-------
The total pressure drop as a function of time determined by correlation given
by Equation (IV-5) is plotted as a solid line on Figures IV-3, IV-4 and IV-5.
Again good agreement between experimental and predicted pressure drops can be
seen.
Using standard pressure drop correlations for packed columns and assuming
constant specific area of the packing and constant intrinsic rate of scale
formation in the column appears to give good agreement between predicted and
experimental pressure drops. The concept of intrinsic rate of scaling is
therefore a useful measure of the scaling taking place under given operating
conditions.
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B. Experimental Results
1. pH Control
Preliminary experiments were carried out to determine the effects of
mechanical deposition of solids in the packed bed scrubber by circulating
a slurry of 15% solids content in the absence of sulfur dioxide or oxygen.
From the increase in pressure drop across the column it was obvious that
some solids deposition was occurring even in the absence of scrubbing or
oxidation. However, with a 6% solids concentration the problem disappeared.
Further experiments were carried out to examine the extent of scaling
in the presence of 5% calcium sulfate solid, but without oxygen in the
flue gas to cause further oxidation. The process water was decanted from
the solid and used to prepare fresh make up slurry to ensure that a
saturated solution was maintained and simulate closed loop operation.
Results from these runs are summarized in Table IV-3. With the inlet pH
at 6.2 - 6.3 appreciable scaling was observed at 100°F and the rate was
a factor of three higher at 125°F. Scaling at 125°F was much less when
calcium sulfate was excluded from the slurry, suggesting that the presence
of calcium sulfate is a factor in calcium sulfite scaling. However, when
the stoichiometry was reduced (Run 33) to control the pH between 5.8 -
6.0 the scaling was reduced to about 15% of that observed at pH 6.2 - 6.3
(Run 31). Thus, pH control below pH 6 has a dramatic effect in reducing
calcium sulfite scale, presumably because it allows much more of the sulfite
to remain in solution minimizing precipitation in the scrubber. Operation
at this lower pH has an adverse effect on efficiency. S02 removal is
15-20% lower but it also provides a lower stoichiometry and more efficient
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TABLE IV-3 SUMMARY OF CALCIUM SULFITE SCALING DATA
Run
2000 ppm P ,
so2
Solid Cone.
5 scfm, L/G = 50
Oxygen
pH Inlet
Make-up used 85% Recycled Process Water
Time on
Line (hrs)
Intrinsic
Scaling ,
Rate (hr )
Stoichio- Temp. SO- Removal
metry (°F) Efficiency (%)
to
K3
30
31
32
5% CaSO.
1% CaCO;
5% CaSO,
1% CaCO,
5% CaSO,
1% CaCO;
6.2-6.3
6.15-6.3
5.7-6.4
28
33
0.013
0.045
0.011
0.94 100 65-75
For pressure
drop from 0.25
to 1.5 in. of
water.
0.94 125 60^.75
For pressure
drop from 0.25
to 2.0 in. of
water.
0.94 125 65-80
For pressure
drop from 0.5
to 3.0 in. of
water
I
5
[T
33
5% CaSO,
0.5% CaCO.
1) 5.8:0-20 hr
2) 6.0:20-38 hr
3) 6.3:38-58 hr
4) 5.8:58-72 hr
72
0.0067
0.42-0.65
125 45-55
-------
utilization of calcium carbonate. The loss in absorption efficiency can
be counteracted by a change in scrubber design or by the use of additives
with higher stoichiometry (described in Section 3).
2. Seed Crystals
If supersaturation is minimized by acceleration of the precipitation process
away from the walls of the containing vessel, scaling can be avoided. In some
systems, this can be achieved by seeding the solution with particles of a
particular type, preferably of the same polymorph as the precipitating species.
Nucleation and growth of crystallites is thereby enhanced, and precipitation
is largely confined to the bulk of the solution. Even where crystallization
occurs on the walls the deposits may be softer and easier to remove.
Seeding was one of the earliest scale control methods to be used in wet
scrubbing plants, and is still applied more or less successfully in modern units.
One of the first wet scrubbing systems for control of sulfur dioxide emissions
from power plants was set up in Fulham, England around 1935. Blockage became
severe enough to force a shut-down after 72 hours. (A scale 2-3 inches thick
was found on portions of the scrubber surfaces). Hard deposits of pure gypsum
were found in an area of the scrubber that had only been exposed to clarified
liquor and it became clear that a crystallization phenomenon or chemical scaling
was involved.
(2)
Lessing's work showed that if supersaturation of calcium sulfate were to
be limited in the scrubber, a high proportion of solids in the recirculating
liquors was necessary to seed the crystallization adequately. Actual times for
the desupersaturation of solutions which had the necessary slight degree of
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supersaturation required to prevent scaling were longer than those found in the
laboratory (with similar seed concentrations). The crystalline shape of the
calcium sulfate was found to seriously limit the effectiveness of calcium sulfate
seeding.
