (Appendix C);
2. Presentation of the kinetic relationship for sequential reactors
(Appendix D);
3. Demonstration of the feasibility of using the single channel
Model B Coulter Counter for flocculation kinetics studies
(Appendix E).
DISCUSSION
The scope of this work included investigation of the stability of
rolling oil emulsions at low concentration (generally 1,000ppm or
less) and the rate of flocculation of such rolling oil emulsions.
Several kinds of rolling oils were used for this purpose. "Typical"
rolling oil compositions were examined as well as laboratory pre-
pared commercial rolling oil formulations and mill-used and mill-
spent emulsions. These different systems were investigated in order
to gain an insight into what produces a stable emulsion, what will
destabilize an emulsion, and finally how rapidly and by what mechanisms
emulsions are flocculated.
hk
-------
PIG. 9. KINETICS OF FLOCCULATION: FRACTIONAL
RESIDUAL TURBIDITY vs. FLOCCBLWION TUB
AT VARIOUS VELOCm GRADIENTS
Oil Cone,
Alua
LiHMJ
pH
-1008 pp.
- 119 PPM
- 34.1 pp.
- 6.0 + 0.2
Triton X-100/Oil - 0.0242
Oil AN - 16
0.1
C.K27.7
G.-59.6
'1
•ec
V 0^-144,
•ec'1
0.0
120
240 360
TDD, SBCS.
480
-------
The Effects of Rolling Oil Composition and Concentration
The commercial rolling oils and. the "typical" rolling oils studied
are primarily composed of mineral oil, lard oil with associated free
fatty acids, and added emulsifiers. Commercial oils may additionally
contain bactericide, rust inhibitors, extreme pressure lubricants,
etc., but the total content of these components is small and their
effect on emulsion stability is probably not of major significance.
As a result, the evaluation of chemical composition effects on emulsion
stability in this work was concerned primarily with variations in acid
number, emulsifier content, and emulsifier type. The oil fraction was
Ii5>-5>0 percent mineral oil and f>0-££ percent lard oil in all cases.
In Table VI are compared "typical" rolling oil emulsions of different
acid number, both emulsifier-free or having an emulsifier (Triton X-100)
to oil ratio of 0.02^2. It is significant that at AN16 a maximum in
zeta potential and near minimum in turbidity were observed in the
emulsifier-free system. In the emulsifier containing system, acid
number had little effect on turbidity. The lower turbidity values
for the emulsified oils is consistent with observations made relative
to the ultraviolet absorption characteristics of the emulsifier
(Cf. Figure U). That the zeta potential is attenuated by added
emulsifier can be explained by the fact that the Triton X-100
nonionic surfactant probably shifts the plane of shear further from
the oil droplet surface. As a corollary, the emulsifier should also
serve to diminish the ease of charge neutralization.
In Figure 3 it is shown that for the AN11 and AN16 cases studied
there is a significant change in zeta potential over an emulsifier
to oil ratio range of 0.1 to O.lu Such an observation would be
consistent with monolayer adsorption of the emulsifier on the oil
droplets. Monolayer adsorption is also indicated by the fact that
the particle size distribution is minimal at an emulsifier to oil
ratio of 0.121 compared to either smaller or significantly higher
ratios. The creaming of the emulsions was also a minimum at the 0.121
ratio. Thus, it would seem that typical emulsifier contents are well
less than that necessary for monolayer adsorption but sufficient to
exert an influence on emulsion droplet stability as indicated by both
zeta potential and particle size distribution measurements.
The zeta potentials of emulsifier containing emulsions in Table VI
are lower than those reported in Figure 3 for equivalent emulsifier
to oil ratios. The only difference in the two cases appears to be
in the method of preparation. The emulsifier was added with gentle
stirring to the emulsions given in Figure 3 whereas in Table VI the
emulsifier was added to the oil before emulsification.
There may also exist differences in emulsion stability as a function
of emulsifier type. For example, it was mentioned in the experimental
section that alum addition to an "A" emulsion could lead to positive
-------
relative streaming current values whereas such an effect was not
observed for a "B" emulsion. This might be explained by the fact
that the "B" oil contains some anionic emulsifier whereas the
several rolling oils familiar to us (about 3-lt) contain only
nonionic emulsifiers.
The'concentration and nature of the particular aluminum species
present in the continuous phase will determine the extent of charge
reduction or reversal observed. These species may well have positive
charges below pH 5.5-6.0. On the other hand, the free fatty acid
molecules exposed at the surface of the oil droplet can ionize to
but a single negatively charge state. Thus, charge reversal in
nonionically emulsified oils may occur by incomplete polyvalent
ion salt formation. Further, the aluminum-fatty acid soaps are
quite stable and only small doses of alum should be required for
this reaction to go to completion. This conclusion is supported
by the fact that only about 5 ppm alum was found to be required to
bring about the complete charge neutralization of a UOO ppm non-
ionically emulsified oil whereas an excess of alum produced emulsion
droplets of a positive zeta potential.
An anionically emulsified system may readily present another case,
however. Firstly, one would not expect an anionic emulsifier to
be replaced by the complex aluminum species (any more than would be
a nonionic emulsifier) because the van der Waals forces holding such
a molecule to the oil droplet surface would be fairly strong. Although
the complex positive aluminum species can react with exposed free fatty
acid, no such "neutralization" of the anionic emulsifier will occur
because most of these materials are soluble salts of sulfonic acids
of large pK values. Thus, for anionically emulsified systems the
zeta potential may diminish to a certain minimum value but remain
negative.
These assumptions concur with the fact that neither a brand "B" nor
a "typical" anionically emulsified oil would yield positive zeta
potentials even with excessive alum additions. It is worth noting,
that in terms of particle size distribution no sensible distinction
was evident between the typical anionically emulsified and non-
ionically emulsified systems. Thus, while the emulsification
properties of these two types of emulsifier are similar, their
response towards alum is distinctly different^.
^Further discussion on the effect of emulsifier type on jar testing
is presented later.
-------
The effect of acid number on jar test results appears to be signifi-
cant. This is amply illustrated in Tables VII, VIII, and DC. The
comparative jar tests (Tables VII, VIII) clearly show a relationship
between the acid number of an emulsifier-free oil and the relative
degree of post-flocculation turbidity. The reasonably good comparison
between the measured post-flocculation turbidity and the calculated
values is surprising.
Recalling that this relationship was:
JTUf* - 1.1 x 10-3 (JTUi)(AN)1«^0
it is evident that acid number plays a significant role in floe-
culation. The detrimental effect of increasing acid number of
emulsifier-free emulsions on their degree of flocculation is postulated
as being due primarily to the higher positive charge the emulsion
particle can assume upon reaction with alum (alum was added before
lime in these tests).
The effects of acid number on the post-flocculation turbidity of
emulsifier containing systems was found to be significant only at
high acid number values. In Table XII, for example, the post-
flocculation turbidities of nonionically emulsified "typical"
rolling oils is seen to vary only slightly for oils of AN's 6-16.
At AN26, however, the post-flocculation turbidities of the identically
treated systems was severalfold higher. Thus at least to AN16 the
acid number effect appears overshadowed by the effect of the emulsifier.
Whether the effect at AN26 is due to the acid number or is a consequence
of an anomalously stable emulsion is not known.
The effects of concentration on post-flocculation turbidity was found
to be somewhat linear for emulsifier-free systems* For example, the
post-flocculation turbidities of emulsifier-free AN11 emulsions at
various concentrations (Table VIII) were found to be such that
In JTUf - $/h In Cj. (approximately)
Unfortunately time did not permit a similar study on emulsifier con-
taining emulsions at optimum dosages and pH.
Additionally the effects of lard oil, mineral oil and emulsifier
contents could not be studied. Such studies at the pilot plant level
may assist in optimizing emulsion breaking efficiency with rolling
performance. It is known, for example, that for a given rolling
operation emulsions can be too "tight" as well as too "loose" for
good performance.
-------
The Effects of pH and Clay
The extensive jar testing program on 1,060 ppm emulsifier containing
"typical" rolling oil was undertaken because the pH effect was essen-
tially unknown. The data given in Figures 7 and 8 would indicate that
the optimum pH of flocculation is somewhat lower than in usual commercial
practice. The reason for this anomalous behavior is unknown.
These jar test results (Figures 6-8 and Table X) provide an insight
into the mechanism of flocculation of rolling oils by alum floe, however.
It is significant, for example, that the maximum in alum floe turbidity
and the minimum in broken oil turbidity occur in the same pH range.
This is consistent with a flocculation model requiring the coagulated
(or destabilized) oil droplets to adsorb on or be occluded in the
alum floe.
The maximum clarification of both the nonionically and anionically
emulsified systems by alum and lime occurs over the pH range where
the microscopically observable system is approaching zero charge
although still positive?. The same analogy holds true for the alum
and lime plus clay situation. The only significant distinction in
the latter case is that the alum-lime-clay floe is less positive than
the alum-lime floe. This would not be unexpected, however, as any clay
adsorbed on the alum floe would tend to attenuate the zeta potential.
The anionically emulsified oil appears to be more efficiently broken
than does the nonionically emulsified oil, however, in spite of their
essentially identical histories, concentrations, particle size
distributions, and as-prepared turbidities (300 vs. 2?5, resp.).
In the lower pH ranges (less than U.5>) where the zeta potential
reflects the charge on the emulsion droplets, the zeta potentials
are negative in the anionically emulsified system but positive in
the nonionically emulsified system. These differences are believed
to be related to both the nature of the various aluminum species
present and their subsequent reaction on the emulsion droplets.
The nature of the aluminum species varies as the pH (see below).
zeta potential measurements for the oil + alum + lime and oil +
alum + lime + clay "systems" probably reflects more the state of the
hydrous aluminum oxide floe and associated oil and clay than any
remaining free oil or clay at pH values above about U.6 to U.8.
This is due to the fact that unflocced oil and clay is of near
micron to submicron size whereas the alum floe is considerably
larger and easier to detect.
-------
In the absence of alum the pH appears to have less of an effect on
emulsion stability than specific conductivity at moderate pH values.
This is demonstrated by reference to Figure 1 and Table IV in which
the zeta potential of a mill spent "B" emulsion was examined in terms
of both pH (over the range 5.8 to 7.7) and specific conductivity.
Considering that only very small amounts of alum are necessary to
neutralize the zeta potential of an "A" emulsion whereas dosages of
125 ppm were required for flocculation would indicate a dominate role
for the alum floe during flocculation. This effect also serves to
explain why an optimum alum dose of 100 ppm was required to floe a
$30 ppm "typical" emulsions whereas the same emulsion at 1,060 ppm
required only 125 ppm alum to achieve the same post-flocculation
turbidity. These differences are also believed to be related to both
the nature of the various aluminum species present and their subsequent
reaction with the emulsion droplets containing the different eraulsifiers,
Proposed Mechanism of Flocculation
The nature of the aluminum species in solution varies with pH. For
simple aluminum salts (e.g., A1(N03)3 or AlCl^) species such as
Al(H20)6+3 will exist in strongly acid media whereas from about
pH k to pH 7 species such as Al8(OH)20+i* will predominate(18). At
higher pH values of 8 to 9, insoluble A1(OH)3 will form. In between
the extremes of pH h and pH 8-9, solid phase positively charged floe
will exist to a limited degree. In commercial practice alum is used
for waste treatment. Since the sulfate ion of alum is capable of
coordination within the coordination sphere, the pH values at which
soluble positively charged hydroxo species, positive floe, and
neutral floe predominate will be shifted to lower pH values such
as observed in this work.
The hydroxo complexes are apparently the most effective flocculants.
Given sufficient time these species undergo hydrolysis to insoluble
aluminates which settle out. For practical reaction times, however,
the pH must be maintained such that insoluble floe (containing the
broken emulsion) will form rapidly enough to permit of reasonable
oil removal rates. The formation rates of these insoluble species
are finite at normal operating pH values. Thus, the soluble species
initially formed will increase in hydroxylation until intermediate
positively charged pinpoint floe is generated. The total surface
area at this small floe diameter is probably very large and highly
sorptive. As precipitation continues, the floe becomes larger and
the unit area charge diminishes so that in the end a visible "neutral"
floe is finally achieved.
-------
In the early to intermediate phases of alum-lime emulsion breaking
(at optimum pH) there will be two species to consider; a positively
charged insoluble alum floe and oil droplets that may be positive,
negative, or neutral. If the emulsion particles are positive as
we postulate for those nonionically emulsified, any long range
coulombic forces will be repulsive. If the emulsion particles
are negative, as we believe they will be for anionically emulsified
systems, a net attraction will result.
The net result may be attraction or repulsion depending on the
magnitude of the residual charge on the oil droplets.
Very short range van der Waals attractive forces will exist between
the hydroxo complexes of aluminum on the emulsion and on the alum
floe particles independent of coulombi charge effects8. Only at
very close approaches will positive emulsion particles be attracted
to the positive alum floe, therefore, whereas the attraction of
positive floe for negative emulsion particles will prevail at all
distances.
If these postulates are correct an anionically emulsified oil should
break more effectively than a nonionically emulsified one under certain
conditions. For example, brand "B" rolling oil contains both anionic
and nonionic emulsifiers whereas "Dn contains only nonionic emulsifiers.
That "B" emulsion should break more easily than "D" was confirmed in
the comparative Jar testing studies (Table V).
The increased mean velocity gradient necessary to break the "D" emulsion
night be explained as being the result of charge repulsion, viz., a
higher velocity gradient was necessary to overcome the long range
potential energy barrier to van der Waals adsorption of the positively
charged oil droplets on the positive alum floe. Such a potential
barrier might also explain the greater induction periods to visible
floe.
"Redispersion" of the "B" emulsion at the higher mean velocity
gradients has as yet no satisfactory explanation. It is possible
that this may result from the anionic emulsifier adsorbing sufficiently
strongly on the alum pinpoint floe that the floe cannot agglomerate
due to anionic repulsion. This explanation seems tenuous without
direct supporting evidence, however.
^CouLombic forces diminish as the square of the distance of charge
separation. Van der Waals forces will vary at about the fourth
power of the separation distance of dipoles.
51
-------
In considering the effect of pH on the aluminum species present
and in turn the effect of these species on emulsion destabilization
and flocculation, it is also necessary to consider the fact that
the pH may not be constant during the process. At equilibrium
a given pH (with experimental error) will result regardless of
whether alum is added first, lime is added first, or both alum
and lime are added simultaneously. However, if alum is added
first, positive floe will form initially and proceed toward
insoluble neutral floe. If alum and lime are added simultaneously
the same process should occur but less time would be available
for reaction of soluble positive aluminum species with the oil
droplets. This is probably why the nonionically emulsified
system was generally most efficiently broken when treated with
alum and lime simultaneously whereas process efficiency in the
anionically emulsified case showed little preference to whether
alum was added initially or added simultaneously with lime (cf.
