United States
Environmental Protection
Agency
Industrial Environmental Research LP4-6 00,7-78-143
Laboratory JHy 1978
Cincinnati OH 45268
Research and Development
SEPA
Hydrolysis of Iron
From Acidic Liquors
ENVIRONMENTAL
PROTECTION
AGENCY
DALLAS, TEXAS
LIBRARY
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RESEARCH REPORTING SERIES
Research reports of the Office of Research and Development, U S Environmental
Protection Agency, have been grouped into nine series These nine broad cate-
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vironmental technology Elimination of traditional grouping was consciously
planned to foster technology transfer and a maximum interface in related fields
The nine series are
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6 Scientific and Technical Assessment Reports (STAR)
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This report has been assigned to the ENVIRONMENTAL PROTECTION TECH-
NOLOGY series This series describes research performed to develop and dem-
onstrate instrumentation, equipment, and methodology to repair or prevent en-
vironmental degradation from point and non-point sources of pollution This work
provides the new or improved technology required for the control and treatment
of pollution sources to meet environmental quality standards
This document is available to the public through the National Technical Informa-
tion Service, Springfield, Virginia 22161
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EPA-600/2-78-143
July 1978
HYDROLOSIS OF IRON FROM ACIDIC LIQUORS
by
Alan D. Randolph
Richard D. Williams
Department of Chemical Engineering
University of Arizona
Tuscon, Arizona 85721
Grant No. R-802390
Project Officer
Mary Stinson
Industrial Pollution Control Division
Industrial Environmental Research Laboratory
Edison, New Jersey 08817
INDUSTRIAL ENVIRONMENTAL RESEARCH LABORATORY
OFFICE OF RESEARCH AND DEVELOPMENT
U.S. ENVIRONMENTAL PROTECTION AGENCY
CINCINNATI, OHIO 45268
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DISCLAIMER
This report has been reviewed by the Industrial Environmental Research
Laboratory - Cinncinnati, U.S. Environmental Protection Agency, and approved
for publication. Approval does not signify that the contents necessarily
reflect the views and policies of the U.S. Environmental Protection Agency/
nor does mention of trade names or commercial products constitute endorse-
ment or recommendation for use.
11
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FOREWORD
When energy and material resources are extracted, processed, con-
verted, and used, the related pollutional impacts on our environ-
ment and even on our health often require that new and increas-
ingly more efficient pollution control methods be used. The
Industrial Environmental Research Laboratory - Cincinnati (IERL-
CI) assists in developing and demonstrating new and improved
methodologies that will meet these needs both efficiently and
economically.
The subject of the report is a laboratory-scale investigation of
retained solids crystallization technology for a continuous
treatment of acidic copper leach liquors. The results show that
the retained solids crystallization effectively removes iron in
the form of iron oxide suitable for iron smelting and regenerates
the sulfuric acid for recycling to the ore leaching process.
This information will be of value both to EPA's enforcement pro-
gram and to the industry itself in arriving at meaningful and
achievable discharge levels. Within EPA's R & D program the in-
formation will be used as part of the continuing program to
develop and evaluate improved and less costly technology to min-
imize industrial waste discharges. Besides its direct applica-
tion to treatment of liquors from sulfuric acid leaching of
iii
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copper ores, this technology may find application in the control
of discharges from such sources as leaching of other non-ferrous
minerals, pickling of steel, electrolytic zinc manufacture, and
titanium pigment manufacture.
For further information concerning this subject the Industrial
Pollution Control Division should be contacted.
David G. Stephan
Director
Industrial Environmental Research Laboratory
Cincinnati
IV
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ABSTRACT
This report presents the results of a project to demonstrate the feasibility
of utilizing retained solids crystallization technology to increase the iron
removal rate from acidic process liquors. The increased iron removal rate
with retained solids results in lower solution concentrations of iron. The
solid product formed is called alpha hematite and is suitable for iron
smelting. Essentially all the acid values are retained in solution for re-
cycling to the process involved.
A 2-gallon bench-scale crystallizer/reactor was built with a feed and with-
drawal system that allowed liquid residence times varying from 0.33 to 1.33
hours and solid residence times varying from 0.33 to 30 hours. Free acid
and ferric iron concentrations up to 20 and 30 g/liter respectively in the
feed were handled at temperatures of 135° to 165°C. In all cases, retained
bed operation resulted in significantly higher iron removal rates than mixed
product removal operation. This is explained in terms of kinetics limiting
behavior. The formation of alpha hematite under retained bed operation and
basic iron sulfate under mixed product removal operation is explained by the
equilibrium phase diagram in light of iron concentrations remaining in solu-
tion with each of these modes of operation. Ferrous iron in the feed could
be oxidized to ferric iron by maintaining an oxygen overpressure in the
system. The oxidation rate of ferrous iron was enhanced when the oxygen
was sparged at the reactor impeller. Alpha hematite crystallizes with some
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sulfate at these mild process conditions but with essentially total rejection
of copper and aluminum ions.
This report was submitted in fulfillment of Grant No. R-802390 by the Depart-
ment of Chemical Engineering of the University of Arizona under the sponsor-
ship of the U.S. Environmental Protection Agency. This report covers the
period, April 1, 1973 to September 30, 1975 and work was completed as of
September 30, 1975.
VI
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CONTENTS
Foreword
111
Abstract v
Figures viii
Acknowledgment ix
I. Introduction 1
Background 1
Review of Iron Hydrolysis 4
Scope of Project 9
II. Conclusions 11
III. Recommendations 13
IV. Experimental Equipment 15
V. Results 22
Effect of Process Variables and Mode of Operation
on Iron Removal 23
Effect of Feed Constitutents other than Acid and
Ferric Iron 25
Discussion of Results 27
VI. References 37
VII. Appendices 38
1. Sample Analysis Reactions 38
2. Summary of Experiments Run and Data Obtained . 39
3. Details of Data Analysis 45
Vll
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FIGURES
Number Page
1. Fe2O3-SOs-H20 Phase Diagram 5
2. Photograph of Reactor Internals,
Including the Removal Screen 16
3. Photograph of Reactor Placement 18
4. Schematic of Reactor and Ancillary Equipment . . 19
5. Effect of Reactor Residence Time on
Iron Removal 24
6. Effect of Temperature on Iron Removal 26
7. Effect of Oxygen Sparging on the Conversion
of Ferrous Iron to Ferric Iron 28
8. Phase Diagram Temperature Dependence
for the Fe2O3-S03-H20 System 30
9. Sulfur Content of Solid Product
versus Growth Rate 32
Vlll
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AC FNOWLE DGME NT
This report is based on data obtained during the thesis research work of
David A. Milligan, currently with the Anaconda Company, New Mexico Operation,
Grants, New Mexico. The authors would also like to acknowledge the contribu-
tions by Anaconda of Mr. Milligan's time to the project while he was a full
time employee of the Anaconda Company's Primary Metals Division-Extractive
Metallurgy Group in Tucson, Arizona.
IX
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I. INTRODUCTION
BACKGROUND
Concern for the potential hazards of metals in effluent
streams has recently become widespread. The importance of
removing these contaminants from waste waters or recirculated
water is indicated by the wide variety of laws, both State
and Federal, which have been designed to minimize the degree
of pollution due to beneficial usage of water and by the
large quantity of money spent by industry in improving water
quality. The policy of both State and Federal regulations is
one which allows the management and control of identifiable
pollution sources to the degree reasonably practicable with
available technology. The task of meeting more stringent
pollution standards is being imposed upon the minerals in-
dustry. Numerous techniques have been proposed or are being
used to meet the environmental problems related to metals in
aqueous waste streams.
The generation of metals in aqueous waste streams carries two
economic burdens. The first is a direct operating cost
associated with the dissolution of the metal and the handling
of the aqueous waste stream. The second cost is an intangible
cost associated with returning these metal-containing wastes
to the environment. Present social conscientiousness toward
pollution is increasing the cost of disposal of metal-contain-
ing wastes and forcing the development of new processes which
do not generate these wastes.
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One of the major metal contaminants of aqueous waste streams
is iron. This results from several factors such as iron's
high reactivity and abundance. Iron is present with nearly
all other metals in their ores. Iron readily reacts with
most acids to form aqueous iron salts. Some of the major
aqueous iron sources are from industrial operation such as
pickling steel, electrolytic zinc manufacture, titanium pig-
ment manufacture, mine waste water, leaching of non-ferrous
minerals and the cementation of less reactive metals upon
metallic iron.