More recently, Research-Cottrell found that an increase in slurry
concentration from 2% to the 4-8% range helped to reduce scaling. The results
(9)
were in part confirmed by TVA in their 4-8 cfm scrubber . Circulation of
1% solids led to rapid scaling, while 4% solids helped to alleviate scaling.
However, delay times were also different in the two cases. Mitsubishi
reports the use of seed crystals as one scale control measure, but does not
present quantitative data on total solids or percentage of sulfite and sulfate
in the scrubber liquor.
There are several practical difficulties in attempting to use seeding as an
effective scale-control technique. These include:
a) Sludging and buildup of seed particles in stagnant zones—
particularly with calcium sulfate hemihydrate
b) Change in the crystalline form of calcium sulfate over
time, and an attendant loss of effectiveness.
c) Reduction of the efficiency of calcium sulfate seeds
in the presence of calcium sulfite.
Seeding Experiments
To investigate the effects of calcium sulfate scaling, we used as a control
run a carbonate slurry which was saturated with calcium sulfate, but contained
no sulfate solid. The scrubber was operated at 125°F with 4% oxygen in the
flue gas, and scaling was quite rapid causing shutdown of the unit in about 12
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hours (Runs 36 and 45). The scrubber was also operated below pH 6 in order to
minimize any possible calcium sulfite deposition. Experiments to show the effect
of different seed crystals - gypsum, hemihydrate and anhydrite - are summarized
in Table IV-4. With 1% gypsum,seed crystals the scaling rate decreased to about
40% of that observed for the control run in spite of the fact that gypsum is
not the thermodynamically stable form at 125°F. Concentrations of gypsum solid
greater than 1% did not appear to reduce the rate of scaling much further, which
is somewhat contrary to literature observations which suggest that the optimum
is found in the range of 3-5% gypsum solids. It may be that a proportion of
the seed crystals in prior work were inactive, due to aging or crystal form.
Experiments were also carried out with calcium sulfate hemihydrate and anhydrite.
Both of these were much worse than gypsum, and hemihydrate gave higher scaling
rates than the control run.
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TABLE IV-4 SUMMARY OF CALCIUM SULFATE SEEDING DATA
2000 ppm Pgo , N2 = 5 scfm, 4% 0^
Run
36
37
35
40
46
Solid Cone.
Sat. soln CaSO,
0.5 CaC03
1% CaSO,
0.5% CaC03
1% CaSO
0.5% CaC03
1% CaSO,
0.5% CaCO
1% CaSO
Qi.5% CaCO.
Additive pH inlet
None 5.6-6.0
Seed crystal 5.6-6.0
gypsum
Seed crystal 5.6-6.0
gypsum
Seed crystal 5.6-6.0
Hemihydrate
Seed crystal 5.6-6.0
Anhydrite
L/G = 50, 85% Recycled Process Water
Intrinsic
Time on Scaling _1
Line (hrs) Rate (hr~ )
12.25 0.027
37 0.0086
32 0.0111
15 0.036
16 0.0200
Stoichio Temp, S02 Removal
metry (°F) Efficiency (%)
0.5-0.75 125 45-55
0.5-0.8 100 50-60
0.5-0.65 125 40-50
0.4-0.6 125 45-55
0.45-0.55 125 45-55
c
-1
a
'
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3. ADDITIVES
Threshold additive treatment, although poorly understood, has become
the most common scale control method used in contemporary desalination
processes. Basically, it has been found that small quantities of certain
polymeric electrolytes inhibit scale growth when present in far less than
stoichiometric amounts. Although many fundamental studies have been carried
out, the mechanism of scale control by threshold addition has never been
adequately explained (and more than one mechanism may be operative).
Phosphate Additives
Sodium hexametaphosphate (Graham's Salt) was one of the first additives
successfully used to treat boiler scale. It is known that hexametaphosphate.
addition greatly increases the negative mobility of the calcium ion, as
though strongly negatively charged complexes were formed. However, the
effective dose of hexametaphosphate is far too low for a sequestering
mechanism to be tenable. It has been shown that surfaces treated with
hexametaphosphate retain their scale retarding properties even when con-
tacted with supersaturated carbonate solutions not containing the phosphates.
This evidence, combined with the observation that crystals formed in the
presence of hexametaphosphate are distorted in structure, suggests a
surface absorption mechanism in which crystal nucleation and growth is
inhibited.
Organic Additives
A range of acrylic acid homopolymers and copolymers was synthesized
and tested for alkaline scale suppression in a pilot plant evaporator (at
around 3 ppm). With a polymer of this type containing 66% acrylic acid,
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it was shown by using conductance measurements to follow ionic concentration,
that the mechanism of scale suppression was an interference with crystal-
lization kinetics which results in a modified crystal habit. During the
evaporation of calcium sulfate solutions in beaker tests, specific conduc-
tance increased as the concentration rose, but then fell rapidly as the
scaling rate exceeded the rate of concentration of the solution. Four
ppm was sufficient to extend the period for initiation of significant
precipitation from 10 minutes to 25 minutes. The presence of the polymer
(at 3 and 4 ppm) introduces an induction period in the scaling rate which
radically decreases the rate of crystallization. Thus, the effect of the
additive is to hold calcium sulfate in solution at levels well above the
equilibrium solubility limit in the absence of the additive.