Table XI). When lime is added initially followed by alum at
the same dosages, however, process efficiency is quite poor for
both emulsifier types. This is probably due to the fact that
negative species such as A1(OH)£ would be formed initially and
proceed toward the insoluble alum floe state. Thus, the unneutral-
ized negatively charged emulsion particles would be repelled by
the negatively charged aluminum containing species in this latter
case and whatever clarification is achieved might be due primarily
to entrapment of the emulsion particles by the growing floe.
Kinetics
Kinetic factors include the rate at which the coagulated emulsion
particles are driven over the potential energy barrier to floccu-
lation. For this reason the effect of velocity gradient was
examined. In Table XII the final results after flocculation
under various velocity gradients were given. More importantly,
however, are the results of the kinetic runs given in Figures 9
and 109.
?For simplicity it was assumed in the kinetics portion of this
work that ln(JTUe/JTUi) = 1.18 In N/NO. This assumption appears
to be valid on the basis of the discussions presented in
Appendices C and E.
-------
1.0
\n
0.5
0.1
0.05
3
5
FIG. 10. FRACTION RESIDUAL TURBIDITY AS A FUNCTION OF
VELOCITY GRADIENT AT CONSTANT FLOCCULATION TIMES
Oil:ANl6, 1008 ppm
On: Alum: Lime = 1008/119/34.1
t = 120 sees.
t = 180 sees.
t = 240 sees.
t - 30O s«e*.
t = 360 sees.
1 1
, , 1 ,
0 10 25 W) 50 60
Gm, SI
,l,i
JC-1
, , 1 ,
75 90 100 110 125
1
1 1 1 1
140 150 160 175
-------
In applying the flocculation kinetics relationships to these data
it should be noted that the linearity in the plots of ln(JTUe/JTUi)
versus time in Figure 9 would tend to support the hypothesis of
constant D during coagulation, at least from t = 60 to t - 180 sec
for the Gjn » iWi sec-* and 59.6 sec-1 cases. For the ^ * 27.7
sec-1 case this does not appear valid. However, if one plots
ln(JTUe/JTUi) versus Gm at constant t, (Figure 10) it is found
that one can expect true coagulation kinetics in the Gm * 27.7
sec-1 case only at about t • 120 sees.
The slope of the 27.7 sec-1 data at t • 120 sec and the slopes of
the 59.6 and Ihk sec-1 cases over t * 60 to t = 180 sees, yield an
average value of -1.09 i 0.05 x 10-U for -(kDNo/l.l8). Thus the
composite rate constant (K#) value is:
K* - kD = 1.3 x 10-7(mg/l)-l
It is significant that the efficiency of coagulation with time
diminished markedly at t = 180 sees. Thus, it may be useful in
pilot plant studies to investigate this efficiency-time factor as
it influences engineering design/cost values. (It must be remembered,
however, that the time required for a given degree of turbidity
removal will be longer in a continuous flow consecutive reactor
system per reactor than in a batch reactor of the same design and
volume.)
Another unanticipated phenomenon observed is that in the Gm = 27.7
sec-1 case the turbidity initially fell and then rose again.
Although this could be due to experimental error, it is believed
that this occurred as a result of (a) delayed growth of alum floe,
and (b) a slow rate of interaction between emulsion and alum floe.
The combination of these factors could readily lead to an inter-
mediate state in which the total number of all particles present
would increase somewhat leading to a short term increase in turbidity.
It is quite apparent that the equilibrium and time rate efficiency of
flocculation are continuously increasing to the upper limit of mean
velocity gradient tested (llJi sec-1). What the upper limit may be
before "redispersion" occurs is not known. However, the quick tests
via the Waring Blendor indicated that very high initial velocity
gradients are beneficial if not maintained for long periods of
time.
The "ideal" flocculation process would probably be such as to provide
for a continuous velocity gradient. For example, in a continuous
flow large volume system the velocity gradient near the inlet can
be quite high, the values at all points in the bulk should be as
large as possible without leading to excessive floe break-up, and
the velocity gradient at the outlet should be sufficiently low so
as to provide for the growth of the large and fragile floe particles.
-------
The ideal system should also minimize the mixing of fully flocculated
with partially flocculated material if possible. This contention is
based on the fact that the optimum velocity gradient will, in
general, diminish with increasing floe size.
55
-------
SECTION V
FIELD STUDIES
The five-stand tandem cold mill at the Ashland, Kentucky Works of the
Armco Steel Corporation is capable of being operated with three
individual cooling and lubricating systems. These coolant systems
include once through water on the first standj recirculated soluble
oil emulsions on the middle stands; and once through water on the
last stand. Therefore, three waste streams can be produced from
this modern mill.
On the first stand, large volumes of water (1,500 gpm and up) are
applied on a once through basis for roll cooling and to remove the
coating oils which are applied at the exit end of the picklers.
These coating oils are typically high viscosity mixtures of mineral
oil and fat for rust and scratch protection during storage prior to
cold rolling. The removal of these oils is desirable to minimize
contamination of the coolant systems on the remaining stands. The
resulting waste is a large volume of water contaminated with both
free and emulsified oils.
On the middle stands of the cold mill, a recirculated coolant system
is provided. This coolant is typically a tight oil-in-water (o/w)
emulsion. Although this system does not result in a continuous waste,
the coolant must be dumped due to contamination from a variety of
sources. This batch dumping may involve as much as U0,000 gallons at
a time of solution containing from 0.5 to 5«0 percent emulsified oil.
On the final stand of the mill, large volumes of water (1,500 gpm and
up) are applied on a once through basis for roll cooling and removal
of the last traces of emulsion carried from the recirculation system
on the strip. The final rinsing step is often necessary to produce a
high quality, clean steel strip the customer demands. The resulting
waste is a large volume of water contaminated with a tight o/w emulsion.
A detailed evaluation of the water requirements on the cold mill by
Armco research scientists and cold mill operating personnel was
conducted. As a result, it was concluded that small concentrations
of soluble oil and suspended solids in the water used to remove pickler
coating oil at No. 1 stand would not harm the cold mill operation. It
was, therefore, decided that the oily rinse water from No. 5 stand
would be reused as rinse water on No. 1 stand prior to treatment. This
approach reduced the size of the oily waste water treatment facility by
one-third and the hydraulic loading by one-half. Design and construc-
tion of the facility were undertaken with this concept in mind.
From the standpoint of capital and operating cost, it was desirable to
make further reductions in oily rinse water volumes. It was the
57
-------
intention to include as part of the grant study an evaluation of two
alternate schemes for reduction of waste volumes. These schemes
included the use of recirculated emulsion on all five stands and the
use of rinse water on No. 5> stand only. Prior to the start up of
the treatment facility, the operating practice for the cold mill was
changed to a completely recirculated coolant system on all five stands.
Not only was this scheme found to be acceptable in terms of product
quality but this method of operating was preferred by operating per-
sonnel. It should be noted that the ability to operate the cold mill
entirely on a recirculated coolant system is a function of the final
destination of the cold rolled product. This mill was able to do so
because all cold rolled material was followed by subsequent pickling
or annealing processes which removed any residual oils. It is prob-
able that the sale of a cold rolled product directly from the cold
mill would require a rinse water system to provide a cleaner strip.
Thus, the operation of a cold mill and the actual and potential markets
it serves are an important factor in determining the waste products.
Therefore, not all cold mills would be provided a choice of coolant
systems to be used.
The fact remains, however, that activities associated with this grant
study were directly responsible for a sizable waste reduction. This
change in operating practice substantially altered the nature of the
wastes to be treated. Instead of the anticipated 1,500 gpm rinse
water flow with periodic batch dumps of concentrated coolant, the waste
consisted of concentrated coolant only. This waste was generated from
a variety of miscellaneous sources such as spills, strip overflow,
hydraulic leakage, and occasional batch dumps of contaminated coolant
systems. Oil concentrations of this waste varied considerably; however,
concentrations were sufficiently high to preclude the use of the
coagulation process for treatment. Quantities were of small enough
volume to permit the batch-wise treatment of these concentrated wastes
with the use of a cationic polyelectrolyte emulsion breaker. A program
was developed for full-scale evaluation of operating parameters for the
treatment of concentrated coolant on a batch-wise basis. Subsequently,
in order to evaluate the alum coagulation process for rinse waters on
a continuous basis, arrangements were made with operating personnel to
convert to a rinse water system similar to that for which the treatment
facility was designed. A program for optimizing and demonstrating the
alum coagulation of rinse waters on a continuous basis was developed
and executed. Although the treatment plant is capable of treating
either concentrated oily wastes on a batch-wise basis or oily rinse
water on a continuous basis, it is important to note that preferred
operation of the tandem cold mill will have a profound influence upon
the quantity and character of the resultant waste and upon the costs
involved.
BATCH TREATMENT OF CONCENTRATED COOLANT
Initial operation of the treatment facility was directed toward the
-------
treatment of concentrated roll coolant generated by the tandem cold
mill recirculated coolant system. The sources of this waste included
spills, strip overflow, hydraulic leakage, periodic dumping of con-
taminated coolant systems, flushing and cleaning water, etc.
The characteristics of this incoming waste received over the period of
testing are shown in Table XIII. It is apparent that large variations
exist in the waste characteristics.
TABLE XIII
SUMMARY OF INFLUENT ANALYTICAL DATA FCR CONCENTRATED COOLANT
Number Average Range
Analysis Samples (ppm) (ppm)
Suspended Solids 21 1,02$ 156 - 3>010
COD 19 31,600 3,200 - 1*6,800
Total Oil 22 11^,600 2,088 - U6,7U2
Emulsified Oil 23 3,U9U 53 - 10,0?8
Jar Testing Program
An extensive jar testing program was conducted to establish optimum
chemical treatment and to evaluate various organic emulsion breaking
chemicals. This program included the evaluation of 27 different
emulsion breakers at various concentrations and temperatures. Most
of the chemicals tested were not effective on the waste, which
initially consisted primarily of Brand "A" rolling oil. A few were
effective only at excessive concentrations. Two cationic polymers
were found to be effective at 1,000 ppm or less. One of these products
was in the developmental stage and not coranercially available. The
other (EBl) was therefore selected as the emulsion breaker to be used
for treatment.
During jar testing it was found that the effective emulsion breakers
were quite sensitive to overdosing and temperature. EBl appears to
have a limited range of concentrations within which it is capable of
breaking the emulsion. At concentrations of 50 percent of the optimum
or 125 to 150 percent of the optimum, this emulsion breaker was con-
sidered completely ineffective. These observations are consistent
with the fundamental studies which indicated that charge reversal can
occur in coagulation processes. In the case of cationic polyelectro-
lytes, it is conceivable that this may lead to emulsion restabilization.
This conclusion was further supported by the fact that overdosing with
emulsion breaker could be compensated for by subsequent dilution with
more waste.
The optimum temperature for EBl was found to be between 100 and 150 F,
59
-------
During the program for batch treatment evaluation, a change in rolling
oils from "A" to "B" was made. Due to the specificity of emulsion
breakers to oil types, several hundred additional jar tests were per-
formed. Three emulsion breakers, including EB1, were found to be
successful at similar concentrations for "B". To avoid introducing
a new process variable into the evaluation program, treatment was
continued with EB1.
Operations Testing Program
With the results of the preliminary jar testing, a program was estab-
lished for full-scale evaluation of the treatment of concentrated
coolant on a batch-wise basis. The following basic procedure for
treating the coolant was established. Wastes were collected in the
60,000 gallon holding pit as they were received from the cold mill sources
and heated with steam to a predetermined temperature (Figure 11). From
the holding pit, the waste was pumped directly to the second of two
rapid mix tanks where the emulsion breaker was added. The treated
waste then flowed by gravity into one of two parallel flocculation
tanks, into the corresponding air flotation tank, and into the receiv-
ing stream. Because of the small volumes, only one of the two parallel
flocculation-flotation tank units was necessary at any time. The
separated oil was skimmed from the flotation tank surface and pumped
to the oil recovery area. This procedure allowed for treatment of an
intermittent stream or treatment of a specified volume or "batch" of
60,000 gallons. The treated waste was permitted to settle until an-
other quantity of waste was accumulated. Then another 60,000 gallons
were treated, which purged the system of the previously treated waste,
and again permitted to settle. The end result was a series of 60,000
gallon batch treatments.
A series of tests was established to evaluate three variables considered
to be of importance. These variables included chemical concentration,
temperature, and flow rate. Because of the sensitivity of chemical
concentration to oil content as indicated by jar tests and because of
the variability of the incoming waste, it was essential to vary chemical
dosage with regard to the influences of the incoming waste. Thus, jar
tests were performed for each batch to be treated, and the optimum
chemical concentration as indicated by the jar test was used as a base
value. To relate the plant-scale performance to jar test results, the
three levels of chemical concentration evaluated were the optimum dose
as indicated by jar test, one-half that optimum dose, and no chemical
addition at all. The two temperature levels evaluated were 100 and
150 F. Flow rates evaluated were 200 gpm and 35>0 gpm. Total oil
removal was used as the basis for evaluating performance under the
various conditions.
A total of 21 tests for which valid analytical data were obtained were
conducted. These represented ten combinations of operating conditions.
The conditions and the results for these combinations are summarized in
60
-------
Waste
Coolant
D
Emulsion
Breaker
60,(XX) gal. pit
Recovered
Oil
Steam
Coils
D
8,000 gal.
Mix Tank
20,000 gal.
"Holding Tank,
To 60,000 gal. pit
D
t
« «
«
8,000 gal.
Floe Tank
n
20,000 gal.
Air Flotation
Tank
I
To Sewer
20,000 gal.
^Clarifier,
Steam
Coils
FIG. 11. SIMPLIFIED FLOW DIAGRAM - BATCH TREATMENT SYSTEM
-------
Tables XIV and XV. As the values for total oil removal indicate, there
is little difference in deviation of values for each variable over the
ranges tested. It is believed that the reason for the lack of a signif-
icant difference in removals lies in the inherent batch technique
employed. Because of the difference from the basic plant design flow,
mixing times were in the range of 25-35 minutes as opposed to approxi-
mately five minutes used in jar testing. Furthermore, treated batches
were permitted to settle for periods ranging from one to four days.
These extended mixing and settling periods were sufficient to dampen
any difference in process parameters.
A statistical analysis of the data permits the following observations
to be made:
1. The addition of no chemical yields significantly lower oil
removal than adding either one-half or the full optimum dose
as determined by jar testing.