New methods of extracting copper from copper ores are being
developed to reduce air pollution costs. One method being
developed is copper concentrate leaching. The success of a
copper concentrate leaching process will be controlled by
the efficient removal of copper while leaving behind an iron
salt. Furthermore, the copper smelting industry is consider-
ing the use of sulfuric acid leaching of smelter slags simi-
lar to processes for extracting zinc from undissolved neutral
leach residue (Gordon and Pickering, 1975) . These slags con-
tain values of copper and zinc. Large-scale leaching of slag
would generate hundreds of thousands of tons of aqueous iron
solutions for disposal each year. The practice of copper
cementation using scrap iron is widespread in the copper in-
dustry for the recovery of copper from heap and vat leach
solutions. In many cases, this waste effluent is unfit to be
returned to the leach operation because of dissolved iron
content. However, the large volumes and dilute concentra-
tion of heap leach liquors (as with mine waste water) makes
iron removal by most processes uneconomical. Another major
source of aqueous iron salts results from the sulfuric acid
pickling of steel. Its disposal has been a problem to the
steel industry for more than 75 years.
Industry presently uses various methods to remove the iron
2
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from these effluent streams. For example, several methods
of complex regeneration of pickle liquors are available
(Nemerow, 1971). Other methods such as evaporation with
crystallization of iron salts, deep well disposal, holding
pond disposal, disposal at sea, and neutralization have been
used by industry. Many processes have been modified at great
expense to prevent the generation of acid-containing iron salt
solutions. Efficient removal of iron from acidic solutions in
either the ferrous or non-ferrous metals industries can im-
prove processing efficiency, reduce pollution, and allow the
recycling of an aqueous stream within the process in which it
was generated. The need for an economic process for the re-
moval of iron is demonstrated by the ferrous industries
failure to develop a widely accepted solution for disposal
of spent pickle liquor and rinse water.
Cations can be removed from aqueous solution by several
techniques, all of which are applicable in one form or another
to the problem of iron removal. Among these methods are ion
exchange, oxidation with addition of base, concentration with
subsequent salt crystallization, and hydrolysis. Hydrolysis
is a favorable technique for cation removal, provided suitably
mild conditions can be found, as no reagent expense is in-
curred and the hydrolysis product can often be recycled or
sold. The fact that no foreign ions are added permits total
recycle of the original aqueous stream.
Hydrolysis results when acidic iron solutions are subjected
to temperatures above 100°C. More research is required to
unify existing information and develop new information on the
rates and process conditions for hydrolysis of iron from
aqueous solutions. Often kinetic factors as well as equilib-
rium relationships determine the composition and physical
nature of the hydrolysis product. Thus there is room for and
incentive to explore hydrolysis reactions and conditions that
3
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will lead to more economical processes for iron removal.
Hydrolysis of iron can generate a variety of solid iron pro-
ducts as well as regenerate acid. The development of tech-
nology to hydrolyze iron to form a solid product for poten-
tial sale while regenerating all the acid consumed by the
dissolution of iron could result in the total recycle of iron-
containing waste solutions within the process. This potential
for total recycle could eliminate a serious threat to the
environment.
REVIEW OF IRON HYDROLYSIS
Hydrolysis is a widely known phenomenon generally described
as the reaction of a component in aqueous solution to form a
different component by reaction with water or hydrated
hydrogen ion or hydroxide ion. Hydrolysis is an essential
part of geologic weathering and bacterial activity. Indus-
trial chemistry studies around the turn of the century indi-
cated that hydrolysis of iron salts in boilers produced iron
oxide and the corresponding acid. The hydrolysis of iron
sulfate to form a coagulant in water treatment was used as
early as 1932. Early studies described the hydrolysis of
ferric iron in sulfate solutions (Arden, 1951) up to tem-
peratures of 100°C (Levy and Quemeneur, 1966). In the early
1960's industrial hydrolysis studies were started in
Australia (Scott, 1962). The Chemical Research Laboratories
of the Commonwealth Scientific and Research Organization,
Australia (CSIRO) generated significant information on the
removal of iron from aqueous solutions by various techniques
including hydrolysis at high temperatures. Continued work by
Haigh (1961) led to significant improvements in the handling
of iron containing wastes from the zinc process of Electro-
lytic Zinc Company of Australia. Kwok and Robins (1973) re-
cently summarized existing metal hydrolysis technology.
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Equilibrium relationships in the water-sulfate-iron system
have been studied at high temperatures up to 200°C by Posnjak
and Merwin (1922). Figure 1 illustrates the several solid
phases that can exist in the Fe2C>3-H2O system below 160°C. In
the narrow region A the stable solid product formed is alpha
hematite (FeaOa) and the equilibrium solution contains very
low concentrations of acid and iron. In region B the solid
product is a mixture of Fe20s and FeaOa.SSOs with an equilib-
rium solution with slightly higher iron and acid concentra-
tions than exist in region A. In region C the solid product
is the basic iron salt 3Fe203»4S03»9H20 and the solution in
Fe203
HoO
Fe203-2S03-H20
SO-
Figure 1. Fe203~S03-H20 phase diagram. (After Posnjak and Merwin)
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equilibrium has even higher concentrations of iron and acid.
This relationship between the solution concentrations of iron
and acid and the stable solid product formed is very impor-
tant. The iron and acid concentrations in solution in a
steady state/continuous process are strongly influenced by
the rate of hydrolysis. Any factor, such as solids retention,
which can influence rates can thus potentially dictate the
stable solid product formed. Significantly, region A is sel-
dom reached in batch processes or at low temperatures. Equi-
librium studies of this system are complicated by the numerous
solid phases and slow approach to equilibrium even at elevated
temperatures. Solids will often be formed which are not the
equilibrium phase. Conversion of these solids to the equilib-
rium phases often takes months.
Several studies have dealt with the reaction rates of iron
hydrolysis, e.g., Scott's (1962) batch hydrolysis studies,
Leak and Fine (1969) , and Nightingale and Benck (1960) . This
information indicates that under proper kinetic conditions
iron can be removed from solution by direct thermal precipita-
tion. In most of these studies mixed iron sulfates and iron
oxides were produced as the solid phase. These solids appear
to be a non-equilibrium product resulting as a consequence of
the interaction of the relative kinetic rates and reaction
times. All studies reviewed were of a batch nature, indicated
improved particle growth by the use of higher temperatures,
and showed more complete removal of iron at higher tempera-
tures. No general agreement is found between authors about
how efficient the removal of iron from solution is at a
particular temperature for a given solution. This observation
indicates the non-equilibrium and kinetics-limited nature of
such reactions.
Gordon and Pickering (1975) give an excellent review of three
industrial processes used for iron removal from electrolytic
6
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zinc acid leach solutions. Similar processes would be
applicable to the treatment of acid liquors from a copper
slag leach process. These processes and their associated
hydrolysis reactions are summarized as follows:
i) Jarosite Process: In this process foreign cations
must be added as reagents (ammonium or sodium) and
the free acid generated in the hydrolysis must be
neutralized to allow the reaction to proceed. The
principal reaction is as follows:
3Fe2 (SCU)3 + 2(NH1|,Na)OH + 10H20 *
2(NHit ,Na)Fe3 (SCU) 2 (OH) 6 + 5H2SO,»
The jarosite solids are themselves not a useful product
and must be disposed of.
ii) Geothite Process: In this process iron (ferrous) is
oxidized to ferric and precipitated as FeOOH (geothite)
at temperatures under 100°C. Again, the free acid
generated by hydrolysis must be neutralized (pH ca.
2.5) in order for the reaction to proceed. Further-
more, the iron must first be reduced to ferrous in a
separate step if appreciable conversion to ferric has
occurred in the hot leach. The overall reduction/
geothite reaction is given as
ZnS + %O2 + 3H2O -> 2FeOOH + ZnSO.,
S + 2H2SO1,
iii) Hematite Process: In this process iron (ferric) is
hydrolyzed directly to a useful product (FezOs) without
reagent addition. However high temperatures (>100°C,
requiring pressure reactors) are required. Acid from
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the original leach liquor is totally regenerated in the
process. The hematite reaction is given simply as
2 o o°c
3H20 Fe203 + 3H2SOi,
The iron oxide product is useful as a quality iron
source provided the occluded sulphate content is
acceptable. Crystallization kinetics and process con-
ditions determine the hematite purity. At high sulfate
levels basic iron sulfates (x FeaOs »ySC>3 'zEzO) are pro-
duced, rather than red hematite. These yellow basic
sulfates consume acid in a recycle process and are not
useful products.