The deposits formed in the presence of polymeric additives generally
have'a rather large organic content. This suggests considerable attrition
of additive material and hence an increase in additive costs. Reports that
acrylic-based polymers yielded self-cleaning scales are found to be probably
dependent on the presence of small amounts (M).2 ppm of aluminum, magnesium
or zinc.)
Scales formed in the presence of both polyelectrolytes and the metal
ions above were found to strip easily from the copper surfaces of a labora-
tory spray evaporator on exposure to air. The lower molecular weight
materials in the range 1000-16000 were most effective. This work indicates
how sensitive scale formation is to precise operating conditions particularly
with respect to trace impurities in the system.
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Selection of Additives for Preliminary Screening
Based on previous experience with anti-scaling additives and consultation
with vendors, the phosphate and polymeric additives listed in Table IV-5 were
selected for initial screening. Each of the additives selected has proven
effective in controlling scaling in particular types of lime systems. Nine
of the additives listed are proprietary products and their physical properties
are described in Appendix 5. The remaining four additives are commonly available
chemicals.
Sodium hexametaphosphate (NaPO_),, a so-called glassy phosphate, is a widely
used constituent of commercial inhibitors for controlling carbonate scale in
potable water treatment. Sodium pyrophosphate is reported to be the major
scale inhibiting constituent in multi-stage distillation plants in Kuwait. A
combination of sodium silicate and sodium hexametaphosphate has proven to be
more effective in reducing deposits in potable waters than either material
alone. Sodium tripolyphosphate is one of the constituents of PD-8, and has been
used widely for scale control in desalination operations.
Initial Screening Tests
Initial exploratory experiments were conducted with the objective of developing
conditions which lead to significant and reproducible scaling so that anti-scaling
\
additives might be assessed in a quantitative manner. Details of the apparatus
and test results are presented in Appendix 2.
In addition to the initial slurry composition, several other variables which
may affect the rate of scaling have been identified. These are: the rates of
addition of SO,, oxidants and sulfuric acid, temperature, residence time in the
V *•
reacting area, and the type of surface. Since an exhaustive investigation of
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TABLE IV-5
ANTI-SCALING ADDITIVES SELECTED FOR PRELIMINARY SCREENING
Trade Name
Lime Treet
Acrysol A-l
Acrysol A-3
Acrysol A-5
Calnox 214 DN
Darex 40
Dequest 2000
PD-8 (formerly
Hagevap LP)
Sodium hexa-
metaphosphate
Calgon CL-14
Manufacturer
Dearborn Chemical Division
W.R. Grace and Company
Rohm and Haas Company
Rohm and Haas Company
Rohm and Haas Company
Aquaness Chemical Co.
Hous ton, Texas
W.R. Grace and Company
Cambridge, Mass.
Monsanto Company
St. Louis, Missouri
Bull & Roberts, Inc.
785 Central Avenue
Murray Hill, N.J.
Calgon Company
Sodium Pyrophosphate
Versenex 80 Dow Chemical Company
Quadrol
10 ppm Metso and
10 ppm sodium
hexametaphosphate
Sodium tripoly-
phosphate
Wyandotte Chemicals Co.
Type of Additive
Mixture of alkaline material
and a synthetic non-ionic
organic polymer
Polyacrylate, MW> 50,000
Polyacrylate, MW>150,000
Polyacrylate, MW>300,000
Aqueous solution of
polyacrylate, MW - 750
Sodium polymethacrylate
Phosphoric acid analog of
EDTA (ethylenediamine
tetracetic acid)
Mixture of sodium tripoly-
phosphate, lignin sulfonate,
and an anti-fearning agent.
Liquid organic formulation,
containing a proprietary polymer
and an organic phosphorous
compound.
Aqueous solution of pentasodium
diethylenetriaminepentacetic acid
N, N, N1, N^tetrakis (2-
hydroxypropyl)-ethylenediamine -
Mixture of sodium silicate (Metsc)
and sodium hexametaphosphate
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100
80
60
>
c
8
£ 40
20
I
4 6 8 10
Weight of Scale Formed on Test Coupons, mg/in.
FIGURE IV-8 SCALE ANALYSIS VERSUS SCALE WEIGHT
12
14
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all these variables at several levels is impractical in screening tests,
those combinations which seemed from experience most likely to produce
rapid and reproducible scaling were chosen.
Preliminary experiments with panel materials showed that mild steel
and brass were attacked by the acidic slurry, and that little scale formed
on Teflon. However, since the primary concern was with anti-scaling
additives, subsequent tests were limited to panels of 304 stainless steel
and PVC coated mild steel, both of which collected scale deposits without
corrosion.
Lime Treet, and Acrysol A-3 and A-5 were found to have definite
negative effects resulting in increased scaling. However, Lime Treet was
selected for bench-scale testing to provide a point of comparison.