TABLE XIV
SUMMARY OF BATCH TREATMENT TEST RESULTS
Average
Chemica^L Temperature Flow Rate No. Total Oil Standard
Combination Dose (F) (gpm) Runs % Removal Deviation
A 0 150 200 1 66
B 0 150 350 1 U9
C % 150 350 2 80 15.5
D 1 150 350 1* 92 9.6
E ^ 100 350 3 95 i^.ij
F 1 100 350 2 86 13.5
0 % 100 200 U 95 ii.5
H 1 100 200 1 89 -
I % 150 200 2 90 10.6
J 1 150 200 I 8h
a 0 = no chemical addition; % - % jar test optimum dose; 1 = jar test
optimum dose
62
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TABLE XV
SUMMARY OF BATCH TREATMENT PERFORMANCE FCR VARIABLES EVALUATED
Average
Number Total Oil Standard
Variable Runs b % Removal Deviation
Chemical Dosea
0 2 58 12.0
5g 11 91 9.0
1 8 88 8.3
Temperature
100 F 10 93 9.1
150 F 9 8? 9.8
Flow Rate
200 gpm 8 92 6.5
350 gpm 11 89 9.7
a 0 • no chemical addition; ^5 = ^g jar test optimum dose; 1 * jar test
optimum dose
k Runs with zero chemical addition not included for averaging purposes
for temperature and flow rate
2. There is no significant difference between applying one-half
the optimum or the full optimum chemical dose.
3. There is no significant difference between waste temperatures
of 100 F and 150 F for total oil removal.
U« There is no significant difference between running a batch at
200 gpm and at 350 gpm for total oil removal.
Based on these observations, it can be concluded that batch treatment
can be performed under the most economical of the conditions studied
without sacrificing performance as measured by total oil removal. This
most economical set of conditions would be one-half the optimum jar
test chemical dose, 100 F, and 350 gpm flow rate. The three runs actu-
ally performed under these conditions yielded the following average
performance data:
1. 9$% total oil removal
2. 92% COD removal
3. 9J# emulsified oil removal
U. 71$ suspended solids removal
5. 18£ bottom water and solids in the scum
6. 1.8£ solids in the scum
63
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7. 1850 gallons of scum per 60,000 gallon run
No other set of conditions yielded appreciably better performance.
The chemical requirement, as determined by jar test for each run, was
found to be highly dependent upon the oil being treated. For "B" the
average optimum chemical dosage was 2.06 gallons/1,000 gallons of
waste. For "A" only 0.60 gallons/1,000 gallons was required.
At $3.?U per gallon for EB1, the cost to treat "A" would be $2.2U/1,000
gallons as compared to $7.70/1,000 gallons for "B". Findings of this
program indicate these cost figures may be halved by using one-half the
jar test optimum without decreasing performance. It is conceivable that
further reductions in chemical requirements are possible. The smaller
chemical requirements indicate that the full-scale plant is somewhat
more efficient than standard bench-scale jar testing techniques.
(Appendix A)
If it is assumed that a 60,000 gallon batch of "B" waste is treated
under the most economical situation, as mentioned above, and will yield
l>8f>0 gallons of 82 percent oil, then a chemical cost of $231 is
required to recover 1,520 gallons of oil. The chemical cost is there-
fore l5.2£/gallon of oil recovered for a product worth approximately
8.5#/gallon as fuel oil. The same analysis for "A" will result in a
chemical cost of U«8#/gallon of oil recovered. It should be emphasized
that these figures are for chemical cost only and do not take into
account the possibility of skimming tramp oils directly from the holding
pit to the oil recovery area nor the non-chemical cost of processing
the scum to an oil quality worthy of the 8.5#/gallon credit. The
designation of the recovered oil as fuel oil is not unrealistic in
light of the fact that centrifugal analysis of Oil Holding Tank samples
consistently showed bottom water and solids of less than one percent
with virtually no solids. This improvement of oil quality over that of
the collected scum was achieved by heating to 175 F and periodic mixing
in the clarifier (cf Figure 11).
Over the course of the testing period, the cold mill averaged 116,000
gallons of waste per week for a 12-turn average week. However, volumes
were substantially reduced toward the end of the period so that a vol-
ume of approximately 100,000 gallons may be anticipated for comparison
of economics. This reduction was brought about as a direct result of
this grant study. Quantities of waste were monitored and related to
cold mill operating personnel. Alarmed at the quantity of coolant be-
ing lost in production, the operating personnel tightened their
maintenance of the cold mill sources and effected the waste reduction.
On the basis of receiving 100,000 gallons per week, the chemical cost
of treating the waste (assuming "B") would be $385 per week or $55/day
on a proportionate basis. Credit received from the recovered oil would
be $2l5/week on the basis previously established or $31/day. Therefore,
the net chemical-cost vs. oil-credit result would be $2i|/day. These
6k
-------
figures will be used for comparison with the cost of treating rinse
water on an alum coagulation basis. It is obvious that a great reduc-
tion in the chemical cost of treating the waste, and therefore a net
positive credit toward the operation, can be achieved if Brand "A"
rolling oil is chosen over Brand "B".
In summarizing the results of the studies of batch treatment of con-
centrated coolant, it is apparent that the treatment plant is somewhat
less sensitive to operating variables than would be indicated at bench
scale. What were believed to be extremes in operating conditions
actually revealed no significant difference in plant performance.
This discrepancy between bench and full-scale requirements is believed
to be a result of the extended periods of mixing and settling provided
by the batch treatment method as compared to the relatively short
periods provided in jar testing procedures. It is inconceivable that
operating conditions can be varied indefinitely without altering plant
performance, as was indicated by reducing chemical addition to zero,
but the results of this study do indicate that the plant as operated
on a batch treatment basis is capable of performing satisfactorily over
a wide range of operating conditions even for the relatively sensitive
treatment technique employing an emulsion breaker.
CONTINUOUS TREATMENT CF RINSE WATERS
In order to evaluate and optimize the alum coagulation process for the
treatment of rinse water, the tandem cold mill was converted to a rinae
water system on No. 1 stand. A re circulated coolant system was main-
tained on the other four stands. The rolling oil utilized during this
period was Brand "C".
Prior to beginning an actual program for evaluating various parameters,
a period of several weeks was utilized for preliminary bench studies
and resolving various minor mechanical problems. A quantity of con-
centrated emulsion (from strip overflow, spills, floor drainage, etc.)
continued to be received despite the conversion to rinse water on stand
No. 1. The original intent was to separate the concentrated emulsions
from the rinse water, if possible, and treat the emulsions on a batch
basis during periods when the rinse water system was not operating.
However, it soon became apparent that the quantities of concentrated
emulsions were too great and associated storage capacity was too small
to permit this total segregation. Hence, a small quantity of concen-
trated emulsion was metered into the rinse water stream and the resultant
waste treated. Figure 12 shows a simplified flow diagram for this system.
Following the establishment of an initial set of operating conditions
and the resultant production of a satisfactory effluent, a program was
developed for determining the effects of operating parameters on
the treatment plant's coagulation process. This program encompassed
chemical dosage rates, chemical addition points, rapid mix detention
-------
Oily Rinse
15,000 gal.
Equalization
Tank
Waste
Coolant
60,000 gal.
pit
8,000 gal.
Mix Tank
8,000 gal.
Mix Tank
, 2-8,000 gal,
Floe Tanks
n
2-20,000 gal.
Air Flotation
Tanks
Scran to
Recovery
To Sewer
FIG. 12. SIMPLIFIED FLOW DIAGRAM - CONTINUOUS TREATMENT SYSTEM
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time, flocculator detention time, flocculator tip speed, air flotation
detention time, and air flotation recylce rate. The order of evaluation
was approximately as given, although dependent parameters such as air
flotation and flocculation detention time necessarily were related due
to the fact that the flocculation chamber and air flotation tank are
physically one un±t, separated only by a baffle under which the waste
stream must pass. (Appendix B) The above parameters were varied over
the maximum practical range. As the optimum condition for each para-
meter was established, the next parameter was evaluated. Operating
cost, effluent water quality, and oily scum quality were the basis for
evaluation.
The analytical work undertaken included: (a) pH, temperature, COD,
total oil, soluble oil, and suspended solids for influent samples; (b)
pH, temperature, COD, total oil, suspended solids, and turbidity for
effluent samples; and (c) bottom water and solids for separated oil
samples.
Preliminary Bench Studies
Bench studies were performed to derive a first approximation of optimum
chemical requirements for alum coagulation of the rinse water and to
evaluate emulsion breaking chemicals capable of being used for treatment
of Brand "C" rolling oil wastes.
A first approximation of optimum operating conditions was derived from
an extensive series of jar tests. For a waste stream consisting of
about one part of concentrated emulsion per 12 parts of rinse water
(average total oil 895 ppm), the combination of 195 ppm alum (Al2(SOU)3*
18H20), U3 ppm lime, 32 ppm clay, and 1 ppm anionic coagulant aid was
found to produce a satisfactory effluent. Jar tests on the waste also
indicated that about 100 ppm alum may be added to the waste stream with-
out requiring lime due to sufficient alkalinity in the water to maintain
a pH of 6.0 or higher. At higher alum dosages, approximately one part
lime for each additional two parts of alum is required. Although this
information was used for a first approximation of the lime requirement,
the actual pH in the rapid mix tank was used to govern the lime require-
ment on an operating basis.
Because of the specificity of emulsion breakers to type of rolling oil,
an evaluation of various emulsion breakers was made for the Brand "C"
oils for purposes of treating batches of concentrated emulsion or as a
possible alum substitute in the coagulation process. EB1 which was
previously used for breaking concentrated emulsions of "A" and "B" was
not well-suited for the "C" oil. A total of 5U polymers were evaluated
as possible replacements for EB1. Of these, eight appeared to work at
lower doses. Further evaluation resulted in the following findings:
1. One emulsion breaker was rejected because it produced a
very sticky, gummy sludge.
67
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2. Four other emulsion breakers were rejected because they were
not effective at low oil concentrations.
3. The three remaining polymers were effective emulsion breakers
at any oil concentration down to 0.£ percent and were still
effective somewhat below 0.5 percent. One of these three was
considerably more economical and was chosen for use for treat-
ing "C" on a batch basis if and when required to do so.
A series of jar tests was carried out to determine whether cationic
polyelectrolytes could economically be substituted for part or all of
the alum. This change would reduce the ash content of the recovered
oil; it might also eliminate the need for lime in the process, which
would eliminate the lime feeding problems (see below) as well as reduce
the hardness and dissolved solids in the plant effluent. It was found
that at least half the alum could be replaced with polymer, but at a
cost of six to seven times the cost of the alum replaced. The sub-
stitution could not, then, be justified.
Operational Problems
Definitive data on the effects of varying operating conditions were
often difficult to obtain due to the influence of uncontrollable
factors which in many cases were more influential than the process
variations themselves.
The inability to maintain complete control over these extrinsic
factors was a major concern during the optimization phase. A discus-
sion of these factors appears to be in order in view of the fact that
they represent a major concern in the design and operation of a
similar treatment facility.
First, problems with operation of the cold mill itself were exhibited.
Wastes from the recirculated coolant system and from the rinse water
on stand No. 1 ideally are collected separately and pumped in separate
lines to the treatment plant where the coolant is metered at a constant
rate into the rinse water for treatment. Segregation of these wastes
was a problem, however. Pumps at the mill periodically were down for
maintenance. As a result, the combined collection of coolant and rinse
water was necessary in order to avoid inundating the cold mill oil
cellar. Thus, concentrated coolant sometimes was received through the
rinse water line and vice versa. This action had pronounced effects
on the oil concentration and treatability of the wastes.
The capacity of the holding pit for concentrated coolant is limited.
Ordinarily, the rate of metering coolant into the rinse water was
sufficient to maintain a satisfactory level in the holding pit. At
times, however, the relative volumes of rinse water and coolant became
unbalanced which resulted in rapid changes in holding pit levels. In
order to prevent inundation of the treatment plant, a level control
activated a pump, at a pre-established level, which conveyed the cool-
ant to the process for treatment. This additional volume of coolant
68
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during pit pump-downs increased the oil concentration of the waste to
the point where system efficiency was substantially reduced. Periods
of low rinse water flows were generally associated with converting to
a completely recirculated coolant system on the cold mill as opposed
to using rinse water on stand No. 1. This change was usually a result
of required maintenance on the rinse water spray system or of cold mill
operating schedules.
When operation was normal, the cold mill pumping system was still a
source of concern. Pumping of wastes is controlled by sump level con-
trols at the cold mill, so that when a specified level is reached the
pumps are activated for a period of sufficient duration to empty the
sumps. Observations revealed that this resulted in a cycle of perhaps
five minutes of pumping at 1,300 gpm and ten minutes off altogether.
Sufficient equalization was not provided to account for the problems
encountered. It is believed that equalization capacity should be a
prime consideration in future designs due to the unpredictable operation
of the cold mill.
Another operational problem was pH control. Difficulties were encoun-
tered in attempting to pump lime slurries with positive displacement
pumps. After many futile attempts, a centrifugal pumping system for
lime was installed. This essentially eliminated the problem. During
the period when lime pumping problems were encountered, the slightest
malfunction was observed to have pronounced effects on the process. It
was determined that when the pH was permitted to fall below 6.0 the
alum floe precipitation did not occur. The reliable pumping of lime
was found to be the single most important process factor. Future in-
stallations should be equipped with continuous pH monitoring of the
effluents, and care should be taken in designing pumping facilities for
lime slurries.
A good deal of emphasis was placed upon effluent turbidity as a measure
of process performance due to its sensitivity and rapid response. This
necessitates reliable instrumentation for turbidity. The turbidimeter
employed in this installation required frequent cleaning to ensure
reliable readings. It is believed that greater flow rates in the tur-
bidimeter sample lines would alleviate the settling of solids and oil
which creates plugging of these lines.
Physical destruction of the delicate floe structure was also observed.
A primary cause of this problem was the hydraulic shock exerted on the
system during a sudden start-up after an extended period when no waste
was being received. Even the influence of a rain shower was observed
to destroy the "air-floe" structure at the flotation tank surface so
that effluent turbidity increased. The problem with sudden start-ups
could be alleviated somewhat if adequate equalization or storage capac-
ity was provided. The influence of rain suggests the need for some
type of shelter for the flotation units.
69
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When all the above factors could be controlled, the process operated
in a highly desirable fashion. However, these difficulties recurred
with such regularity and frequency as to represent significant prob-
lems. It is felt that these factors must be weighed with even more
concern than the process parameters themselves when considering future
treatment plant designs.