The iron hydrolysis process of this study can be compared
against a backdrop of the above three commercially-used iron
removing processes. The process studied was iron hydrolysis
as hematite and the process limits (kinetic as well as
equilibrium) separating the hematite and basic iron sulfate
fields were delineated. The physical form and filterability
of the hydrolysis product is important in any process. These
factors are often determined by kinetic rather than equilib-
rium relationships. The nub of the current research was to
exploit these kinetic factors by carrying out the hydrolysis
reaction on a dense bed of retained hematite crystals to im-
prove the yield, lower the severity of process conditions
and improve the purity and filterability of the hematite
product.
As the reaction temperature increases, the operating pressure
increases dramatically. Operating and capital costs for a
hematite process are.significantly lower at lower pressures
and temperatures but the approach to equilibrium is poor.
The current study has developed a process which approaches
8
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equilibrium at lower temperatures and provides a significant
improvement in the technology of iron removal from iron con-
taining waste streams.
SCOPE OF PROJECT
The iron hydrolysis research supported under EPA Grant
R802390 had the following objectives and goals.
To design and fabricate a Mixed Suspension, Mixed
Product Removal (MSMPR) reactor/crystallizer suitable
for studying the continuous pressure hydrolysis qf
iron from acidic aqueous solutions.
To design and implement a technique for separating the
residence times of solids (product) and mother liquor
in the pressure reactor/crystallizer in order to study
the hydrolysis reaction on a dense bed of retained
product (hematite) crystals. Such retained-bed crystal-
lizer technology, common practice in the inorganic
fertilizer industry, has never been applied to a
pressure crystallizer.
To study the conditions of temperature, acid concentra-
tion, feed composition and liquid/solids residence time
that yield a dense-phase filterable hematite (Fe^Oa)
product at high yield. To determine conditions that
produce hematite (and thus regenerate acid) and compare
to conditions that produce a basic iron sulfate pro-
duct (and thus consume acid).
To study the efficiency of separation of the hydrolysis
product from valuable ions (e.g. copper remaining in
solution) and the feasibility of simultaneously oxidizing
ferrous ions to ferric ion in the hydrolysis reactor/
crystallizer.
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All these major research goals were accomplished during
the experimental phase of this project and a significant
process improvement for pressure hydrolysis of iron was
achieved.
10
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II. CONCLUSIONS
A continuous-flow pressurized reactor/crystallizer with a
separation of residence times of solids vis-a-vis liquids was
successfully applied to the hydrolysis of hematite (FeaOa)
from an acidic aqueous synthetic copper leach liquor. The
following specific results were obtained from this study:
Iron was removed as the red alpha hematite when the
process was conducted using the retained bed crystal-
lizer mode. Otherwise identical conditions, but with
a mixed discharge mode of operation, resulted in pro-
duction of yellow basic iron sulfate crystals. Yields
were necessarily increased when producing the red iron
oxide form since hematite only exists at the lowest Fe
concentrations in the iron oxide/iron sulfate phase
diagram.
Free acid values originating with the sulfate ion
(sulfuric acid) were regenerated.
Alpha hematite crystallizes with some sulfate (ca. 0.06
S/Fe mole ratio) at these mild conditions but with
essentially total rejection of copper and aluminum ions.
Free acid and ferric ion concentrations up to 20 and 30
g/liter respectively in the process feed (appropriately
higher acid concentration in the mother liquor, depend-
ing on original sulfate concentration) were handled with
this process at temperatures of 135° to 165°C while
still maintaining high yield.
Ferrous ion was oxidized to ferric ion by maintaining a
11
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partial pressure of oxygen in the hydrolysis reactor.
Oxidation of ferrous ion did not limit the hydrolysis
reaction when oxygen was sparged in at the reactor
impeller.
An invention disclosure covering use of the retained-
bed process in the iron hydrolysis reaction has been
filed.
In short, this research study has revealed that iron can be
effectively removed from acidic solutions in an environ-
mentally acceptable manner by pressure hydrolysis in a re-
tained-bed crystallizer.
12
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III. RECOMMENDATIONS
1. Further bench-scale studies of the iron hydrolysis process
should be carried out to elucidate key process features and
obtain a reaction engineering model for quantitative design
studies and scale-up. Specifically, the following areas
should be studied in depth.
i) A kinetic model for crystal growth rate in terms of ion
concentrations and crystal surface area should be de-
veloped. Such a growth kinetics model is essential if
process yields are to be calculated in scale-up compu-
tations.
ii) A kinetics model for crystal nucleation rate in terms
of agitation, solids concentration and iron concentra-
tion should be developed. Such a nucleation kinetics
model is essential if crystal-size distribution (CSD) of
the product hematite crystals is to be studied as a
function of crystallizer configuration.
iii) Rigorous process simulation of CSD and yield should be
carried out using the kinetics obtained in i) and ii)
above.
iv) The nature of the ion impurities in the hematite
product should be studied in depth, as to whether the
impurities occur by solid solution, occlusions, co-
precipitation or lattice substitution.
v) Several typical hydrolysis runs should be duplicated
using typical plant-generated copper leach solutions to
verify the principal reactions and check the possibility
13
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of any deleterious ions present in the more complex
industrial liquor.
2. Parallel development of the hydrolysis reaction on a
pilot scale should be conducted by an industrial concern
who might be interested in applications of this technology.
14
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IV. EXPERIMENTAL EQUIPMENT
The principal equipment for the hydrolysis of iron in acidic
solutions is the pressure reactor/crystallizer. Selection of
the laboratory pressure reactor/crystallizer was determined
based on minimum size suitable for proper sampling with con-
tinuous operation in a corrosive high temperature environ-
ment. Several inlets and outlet connections were required for
proper control of the temperature, pressure, inlet flow, out-
let flow and level. Internal structures provided flow
pattern control, agitation, heating, and different residence
times of solids vis-a-vis liquid phases.
A commercially available nominal 2 gallon 316 stainless steel
autoclave was chosen as the base unit for the pressure reactor/
crystallizer. Heating was provided by high pressure steam
both in a jacket and in an internal heating coil wrapped tightly
around an internal 7.6-centimeter diameter draft tube. A
variable speed 7.6-centimeter propeller was chosen to provide the
necessary agitation. Within the reactor, a multilayer metal
filter mesh was supported on a stainless steel block attached
to an outlet solution line. The filter mesh of approximately 2
microns retaining capacity was installed to provide liquid re-
moval without allowing solid removal. This removal screen, shown
in Figure 2, was the critical process element in this study
enabling a separation of solids versus liquid residence times
and providing for build-up of a dense bed of product crystals.
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Figure 2. Photograph of reactor internals,
including the removal screen.
16
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The reactor was provided with a hydraulic jack to raise and
lower the bottom half of the flanged reactor container
efficiently for inspection, repairs and clean up. Figure 3
illustrates the general placement of the pressure reactor/
crystallizer in the hydrolysis laboratory.
Ancillary equipment included a complex feed and withdrawal
arrangement as shown in Figure 4. A 100-liter feed tank was
chosen to provide approximately 12 reactor volumes of feed.
A positive displacement pump feeds the reactor through a set
of valves providing a direct inlet into the reactor or flushing
through the solids screen into the reactor. Thus the feed/
discharge arrangement provides for alternate periods of feed
flow into the reactor through the screen and liquor discharge
out through the screen. The inlet pumping system is con-
structed to relieve excess pressure into the pump inlet piping
and to prevent pulsations in the feed piping. As a fail-safe
measure, the feed pump is shut off during the liquor discharge
cycle.
Nitrogen or oxygen is supplied through a high pressure regula-
tor to the reactor for pressure control within the reactor at
approximately 10 atmospheres. Gas is removed through a con-
denser vent system which returns the condensed water vapor,
allowing the nitrogen or oxygen to escape at room temperature.
For reasons of safety, a pressure relief valve was installed
on the reactor and set at approximately 12 atmospheres. The
vent piping from this pressure relief valve is piped directly
to the drain through corrosion resistant piping to prevent
flashing into the room and plugging due to corrosion products.