Darex, Versenex 80, Quadrol, sodium silicate-sodium hexametaphosphate,
sodium pyrophosphate and sodium tripolyphosphate exhibited small effects
(either positive or negative) and were felt to be of little use compared
to the more promising additives examined in the bench scale scrubber.
X-ray Analysis/Scale Deposits
The X-ray diffraction analyses of scale deposits removed from coupons in
the screening runs show a general trend in composition as a function of thickness,
noted graphically in Figure IV-8. As a rule, sulfate deposits appear first as
2
the scale is initially formed and until it grows to a certain thickness (<1 mg/in )
Carbonate and then sulfite deposit after this thickness has been reached. The
deposition of calcium carbonate, which is not normally found in bench scale or
full scale scrubber experiments is probably due to the use of a very large excess
of carbonate in the screening experiments. Thus, the growing scale appears
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to collect carbonate solid from the slurry. As the scale thickness increases,
sulfite scale predominates with some carbonate deposition and essentially no
further sulfate deposition. This suggests that sulfate is an important precursor
or initiator of overall scale formation.
Bench Scale Scrubber Tests
Five of the more promising additives identified in the screening program
were tested under conditions of the control run which was conducive to high rates
of calcium sulfate scaling. Results in Table IV-6 show that the qualitative
agreement between the screening and bench scale scrubber tests are quite good.
Details of the bench scale scrubber tests are summarized in Table IV-7. One
additive "Lime Treet" was found to give poor results as it did in the screening
experiments and the run was terminated. All the other additives showed im-
provements over the control run. Bequest 2000 and PD-8 were both tested at
two concentrations, sodium hexametaphosphate at one concentration. These three
additives tended to lower the pH which was compensated by increasing stoichiometry
(with a consequent improvement of 15-20% in SO. removal efficiency). The best
additives found were Calnox 214 DN and Calgon CL-14 which reduced the rate of
scaling by almost 75% of that observed in the control experiment. Again, a
buffering effect and improvement in S0« removal efficiency was observed with
these two additives. Calnox 214 DN was tested at three concentrations. The
highest concentration, 0.25 ml/1 was found to be the most effective. A
synergistic effect was also found when Calnox additive was used with gypsum
seed crystals. The scaling was further reduced although it was observed that
3% gypsum showed a higher rate of scaling than the 1% gypsum slurry.
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Table IV-6 COMPARISON OF ANTI-SCALING ADDITIVES IN
SCREENING AND BENCH-SCALE SCRUBBER TESTS
Additive
Calnox 214 DN
Calgon CL-14
Sodium hexameta-
phosphate
PD-8
Dequest 2000
None (control)
Lime Treet
Hours of
Scrubber
Operation
48.5
44.5
31
27.5
20
12.25
3
Scale Weight
2
mg/in (screening tests)
PVC S. Steel
1.4
3.5
0.6
1.5
0.7
7.5*
22.6
2.0
2.2
0.8
0.9
0.8
7.2*
11.8
Type of Additive
Sodium polyacrylate
(M.W. about 750) +
lignosulfonates
Aminomethylenephos-
phorate (AMP)
Sodium tripolyphosphate
+ lignosulfonates
Polyacrylate
Synthetic Polymer
* average of 3 runs.
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TABLE IV-7 SUMMARY OF ADDITIVE DATA
2000 ppm P , N = 5 scfm,
S02 2
Co
Oi
!J>
irthur D Lii
Run
45
38
39
41
42
43
44
47
48
49
Solid Cone.
Sat. soln CaSO,
0.5% CaC03
Sat. soln CaSO,
0.5% CaC03
Sat. soln CaSO,
0.5-1.0% CaC03
Sat. soln CaSO,
0.5-1.0% CaC03
Sat. soln CaSO,
0.5-1.0% CaC03
Sat. soln CaSO,
0.5% CaC03
Sat. soln CaSO,
0.5% CaC03
s
Sat. soln CaSO,
1% CaC03
Sat. soln CaSO,
1% CaC03
1% CaSO,
1% CaC03
Additive
None
Deques t 2000
(.005, .05 ml/1)
Sodium hexameta-
phosphate (0.02 g/1)
PD-8
(0.02, 0.04 g/1)
Calnox 214 DN
(0.025 ml/1)
Calgon CL-14
(0.05 ml/1)
Lime Treet
(0.005 g/1)
Calnox 214 DN
(0.01 ml/1)
Calnox 214 DN
(0.017 ml/1)
Gypsum
Calnox 214 DN
(0.05 ml/1)
4% 02 , L/G = 50, 85%
pH Inlet
5.6-6.0
5.6-6.0
5.6-6.0
5.6-6.0
5.6-6.0
5.6-6.0
Lime Treet,
3 hours due
5.6-6.0
5.6-6.0
5.6-6.0
Time on
Line (hrs)
13
20
31
27.5
48.5
44.5
insoluble in
to excessive
19.5
22
70
Recycled Process Water
Intrinsic
Scaling _, Stoichio-
Rate (hr ) metry
0.0246
0.0174
0.011
0.0124
0.0071
0.0077
alkaline
settling
0.0164
0.0145
0.00457
0.5-0.7
0.6-1.4
0.6-1.2
0.7-1.4
0.6-1.0
0.6-0.80
solution. Run
of solids.