Chemical Dosage Rates
A first approximation of chemical dosage rates was derived from an
extensive series of jar tests. For a waste stream consisting of one
part of concentrated coolant per 12 parts of rinse water, the combi-
nation of 195 ppm alum, 1*3 ppm lime, 32 ppm clay, and 1 ppm coagulant
aid was found to be a conservative chemical program to produce a sat-
isfactory effluent. Characteristics of the incoming waste are given
in Table XVI.
TABLE XVI
INFLUENT ANALYTICAL DATA - RINSE WATER a
Analysis No. Samples Average Range
Suspended Solids (ppm) 22 80 16 - 280
COD (ppm) kO 5,556 1,OUO - 37,600
Total Oil (ppm) 37 895 113 - 3,03U
Emulsified Oil (ppm) 37 2lUj. 30 - 850
pH 57 7.07 6.2 - 8.1
Temperature (F) 57 91 79 - 100
BCD (ppm) 5 383 165 - 5UO
a Samples are for combined waste of rinse water and concentrated
emulsion at a ratio of approximately 12:1.
The initial chemical program was applied to the full-scale facility and
a satisfactory effluent was attained. Typical performance data is given
below:
Effluent Turbidity 5-1.5 JTU
Effluent Total Oil 52 ppm
Effluent COD 765 ppm
Total Oil Removal 92%
COD Removal
Chemicals were systematically reduced and the resulting treatment
observed. It was found that chemical additions could be reduced to
175 ppm alum, 36 ppm lime, 16 ppm clay, and 0.5 ppm coagulant aid
without diminishing effluent water quality. This reduction probably
reflected the conservative nature of the recommended jar test program.
70
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Attempts to further reduce chemical requirements met with varying
success. Although an acceptable effluent was sometimes observed
with lower chemical concentrations, the combination above is believed
to be necessary in order to cope with influent oil concentrations and
other variables. This combination of chemicals represents a chemical
cost of $.051/1,000 gallons of waste. The slight savings in chemical
cost afforded by reduced chemical addition is not believed worth the
reduction in treatment reliability. For example, the combination of
l50, 2k., 16, and 0.5 ppm was observed to give sporadic treatment at a
cost of $.0146/1,000 gallons. At the waste volumes encountered, this
represents only $3.60/day savings in chemical costs.
Thus, for this particular application and mode of operation, the
desirable chemical program is 175 ppm alum, 36 ppm lime, 16 ppm clay,
and 0.5 ppm coagulant aid. It should be noted that any alteration in
concentrated coolant to rinse water ratio would substantially alter
this set of chemical requirements.
Chemical Addition Points and Rapid Mixing Time
Chemicals were capable of being added to the waste at a variety of
locations. Alum, lime, and clay could be added to the first or second
rapid mix tank. Coagulant aid was capable of being added to either the
second mix tank or the entry line to the flocculation chamber. Pre-
vious to this evaluation, alum and line were added to the first mix
tank, clay to the second, and coagulant aid following the second prior
to flocculation.
This scheme was based on jar testing which had indicated the necessity
of 5 minutes mixing between alum and clay additions and 5 minutes mix-
ing between clay and coagulant aid. Rapid mix tanks were sized based
on this concept. However, flow rates were a maximum of approximately
500 gpm over an extended period as opposed to the 1500 gpm design flow.
Thus, detention times in the mix tanks were a minimum of 15 minutes on
a continuous basis. These extended mixing periods must be taken into
account in interpreting results of changes in points of chemical
addition.
Various schemes for applying the four chemicals to the waste were under-
taken. These schemes appear in Table XVII.
71
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TABLE XVII
CHEMICAL ADDITION SCHEMES
Scheme Rapid Mix #1 Rapid Mix #2 Line to Floe Tank
1 A,L C P
2 A,L,C P
3 A,L,C - P
b - A,L,C P
5 A C,L P
6 - A,L,C,P
7 A,L C,P
A = alum, L - lime, C « clay, P • polymer
Sharp contrasts in effluent quality for the different schemes were not
found. General trends indicated, however, that schemes 6 and 7 were
not capable of providing a consistently good effluent. These two
schemes are distinguished by the fact that a period of mixing was not
provided between the addition of clay and of coagulant aid. This
observation is consistent with jar test findings and recommendations of
polymer suppliers which indicate that coagulant aid functions most
efficiently if added after other chemicals have been applied.
The lack of a significant difference for the separate addition of alum
and lime indicates that this method offers no advantage. It was be-
lieved that pH adjustment for optimum flocculation after providing a
period for emulsion breaking by the alum would prove beneficial. This
idea was not substantiated, however.
The inability to note significant differences in treatment for other
schemes leads to the conclusion that simultaneous addition of clay and
alum does not appear to reduce treatment performance. It is possible
that the total system serves to buffer effluent quality and renders it
insensitive to subtle changes such as chemical addition points.
Although a period of mixing is required between the alum-clay addition
and the coagulant aid addition, no significant difference between one
mix tank or two mix tanks could be detected. However, this fact is
necessarily qualified in light of the extended rapid mix detention times
brought about by less than design flow rates. It is possible that a
difference may be detected between f> and 10 minutes, whereas none was
noted between approximately IfJ and 30 minutes.
In view of these findings, satisfactory treatment can be obtained with
a more economical scheme than that of the design for the flow rates
that were received. Such a scheme would consist of a single rapid mix
tank with provisions for adding alum, lime, and clay to the mix tank
and coagulant aid to the line leading from the mix tank to the floccu-
lation chamber. This practice would eliminate the need for one mix
72
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tank, one mixer, and miscellaneous chemical piping. Therefore, in
terms of economics as well as treatability, this scheme must be
considered most appropriate. However, it is also possible that two
rapid mix tanks of less detention time would provide a more economical
scheme.
Flocculation - Air Flotation Detention Time
Following chemical addition and rapid mixing, the waste flow splits
into parallel flocculation - air flotation tanks. Detention time for
flocculation and air flotation are interrelated. Prior to this eval-
uation of detention times, flow was diverted through both parallel
units. Design detention times for flocculation and flotation were
10 and 20 minutes, respectively. However, the maximum flow rate
attained on a continuous basis resulted in actual detention time of
32 and 57 minutes.
In an effort to evaluate the significance of detention time and at the
same time approach design conditions, the entire flow was diverted
through one of the parallel flocculation - flotation units. Under this
condition, maximum flow rates attained on a continuous basis ($00 gpm)
resulted in actual detention times of 16 minutes for flocculation and
2$, 28, or 33 minutes for flotation, depending upon the recycle rate.
The immediate effect of this halving of detention time was a gradual
but significant rise in effluent turbidity. However, after several
turnovers of the system, effluent quality returned to an acceptable
level. This period of adjustment was believed to be another exempli-
fication of the need for a fairly stable hydraulic loading. Nonethe-
less, an acceptable effluent was attained with the lesser detention
times. Further reductions in detention time for the flocculation -
flotation units maybe possible.
Optimum detention time in terms of operation of a facility is academic
since this is a function of incoming flow rate. Intuitively, the
longer the detention the better will be the results. From a design
standpoint, it can only be concluded that a combined flocculation -
flotation detention time of Ul minutes has proven to be adequate and
that lesser detention maybe possible.
Flocculator Tip Speed
Flocculator tip speeds prior to the evaluation of this parameter were
maintained at 0.6 feet per second (fps). When flow was still being
diverted through both flocculation units, an experiment to test the
effects of no mechanical flocculation was undertaken using one unit
as a control. After several days of operation no significant dif-
ference in effluent quality could be detected between the unit
employing a flocculator tip speed of 0.6 fps and that having the
flocculator turned off altogether.
73
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When detention time was halved by diverting the entire flow through
one unit, four flocculator tip speeds were tested for performance.
These were 0, 0.6, 1.8, and 3.6 fps. No difference in performance
for any of these conditions was evident. Satisfactory treatment was
achieved whether the flocculators were turned off completely or whether
they were operated at maximum speed.
There are two possible explanations that might refute the conclusion
that the flocculators are not needed. First, there is the matter of
detention time. It is possible that the relatively long period pro-
vided for settling as a result of the decreased flows is sufficient
to compensate for operation of the flocculators, whether beneficial
or detrimental. Secondly, the hydraulics of the system may be such
as to negate flocculator performance. The flocced waste must pass
under a baffle and through a perforated distribution plate where
dissolved air is introduced. This action may be violent enough to
break up the delicate floe, regardless of how efficient the flocculator
has been.
Nevertheless, because effective treatment is obtained with or without
mechanical flocculation, the most economical condition is with the floc-
culators turned off. In designing a similar treatment facility, it
appears that the flocculating equipment could be omitted and a corre-
sponding savings realized. However, careful consideration of this
facet should be given, particularly in relation to decreased detention
times for the flocculation - flotation units.
Air Flotation Recycle Rate
The recycled portion of the effluent which is pressurized and saturated
with dissolved air had been maintained throughout the program at 100 gpm.
A specific percentage of recycle was unattainable due to the variation
in waste flow rate. A recycle rate of 100 gpm results in a range from
100 percent to the flow for periods of no incoming waste to 1|0 percent
of the flow at the maximum flow rates experienced over an extended period,
Although the recycle system is capable of a range from 0 to 350 gpm,
maintaining recycle flows below 100 gpm is impractical due to the rela-
tive insensitivity of the controls below this level.
During the period when flow was diverted through both flotation units
and detention times were maximum, an experiment to test the need for
the air was undertaken using one unit as a control. Within an hour
after the recycle system was turned off in one unit, the effluent tur-
bidity of that unit began to soar while the control unit maintained an
excellent effluent quality. After several turnovers of the system,
the air was turned back on for the test unit. Effluent quality improved
immediately and returned to the original satisfactory quality. The
treated waste during this test was observed to consist of an excellent
floe, but one which remained suspended. The system as operated at this
facility definitely requires air for efficient separation of the oily
floe despite the extended detention times.
-------
For operation through one unit of the system, an evaluation of the
recycle rate was undertaken. The rates evaluated were 100, 200, and
300 gpm. These represent percentages of waste flow of 20, UO, and 60
percent at the maximum flow rates experienced over an extended period.
The results of these tests indicated that no significant difference in
effluent quality was observed for the three recycle rates. All were
capable of producing a consistently good effluent.
The failure to detect differences in effluent quality for these large
variations in recycle rate may be the result of compensating factors.
One might expect the separation of oily floe to improve as the amount
of dissolved air is increased. However, at the same time the detention
time is decreased by the added recycle rate. For example, increasing
the recycle from 100 to 300 gpm decreases detention time in the flota-
tion unit by approximately 8 minutes. Furthermore, the influence of
increased hydraulic agitation brought about by the higher recycle rates
is not known. Other observations on the effects of hydraulic influences
suggests that this may be a matter of significant importance.
The results of the evaluation of air flotation recycle rates indicates
that some dissolved air is necessary for separation of oily floe but
that the recycle rate required is independent of effluent quality, at
least above 20 percent. The most economical condition, therefore, is
the smallest quantity of recycle capable of being controlled or 100 gpm
in the case of this installation.
Demonstration Program
Following the optimization of a number of treatment variables, an
extended demonstration phase was undertaken to establish treatment
efficiency and costs over a relatively long period. The conditions
which were established and maintained throughout the demonstation
period for most effective treatment are summarized as follows:
1. Chemicals
Alum - 175 ppm, added to second rapid mix tank
Lime - A sufficient quantity to maintain the pH between
6.5 and 7.5, added to the second rapid mix tank
Clay - 16 ppm, added to the second rapid mix tank
Coagulant Aid - 0.5 ppm, added between the second rapid
mix tank and the flocculator
2. Flocculation - One flocculation tank on line with flocculator
mechanism off
3. Air Flotation - One air flotation tank on line with the
recycle rate at 100 gpm
75
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U. Oil Recovery Area - Clarifier temperature l£0 P with mixer
operated eight consecutive hours out of each twenty-
four, and holding tank temperature 180 F
Skimming of the oily floe from the surface of the air flotation tank
was performed mechanically according to an intermittent on-off cycle.
Operating experience indicated that a cycle of five minutes of oper-
ation out of each hour, the required time for one flight to travel the
length of the flotation tank, minimized the amount of water skimmed to
the oil recovery area, in addition to preventing dewatering of the scum
to the point of being unmanageable.
Under these conditions, an extensive sampling and analysis program was
established around the treatment facility. The results of this pro-
gram are reported in Table XVIII,10
As shown in Table XVIII, the treatment facility is quite efficient in
the removal of COD, BOD, and total oil with average removals in the
range of 90 percent. Suspended solids removals, however, were found
to be quite variable and in general quite low. The poor suspended
solids removal is attributed to rapid variations in hydraulic loading
which result in removal and carryover of solids previously deposited
on the air flotation tank bottom and side walls. Substantial improve-
ments in removal efficiencies for suspended solids in addition to COD,
BOD, and total oil would be expected if additional equalization facil-
ities were provided upstream of the treatment plant.
The oily scum produced in the air flotation tank was found to be quite
variable in volume and composition ranging from 2,000 to 12,000 gpd
and 10 to 90 percent oil by volume. These wide variations are primar-
ily a result of variations in hydraulic loading and waste composition.
Additional equalization would probably result in a more uniform scum
quantity and quality.
In an attempt to determine whether the oil could be upgraded chemically,
a bench scale test was performed employing the addition of an oil
soluble polymer. Varying proportions of the polymer were added to oil
samples and placed along with a blank into a drying oven at 175 F for
2k hours. Each sample was then observed to consist of a substance
having no fluid characteristics whatsoever. No stratification was
observed in these solid samples. It was apparent from this experiment
that any concerted efforts to improve oil quality by simple dewatering
(gravity separation) would result in an undesirable and unmanageable
end product.
The oil recovered from the facility was found to be of insufficient
10 All analysis were performed in accordance with 12th Edition of
Standard Methods for the Examination of Water & Waste Water.
76
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DIGITALLY
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quality for use as fuel oil due to its high concentrations of water and
solids. The average water and solids concentrations observed during
this period were 8 percent and 2 percent by volume, respectively, as
determined by centrifugal separation. The viscosity of the oil layer
following centrifuging was observed to be extremely high, indicating
that a large amount of solids remain with the oil. This observation
coupled with the previously reported observation that complete dehy-
dration of the scum results in the formation of a semi-solid material
indicates that the simple application of heat and gravity separation
is insufficient to produce an oil suitable for fuel oil use.
Although fuel oil quality appears to be unattainable, the recovered oil
was in demand for use as road oil. The 107,000 gallons of recovered
oil produced during this seven week period were given to nearby munici-
palities and counties for use on local roads.
During the demonstration phase of the program, the average rinse water
flow when the mill was operating was 350 gpm. For a 12-turn operating
week, this would total 2.02 million gallons. An additional 3¥>,000
gallons-" of concentrated coolant could be expected for a total quantity
of waste of 2.36 million gallons to be treated. At a chemical cost of
$0.0^1/1000 gallons of waste, the weekly chemical cost would be $121 or
$17/day.