For temperature control, a thermocouple lead was placed into
the reactor through a hollow probe and extended approximately
17
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Figure 3. Photograph of reactor placement
18
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19
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12 centimeters into the liquid. The thermocouple and
controller were calibrated against a standard mercury
thermometer from 0° to 100°C. The temperature controller
opens the supply of steam to the reactor. The steam supply
is provided with a block valve upon entering the room. The
steam passes through a steam regulator which lets down the
pressure to the desired level. A bypass valve has been pro-
vided to allow operation at the highest possible steam
pressure. However, maximum hydrolysis temperature is limited
to the condensing temperature of the 8-atmosphere high pressure
steam supply. The control solenoid valve controls the steam
flow to both the internal heating coils and the reactor
jacket. The control valve and the steam trap on the reactor
jacket have been provided with bypass valves to facilitate
fast startup. Separate steam traps (shown as one in the
diagram) have been provided for both the reactor jacket and
heating coils.
The level within the reactor/crystallizer is controlled with
a direct contact tuned capacitance probe control. At high
level, the control shuts off the feed pump, closes the feed
valve and opens the outlet valve. Once the level falls below
the probe, the feed pump starts, the feed valve opens and the
outlet valve closes. This method of control provides a fail-
safe method of control if the outlet piping should become
blocked. The platinum wire level probe enters the reactor
through a sealed, heavy-walled, 6.5 milimeter glass tube ex-
tended approximately 9 centimeters below the top of the upper
flange face through a Swagelok fitting with teflon ferrule.
Care must be taken to prevent uneven stress on the glass
during tightening.
The screen removal system provides two modes of operation. The
first provides mixed magma withdrawal. The second provides
classified solution withdrawal. Combina;ion of these two modes
20
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can provide differing residence times for the solid and the
liquid phases within the reactor. Withdrawal from the re-
actor is complicated by the presence of solids and flashing.
To prevent flashing and to obtain quench cooling to insure the
best sampling possible, the effluent is cooled from an operat-
ing temperature near 150°C in approximately 3 seconds by dis-
charging through a water cooled heat exchanger. The exiting
temperature of the solution is normally below 50°C. The with-
drawn solution or slurry passes through a set of control
valves. The first valve is a ball valve which determines the
mode of operation, i.e. mixed slurry or liquid discharge.
The next two valves (shown as one in the diagram) in the series
are pneumatic pilot valves which act in a full open or closed
position. Two valves are required to prevent weeping in this
difficult service. If throttling is necessary it is done by
the ball valve. The pressure drop from the reactor to atmos-
pheric pressure is mainly absorbed in the outlet piping. The
liquid level control system has been sized to provide a fill
cycle of approximately 3 to 6 minutes and an effluent with-
drawal cycle of 5 to 10 seconds. A pneumatic delay system was
designed into the pneumatic valve control system for the outlet
valves.
The reactor effluent was saved for blending into later feed
solutions. Effluent solutions were not recycled more than
twice to prevent impurity buildup. Sample analyses performed
are summarized in Appendix 1.
21
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V. RESULTS
A series of experimental runs using the equipment described was
made in order to determine the steady state effect on iron re-
moved of temperature, sulfuric acid content of the feed, iron
content of the feed, and crystallizer residence time. These
effects were determined for batch, continuous mixed-suspension-
mixed-product removal (MPR) and continuous mixed-suspension-
classified-product removal (CPR) modes of operation. In addi-
tion to this main thrust of the work, runs were also made in
order to determine the feasibility of simultaneous oxidation
of ferrous iron to ferric iron and to determine whether for-
eign irons such as copper and aluminum, if present in the feed,
would be excluded from the solid iron product formed. In these
runs temperature was varied in the range of 135° to 165°C, sul-
furic acid content of the feed varied from 1 to 20 g/liter,
iron content varied from 5 to 30 g/liter, liquid residence
time varied from 0.33 to 1.33 hours and solid residence time
>aried from 0.33 to 30 hours. A detailed summary of experi-
mental levels used in each run is given in Appendix 2. The
results obtained for these runs include effluent solution con-
centrations of total iron and ferrous iron (if present), sul-
furic acid, copper (if present) and alumina (if present). The
effluent solid slurry concentration was measured and the solids
were analyzed for content of sulfate, iron, copper (if present)
and alumina (if present). The density of the solid phase was
determined. In addition the particle size distribution was
determined using a Model T Coulter Counter. These data were
used to calculate the iron removed (g/liter), molar ratio of
-------
sulfuric acid produced in solution to iron removed from solu-
tion, the molar ratio of sulfur to iron in the solid, and the
average particle growth rate (cm/hr). This information as
well as the raw data is presented in Appendix 2. Details of
these calculations are given in Appendix 3. The first experi-
mental results discussed will be for the primary experiments
demonstrating the effect of the main process variables on
iron removal. Then the oxidation of ferrous iron in the feed
to ferric iron and the effect of foreign ions will be pre-
sented.
EFFECT OF PROCESS VARIABLES AND MODE OF OPERATION
ON IRON REMOVAL
The effect of increasing residence time on grams per liter of
iron removed is shown in Figure 5. For MPR operation an in-
crease in residence time resulted in an increase -in iron re-
moval as expected. When the reactor/crystallizer was operated
in the CPR mode, but with the same temperature and feed compo-
sition as used in the MPR mode, a dramatic increase in iron
removal is observed for a given residence time. This is shown
in Figure 5 for three different acid-iron feed combinations.
Another way of interpreting these results is that in order to
obtain the same iron removal in the MPR mode as obtained in
the CPR mode, significantly higher residence times would be re-
quired. For a given treatment rate this would result in much
larger equipment. The cause of this increase in iron precipi-
tation rate is the increase in solids slurry concentration re-
sulting from retaining the iron bed in the CPR mode. Thus,
the hydrolysis reaction of iron in such acidic liquors is
seen to be kinetics limited. It should also be noted that CPR
iron removal is increased as residence time is increased,
allowing for a closer approach to equilibrium. Thus, addition-
al solids retention beyond that occurring with these runs may
be advantageous. The effect of concentrations of acid and iron
23
-------
£"- in
"S'^cf N'
a> > W
"- ~
D I
I.I.I
I i I
q
c\i
CO
CM UJ
- o
UJ
Q 9
- (/)
UJ
oc
3 8
a:
CVJ
d
CD N- CD in ^" rO C\)
(J94!l/iu5) Q3AOW3d NOdl
C
O
c
o
-------
on removal rate is also apparent from this graph. Increases in
iron concentration result in increases in removal rate while
increases in acid concentration result in decreases in rate.
This is because acid is a product in the reaction and thus
tends to reverse the reaction while iron product (surface area
of hematite) promotes the reaction.
The effect of temperature on iron removal is illustrated in
Figure 6. In all cases iron removal is increased with tempera-
ture as would be expected from Arrhenius exponential reaction
rate dependence. It is again apparent from these results that
the operation is kinetically, rather than equilibrium con-
trolled, since an increase in temperature for an exothermic
reaction, such as precipitation, at equilibrium would result
in a decrease in iron removal rather than the observed in-
creases. Again the effect of the retained bed on enhancing
reaction rate is illustrated since in every case the CPR mode
resulted in increased iron removal of over 100%. The effects
of acid concentration and iron concentration on iron removal
were again negative for acid concentration and positive for
iron concentration. The implication of the observed tempera-
ture effect is that in the CPR mode a given iron removal rate
can be achieved at a lower temperature than in the MPR mode.
This results in a lower operating pressure and less stringent
pressure requirements in fabricating the process equipment, as
well as higher throughput rates at equivalent yields for the
CPR vis-a-vis MPR modes of operation.
EFFECT OF FEED CONSTITUENTS OTHER THAN
ACID AND FERRIC IRON
Frequently industrial acid solutions contain ferrous rather
than ferric iron and/or additional ions such as copper or
aluminum. Several experiments were conducted in order to
determine whether ferrous iron could be simultaneously oxidized
to the ferric oxidation state during the iron removal process
25
-------
o
o
CM
O
00
o
vo
o
in 01
10
O &.
^- (I)
i a.
o
CO
o
CM
(0
o
E
Oi
c
o
c
o
0)
!-l
d
-P
fd
^4
(U
a,
E
(U
-P
u
0)
O
O
(D
M
d
o
26
-------
and whether foreign ions would be rejected from the solid iron
product formed.
To answer the first of these questions ferrous iron was intro-
duced into the feed as a part of the total iron content. The
total iron and ferrous iron concentrations in the feed for
these runs are given in Appendix 2. Two of these runs were
in the batch mode with oxygen provided either on the surface
of the solution or sparged through a tube near the impeller.