0.65-0.85
0.4-0.7
0.6-0.9
Temp. S02 Removal
(°F) Efficiency (%)
125 50-60
125 55-70
125 50-70
125 60-75
125 60-70
125 55-70
terminated after
125 60-70
125 45-65
125 55-65
o
50 3% CaSO.
1% CaCO,
Gypsum
Calnox 214 DN
(0.1 ml/1)
5.6-6.0
51.5
0.00622
0.8-0.9 125
55-70
-------
C. Implications of the Scaling Model for Scrubber Operation
The scaling model works quite well for the packed bed scrubber and it
should be possible to extend this model to other types of scrubber. For
example, an intrinsic scaling rate for the TCA pilot plant scrubber can be
estimated from the time to catastrophic scaling and the change in void
fraction of the grid using Equation IV-3:
e - e ,
o , -1
Similarly the time to catastrophic scaling can be estimated by using an
appropriate form of the Leva correlation to relate pressure drop to void
fraction. Changes in pressure drop can then be used to define intrinsic
rates of scaling and predict the time on line before flooding of the column
occurs due to scale build-up.
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V. SULFITE OXIDATION
A. Introduction
Scale formation incorporating calcium sulfate is a major problem in
limestone wet scrubbers due to the oxidation of calcium sulfite to calcium
sulfate by oxygen in the flue gas. The extent of oxidation in different
scrubbers is quite variable and the conditions controlling oxidation and the
reaction mechanism are not well understood. Oxidation of calcium sulfite under
controlled laboratory conditions has not been widely studied although a related
reaction, oxidation of sodium sulfite, has been extensively investigated. However,
in spite of this detailed study over many years, the mechanism of the latter
reaction is still not completely understood.although it appears to be very
rapid, and mass transfer effects can be important. Most literature data refer
to oxidation of relatively concentrated sodium sulfite solutions in alkaline
solutions, whereas the present interest is in oxidation of sulfite in calcium
carbonate/calcium sulfite slurries at neutral or acid pH.
The objectives of this limited present program were:
1. to investigate the variables controlling the oxidation
of dilute sodium sulfite solutions, particularly as a
function of pH.
2. to investigate the variables controlling the oxidation
of calcium sulfite using the previous results as a
baseline.
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A continuous stirred tank reactor (CSTR) was chosen to carry out the ex-
periments because this type of reactor is simpler to treat in terms of the
kinetic expressions, although the interfacial contact area between the liquid
and gas phases is not well defined. Our experiments with sodium sulfite were
intended to provide an appropriate baseline which could be compared with
literature data and with the calcium salt data.
Absorption in Stirred Vessels
It is possible to distinguish three separate regions of overall absorption
rate characterized by the dependence of the reaction rate on the rate of stirring,
or rate of impeller rotation. Figure V-I shows these three zones. In Region 1
the oxidation rate is slow and strongly dependent on the gas flow rate, essentially
independent of the agitation rate. In Region 2 the oxidation rate is increasing
rapidly; the agitation rate is more important and the gas flow rate is of less
importance. In Region 3 the reaction rate has become independent of both
agitation rate and gas velocity. The mass transfer rate, therefore, has been
increased by the intensity of agitation, until it has become equal to, or faster
than, the chemical reaction rate. At this point, the chemical reaction is rate-
controlling, and the overall rate can be increased only by changing factors
such as temperature, catalyst concentration and reactant concentration, which
would increase the chemical reaction rate.
It is possible to distinguish the system under mass transfer control from
that of the chemical reaction rate control by measuring the overall reaction rate
under conditions of varying impeller speed, gas flow rate, and reactant concen-
tration. For the system under mass transfer control, gas absorption rates will
be strongly dependent on agitation rate, on gas partial pressure, and perhaps
on gas flow rates, and will be substantially independent of temperature and
reactant concentration.
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L
I
I
a
o
T3
>r4
X
O
Oxidation rate computed
from 0 transfer rate
Region 3
Impeller Speed (RPM)
Figure V-l Oxidation Rate vs. Impeller Rotation
& Representative Curve for a Stirred Vessel)
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B. Evaluation of Literature Data on Sodium Sulfite
A detailed comparison of the results with available literature data is
difficult because of the wide range of techniques and reaction conditions
employed by other workers. We have reexamined much of the literature data and
tried to assess its salient characteristics relative to the work in progress.
Much of the literature is concerned with use of the reaction to determine mass
transfer characteristics in various types of equipment, rather than the basic
kinetics of the reaction. Therefore, these experiments tend to concentrate on
higher oxygen concentrations (air) and relatively high sulfite concentrations.