The experience with disposal of recovered oil from rinse water treatment
indicated that neither a credit nor an added cost can be applied. The
quality of the oil that can be attained (10 percent bottom water and
solids and 2 percent solids) is not adequate to permit its use as fuel
oil. However, private concerns disposed of the oil at no cost as rapidly
as it was recovered for use as road oil. Hence, no cost for incineration
of the oil was incurred.
COMPARATIVE ECONCMICS
In view of the fact that this tandem cold mill was operated successfully
on either a completely recirculated coolant system or on a combination
rinse water-coolant system, it is of interest to relate the treatment
costs associated with the two operations. The cost data developed for
each are necessarily based upon average operational information gathered
during the field studies. However, it should be noted that the assump-
tions made in deriving these cost data are subject to drastic variations
This apparent increase in coolant volume is a result of contamination
of the coolant with rinse water and does not mean that more coolant
was necessarily lost when operating on the rinse water system. This
314,0,000 gallons of coolant is actually less concentrated than the
100,000 gallons received when operating on a completely recirculated
system.
79
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which may have significant effects on these figures. Of primary signif-
icance are the waste volumes received at the treatment plant and the
type of rolling oil being utilized by the cold mill.
As mentioned previously, the net chemical cost vs. oil credit for the
batch treatment of concentrated coolant consisting of "B" oil is
$2U/day based upon the following assumptions:
Waste Volume 100,000 gallons/week
Chemical Cost $3.85/1,000 gallons
Oil Recovered 2,5UO gallons/week
Oil Value 8.5#/gallon
Total Chemical Cost $385/week
Total Oil Credit $2l6/week
Net Cost $l69/week or $21*/day
For the use of Brand "A", a net credit of $l5/day can be realized when
considering the chemical cost and value of the recovered oil based upon
the following assumptions:
Waste Volume 100,000 gallons/week
Chemical Cost $1.12/1,000 gallons
Oil Recovered 2,5UO gallons/week
Oil Value 8.50/gallon
Total Chemical Cost $112/week
Total Oil Credit $2l6/week
Net Credit $10U/week or $l5/day
During the course of the 13 week batch treatment program, a total of
117,321 tons of steel was rolled by the tandem cold mill for an average
of 9,000 tons/week. On this basis, the chemical treatment cost per ton
of rolled steel is $.019/ton for "B". Treatment of "A" would show a
credit of $,Oil/ton of steel.
The combined chemical cost and oil credit for the treatment of oily
rinse waters consisting of "C" rolling oil is $l?/day based upon the
following information gathered during the demonstration phase of the
rinse water coagulation programs
Waste Volume 2,360,000 gallons/week
Chemical Cost $0.051/1,000 gallons
Oil Recovered 15,300 gallons/week
Oil Value None
Total Chemical Cost $121/week
Total Oil Credit None
Net Cost $121/week or $17/day
Although at first glance it would appear that rinse water treatment of
"C" is more economical than batch treatment of "B", it must be pointed
80
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out that significant waste volume variations for rinse water are much
more likely to occur than for batch treatment since the rinse water
volume is strongly related to the preferred quantity utilized on the
mill. An increase in chemical cost would not be directly proportional
to increases in rinse water volume since the chemical requirement would
probably be dependent more upon oil concentration than waste volume,
and the oil lost on the mill would be essentially constant for a given
production rate. However, the chemical dosage rate would have to be
closely coordinated with rinse water volume in order to prevent unnec-
essary overdosing due to volume increases which contained lower oil
concentrations. Furthermore, it must be pointed out that regardless
of how economically the plant can be operated for treatment of rinse
waters, the system cannot show improvement over the net credit result-
ing from batch treatment of Brand "A".
While the discussion of costs has been devoted to differences in
chemical requirements and the value of the recovered oil, these dif-
ferences must be viewed in the overall cost of operations. Other
costs associated with the treatment plant were noted to remain
essentially fixed whether the plant was handling concentrated coolant
or rinse water. Forty man-hours per week were required for operation
under both systems and averaged $927 per month or about $Q.02/ton.
Labor supervision, repairs, and utilities are each estimated to add
an additional $0.01/ton. Hence, a fixed operating cost of approximately
$0.05/ton is required irrespective of the nature of the waste or cold
mill operating procedure. The costs for the various treatment methods
are summarized in Table XIX.
TABLE XII
TREATMENT COST PER TON CF STEEL ROLLED ($)
Labor
Supervision
Repairs
Utilities
Chemicals
Oil Credit
TOTAL COST
Batch
Treatment
Brand "B"
.02
.01
.01
.01
•OU5
(.025)
.0?
Batch
Treatment
Brand "A"
.02
.01
.01
.01
.015
(.025)
.ou
Rinse Water
Treatment
Brand "C«
.02
.01
.01
.01
.01
--
.06
Although approximately a twofold difference is represented in the range
of treatment costs, the maximum difference is only $0.03/ton. From
this standpoint, therefore, it appears relatively unimportant as to the
preferred operation of the cold mill.
81
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Economics aside, the total quantity of pollutants to the stream over
a given time period must be considered. Here the batch treatment
prevails over rinse water. Even though the concentration of oil in
the effluent is greater for the batch treatment, the relative volumes
result in approximately three times as much oil to the stream over a
given time period when using rinse water.
A breakdown of the capital costs for the installed waste oil treatment
plant is given in Table XX. As has been indicated throughout this
discussion and based upon the findings of this study, a number of
changes are recommended if this plant were to be built again. Even
though operation of the cold mill on a recirculated coolant system
was found to be acceptable for current production requirements, the
possibility of going to a rinse water system to provide a quality cold
rolled product at some future date dictates that the corresponding treat-
ment facility be provided with a similar flexibility. Furthermore, despite
the fact that flow rates for a rinse water system were much less than
the 1,^)0 gpm plant design rate, the plant would still be built for the
Ij500 gpm flow as a result of management decision to incorporate this
magnitude of flexibility into the cold mill operation.
Several components of the described treatment system were implicated
as possibly being unnecessary. Among these were portions of the rapid
mixing equipment and the flocculation equipment. Since it was not
shown that these components would not be required at the maximum design
flow, such equipment would still be incorporated into another similar
facility contingent upon pilot plant findings. Aside from these uncer-
tain changes, other alterations definitely would be made. Reductions
in capital cost could be made by elimination of the incinerator and its
foundations and by minimizing chemical piping costs. However, increased
costs would result from the installation of larger equalization facili-
ties, including pumps, additional needed instrumentation and controls,
and more adequate piping insulation. The net result would be no signif-
icant change in capital cost within the bounds of estimated changes for
these additions and deletions.
TABLE H
TREATMENT PLANT CAPITAL COSTS
Yardwork $ JjQ,700
Structures 117,200
Permanent Equipment 1*95,700
Distribution Piping U9,100
Electrical 118,600
Indirect Costs 208,000
Contingency UOi7QO
TOTAL $1,070,000
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All factors considered, if the choice between using a recirculated
coolant system and a rinse water system on the first stand is avail-
able, treatment considerations definitely favor the recirculated
coolant system. If the decision to use a recirculated system can be
made prior to installation of the treatment facility, a significant
savings in capital expense can be realized. This savings would result
from the installation of a relatively inexpensive batch treatment
system in place of the more complicated and costly continuous system
described in this work.
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SECTION VI
ACKNOWLEDGEMENTS
The research studies and field activities reported herein were carried
out by the following employees of the Armco Steel Corporation:
Mr. J. E. Barker the Project Director; Mr. Grant A. Pettit the Technical
Consultant; Dr. John C. Hogan the Research Chemist; Mr. Verlin W. Foltz
the Project Chemist; Mr. Ronald J. Thompson the Project Engineer; and
Mr. Bruce A. Steiner the Assistant Project Engineer.
The employees of the Ashland Works of Arroco Steel Corporation are
recognized for their cooperation throughout the course of this study.
Kerns United is recognized for their cooperation in supplying infor-
mation which was of value in the identification of the composition
of typical rolling oil formulations. The Brooks Oil Company is, also,
recognized for supplying proprietary information and materials which
were of value during the experimental studies.
The partial support of this study by the Federal Water Pollution Control
Administrations, Project 12010 EZV, and the advice and assistance of
Mr. Robert L. Feder, Project Officer, and Dr. S. A. Hannah of the Taft
Sanitary Engineering Center are hereby acknowledged.
-------
SECTION VII
REFERENCES
1. Kawamura, S. and Tanaka, Y. "Applying Colloid Titration Techniques
to Coagulant Dosage Control," Water and Sewage Works, 113:3^8
(1966)
2. Private communication with Kern United Co.
3. Private communication with Dan Dick, Armco Steel Corp., Research
& Technology Division
U. Korpi, G. K. "Electrokinetic and Ion Exchange Properties of
Aluminum Oxide," p. 83 Dissertation 65-9UUU, Feb. 1965
5. Mysels, K. J. "Introduction to Colloid Chemistry," Cap. 17,
Interscience Publ., Inc., New York (1959)
6. Stumm, w. and Morgan, J. J. "Chemical Aspects of Coagulation,"
Jour. AWWA p. 971, Aug. 1962
7. Ludwig, J. R. "Changes in Cold-Reduction Lubricant Solution During
Rolling, " Blast Furnace and Steel Plant, p. 6iil-65l, Aug. 1969
8, Private Communication V. W. Foltz, Armco Steel Corp., Research &
Technology Division
9. Slezak, M. W., "Coagulation of an Oil-Water Emulsion Waste,"
M. £. Thesis, McMaster University, Dec. 1968
10. Harris, H. S., et al, J. Eng. Div.. Am. Soc. Civil Engs. , SA6,
p. 95-111, Dec. 19^5
11. S. A. Hannah, J. M. Cohen, and 0. G. Robeck "Floe Strength-
Measurement by Particle Counting," Presented in part at the 1965
Annual Meeting American Water Works Association, Portland, Oregon,
Ju]y, 1965
12. Mima, H. and Kitamori, N. "Stability and Flocculation of Oil
Droplets in Dilute Emulsions," J. Pharm. Sci., 55, No. 1, p.UU-l*8,
Jan. 1966
13. Higuchi, W. J., Okada, R., and Tembergen, A. P. "Aggregation in
Oil-in-Water Emulsions," J. Pharm. Sci.. 5l, No. 7, p. 683-687,
July 1962
lU. Ritchie, A. R., "Certain Aspects of Flocculation as Applied to
Water and Sewage Purification," Dissertation, University of London,
Oct. 1955
87
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15. Morgan, J. J. et al, "Flocculation Behavior of Dilute Clay- Polymer
Systems," Progress Report WP 009U2-02, 15 Aug. 1966
16. Camp, T. R. , "Flocculation and Flocculation Basins," Trans. ASCE,
120. p. 188 -
17. Zeta-Meter Manual, Zeta-Meta, Inc., New York, N. Y. (1968)
18. Riddick, T. M. "Control of Colloid Stability Through Zeta Potential/'
Vol I, Ch. k,9, Livingston Publ. Co., Wynnewood, Pa. (1968)
19. Harris, H. S., et al "Orthokinetic Flocculation in Water Purifi-
cation," Sanitary Eng. Div. , SA6, p. 95 ff, Dec. 1966
20. Teot, A. S. and Daniels, S. L., "Flocculation of Negatively Charged
Colloids by Inorganic Actions and Anionic Polyelectrolytes,"
Environmental Sci. and Tech., 3, No. 9, 783-868 (1969)
ADDITIONAL LITERATURE SURVEYED
Anon., Ind. Water Eng., 5, 22 (July, 1968)
Anon., Lubrication. Ii3, la (1957)
Balden, A. R., J. Water Pollut. Contr. Fed.. 1O_, 1912 (1969)
Funk, B. W., U. S. Patent 3,23U,UU6 (March 29, 1966)
Kovacs, 0. L., (To New Canadian Processes Ltd.), U. S. Patent 3,301,779
(Jan. 31, 1957)
Kovacs, L. , U. S. Patent 2,806,868 (Sept. 17, 1957)
Kovacs, L., U. S. Patent 2,807,531 (Sept. 21|, 1957)
Leidner, R. N., J. Water Pollut. Contr. Fed.. 38, 1967 (1966)
"Manual on Disposal of Refinery Wastes," Vol. I, American Petroleum
Institute, New York, N. Y. (1957)
Operating Practices Committee - Cold Mills, Amer. Iron and Steel Inst.,
Internal Report
Page, L. J., Proc. Ind. Waste Conf . . Purdue Univ.. 50, No. 2,33 (1966)
Tiberio, Q. N., Ibid. . 50, No. 2, UUO (1966)
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SECTION VIII
GLOSSARY
LIST OF TERMS
1. As-prepared emulsion - emulsions prepared in the laboratory.
2. Commercial rolling oil - rolling oil formulations, prepared by
vendors and usable on commercial cold mills.
3. Coolant - Same as mill used emulsions.
lj. Creaming - the partial separation of emulsion which in a very
real sense is analogous to the separation of cream from un-
homogenized whole milk.
5. Mill spent emulsion - recirculated emulsions used on the mill
for such a period of time as to be unsuitable for rolling.
6. Mill used emulsion - recirculated emulsions used on the mill
but still suitable for continued use.
7. Optimum - the value of a given parameter or set of parameters
leading to maximum treatment efficiency as determined by observation.
8. Rinse water - water utilized for purposes of removing undesirable
substances from the steel strip at stands No. 1 and/or No. £.
Generally, rinse water contains less than 1,000 ppm total oil.
9. Tramp oil - unemulsified oil which usually floats out of the
emulsion at a fairly rapid rate on standing. Tramp oil may
either be broken previously emulsified oil or extraneous oil
which has leaked into the system. Tramp oil is usually dark
in color.
10. "Typical" rolling oil - rolling oil formulations, with or with-
out emulsifier which are prepared from mineral oil and lard oil
in proportions typical of commercial rolling oils.
11. Virgin rolling oil - rolling oil emulsions (usually commercial)
which have not been used on the mill.
39
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LIST OF ABBREVIATIONS
A area of flat plate paddles (ft. )
AN acid number of oil, ASTM D117, Comm. D-2, p. UO, 9th Ed.,
Aug. 1969
Cd drag coefficient (dirnensionless), for flat plates Cd-1.2
C^ Concentration of 1,060 ppn emulsion after dilution
C. as prepared concentration in ppm
D particle size distribution function (diraensionless)
fps feet per second
Gn mean velocity gradient (sec" )
JTU measured turbidity in Jackson Turbidity Units
JTUd measured turbidity of diluted emulsion
JTUe measured turbidity after partial flocculation (5 min.)