It was empirically found that the oxidation rate is nominally
first order in ferrous concentration with surface oxidation
and second order when sparging was used. This is illustrated
in Figure 7 where a function of unreacted ferrous iron is
plotted against time. A linear semilog plot of remaining
ferrous iron with time, as obtained with surface oxygenation,
is indicative of a first order process while a linear recipro-
cal plot of remaining ferrous iron with time, as obtained
with sparging, is indicative of a second order process. Over-
all iron removal rates in these runs were equivalent to those
experienced with only ferric iron in the feed indicating the
feasibility of simultaneously oxidizing ferrous to ferric while
producing a solid iron product.
The solid composition data reported in Appendix £ indicates
that for runs with copper in the feed the solids analyzed less
than 0.06% copper. The one run with aluminum in the feed
resulted in 0.66 percent alumina in the solid product. These
results indicate that foreign ions such as copper and aluminum
will be essentially excluded from the solid product.
DISCUSSION OF RESULTS
If the idea is accepted that the hydrolysis of ferric ion to
alpha Fe2O3 is kinetically limited by crystal surface area in
the temperature range 135° to 165°C, then the experimental results
can be adequately explained in terms of twa principal reactions
27
-------
ONINIVIflBd (II) NOdl
LJ
I i I i I i I
UJ
§
h-
c
o
H
W
SH
QJ
>
C
O
O
Q)
X! .
-P C
O
C H
O -H
tn U
C -H
H H
tn M
M 0)
to m
a
w o
-p
c
0) C
C7i O
>i M
X-H
o
w
4-4 3
o o
D 0)
0) tn
w o
-H
(9NINIVlAI3d(II) NOdl)
28
-------
and the Fe203-S03-H2O phase diagram (Figures 1 and 8)
The hydrolysis reactions for alpha hematite and the most
probable first-occurring basic iron sulfate are respectively
given below.
Fe2(SOi,)3 + 3H20 hematite Fe20
crystal
surface
3Fe2 (80^)3 + 14H20 - > 3Fe203 «4S03 «9H20
+ 5H2S0lt
Several observations can be made regarding these overall re-
actions.
i) The hematite reaction regenerates 1% moles of sulfuric
acid for every mole of ferric ion precipitated.
ii) The basic iron sulfate hydrolysis reaction regenerates
5/6 moles of sulfuric acid for every mole of ferric ion
precipitated. Thus the basic iron sulfate reaction
gives a net consumption of acid in the overall leach/
hydrolysis process as evidenced by the sulfate level in
the product.
iii) The hematite reaction is influenced by the amount of
product hematite crystals in the mixed reaction environ-
ment.
From the phase diagram (Figure 1) it was seen that the hematite
phase exists only at low Fe concentrations in the liquid phase.
In this experimental study the hematite solid phase was only
produced in the CPR mode of removal which greatly increased the
29
-------
FtfO,
Figure 8. Phase diagram temperature dependence for the
-SOs-H2O system. (After Posnjak and Merwin)
30
-------
magma density (and hence surface area) of product crystals.
The rationale explaining the experimental results is that the
retained bed mode of operation created a large surface area of
crystals, thus increasing iron removal from the feed solution
as shown in Figure 5. The resulting lower Fe concentrations
in the mother liquor then gave rise to the desirable red
hematite form of precipitate. When the crystallizer was
operated in the MPR mode a lower yield occurred (see Figure 5)
which resulted in higher Fe concentrations in the mother liquor
with yellow basic iron sulfate as the stable solid phase.
Significantly, these same run conditions of feed concentration
and temperature would produce basic iron sulfate and/or a
mixture of hematite/basic iron sulfate when operated in the
batch mode. Once formed, the basic iron sulfate crystals
convert slowly to hematite even when equilibrium end conditions
favor the latter crystal species. This observation accounts
for the great variability in yield and product composition re-
ported in previous batch studies.
The dramatic difference in solid product produced under other-
wise identical conditions, but with CPR or MPR mode of opera-
tion, is demonstrated by either black or red color of product
crystals. X-ray analysis of crystal product revealed that the
hematite crystals produced were not pure, but contained ca.
0.05 to 0.07 mole ratio of sulfur to iron. These hematite crys-
tals appeared dark under polarized light, indicating the
possibility that the sulfate contamination was present in the
hematite phase as a solid solution. However, Fe203 product
with this level of sulfate contamination should still prove
acceptable as a high-grade feed material in an iron smelting
process.
The change in solid phase with decrease in iron concentration
can also be seen in terms of the decreased crystal growth rate
as shown in Figure 9. Thus, as the crystal magma concentra-
31
-------
1
d d c> o o
anos NI NOdi/anos NI
O
co
CVJ
E
CNJ
fe
01
10 CD
O
O
0
-P
u
o
04
o
co .
QJ
M-l -M
O (0
M
-P
(D -P
C O
O M
O tn
S-l 10
m cn
d OJ
to >
(U
S-l
H
Cn
32
-------
tion increases, the crystal growth rate decreases due to a
drop in solution driving forces (Fe concentration) and the
red Fe203 phase is produced. Figure 9 shows significant
amounts of sulfur in the product near the transition point;
this is undoubtedly due to the presence of both crystal phases
under these conditions. Further work should be done to de-
termine if the residual sulfur contamination could be lowered
with lower crystal growth rates or if it is present as a
solid solution and is thus a function of the acid concentra-
tion in the mother liquor.
X-ray analysis of the product for the runs with copper and
aluminum ion present revealed that copper was almost totally
rejected while aluminum was slightly attached. Typical copper
weight percentages in the solids of 0.02% to 0.03% were ob-
served. This contamination level came from a mother liquor
concentration of ca. 4 g/liter as copper. Total yield of iron
(as Fe2C>3) was of the order of 15 g/liter or less. Thus, over-
all rejection of copper was of the order of >99.9%. The
amount of copper reported in the solids was virtually independent
of the concentration of copper in the feed, indicating a constant
entrapment in the crystal lattice rather than entrainment as
solid solution.
An alumina concentration of 0.66% in the solids was reported
for a feed composition of 4 g/liter, indicating a 2% loss of
alumina. This level of aluminum contaimination would probably
be acceptable for use of the Fe2O3 as an iron ore feed. Not
enough information on aluminum leach rates is available to
predict whether or not alumina would build up in the system
with this rate of removal in the iron product.
The hydrolysis runs made with mixtures of ferrous and ferric
ions in the feed solution gave total solids yields comparable
to those with 100% ferric ion provided the O2 for oxidation was
33
-------
sparged into the vessel near the impeller. Somewhat lower
yields were experienced when ferrous oxidation was attempted
by only maintaining a partial pressure of 02 (instead of N2)
over the liquid interface. These results indicate that oxida-
tion of ferrous ion is limited by liquid diffusion of 02 ,
rather than kinetically controlled by the oxidation step.
These results indicate the necessity of O2 sparging and good
agitation if the reaction was to be carried out on a larger
scale.
The conclusions that can be drawn from the exploratory hy-
drolysis runs presented in this report is that the hydrolysis
of ferric ion to hematite is kinetically limited in the range
135° to 165°C by diffusion/surface reaction of Fe2C>3 onto/into the
hematite crystal lattice. This limitation can be overcome
by conducting the reaction on a dense bed of retained product
crystals, thus lowering the severity of process conditions
and/or increasing throughput. High yields are characteristic
of retained bed operation and, in fact, are necessary to lower
the iron concentration to a region where hematite, rather than
a basic iron sulfate, is the stable solid phase. A process
model for this system which can predict the yield of iron has
not been assembled from these data. Undoubtedly further data
will be required to describe the nucleation-supersaturation and
growth rate-supersaturation kinetics of the system. It is
essential that both growth and nucleation rate kinetics be in-
cluded in any process model in order to be able to predict the
effects of solids retention on the crystal-size distribution
and hence to predict surface area and yield.
The hematite crystals produced in the hydrolysis reaction con-
tain substantial sulfate contamination; however, aluminum and
copper ions are rejected, the latter almost totally. The
hematite product obtained in this study should be acceptable
as a high grade feed to an iron smelter.
34
-------
The process conditions and results of this study should be
compared with the conditions and needs of iron removal from
waste pickling liquor streams to evaluate whether this tech-
nology is applicable to the ferrous industry. It is conceiv-
able that iron hydrolysis together with evaporation (with or
without additional sulfate removal) would be applicable to
the regeneration and total recycle of sulfuric acid type waste
pickle liquors.