Reaction Order
When sulfite is present in a large excess, then the reaction might be expected
to become proportional to some power of the oxygen concentration but independent of
sulfite concentration. The reverse should be true when oxygen is in large excess and
the rate can become proportional to the sulfite concentration. In reality, these
represent limiting cases and the reaction order becomes variable and much more
complex. Fractional orders are possible due to changes in the potential rate
controlling steps. For the heterogenous reaction of oxygen (gas) with sodium
sulfite solution in a continuous stirred tank reactor, the situation is further
complicated by the mass transport limitations. These include absorption of
oxygen into the liquid and diffusion of oxygen and sulfite ions in the liquid
phase. The process of oxygen absorption is itself complex and is dependent on
such parameters as gas flow rate, oxygen partial pressure, vessel design and
the interfacial area between the gas and liquid phase. The latter is controlled
by the impeller design and speed of rotation.
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We have summarized in Table V-l the literature data showing the order of
reaction and the range of reaction conditions under which these parameters were
measured. It can be seen that the results vary between 0 and 2 for both oxygen
and sulfite ion. The range of sulfite concentration studied varies from 1 to
_3
10 moles per liter. However, in general the results follow the expected
direction. For example, Astarita?found the reaction was second order in oxygen
at high sulfite concentrations and zero order at low sulfite concentrations. Most
workers found that the reaction was first order with respect to sulfite, but at
high sulfite concentrations (>0.5 mole/1) the reaction rate becomes zero order
in sulfite. At the other extreme, Rand and Gale found the reaction approximately
_3
second order in sulfite at very low sulfite concentrations (10 mole/1).
Most of the literature data refers to reaction with air (i.e., dissolved
oxygen in equilibrium with air). We are concerned with a much lower oxygen
partial pressure (4%); this would tend to make the overall reaction rate more
dependent on the oxygen absorption rate. However, the range of sulfite concen-
tration of interest is also somewhat lower than the ranges previously investigated.
Thus, it is difficult to predict, a priori, from the literature data, what the
pH Effects
Data in the literature, ' although not comprehensive, indicate that the
overall oxidation rate of sodium sulfite decreases with pH. There is disagree-
ment on the magnitude of the effect; however, if should be noted that the
two authors used widely different measuring techniques.
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TABLE V-l SUMMARY OF LITERATURE DATA ON REACTION ORDER
Reference
Relth1
f
Astarita et a]/
De Vaal & Okeson"
(catalized)
Srivastava et al
Rand & Gale
pH 6.5-7.7
pH 4.2-6.4
Cooper et al
Fuller & Christ7
Westerterp et al
(catalyzed)
8
SO Cone.
mole/ft
0.8
.06
.25
.25-1.0
0.8
>0.4
io~-
0.1-1
<1.5 x 10
0.02
0.01
~2
02 Cone.
mole/Jl
Reaction
~4
6 - 24 x 10
D0>0.8 Mg/£
(in = with air)
Sat. soln.
Order
0~
0
1
2
2.3
1
1
1
0
1
0
0
1
0
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C. Experiments with Sodium Sulfite
(Mass Transfer Effects)
The literature data of Cooper, Fernstrom and Miller on the oxidation of
sodium sulfite in aqueous solution catalyzed by cobalt naphthenate found that
there was no evidence of chemical reaction rate control, even up to fairly high
agitation power inputs. The overall reaction rate was independent of sulfite or
sulfate concentration, and strongly dependent on agitation rate and gas flow rate.
All the data we have obtained for sodium sulfite also appeared to show the
reaction was dependent on impeller rotation rate, even at the highest rates
0\/700 rpm) , obtainable in the apparatus, although there is perhaps some indi-
cation that we are close to the chemical, reaction rate limit in a few experiments
(e.g., the data in Figure 5, Appendix 2).
As further confirmation, we have used the equation proposed by Westerterp
8
and co-workers to find what they define as the "critical impeller speed".
Westerterp and co-workers have shown that there is a critical impeller speed
below which both gas velocity and impeller speed will affect the mass transfer
coefficient. Above this critical impeller speed the mass transfer coefficient
is independent of the gas velocity. This critical impeller speed is defined by
n in the equation shown below. It is notable that n > defined by:
nQD = [og/p]0'25 [A+ G (T/D)]
where D = impeller diameter, ft.
l>g/p] °'25 • 1950 ft/hr. for water at 25°C
A + B are constants and functions of impeller shape, for a
flat bladed turbine A = 1.22, B = 1.25
T = Tank diameter, ft.
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is usually quite high and calculation from this equation gives a value in our
apparatus of n = 726 rpm. This is approximately the upper limit of rotation
rate in the present apparatus; therefore, all the results obtained are below
this critical impeller speed, and should be a function of the gas velocity and
impeller speed.
On this basis, our results should be comparable with the work by Cooper, et
al. , in a batch stirred tank reactor with a similar configuration to the present
equipment. Using their data (stirring speed of 600 rpm, sodium sulfite concen-
tration =0.1 moles/1, oxygen concentration = 21%), we derive a mass transfer
-2 3
coefficient, K a = 1.92 x 10 Ib moles/ft hr. atm.