JTUj measured turbidity after 5 min flocculation
empirically calculated post-flocculation turbidity
measured as-prepared, initial untreated, or "influent" turbidity
empirically calculated value as-prepared turbidity
JTU0 turbidity of original emulsion, subsequently diluted
k flocculation rate constant (N units)
K ratio of rotating velocity of fluid to rotating velocity of
flat plate paddles, equal to 0.32 according to Camp (16)
K2 a constant equal to W/u (sr (sec/ft^)
# ,
K composite rate constant equal to kD (N"1 units)
m number of sequential reactors
n revolutions per minute of stirrer shaft
N number or mass of particles per unit volume at time, t
90
-------
N number or mass of particles per unit volume at t=o
ppm parts per million of solute or dispersed phase by weight
rb distance from center of paddle to center of rotating shaft (ft)
a stirring tip speed of paddle (ft/sec)
S.C. specific conductivity (micromhos)
t time (sees)
t/m retention time per sequential reactor (sees)
2
u absolute viscosity (ib-sec/ft )
V volume of reactor vessel
W power dissipated per unit volume per unit time (ft-lbs/sec-ft^)
Zp zeta potential as determined directly on emulsion (mv)
Zpi zeta potential of untreated reference emulsion state (mv)
Zpf zeta potential of emulsion in absorption studies after treatment
with alum (mv)
Zpo zeta potential at "infinite dilution" (less than 30 ppm), in mv
91
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SECTION IX
APPENDICES
APPENDIX A - RESEARCH AND DEVELOPMENT WORK
Description of Problem
A process is required to treat oily waste water from the new five-
stand tandem cold mill at the Ashland Works. This process must be
capable of treating 1,600 gallons per minute (gpm) of water containing
up to 1,000 parts per million (ppm) of pickler coating oil from stand
No. 1 and 1,600 gpm of water containing up to 1,000 ppm emulsified
rolling oil from stand No. 5>. The process must work at any concen-
tration up to 1,000 ppm of any type or brand of oil and be capable of
handling other miscellaneous oily wastes, including periodic dumps of
spent rolling emulsion from the recirculating coolant systems containing
up to £ percent emulsified oil.
Review of Earlier Work
Several years ago, a great deal of experimental work was carried out
on the treatment of aqueous wastes from the four-stand tandem cold
mill at Middletown Works* Since coolant was being recirculated on
all four stands of this mill, the only waste studied was the spent
coolantj namely, water containing 1-5 percent emulsified oil.
The following scheme was developed:
(a) Continuous skimming of floating "tramp" oils from the emulsion in
the recirculating system.
(b) Clarification of the oily skimmings at elevated temperature
(e.g.,160 F) to reduce the moisture and solids contents to low
levels.
(c) Cracking discarded emulsion with an organic polyelectrolyte
emulsion breaker, with the oily "sludge" thus produced being
sent to the skimmings clarifier and the clear water being sent
to the sewer. The above process was tested on a pilot scale
at that time and has been recently applied in a full-scale plant.
Selection of a Treatment Process for the Ashland Cold Mill Wastes
An attempt to apply the previously-studied polyelectrolytes to wastes
containing 1,000 ppm emulsified oil was unsuccessful; no breaking of
the emulsion took place. Further investigation showed that these
chemicals were not effective at oil concentrations below about 1 per-
cent (10,000 ppm).
A new series of organic polyelectrolytes, claimed to be effective at
low oil concentrations, were obtained and applied to this waste with
93
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limited success. An acceptable water layer could be produced if the
emulsion was treated and held at 100 F for 2i| hours, then allowed to
stand for another 2 or 3 days at room temperature. This approach is
obviously not practical for treatment of a 1,600 gpm stream.
Another process which showed promise consisted of:
(l) Production of a batch of oily sludge by cracking a concentrated
emulsion (e.g., 5 percent) with an organic polyelectrolyte.
(2) Treatment of the dilute waste emulsion by mixing with this sludge,
then allowing the sludge to separate.
Although several series of experiments along these lines were success-
ful, it became apparent that a considerable amount of research would
be required to develop a reasonable understanding of the more funda-
mental factors involved in optimizing a process of this sort. This
approach was abandoned in favor of coagulation.
It was found at bench scale that coagulation, using a ferric iron salt
or alum as the primary coagulant, could be applied successfully to
virtually any cold mill waste water containing a very broad range of
concentrations of almost any type of oil. This process is readily
adaptable to continuous operation and functions relatively quickly.
It appeared that this process could be developed and applied to a
full-scale plant within a reasonable period of time. Therefore, this
line of investigation has been studied in considerable detail as
reported in the following pages.
Description of Coagulation Process
The process of chemical coagulation of dilute oil-in-water emulsion
is not completely understood. Several mechanisms appear to be involved;
and as the conditions under which coagulation takes place are changed,
the relative importance of each mechanism probably changes also.
The emulsion is first destabilized by the addition of a positively
(and multiply-) charged ion called a primary coagulant. Ferric salts,
alum, and cationic polyelectrolytes perform this function, which is
essentially one of reduction of electrical charges on the oil droplets.
Oil droplets are also absorbed on a precipitate of ferric or aluminum
hydroxide; coagulation appears to work best in the pH range in which
the hydroxide is least soluble. At this stage of the process a "pin-
point" floe is normally visible, though the water is still very turbid.
The process is then carried from coagulation to flocculation by the
addition of one or more "coagulant aids" to consolidate the insoluble
materials into large floe aggregates. This function is performed by
certain clays and long-chain organic polymers. Best results are often
obtained by treating first with a clay, then with a polymer. The clay
appears to promote floe growth from pinpoint up to a clearly visible
size by adsorption. Further floe growth may then be induced by adding
9h
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the polymer; the main action of the polymer appears to be mechanical
entrapment, although the slight superiority of an anionic polymer over
a nonionic suggests a function as co-coagulant working with the
cationic materials.
Experimental Procedures
Nearly all of these studies were carried out by the jar test technique,
in which a series of Identical test specimens are stirred with a gang
mixer while different quantities of chemicals are added to the specimens.
Since paddle sizes, container sizes, mixing times and mixing speeds are
identical, the effects of the variations in chemical dosages can be
determined visually.
In this case, a 500 ml specimen in an 800 ml beaker gave a specimen
volume of about 3-3/U" in diameter by 3" deep. The mixer was equipped
with four 1" wide (top to bottom) paddles, each sweeping a 2-1/2"
diameter volume; speed was continuously variable up to 162 rpm (approxi-
mately 1-3/U ft. per second maximum tip speed). The mixer was operated
at maximum speed when "rapid-mixing11 and at about UO rpm when flocculating.
Evaluation & Selection of Primary Coagulants
Four chemicals were evaluated as primary coagulants; namely, ferric
sulfate, aluminum sulfate (alum), and two cationic polyelectrolytes.
Eleven other cationic polymers were examined but did not perform satis-
factorily.
One of the cationic polymers, EB-2, appeared to be more economical than
the other polymers (based on performance). In some cases, EB-2 could
be substituted for a major portion of the inorganic coagulant at a
savings in total chemical costs; this would also appreciably reduce the
ash of the reclaimed oil. In other cases, EB-2 could not be substituted
for any portion of the inorganic coagulant.
Of the two inorganic coagulants, alum is preferable to iron salts. At
concentrations normally used for coagulation, a good alum floe can be
developed at a pH of 6, whereas an iron floe develops best at high pH
values; therefore, less lime is required for pH control of alum coagu-
lation. Also, an alum floe tends to float after coagulation of even
trace amounts of oil, whereas an oily iron floe may float or sink;
either floe could be recovered by dissolved air flotation, but an
oily alum floe would normally float at all times, even if for some
reason the air flotation equipment were out of operation.
All further work was carried out using alum and, where economical,
EB-2 as the primary coagulants. Lime was also added along with the
alum, where needed, to keep the pH above 6.
-------
Evaluation & Selection of Coagulant Aids
Two types of coagulant aids were studied, clay and organic polymers.
Bentonite clays from six different suppliers were examined; but by
far the best performance was obtained by using one of the montmorillonite
type.
Six different coagulant aids of the nonionic and anionic polyacrylamide
type were applied to this process. When evaluated on an equal-cost
basis, the anionic performed best with the nonionic performing nearly
as well in most cases.
All further work was carried out using clay as necessary to promote
the growth of the floe from the "pinpoint" size produced by the primary
coagulants up to a size which could be successfully "blossomed" with
a long chain polymer. Only minimum amounts of clay were used, however,
since clay obviously adds to the ash of the reclaimed oil.
Development of the Coagulation Program
The next step in this investigation was to determine the important
parameters in the five-chemical program: cationic polyelectrolyte -
alum - lime - clay - nonionic or anionic polyelectrolyte. Since
these investigations involved over 300 separate experiments, the data
presented in this report are limited to those selected to illustrate
important points.
The results of these experiments indicate:
(l) In some cases, EB-2 can be substituted for a portion of the alum.
This seems to depend on brand of oil (type of emulsifier).
(2) Where used, EB-2 addition should be followed by five-minute
rapid paddle-type mixing prior to alum addition.
(3) Addition of lime, if used, and alum should be followed by another
five minutes of rapid paddle-type mixing. Propeller mixing, even
for the first one minute, resulted in the eventual production of
an inferior floe (smaller, more poorly agglomerated, and slower
to rise).
(U) Addition of clay should also be followed by five-minute rapid
paddle-type mixing prior to addition of polymer coagulant aid.
(5) If clay is added at the same time as the alum, the effectiveness
of the clay is slightly reducedj the five-minute mixing times
between cationic polyelectrolyte and alum additions and between
clay and polymer coagulant aid are critical, however.
(6) Addition of nonionic or anionic polymer should be followed by a
short rapid mix and then five minutes slow mixing (flocculation).
96
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(?) Proper dosage of chemicals depends so much on type and concentration
of emulsified oil (amount of emulsifier) that treatment of No. 1
and No. 5 stand wastes as two separate streams would be highly
desirable. (Table Al)
Proper dosage also depends on history of emulsion (Table AI) and
analysis of water (Table All) as well as concentration of oil.
Final adjustment of dosages will also depend on ultimate disposal
of oil; the plant should be optimized for minimum operating costs
if the oil is to be destroyed, but for minimum ash in oil if oil
can be used for fuel. The plant should be capable of feeding the
following dosages:
Probable Maximum Dosage Required, ppm
Stand No. 1 Waste Stand for No. 5 Waste
Cationic Polyelectrolyte 10 10
Lime - 150
Alum * 30 500
Clay 50 50
(Nonionic Polymer 1 2
( or
(Anionic Polymer 1 1-1/3
* Computed as ASOj . 18 HO
Application of Dissolved Air Flotation
Several typical coagulation experiments, involving both pickler coating
and rolling oils, were concluded with the application of dissolved air
flotation to the final floe. In every case in which a satisfactory floe
had been formed, the floe was floated successfully with a "too fast to
measure" rise rate. It was found, however, that the final flocculation
step (with nonionic or anlonic polymer) had to be carried out in the
same container in which flotation was to be applied; the floe was too
fragile to withstand a transfer to another container.
Disposal of Oily Sludge
The sludge produced by this process can exhibit a very broad range of
properties (moisture content, viscosity, ash, etc.).
At one extreme, a sample of stand No. 1 waste required very little
chemical treatment to produce a sludge which dried relatively quickly
to a product suitable for use as boiler fuel.
At the other extreme, a sample of stand No. 5 waste which required 350
ppm alum and 50 ppm clay to coagulate (at 1,000 ppm oil) produced a
sludge which was 90 percent water, could be dried in five days at 160 F
to 25 percent water, had an ash of 9 percent (based on dry oil), and
at room temperature looked remarkably like ordinary axle grease,
97
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It is not known at what point the recovered oil will actually fall
between those extremes. It appears advisable at this time to plan to
clarify at 160-180 F with mixing and to prepare to dispose of recovered
oil by incineration.
Future studies will be directed to producing a type of recovered oil
suitable for fuel usage.
Handling of Concentrated Waste Emulsion
Certain brands of rolling oil, such as those in use at Middletown Works
which were referred to earlier in this report, are relatively easy to
break from emulsion at the 1-5 percent oil level. Others, such as the
brand "B" product currently used at Ashland, do not respond to this
treatment.
It would appear necessary, then, to provide a tank to contain a "dump"
from the recirculatiiig system until it can be treated. If the emulsion
can be broken at ambient temperature by addition of a polyelectrolyte,
the emulsion could be cracked in this tank, the phases allowed to
separate, the clear water sent to the sewer, and the oily sludge sent
to the oil recovery area. Otherwise, the emulsion could be bled
slowly into the stand No. 5 waste stream, where treatment by coagu-
lation would take place. This flexibility of treatment schemes is
necessary for maximum efficiency of operation.
Control Testing Procedures
Two simple and basic control tests for this process are the pH through-
out the process (not less than 6.0) and the turbidity of the effluent
to the sewer.
To date, however, no procedures have been developed to replace jar
testing for determination of chemical dosage. It may prove difficult
to train a plant operator to jar test a five-chemical program.
Good control testing is important, because underdosing of a single
chemical will almost always hurt both water and oil quality; overdosing
will in some cases hurt water quality, and in most cases will raise
the ash content of the oil as well as increase operating costs.
Methods are needed which will minimize or eliminate operator judgment
in determining dosages of at least two of the chemicals.
Summary and Conclusions
Coagulation appears to be the best approach to the treatment of waste
of the type that will be produced by the Ashland Works tandem cold
mill.
In order to be capable of handling the large variety of wastes which
may be sent to this treatment plant, the plant should be designed for
98
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a five-chemical program:
(1) Cationic polyelectrolyte - as a primary coagulant, where economical
to substitute for part of the alum.
(2) Alum - as the main primary coagulant.
(3) Lime - for pH control at large alum dosages.
(U) Clay - as a coagulant aid to promote growth of the floe.
(5) Nonionic or anionic polyelectrolyte - to further promote growth
of floe.
The selection of chemicals to be used and dosages at which to apply
them depend upon type of oil (amount of emulsifier), brand of oil
(nature of emulsifier), concentration of oil, history of emulsion,
water analysis, and ultimate disposal of reclaimed oil.
Since the stand No. 1 waste requires far less in the way of treatment
chemicals than the stand No. 5 waste, it would be advisable to keep
the streams separate and treat them separately.
Addition of cationic polyelectrolyte should be followed by five minutes
of rapid paddle-type mixing. Addition of alum (and lime added simul-
taneously, if needed) should also be followed by five minutes rapid
mixing. Addition of clay should also be followed by a five minute
rapid mix. It is possible to add the clay with the alum and still
achieve good coagulation, but it is important to mix a full five
minutes after both cationic polyelectrolyte and clay. Addition of
nonionic or anionic polymer should be followed by five minutes of
slow mixing in the same vessel in which dissolved air flotation will
take place.