Finally, the simultaneous oxidation and hydrolysis of ferrous
ion in this process appears to be feasible if the Oa for the
oxidation is well-dispersed in the reactor.
35
-------
VI. REFERENCES
1. Arden, T.V. The Hydrolysis of Ferric Iron in Sulphate Solu-
tion. Chem. Soc., 1951; 350(1951).
2. Gordon, A.R., and R.W. Pickering. Improved Leaching Tech-
nologies in the Electrolytic Zinc Industry. Met. Trans.,
6B; 43(1975).
3. Haigh, C.J. Aust. J. Appl. Sci., 12; 407(1961).
4. Kwok, O.J., and R. G. Robins. International Symposium on
Hydrometallurgy. Ch. 39, AIME, 1975.
5. Leak, V.G. and M.M. Fine. Induced Oxidation - Precipitation
of Iron From Aqueous Solutions of Mn S04~Fe 804. Ind.
Eng. Chem. Fund., 8_; 411(1969).
6. Levy, L.W., and E. Quemeneur. Bull. Soc. Chim. Fr., 1966;
1947(1966).
7. Nemerow, N.L. Liquid Waste of Industry. Theories, Prac-
tices and Treatment. Addison-Wesley Pub. Co., Reading,
Mass., 1971.
8. Nightingale, E.R. and R.F. Benck. Anal. Chem., 32, 566
(1960).
9. Posnjak, E. and H.E. Merwin. The System: Fe2C>3 - 803 - H2O.
J. Amer. Chem. Soc., 44; 1965(1922).
10. Scott, T.R. Alumina by Acid Extraction. J. Metals, 14; 121
(1962) .
36
-------
VII. APPENDICES
APPENDIX 1. SAMPLE ANALYSIS REACTIONS
Sulfuric Acid Titration
2KI + FE2(S04) > 2FeSC
2Na2S203 + I2
2NaOH + H2S04
K2S04
Na2S406 + 2NaI
Na2S04 + 2H2S04
Iron Titration
Pb + Fe2(S04)
6FeS04 + K2Cr207
2FeS0
3Fe2
+ Cr2(SO4) + 7H2S04
3 3
Sulfate Determination
Fe2(S04) + 3BaCl2
3
2FeCl3 + 3BaSO
Copper and Alumina were determined by atomic adsorption spectro-
scopy.
37
-------
APPENDIX 2. Sl'MMARY OF EXITRIMFNTS RI'N AND DATA OBTAINED
TABLE 2.1 EXPERIMENTAL LEVELS USED
Run
Number
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24
25
26
27
28
29
30
31
32
33
34
35
36
37
38
39
40
41
42
43
44
45
46
47
48
49
50
51
52
53
54
55
56
57
58
59
60
61
62
63
64
65
66
67
68
69
70
Mode of
Operation
R\TCH
BATCH
BATCH
BATCH
BATCH
BATCH
BATCH
MSMPR
MSMPR
BATCH
BATCH
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSCPR
MSCPR
MSCPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSCPR
MSCPR
MSCPR
MSCPR
MSCPR
MSCPR
MSCPR
MSCPR
MSMPR
MSCPR
MSCPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSCPR
MSCPR
MSCPR
MSCPR
MSCPR
MSCPR
MSCPR
BATCH
MSCPR
BATCH
MSCPR
MSCPR
MSMPR
MSMPR
MSCPR
MSCPR
Agitation Temp,
(RPM) (*C)
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
726
256
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
535
146.6
164.4
161.1
162.8
172.2
162.8
162.8
162.8
165.8
162.8
148.9
154.4
154.4
154.4
154.4
154.4
154.4
143.3
162.8
143.3
143.3
143.3
162.8
162.8
154.4
154.4
143.3
154.4
154.4
143.3
143.3
143.3
143.3
160.0
162.8
162.8
154.4
143.3
154.4
154.4
135.0
135.0
162.8
154.4
162.8
162.8
154.4
162.8
154.4
143.3
154.4
154.4
154.4
154.4
154.4
143.3
162.2
143.3
154.4
154.4
143.3
154.4
154.4
154.4
154.4
143.3
143.3
143.3
143.3
143.3
Retention Time
Solution Solid
i!r-) Pir)
r/25
1.12
2.08
6.75
4.00
0.75
3.83
0.32
0.32
50.25
9.25
0.33
0.33
0.65
0.65
0.65
1.31
0.33
0.32
0.66
0.66
1.32
1.29
0.65
0.33
0.61
0.43
0.33
0.65
0.66
0.66
0.66
0.33
0.32
0.65
0.34
0.33
0.49
0.72
1.47
0.48
0.50
0.52
0.33
0.33
0.33
1.31
0.32
0.33
0.33
1.31
0.65
0.33
0.33
0.31
0.32
0.50
0.32
0.31
0.32
0.33
3.00
0.31
2.00
0.31
0.31
0.33
0.33
0.31
0.31
1.25
1.12
2.08
6.75
4.00
0.75
3.83
0.32
0.32
50.25
9.25
0.33
0.33
0.65
0.65
0.65
1.31
0.33
0.32
0.66
0.66
1.32
1.29
0.65
5.40
11.74
4.24
0.33
0.65
0.66
0.66
0.66
0.33
0.32
0.65
1.45
4.19
5.19
3.79
11.01
8.63
20.43
14.04
0.33
24.00
24.00
1.31
0.32
0.33
0.33
1.31
0.65
0.33
0.33
3.89
5.39
10.40
12.36
6.98
28.47
17.67
5.01
5.34
3.20
0.33
0.33
3.50
5.44
Feed Composition
H;SO» Total Fe
Upl) (apl)
10.00
10.00
10.00
10.00
20.00
10.00
30.00
6.65
' 6.65
6.65
8.70
7.99
8.31
9.60
9.60
9.31
9.63
9.31
9.39
9.58
9.58
9.11
9.11
9.39
9.30
9.22
9.58
8.59
9.44
10.56
11.38
11.38
9.61
11.07
10.36
9.05
12.04
9.08
9.11
9.20
8.80
9.47
10.09
9.79
5.00
5.50
9.07
3.52
3.34
3.31
3.30
3.33
3.35
3.72
5.45
6.23
5.93
6.14
5.87
12.13
12.33
11.80
10.34
12.16
11.85
0.78
0.84
4.00
4.22
5.80
30.00
30.00
30.00
30.00
15.00
15.00
50.00
5.88
5.88
5.88
5.47
4.87
5.07
7.26
7.26
7.69
7.66
7.94
7.86
7.85
7.96
7.94
7.94
7.94
7.86
7.91
7.95
15.44
14.70
14.34
13.83
13.83
15.74
15.20
15.30
15.71
16.24
16.19
15.88
15.62
15.43
7.84
7.70
8.01
0.00
7.72
15.36
10.44
10.61
10.35
10.39
10.52
10.32
10.31
10.46
11.00
10.80
10.47
10.38
16.05
16.24
15.84
16.03
15.97
16.06
15.59
15.75
16.76
16.85
16.40
38
-------
TABLE 2.2 FERROUS AND FOREIGN ION FEED LEVELS
Run
No.
60
61
62
63
64
65
66
67
68
69
70
Ferrous
(gpi)
10.16
10.36
10.24
10.34
10.16
10.30
0.00
0.00
11.45
11.54
0.00
Copper
(gpi)
0.00
0.00
0.00
0.00
0.00
0.00
4.12
4.10
2.02
2.04
0.00
Alumina
(gpi)
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
4.00
Oxygen
Pressure
(atm)
4.83
6.20
4.83
5.05
5.03
5.03
0.00
0.00
4.63
4.56
0.00
39
-------
TABLE 2.3 AVERAGE EFFLUK.NT SOLUTION DATA
(1)
Solution Opposition
Run
No.