O
From the equation below, derived from Cooper's and other data, we can
calculate the absorption (or oxidation) rates to be expected in our apparatus.
K a = 6.6 x 10~6 V °'67 (P/VT)0'76
S o Li
where VQ = superficial gas velocity in vessel ft/hr. (based
on empty cross section)
P = power input to liquid gas mixture in vessel from
impeller, ft Ib f/min.
V = volume of liquid in vessel, cu ft.
The calculated values are shown in Table V-2, together with the measured oxidation
rate. It can be seen that the measured values are higher than the calculated
values by a factor of 2, reasonable agreement considering the differences in
apparatus and experimental technique. The equation also predicts the increased
oxidation rate due to higher gas flow rates at 4% oxygen.
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TABLE V-2 COMPARISON OF CALCULATED AND MEASURED OXIDATION RATES
Oxygen %
4
20
4
Gas Velocity Rate of Oxidation (Ib moles/ft hr)
ft /hour Calculated Measured
30
30
60
0.57 x 10
-3
2.8 x 10
-3
1.0 x 10
-3
1.3 x 10
-3
5.2 x 10
-3
2.0 x 10
-3
TABLE V-3 EFFECT OF GAS FLOW RATE ON OXIDATION RATE
Nominal Sulfite
Concentration
(mmoles/£)
Oxygen
Concentration %
Oxidation Rate at 10
Compared with 5 &/min
10
50
100
50
100
4
4
4
20
20
no difference
higher at 10
higher at 10 £/min
no difference
no difference
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All the experiments were dependent on the impeller speed, but not all
experiments were found to be dependent on gas flow rate. Results which were
obtained for different solutions are summarized in Table V-3.at gas flow rates
of 5 liters/minute and 10 liters/minute respectively. With 20% oxygen in the
gas stream and nominal inlet concentrations of 0.1 molar and 0.05 molar sulfite
solution no effect of gas flow rate was observed. A significant difference was
observed for the same two sulfite concentrations at 4% oxygen in the gas stream;
however, when the sulfite concentration was reduced to 0.01 molar with. 4% oxygen,
then no effect of gas flow rate was observable.
Effect of Oxygen Concentration
Thus, the experiments with sodium sulfite have shown that even at low oxygen
concentration in the gas phase, equivalent to flue gas, and low concentrations
of sulfite in solution, the reactions may still be mass transfer controlled.
Therefore, one would expect the reactant concentration, temperature, and catalyst
concentration to have little or no effect on the overall reaction rate. Certainly,
the oxygen partial pressure in the gas phase must be considered one of the most
important variables. Experiments were carried out [Figure 3, Appendix 2] up to
40% oxygen which did not show any limit to the reaction rate in sodium sulfite
solution-
The experiments showed that the data could be represented on a log-log
plot (Figure V-2) by the equation below,
RQX = 78.7 x 10"6 [P (O^]1'24 moles/1.sec)
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30
20
10
u
0)
en
to
0)
r-{
i
vO
O
0)
a
o
•rl
0.03
0.05
0.1
0.5
Figure V-2. Log Oxidation Rate vs. Log Partial Pressure of Oxygen
for 400 mmoles/1 Na.SO Feed (Impeller Rotation = 375 RPM)
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showing that the data is approximately proportional to the oxygen concentration,
although apparently not an integral number. This suggests that oxidation could
be more rapid in the hold tank where the liquid is in contact with 20% oxygen,
than in the scrubber where it is only in contact with the 4% oxygen present in
the flue gas.
Effect of pH
Data on the oxidation rate of sodium sulfite as a function pH (changed by
sulfuric acid addition)was plotted to evaluate the effect of hydrogen ion con-
centration and/or sulfite concentration. Although the change in oxidation rate
as a function of pH is quite noticeable, in terms of hydrogen ion concentration,
it leads to a relatively low exponent. It can be seen in Figure V-3 that the
oxidation rate is inversely proportional to hydrogen ion concentration over the
range from pH 9 to pH 3. There is no .sign of any upturn at the lower end of
the curve.
Effect of Impeller Speed
These data were then combined with the data showing the effect of impeller
speed (I) at various sulfite concentrations. It was found that all this data
could be consistently expressed by the equation given below (Figure V-4) which
shows that the oxidation rate is proportional to impeller speed and inversely
proportional to hydrogen ion concentration (or perhaps directly proportional
to hydroxyl ion concentration). It was found that the most consistent set of
RQX = 3.1 x 10"10 [I]1>12/[H+]Q'12
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20
10
u
0)
CD
0)
rH
i
vo
o
rH
X
(moles/1)
-1
Table 1 76 mmoles/1
i
Table 3 10 mmoles/1
10"
10
6
10
10
=r
-t
D
c:
Figure V-3. Log Oxidation Rate vs. Log [1/H ] for Na-SO Feed at 4% Oxygen
(Impeller Speed = 600 RPM), Tables 1 and 3.
o
-------
1.0
0.5
0.2
0.1
CM
i-H
--- 50 mmoles/1 - Table 8
10 mmoles/1 - Table 9
100 mmoles/1 - Table 7
Impeller Speed (RPM)
100
200
500
1000
Figure V-A.Log Rate/[— ] vs. Log Impeller Rotation for
H
Na2SO Feed at 4% Oxygen, Tables 7,8, and 9.