Scum recovered from the air flotation units should be transferred to
a heated (160-180 F) mixer-clarifier providing a few days retention
time. '*ftiether the product reclaimed oil will be suitable for use as
boiler fuel or will have to be destroyed by incineration can only be
determined after start up.
A separate tank should be provided to contain the periodic "dumps"
from the recirculation system. This waste may then be handled one of
two ways:
(a) If the emulsion is of the type which can be broken with a
polyelectrolyte emulsion breaker, the emulsion can be broken
and oil separated in the tank, with clean water going to the
sewer and oily sludge to the oil receiving area.
99
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(b) If the emulsion is of the type which does not respond to poly-
electrolyte emulsion breakers, the emulsion can be disposed of
by bleeding into the stand No. 5 waste stream and coagulating.
A few additional simple precise control testing procedures would be
very helpful.
Recommendations for Further Work
1. It would be highly desirable to have some testing methods for
plant control which are more objective and precise than visual
inspection of the floe and water during jar tests. If coagulant
dosage could be based on a colloid titration technique, for
example, and lime dosage based on coagulant dosage, only clay
and anionic polymer dosages would have to be chosen by operator
judgment based on jar tests.
2. Whether oil can be reclaimed for use as boiler fuel cannot be
determined until after plant start up. At that time, data should
be collected on oil concentrations in the wastes, the amounts of
"miscellaneous" oils coming into the system, brands of oil in
use, frequency of coolant "dumps", etc. Then the chemical dosages
in the plant should be optimized to give minimum ash in the oil
(if it can be used as fuel) or minimum plant operating costs
(if oil is to be destroyed by incineration).
100
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EFFECTS OF VARIOUS PARAMETERS
ON MINIMUM CHEMICAL DOSAGES REQUIRED TO TREAT
WASTES CONTAINING 1.000 PPM OIL
Table AI - Various Sources and Types of Oil
Chemical Dosages, ppm
Specimen
Brand "A"
Rolling Oil
Brand "A"
Rolling Oil
Middletown Cold Mill
Brand "B"
Rolling Oil
Brand "B"
Rolling Oil
Ashland Cold Mill
(Slow Speed)
Same, but mill running
at faster speed
EB-2
7
10
Nonionic
Alum* Lime Clay Polymer
20
20
60
70
335
10
12$
10
10
20
20
Pickler Coating Oil
Ashland Cold Mill Stand #1
(Sample fortified with new
oil to build cone, to 1,000
ppm)
10
20
0.5
* Expressed as A12(SO^)
101
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EFFECTS OF VARIOUS PARAMETERS
ON MINIMUM CHEMICAL DOSAGES REQUIRED TO TREAT
WASTES CONTAINING 1,000 PPM OIL
Table All - Various Types of Water
Specimen:
Water
Lime - Soda
Softened
1 part tap #
1 part distilled
1 part tap
2 parts distilled
1 part tap
5 parts distilled
Used Brand "B" Rolling Oil,
Ashland Cold Mill operating at
slow speed; emulsion from recir-
culating system diluted to 1,000
ppm with water from various
sources.
Chemical Dosages, ppm
Lime
10
15
30
10
Alum
70
100
100
80
Clay
20
Nonionic
Polymer
20
* Middletown, Ohio, City Water
102
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APPENDIX B - DETAILED ENGINEERING REPORT & DRAWINGS
INTRODUCTION
This Waste Water - Waste Oil Treatment Plant is primarily designed
to prevent stream pollution due to contaminated water and coolant
(rolling oil), resulting from the new five-stand Tandem Cold Mill.
As described below, this treatment plant is designed as a two-step
treatment process. The first step consists of coagulation and
flocculation of oil, followed by an air flotation process. This
separation of oil and water takes care of the problem normally asso-
ciated with cold mill operations.
GENERAL DESCRIPTION
The Waste Water - Waste Oil Facilities are designed for a capacity
of 1,600 gpm and will consist of underground pipe lines to carry the
emulsified oil from the new five-stand Tandem Cold Mill to the oil
treating facilities located southwest of the existing Maintenance
and Diesel Repair Shops Building No. 729.
Spent rinse water from the five-stand Tandem Cold Mill will be pumped
to an equalization tank, thence by gravity to the rapid mix unit
(2 stage), then by gravity to the flocculation unit and by gravity to
the air flotation unit. Cationic polyelectrolyte, clay, alum, and lime,
will be fed into the influent of the rapid mix unit in proportion to
the flow. Coagulant aid will be fed into the influent of the floccu-
lation unit.
The periodic blow-down of spent coolant from stands No. 2, 3, and U
of the Tandem Cold Mill will be pumped into a new 60,000-gallon holding
pit. The holding pit will be provided with a 500 gpm recirculating
pump, and suitable valving and pumping equipment to transfer the combined
waste oils to the equalization tank at a rate of 20 gpm. When this
occurs additional chemicals must be added to the influent of the
rapid mix tank to compensate for the additional flow.
The effluent from the air flotation system will be sent to White Oak
Creek. The bottom sludge, consisting of dirt and water, from both
systems will be periodically drained by gravity to dumpster boxes
for disposal at the dump.
The scum from the unit will be drained to a 800-gallon collecting
tank and pumped to a waste oil treatment and disposal system.
The chemical storage, dilution and feeding facilities required to
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support the above operations are, a cylindrical acid resistant liquid
alum storage tank, a cylindrical cationic polyelectrolyte storage
tank and two cylindrical dilution tanks for each of three chemicals;
coagulant aid, lime and clay. Each dilution tank will hold a 2U-hour
supply and will be equipped with suitable mixing equipment to dissolve
the chemicals in water and keep them in suspension. Suitable pro-
portioning pumps and spares will be provided to feed the chemical
solutions and slurries to the rapid mix units and flocculators.
Sufficient dry chemical storage space will be provided to handle a
20-day supply of coagulant aid, lime and clay.
An enclosure will be provided for all chemical storage, tanks and
feeding equipment, the 60,000-gallon holding pit and oil treatment
facilities.
Water for chemical mixing will be supplied from the recycle system
of the air flotation tanks. The initial filling will be from the
city water line.
The waste oil treatment and disposal system will consist of pumping
the scum from the 800-gallon collecting tank to a heated clarifier.
The separated oil will flow by gravity to a 20,000-gallon oil holding
tank. From the holding tank the oil will flow by gravity to a 300
gph incinerator for disposal.
CONTROL
The influent lines to the equalization tank will have a flow meter
to control the three-way solenoid operated valves on the suction side
of all chemical feed pumps. These three-way solenoid operated valves
admit either chemical or water to the pump suction as required. The
capacity of all chemical feed pumps is adjusted manually. Scum
pumps will be float controlled. The cationic polyelectrolyte, alum
and oil storage tanks will have a liquid level gauge. All recycle
pumps, storage pit effluent pumps and the alum unloading pumps will
be manually controlled. All other functions are by gravity. Temper-
ature of the heated clarifiers will be automatically maintained by
temperature control valves.
Effluent to White Oak Creek will be monitored automatically and
continuously with an effluent turbidimeter.
ENCLOSURE
The Waste Water - Waste Oil Treatment Enclosure will be 1|0 feet wide
by 100 feet long by 20 feet high for a length of 80 feet, and U6 feet
high for a length of 20 feet. The enclosure will contain the oil
holding pit, alum storage tank, cationic polyelectrolyte storage tank,
10U
-------
storage area and mixing tanks for coagulant aid, lime and clay, plus
the oil recovery equipment.
MAJOR EQUIPMENT - MECHANICAL
Equalization Tank
The equalization tank will be designed for 1,600 gpm at a retention
time of 10 minutes and will be capable of 2U-hour continuous operation.
The tank will be 12 feet diameter by 19 feet vertical sidewall depth
including 12 inches freeboard.
There will be a recirculating pump to equalize any variations in oil
concentration. The effluent will flow by gravity to the rapid mix
tanks.
Rapid Mix Tanks
The pair of rapid mix tanks will each have a capacity of 7,850 gallons
and will be 11 feet in diameter by 12 feet vertical sidewall, including
12 inches freeboard, with a U5 degree cone bottom. Influent and
chemicals will enter the tank through a standpipe. Each tank will be
equipped with a double turbine, 60 rpm mixer. The rapid mix tanks
will be operated in series. Effluent from the No. 1 tank will flow
to the No. 2 tank by gravity and from the No. 2 tank to the floeculation-
air flotation tanks by gravity.
Flocculation-Air Flotation Tanks
Each of the two flocculation tanks will be designed for 800 gpm at a
retention time of 10 minutes, with paddle type mixers sufficient to
promote the growth of floe. Each tank will be 11 feet long by
10 feet wide by 7 feet 2-inch vertical sidewall, including 6 inches
freeboard, with a vee bottom. Each of the two air flotation tanks
will be designed for 800 gpm at a retention time of 20 minutes for
influent plus 25 percent recycle and will be capable of 2U-hour
continuous operation. Each tank will be 10 feet wide by Ik feet long
by 7 feet 2-inch vertical sidewall depth including 6 inches freeboard.
The bottom of the tanks will be a U5 degree vee with a screw conveyor
for collection of sludge.
There will be a skimmer operating over the entire top of the tank for
removal of the accumulated oil. The scum will be drained by gravity
to the 800-gaUon collecting tank. A recycle pump will be used to
withdraw a portion of effluent and pump it to a pressure tank where
air will be added. The recycle stream and raw flow will be blended
at the inlet manifold.
Holding Pit
The holding pit will have a capacity of 60,000 gallons. The pit will
105
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be 18 feet wide by 38 feet long by 13 feet effective depth, reinforced
concrete, with the bottom sloped to a clean-out area within the pit.
The pit will have a 500 gpm recirculating pump. Effluent will be
pumped alternately by one of two 20 gpm pumps to the equalization
tank. Any solids which may accumulate on the bottom will be removed.
All pumps will be of the submersible type with provisions made for
easy removal for maintenance.
Scum Tank
The scum from the air flotation tanks will flow by gravity to the
800-gallon scum tank. Effluent from this tank will be pumped alternately
by one of two 25 gpm pumps to the oil recovery system.
Heated Clarifier
The heated clarifier will have a capacity of 20,000 gallons and will
be 12 feet in diameter by 22 feet vertical sidewall, with a U5 degree
cone bottom and drain valve. The clarifier will be insulated.
There will be a 6 feet diameter by 7 feet deep, steam heated, plate
coil, center well with a h feet by k feet, 5.2 to 20.6 rpm paddle
type mixer in the center. Effluent will overflow a weir at the top
of the clarifier to the holding tank. Access platforms will be provided
to the top of the clarifier and holding tanks and also monorails and
manually operated hoists for maintenance of the mixers and coils.
Holding Tank
The oil holding tank will have a capacity of 20,000 gallons and will
be 12 feet in diameter by 22 feet vertical sidewall, with a U5 degree
cone bottom and drain valve. The holding tank will be insulated and
heated with a 6 feet diameter by 7 feet deep, steam heated, plate
coil, center well. Oil will be pumped from the tank with a 0 to 300
gph pump to the incinerator for ultimate disposal. In the future,
the holding tank can be converted to a second clarifier if it should
prove desirable, and two additional 10,000 gallon holding tanks
added.
Incinerator
The incinerator will have a capacity of 300 gallons per hour and
designed to consume the scum from the 20,000-gallon holding tank.
Anticipated operation of the incinerator will be intermittent. The
incinerator will be complete with a target wall for personnel
protection.
106
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PAGE NOT
AVAILABLE
DIGITALLY
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APPENDIX C - KINETIC RELATIONSHIPS EMPLOYED
The basic relationship for flocculation kinetics is (18):
= exp -
o
where :
N « number or mass of particles per unit volume at time, t
N » number or mass of particles per unit volume at t « o
k = rate constant in N~ units
D - particle size distribution function
C^ • mean velocity gradient in sec"
t * time in sees
The log. form of this relationship is:
C
in (N/N0) - -kDN0Gnt
The nature of this equation is such that a working engineering
relationship may be attainable if the values of N and NO can be
expressed in terms of some measureable parameter such as turbidity.
For example, in this work it was noted that for a diluted as-prepared
typical rolling oil emulsion:
. -o . JTUd Cri . N
1.18 In _ 2. - In _£ - In «
JTU0 C± W0
where :
JTU0 - turbidity of 1,060 ppm emulsion
JTUd - turbidity of diluted 1,060 ppm emulsion
C^ » concentration of 1,060 ppm emulsion
Gd » concentration of diluted 1,060 ppm emulsion
12?
-------
Therefore, if D remains constant during coagulation,
1.18 In (JTUe/JTU ) - -kDNQGt
where JTUe and JTl^ are the "effluent" and "influent" turbidities,
respectively.
The mean velocity gradient, G^,, is defined as:
(wm/u) -
where
W - the mean power dissipated per unit volume per unit time in the
system
u • absolute viscosity
According,,to Camp (16) if W is expressed in ft-lbs/sec -ft3 and u in
Ib-sec/ft ;
w m 239 Cd (1-K)3 (n)3 Sum (A(rb)3)
_
where :
Cd * 1.2 and is based on the drag coefficient for flat plates
A - Area of vertical paddles (ft2)
r. - distance from center of paddle to center of rotating shaft (ft)
K - 0.32
n • revolutions per minute of the stirring shaft
V - volume of the reactor vessel in ft3
Obviously, the tip speed Is related to n through the relationship
s (ft/sec) • n x U (pi) rb (r^ = 1/2 length of paddle)
If all the paddles are of identical size one obtains
where K2 includes the volume of the system, area and number of paddles,
the absolute viscosity of the system and conversion constants.
Because of the tight restrictions required to compare systems on the
basis of stirring tip speed, results should preferably be reported in
terms of mean velocity gradient.
128
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APPENDIX D - KINETIC RELATIONSHIP FOR SEQUENTIAL REACTORS
The flocculation kinetics relationship for a sequential continuous
flow reactor system is (18):
* - d*wxyi0 t/m) -">
where:
N - number of particles per unit volume at time t
N0 • number of particles per unit volume at t • 0
k - rate constant in N"1 units
D « particle size distribution function
Gm • mean velocity gradient in sec"**
m - number of sequential reactors
t • time in seconds
t/m » retention time per reactor, sees
The flocculation kinetics relationship for the batch reactor type
appears in Appendix C.