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
11
23
24
25
26
27
28
29
30
31
32
33
J4
35
36
37
38
39
40
41
42
43
44
45
46
47
48
49
50
51
52
53
54
55
56
57
58
59
60
61
62
63
64
65
66
67
68
69
70
HjSO,
10.00
22.25
28.26
27.48
27.73
18.68
40. 3^
8.11
11.90
Reactor
18.16
8.80
8.80
10.81
11.70
11.36
12.35
9.91
11.44
No Soli
10.44
11.04
15.96
13.22
14.49
20.98
14.38
13.27
13.85
13.08
Outlet
12.98
11.25
Ferric
30.00
16.62
16.39
16.21
9.93
8.33
48.67
1.fi4
4.61
Leaked - - - - -
5.68
4.56
4.76
7.56
5.52
6.60
6.17
7.71
6.80
7.38
7.03
5.84
6.18
4.54
4.64
4.41
12.74
11.92
13.52
Plugged in Reactor
13.62
14.75
Ferrous
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
14.37 12.43 0.00
16.68 10.66 0.00
23.25 6.20 0.00
20.85 4.12 0.00
16.89 10.90 0.00
21.32 7.68 0.00
19.64 12.37 0.00
14.02 12.03 0.00
12.05 6.33 0.00
20.52 6.16 0.00
10.88 7.02 0.00
Attempted Converison of Basic Iron
17.16
10.79
9.41
6.74
12.06
11.35
10.26
8.37
16.17
16.65
21.23
15.45
20.26
21*49
16.72
34.69
22.36
30.21
20.25
14.54
8.08
8.05
17.21
17.57
10.64
5.29
6.43
7.97
4.56
5.34
5.71
7.08
5.26
6.94
5.02
6.29
4.81
4.28
5.74
6.69
3.65
7.38
6.40
7.45
10.79
6.86
6.36
10.60
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
7.03
7.28
0.54
6.74
0.67
4.31
0.00
0.00
2.81
3.50
0.00
in qpl
: r
Copper
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
Sulfate to a-Fe2Oj
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
4.37
4.17
1.98
2.24
0.00
Alumina
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
4.01
Moltir Rnt io
-------
TABLE J.4 AVERAGE EFFLUENT SOLID DATA
Run
No.
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24
25
26
27
28
29
30
31
32
33
34
35
36
37
38
39
40
41
42
43
44
45
46
47
48
49
50
51
52
53
54
55
56
57
58
59
60
61
62
63
64
65
66
67
68
69
70
SO,
No
NO
No
40.89
38.00
37.38
40.00
40.56
40.18
4.37
33.83
40.42
40.52
42.00
37.50
41.31
41.31
40.08
41.62
No
41.90
41.55
41.88
41.82
42.00
28.44
41.68
42.03
42.09
42.02
Outlet
41.53
41.45
41.51
41.73
41.08
40.93
41.42
41.29
39.79
41.83
41.79
33.21
41.61
Iron
Solid
Solid
Solid
37.67
37.67
38.20
37.67
34.10
34.25
61.00
36.40
35.03
34.70
34.40
34.40
36.17
34.46
34.41
34.49
Solid
34.32
34.50
34.80
35.18
37.67
44.01
34.62
35.67
34.23
33.94
Plugged
22.81
34.33
34.33
34.20
34.36
34.51
34.02
34.32
35.33
34.54
34.58
40.42
34.15
Attempted
N
41.79
41.79
41.71
41.45
41.30
41.76
41.57
41.44
32.40
14.94
7.99
15.24
13.58
9.97
9.27
5.48
7.66
9.78
7.13
33.47
40.80
37.48
10.75
11.98
34.12
34.30
34.22
34.16
34.37
34.36
34.23
33.80
40.03
53.91
57.51
53.40
54.37
57.30
57.52
59.41
59.14
58.01
S8.36
39.41
34.07
35.56
55.87
51.80
Copper
Formed
Sample
Sample
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
Sample
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
in
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
Conversion
N
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.02
0.03
0.03
0.06
0.00
Alumina
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
Reactor
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.00
0.66
Density
(gm/cm)
__
2.93
3.03
3.05
2.96
2.94
2.96
4.14
3.16
2.95
2.94
2.89
3.04
2.90
2.92
2.96
2.91
2.90
2.91
2.89
2.90
2.89
3.34
2.90
2.89
2.89
2.89
2.91
2.91
2.91
2.90
2.92
2.93
2.91
2.92
2.98
2.90
2.90
3.18
2.91
of
n
2.90
2.90
2.90
2.92
2.92
2.90
2.91
2.91
3.21
3.78
4.05
3.78
3.83
3.96
3.97
4.10
4.03
3.96
4.04
3.18
2.93
3.04
3.93
3.89
Solid
Molar Ratio
S0»/Fe
._
0.63
0.59
0.57
0.62
0.79
0.68
0.04
0.54
0.68
0.68
0.71
0.63
0.67
0.70
0.68
0.70
0.71
0.70
0.70
0.69
0.65
0.38
0.71
0.68
0.71
0.72
0.71
0.70
0.70
0.70
0.70
0.69
0.71
0.70
0.66
0.70
0.70
0.48
0.71
Basic
"
0.71
0.71
0.71
0.70
0.70
0.71
0.71
0.71
0.47
0.16
0.07
0.17
0.15
0.10
0.09
0.05
0.08
0.10
0.07
0.49
0.70
0.61
0.11
0.13
Concentration
in Slurry
(gpl)
0.00
38.8
39.4
36.6
13.5
19.4
1.46
6.57
3.71
Reactor Leaked
0.00
0.89
0.89
0.00
2.53
3.03
2.88
0.67
3.06
1.69
2.64
6.04
5.01
85.02
7.43
121.5
0.06
7.87
4.90
0.68
3.01
8.05
13.59
121.5
121.5
121.5
121.5
121.5
121.5
149.7
121.5
2.91
Iron Sulfate to
N n M
12.90
15.03
12.20
6.97
16.98
15.09
13.46
9.57
195.9
146.7
108.5
160.4
153.1
172.6
181.4
14.49
141.0
13.65
199.4
231.5
14.56
19.95
121.5
173.3
(wth Rate
A\vracnV2)
0
--
0.63E-8
0.35E-8
0.14E-7
0.89E-3
0.22F-2
0.20E-2
0
0.15E-2
0.14E-2
0
0.10E-2
0.90E-3
0.26E-2
0.20E-2
0.14E-2
0.80E-3
0.63E-3
0.11E-2
0.13E-2
0.16E-3
0.85E-3
0.12E-3
0.23E-2
0.12E-2
0.12E-2
0.13E-2
0.24E-2
0.22E-2
0.40E-2
0.30E-3
0.27E-3
0.21E-3
0.16E-3
0.31E-4
0.15E-3
0.60E-4
0.62E-4
0.31E-2
aFe2O3
*
0.24E-1
0.35E-2
0.33E-2
0.31E-2
0.43E-2
0.18E-2
0.34E-2
0.63E-2
0.14E-3
0.77E-4
0.13E-3
0.64E-4
0.98E-4
0.60E-4
0.37E-4
0.59E-8
0.79E-4
0.59E-8
0.50E-4
0.21E-3
0.30E-2
0.22E-2
0.10E-3
0.51E-4
41
-------
TABLE 2.5 REDUCED RUN DATA-IRON REMOVAL
Run
No.
8
12
16
17
18
19
21
22
23
24
25
26
27
28
29
32
33
34
35
36
37
38
39
40
41
42
43
44
47
48
49
50
51
52
53
54
55
56
57
58
59
Mode of
Operation
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSCPR
MSCPR
MSCPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSCPR
MSCPR
MSCPR
MSCPR
MSCPR
MSCPR
MSCPR
MSCPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSMPR
MSCPR
MSCPR
MSCPR
MSCPR
MSCPR
Solution
Residence Time
0.33
0.33
0.66
1.32
0.33
0.33
0.66
1.32
1.32
0.66
0.33
0.66
0.33
0.33
0.66
0.66
0.33
0.33
0.66
0.33
0.33
0.33
0.66
1.32
0.33
0.33
0.33
0.33
1.32
0.33
0.33
0.33
1.32
0.66
0.33
0.33
0.33
0.33
0.33
0.33
0.33
Feed Composition
Temperature
CO
162.8
154.4
154.4
154.4
143.3
162.8
143.3
143.3
162.8
162.8
154.4
154.4
143.3
154.4
154.4
143.3
143.3
162.8
162.8
162.8
154.4
143.3
154.4
154.4
135.0
135.0
162.8
154.4
154.4
162.8
154.4
143.3
154.4
154.4
154.4
154.4
154.4
143.3
162.8
143.3
154.4
H2SOi,
(gpD
5.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
io;oo
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
5.00
5.00
5.00
5.00
5.00
5.00
5.00
5.00
5.00
5.00
5.00
5.00
Iron
(gpD
5.00
5.00
7.50
7.50
7.50
7.50
7.50
7.50
7.50
7.50
7.50
7.50
7.50
15.00
15.00
15.00
15.00
15.00
15.00
15.00
15.00
15.00
15.00
15.00
15.00
7.50
7.50
7.50
15.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
10.00
Iron
Removal
(gpD
2.89
0.00
0.68
1.26
0.00
0.64
0.22
0.30
1.53
1.32
3.02
3.02
2.79
1.81
2.58
1.31
0.56
3.30
4.76
9.13
6.91
4.74
7.77
2.92
3.10
1.32
1.48
0.73
4.13
4.24
3.11
1.37
4.81
4.14
3.63
2.45
5.26
3.84
5.60
4.14
5.53
42
-------
TABLE 2.6 REDUCED RUN DATA-MOLAR RATIOS AND
PARTICLE GROWTH RATES
Run No.
8
12
16
17
18
19
21
22
23
24
25
26
27
28
29
32
33
34
36
37
38
39
40
41
42
43
44
47
48
49
50
51
52
53
54
55
56
57
58
59
Growth
Rate/2
(oVhr.)
0.22E-2
0.15E-2
0.90E-3
0.26E-2
0.20E-2
0.14E-2
0.80E-2
0.63E-3
O.llE-2
0.13E-2
0.16E-3
0.85E-3
0.12E-3
0.23E-2
0.12E-2
0.13E-2
0.24E-2
0.40E-2
0.30E-3
0.27E-3
0.21E-3
0.16E-3
0.31E-4
0.15E-3
0.60E-4
0.62E-4
0.31E-2
0.24E-1
0.35E-2
0.33E-2
0.31E-2
0.43E-2
0.18E-2
0.34E-2
0.63E-2
0.14E-3
0.77E-4
0.13E-3
0.64E-4
0.98E-4
Malar
H2SO,, Generated/
Fe in Solid
0.37
1.63
1.07
1.04
1.49
1.17
0.85
1.21
1.88
1.24
0.89
2.05
0.87
1.05
0.93
3.96
1.82
0.77
0.84
0.71
0.84
0.85
1.60
0.88
0.97
3.86
0.62
1.00
0.80
0.83
0.82
0.85
0.88
0.85
0.82
1.18
1.44
1.51
1.27
1.47
Ratio
SO., in Solid/
Fe in Solid
0.69
0.67
0.66
0.70
0.68
0.70
0.71
0.70
0.70
0.69
0.65
0.38
0.71
0.49
0.71
0.71
0.74
0.71
0.69
0.69
0.71
0.70
1.62
0.70
0.70
0.48
0.71
0.71
0.71
0.71
0.71
0.70
0.71
0.71
0.71
0.47
0.16
0.71
0.17
0.15
Total S/
Fe in Solid
1.06
2.30
1.73
1.74
2.17
1.87
1.56
1.91
2.58
1.93
1.54
2.43
1.58
1.54
1.64
4.67
2.56
1.48
1.53
1.40
1.55
1.55
2.22
1.58
1.67
4.34
1.33
1.71
1.51
1.54
1.53
1.55
1.59
1.56
1.53
1.65
1.60
2.22
1.44
1.62
Predicted S/
Fe in Solid
1.50
1.50
1..50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
1.50
43
-------
APPENDIX 3. DETAILS OF DATA ANALYSIS
3.1 Reactor-Crystallizer Residence Times:
A. Liquid Phase
Residence time = voltric solution capacity
volumetric solution flowrate
_ (Reactor volume-retained solids volume)
(Flowrate of feed) (Feed density)
(reactor solution density)
B. Solid Phase
Assuming steady state operation, then
average solids withdrawal rate = solids production rate
. , . . solids capacity
Residence time = =~r-= ir_. .,<*= r
solids withdrawal rate
_ solids capacity
solids production rate
3.2 Densities
A. Liquid Phase - Liquid phase specific gravity for various
solution concentrations and temperatures were interpo-
lated from data available in Perry's Chemical Engineer-
ing Handbook.
B. Solid Phase - Solid phase density was estimated by
linearly interpolating the density measured for pure
Fe2O3 (4.28 g/cc) and pure 3Fe2O3«4SO3«9H20 (2.91 g/cc)
based on the measured sulfur content of the solid phase
produced.
3.3 Solids Concentration
A. Mixed Suspension Mixed Product Removal (MSMPR)
For this operational mode since the effluent stream is
representative of the reactor-crystallizer contents the
solids density is most easily calculated by measuring
44
-------
the amount of iron removed from solution and the iron
content of the solids, from which;
... Fe removed (g/liter)
Solid concentration (g/liter) - Fe fraction in solid
B. Mixed Suspension Classified Product Removal (MSCPR)
In this mode since the effluent stream only contains
solids on an intermittent basis, the average solids
concentration may not be estimated by the above rela-
tionship. Instead the following equation is used.
Solids concentration (g/liter)=
W
W 100-W
psd psl
where W = weight percent solids in the reactor
Psd = solids density ( g/liter)
Psl = solution density (g/liter)
3.4 Particle Dynamics
The Coulter Counter gives the number density of particles
in each of 13 different size ranges. With the average di-
ameter of each channel and the number densities the area
and volume of the solids may be determined by numerical
integration. From this the surface area per gram of solid
is calculated by,
Specific surface area _ surface area of solid
(area/mass) of the solid ~ (Volume of solid)(solid
density)
Then with the solid generation rate defined by,
Solid generation rate _ (flowrate) (iron removed) )
(mass/time) (weight fraction iron in solid)
The solids growth rate is determined by,
Growth rate of solids _ 2 x solid generation rate
(length/time) ~ (Reactor volume)(solid density)
(solid concentration)(Specific
surface of solid)
45
-------
TECHNICAL REPORT DATA
(Please read Instructions on the reverse before completing)
1 REPORT NO.
EPA-600/2-78-143
4. TITLE ANDSUBTITLE
HYDROLYSIS OF IRON FROM ACIDIC LIQUORS
5. REPORT DATE
July 1978 issuing date
6. PERFORMING ORGANIZATION CODE
3. RECIPIENT'S ACCESSIO(*NO.
7 AUTHOR(S)
Alan D. Randolph
Richard D. Williams
8. PERFORMING ORGANIZATION REPORT NO.
PERFORMING ORGANIZATION NAME AND ADDRESS
Department or Chemical Engineering
University of Arizona
Tucson, Arizona 85721
10. PROGRAM ELEMENT NO.
1BB610
11. CONTRACT/GRANT NO.
R-802390
12. SPONSORING AGENCY NAME AND ADDRESS
Industrial Environmental Research Laboratory
Office of Research and Development
U.S. Environmental Protection Agency
Cincinnati, Ohio 45268
13. TYPE OF REPORT AND PERIOD COVERED
Final
14. SPONSORING AGENCY CODE
EPA/600/12
15. SUPPLEMENTARY NOTES
Project Officer: Mary Stinson, HERLEdison, NJ
16. ABSTRACT
Laboratory studies demonstrated the feasibility of a newly invented retained
solids iron crystallization process for treatment of synthetic acidic copper leach
liquors. Retained solids crystallization removes iron in the form of ferric oxide
suitable for iron smelting and regenerates sulfuric acid for recycling to the leach-
ing process. This process provides high yield iron removal under less severe
process conditions (such as lower pressure and temperature) than the conditions re-
quired for present continuous crystallization technology. A 2-gallon continuous
pressurized reactor was built for the study. Free acid and ferric iron concentra-
tions up to 20 and 30 g/1, respectively in the feed were processed at temperatures of
135°C to 165°C. Ferrous iron was oxidized to ferric by overpressure of oxygen in
the system and the oxidation rates were enhanced by oxygen sparging. Iron oxide was
contaminated with a slight amount of sulfate ion but was practically free of other
metal ions. Free acid values as sulfuric acid were regenerated. This research
study was motivated by removal of iron from copper leach solutions but also has
important implications in other hydrometallurgical leach systems.
17.
KEY WORDS AND DOCUMENT ANALYSIS
DESCRIPTORS
b.lDENTIFIERS/OPEN ENDED TERMS
c. COS AT I Field/Group
Sulfuric acid*,
Reclamation,
Waste treatment,
Crystallization
Hydrolysis
Copper leach liquor*
Retained solids
Ferric oxide*,
Iron removal,
Sulfuric acid reclamatior
Ferrous sulfate*
13B
13 DISTRIBUTION STATEMENT
Release to public
19. SECURITY CLASS (ThisReport)
Unclassified
21. NO. OF PAGES
56
20. SECURITY CLASS (Thispage)
Unclassified
22. PRICE
EPA Form 2220-1 (9-73)
: 1978 757-140/1397 Region 5-11
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