129
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APPENDIX E
USE OF SINGLE CHANNEL COULTER COUNTER FOR FLOCCUIATION KINETICS
It is generally assumed that during flocculation, coagulated particles
or aggregates thereof can combine irreversibly to form larger aggregates
and that particles and aggregates of virtually neutral charge have
an equal probability of flocculating. On this bases one should be
able to determine flocculation kinetics by following the decrease
in the number of particles in the system with tljne after "coagulant"
addition. For poly-dispersed systems kinetics data should be obtain-
able by monitoring the number of particles in a small and narrow size
range wherein flocculation of particles having diameters less than
the upper limit of the range necessarily result in new particles or
aggregates having mean diameters greater than that upper threshold.
For this purpose a Coulter Counter (Model B) was employed.
The flocculation studies were conducted on emulsions of a laboratory
prepared commercial rolling oil treated with alum. These results
appear in Figures E1-E3. The values for the final slopes necessarily
include the relatively small but unknown mean velocity gradients.
Further such experiments were run with both alum and lime addition
to produce alum floe. In these instances the rate of particle loss
within the 6.^2 to 9.U1 micrometer range increased rapidly from the
typical values given in Figures E1-E3 (about -10-k sec-1) to values
of -0.6 sec'1 and greater.
Because of the unknown velocity gradient under these conditions, a
more sophisticated reactor-unit was sought. Such a unit was obtained
on loan from Dr. S. A. Hannah of the FWPCA, Taft Engineering Center,
Cincinnati. This equipment was previously used to measure floe
strength and has been described elsewhere by Dr. Hannah. (11)
Unfortunately the instrumental noise encountered in attempting to use
this piece of equipment was too excessive for meaningful work. While
this problem probably could have been overcome with a relatively
small expenditure of time, too little time remained within the grant
period to perform enough reaction studies to be of value. On the
other hand it is clear from what has been presented above that the
single channel Coulter Counter can be used for flocculation kinetics
studies.
131
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FIG. E-l. FLOCCULATION OF 200 PPM EMULSION WITH
ALUM VIA COULTER COUNTER TECHNIQUES
(00.U5 meqs alum/g emulsion)
Range
Range
Range
U.10 - 0.93 um
5.93 - 7.03 urn
7.03 - 7.86 urn
10-14 sec'1
8
s
x 10" sec
10
0
Ik x 103
TIME, SECS
132
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10
FIG. E-2 FLOCCULATION OF UOO PPM EMULSION WITH ALUM VIA
COULTER COUNTER TECHNIQUES
(@ 0.225 neqs alum/g emulsion)
O Range 5.1U - 7.U1 um
D Range 7.U1 - 8.81 \m
A Range 8.8l - 9.83 um
8
3
103
slope =• -1.7U x 10" sec
-1
10
12 x 103
TIME, SECS
133
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FIG. E-3.
FLOCCULATION OF 1,000 PPM EMULSION WITH ALUM
VIA COULTER COUNTER TECHNIQUES
(@ 0.09 meqs alum/g emulsion)
o Range 6. £2 - 9. la urn
Range 9.hi - 11.5
urn
sec
-1
103
6 8
TIME, SECS
10
12
1U x 103
13U
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APPENDIX F - SAMPLES AND REAGENTS
Alum - 99.7 percent purity Al2(SOii)-j . 18 H20 prepared at 10,000
ppm in distilled deionized water
Line - Reagent grade Ca(OH)2 prepared at about 950 ppm level in
distilled deionized water and filtered hot over 0.2 um
Millipore Filter pads. Concentration was determined for
given stock solutions by titration against standard
0.00551N hydrochloric acid with phenolphthalein as
indicator.
Clay - A montmorillonite prepared at $,000 ppm and dispersed
1.0 min. via a single speed Waring Blendor.
Mill Mix Water-Lime softened Ohio River water, departiculated via
filtration over 0.2 um Millipore Filter pads. This
water had a specific conductivity (s.c.) of about $00
micromhos and a pH of approximately 7.U.
Deionized Water - pH 5.2 to 6.2 (depending on C02 content), and s.c.
ca. 10 micromhos.
C02-Free Deionized Water - Boiled deionized water cooled in stoppered
flask, pH 6.78, s.c. ca. 6 micromhos.
Electrolyte for particle size determination -1.0 weight percent
reagent grade NaCl in deionized water departiculated to
0.2 um and having a pH of ca. 6 and an s.c. ca. 20,000
micromhos.
Stock "typical" rolling oil solution (emulsifier-free) - £0 i 0.5
percent in lard oil (of various acid numbers),
SO- 0.5 percent in mineral oil dissolved in chloroform to
a total concentration of 200 g per liter of solution.
"Typical" rolling oil emulsion (emulsifier-free) - 20 ml aliquot of
stock "typical" rolling oil solution evaporated in 500 ml
volumetric flask at 110 C for 2 hrs., dispersed in 200-300
ml deionized water via a Wrist Action Shaker for 1-16
hours (followed by dilution to 500 ml with deionized
water and emulsification via a single speed Waring Blendor
for 2,0 _ o.05 minutes. For emulsifier containing emulsions,
the emulsifier solution was added to the chloroform evap-
orated oil phase before dispersion.
135
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Stock emulsifier solutions - 2.07 to 2.58 g Triton X-100 nonionic
and Atlas G-3300 anionic emulsifiers dissolved in 1.00
liters deionized water.
PVSK standard solution - 9.till x 10~%, prepared by direct weighing
of 2.0030 g PVSK dissolved in 1.00 liters deionized water.
"Cat-Floe" standard solution - 9.708 x 10"%, prepared by dissolving
ca. 9.7 g Calgon "Cat-Floe" (dimethyldiallylairanonium chloride)
in deionized water and standardized by titration against
PVSK with Toluidine Blue as indicator.
Toluidine Blue standard indicator solution - U.80 x 10"-3 weight percent
prepared by dissolving 0.01*80 g o-toluidine blue reagent in
50.0 ml deionized water.
Standard KC1 solutions - 0.1006N and 0.002012N prepared by dissolving
2.7500 g reagent grade KC1 in 500 ml deionized water and
dilution of 10.0 ml thereof to 500 ml, respectively.
Commercial rolling oils - Obtained from various companies having
vendors identification letters "A", "B", "C", "D".
Commercial rolling oil emulsions - Prepared by weighing 0.50 grams of
as-received commercial rolling oil into 500 ml volumetric
flasks, dispersion in 250 ml deionized water, subsequent
dilution to 500 ml and emulsification via a single speed
Waring Blendor for 1.00 - 0.02 minutes.
Spent rolling oil emulsion - Obtained from Ashland 5 stand tandem
cold mill. Four samples were taken: a brand "B" at
6,000 ppm emulsified, 7,500 ppm tramp; a brand "B" un-
analyzed (about 10,000 ppm total); and two brand "C" un-
analyzed, one having been on the mill two hours, the other
being the same emulsion after 2 days on the mill.
136
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APPENDIX 0 - EXPERIMENTAL EQUIPMENT
In addition to the usual laboratory equipment the following equipment
was employed:
Ultracentrifuge - International Equipment Co., Boston, Mass., Model
V, Size 2 with No. 103 multispeed attachment and
No. 296 four-place angle head.
Jar Test Stand - Nalco Chemical Co., Chicago, Illinois, Four-place
1.0 x 2.5 in. vertical paddle on 0.2$ in. diameter
shaft; variable tip speed 0-1.8 fps.
pH Meter - Beckman, Inc., Fullerton, Calif., Model 76
Turbidimeter - Hach Chemical Co., Ames, Iowa, Model 2100, standardized
with 75 JTU polystyrene standard.
UV Spectrophotometer - Beckman Inst. Co., Fullerton, Calif., Model DU.
Coulter Counter - Coulter Electronics Div., Hialeah, Fla., Model B,
standardized with 3.U9 micrometer (urn) latex spheres;
k of 30 urn aperture, 0.1052/M*; for lUO um aperture
k • 1U.52/M*
Relative Streaming Current Detector - Water Association, Inc.,
Framingham, Mass., U cycle per second, polyethylene
block and piston.
Zeta-Meter - Zeta-Meter, Inc., New York, New York, Standard Model,
K factor - 90.7 i 0.2 @ 25 C by standardization with
0.0020lIiN KC1. The conversion factor from electro-
phoretic mobility to zeta potential was 13.1 based
on an estimated temperature of measurement of 2h C
(17).
137
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BIBLIOGRAPHIC:
Armco Steel Corporation, Treatment of Waste Water - Waste
Oil Mixtures, FWPCA Publication (12010EZV 02/70), May, 1970,
137 pages, 25 figures, 22 tables, 32 references, 7 appendices.
ABSTRACT:
Cold reduction of steel strip results in the production of
large quantities of waste water containing variable amounts of
oil. A five stand tandem cold mill located at Armco Steel Cor-
poration's Ashland, Kentucky Works produces 200 to 500 gpm of
waste water containing 400 to 4,000 ppm of oil. The COD of the
waste varies from 400 to 20,000 ppm.
A treatment process and facility was developed, constructed,
and demonstrated, on full-scale, for the treatment of cold mill
wastes. The treatment process utilized chemical coagulation to
break the emulsions. The chemicals employed included alum, 11 met
clay and organic coagulant aid. The process consisted of the fol-
lowing treatment steps; equalization, chemical addition and rapid
mixing, flocculation, and dissolved air flotation. A number of
BIBLIOGRAPHIC:
I Armco Steel Corporation, Treatment of Waste Water - Waste
I Oil Mixtures, FWPCA Publication (12010EZV 02/70), May, 1970,
I 137 pages, 25 figures, 22 tables, 32 references, 7 appendices.
I ABSTRACT:
| Cold reduction of steel strip results in the production of
I large quantities of waste water containing variable amounts of
' oil. A five stand tandem cold mill located at Armco Steel Cor-
| poration's Ashland, Kentucky Works produces 200 to 500 gpm of
waste water containing 400 to 4,000 ppm of oil. The COO of the
I waste varies from 400 to 20,000 ppm.
I A treatment process and facility was developed, constructed,
I and demonstrated, on full-scale, for the treatment of cold mill
I wastes. The treatment process utilized chemical coagulation to
i break the emulsions. The chemicals employed included alum, lime,
' clay and organic coagulant aid. The process consisted of the fol-
I lowing treatment steps; equalization, chemical addition and rapid
mixing, flocculation, and dissolved air flotation. A number of
BIBLIOGRAPHIC:
Armco Steel Corporation, Treatment of Waste Water • Waste
Oil Mixtures, FWPCA Publication (12010EZV 02/70), May, 1970,
137 pages, 25 figures, 22 tables, 32 references, 7 appendices.
ABSTRACT:
Cold reduction of steel strip results in the production of
large quantities of waste water containing variable amounts of
oil. A five stand tandem cold mill located at Armco Steel Cor-
poration's Ashland, Kentucky Works produces 200 to 500 gpm of
waste water containing 400 to 4,000 ppm of oil. The COD of the
waste varies from 400 to 20,000 ppm.
A treatment process and facility was developed, constructed,
and demonstrated, on full-scale, for the treatment of cold mill
wastes. The treatment process utilized chemical coagulation to
break the emulsions. The chemicals employed included alum, lime,
clay and organic coagulant aid. The process consisted of the fol-
lowing treatment steps; equalization, chemical addition and rapid
mixing, flocculation, and dissolved air flotation. A number of
ACCESSION NO.
KEY WORDS:
Waste Water Treatment
Industrial Wastes
Steel Wastes
Emulsions
Coagulation
Flocculation
Oily Wastes
Flotation
Emulsifiers
Zeta Potential
ACCESSION NO.
KEY WORDS:
Waste Water Treatment
Industrial Wastes
Steel Wastes
Emulsions
Coagulation
Flocculation
Oily Wastes
Flotation
Emulsifiers
Zeta Potential
ACCESSION NO.
KEY WORDS:
Waste Water Treatment
Industrial Wastes
Steel Wastes
Emulsions
Coagulation
Flocculation
Oily Wastes
Flotation
Emulsifiers
Zeta Potential
-------
treatment variables were studied in the laboratory and in the field
in order to establish process kinetics and optimum treatment
efficiency.
Zeta potential, streaming current, and particle size distribu-
tion were used in laboratory studies to describe the effect of the
following variables on process kinetics; acid number, initial Oil
concentration, type of emulsifier, chemical dosage, order of
addition of chemicals, reaction time, mixing, and final pH. Based
on these studied, a hypothesis of the emulsion breaking mecha-
nism was proposed.
Oil, COD, and turbidity were used in field studies to establish
the effect of the following variables on treatment efficiency; chem-
ical concentration, order of chemical addition, chemical mixing
time, flocculation mixing time and speed, and air flotation time and
recirculation rate. Based on these studies, optimum design criteria
and operating costs for this process were presented.
This report was submitted in fulfillment of Grant No. WPRD
169-01-68 between the Federal Water Pollution Control Admini-
stration and the Armco Steel Corporation.
treatment variables were studied in the laboratory and in the field
in order to establish process kinetics and optimum treatment
efficiency.
Zeta potential, streaming current, and particle size distribu-
tion were used in laboratory studies to describe the effect of the
following variables on process kinetics; acid number, initial oil
concentration, type of emulsifier, chemical dosage, order of
addition of chemicals, reaction time, mixing, and final pH. Based
on these studied, a hypothesis of the emulsion breaking mecha-
nism was proposed.
Oil, COD, and turbidity were used in field studies to establish
the effect of the following variables on treatment efficiency; chem-
ical concentration, order of chemical addition, chemical mixing
time, flocculation mixing time and speed, and air flotation time and
recirculation rate. Based on these studies,optimum design criteria
and operating costs for this process were presented.
This report was submitted in fulfillment of Grant No. WPRD
169-01-68 between the Federal Water Pollution Control Admini-
stration and the Armco Steel Corporation.
treatment variables were studied in the laboratory and in the field .
in order to establish process kinetics and optimum treatment I
efficiency. i
Zeta potential, streaming current, and particle size distribu- |
tion were used in laboratory studies to describe the effect of the ,
following variables on process kinetics; acid number, initial oil I
concentration, type of emulsifier, chemical dosage, order of i
addition of chemicals, reaction time, mixing, and final pH. Based '
on these studied, a hypothesis of the emulsion breaking mecha- |
nism was proposed. .
Oil, COD, and turbidity were used in field studies to establish i
the effect of the following variables on treatment efficiency; chem- '
ical concentration, order of chemical addition, chemical mixing I
time, flocculation mixing time and speed, and air flotation time and '
recirculation rate. Based on these studies, optimum design criteria |
and operating costs for this process were presented.
This report was submitted in fulfillment of Grant No. WPRD i
169-01-68 between the Federal Water Pollution Control Admini- |
stration and the Armco Steel Corporation. ,
------- |