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data was obtained by assuming the results were independent of sulfite ion
concentration, although the data do not fall completely on one line.
Further data to study the effect of pH changes at a higher concentration
of sulfite where the pH was again varied by the addition of sulfuric acid,
appeared to be relatively independent of hydrogen ion concentration, i.e.,
a lower exponent was observed. This is again indicative of the complicated
nature of the reaction showing that small changes in reaction conditions
have subtle effects on the rate controlling steps.
Summary
It appears that results obtained in the CSTR are comparable with those
obtained in the literature under similar conditions. Mass transfer effects
are important over the range of conditions studied and oxygen concentration
is the most important variable. Results are essentially independent of
sulfite concentration, but proportional to pH (from pH 3 to pH 9) which
could be interpreted as a dependence on hydroxyl ion concentration.
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D. Experiments with Calcium Sulflte
A summary of the results with the calcium solutions and slurries is
shown in Table V-4 with representative results of sodium sulfite for compari-
son. The first set of data was obtained on a saturated solution of calcium
sulfite after solid had been allowed to settle. At pH of 7-8 the concen-
tration of sulfite in solution is about 0.1 mmole per liter and the oxidation
rate is very low—less than 0.1 micromoles per liter per second. At a
pH of 3-4, the oxidation rate is higher by at least an order of magnitude.
The concentration of sulfite in solution (i.e., bisulfite ion) is 15-20
mmoles per liter. A log-log plot of the data from Table 10 Appendix 2 is
shown in Figure V-5. The oxidation rate was found to be proportional to total
sulfite concentration (ST) (effectively bisulfite with pH range of interest)
and independent of pH according to the equation:
Rox = 227 x 10"6 [S^1'16
The secdnd set of data with a calcium sulfite slurry of about 0.2% gave
oxidation rates of 1-2 micromoles/liter-sec. across the pH range 3-8, about
equal to the rate for the saturated solution at low pH, but much higher than
the saturation solution at high pH, showing that the solid is able to dissolve
fast enough under these conditions to cause an increase in the overall
oxidation rate.
The next group of experiments were carried out with calcium carbonate/
calcium sulfite mixtures. With the saturated solutions at pH 7-8, the reaction
rate appeared somewhat higher than for the sulfite alone, despite the lower
solubility of sulfite in the mixture. However, these data are close to the
limits of experimental determination. At the lower pH, the oxidation rate was
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TABLE V-4 SUMMARY OF OXIDATION RATES FOR CALCIUM
Oxidation rate (x 10 moles/1-sec)
Reactant pH 7-8 pH 3-4
CaSO. saturated solution >0.1 1-2
CaS03 slurry (0.2%) 1-2 1-2
CaC03/CaS03 saturated solution 0.2-0.3 1-2
CaCO_/CaSO- saturated solution + CuSO, 0.3
CaCO./CaSO, saturated solution + Calnox 0.5 1
CaC03/CaS03 slurry (0.2%) 2 5
CaC03/CaS03 slurry (0.2%) + CuSO, 2 5-6
Na2S03solution (10 mmoles/1) 3-4 1-2 (H2SO,)*
3-4 (S02)
pH adjusted with H2SO, and SO-,respectively.
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10.0
5.0
1.0
• o
0)
CO
• CO
0>
vO
o
0.5
0.1
• • < • j « • L
CSTR Sulfite Concentration (mmoles/1)
1.0
5.0
10.0
50.0
100.0
FigureV-5. Log Oxidation Rate vs. Log CSTR Sulfite Concentration
for CaSO_ Supernatant Feed (4% Oxygen, Impeller Speed
625 RPM).
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the same as for the saturated sulfite solution. The addition of CuSO, as
a catalyst and the anti-scale agent, Calnox, in different tests did not ap-
preciably alter the rates for mixed solutions. A slurry of calcium carbonate
and calcium sulfite was used to determine the effect of solids dissolution
in a mixed system. Rates were only slightly higher for sulfite slurry alone
at pH 7-8, but significantly greater at pH 3-4. These latter values were
the highest observed in all of the calcium sulfite systems; again, the
addition of CuSO^ did not greatly increase the oxidation rate in the mixed
slurry.
It is noticeable that all the calcium solutions gave higher oxidation
rates at the lower pH's. This is in direct contrast to the results for
sodium sulfite also shown on Table V-4, where the oxidation rate was about
3-4 micromoles/liter-sec, at pH 7-8, higher than any of the calcium results
observed at that level. On the acid side, the sodium data varied from
about 1-4 micromoles/liter-sec., close to the same order as the results
obtained for calcium solutions.
Summary
Saturated calcium sulfite solutions (no solid present) showed very low
oxidation rates above pH 7, but reached rates an order of magnitude higher
at pH 4. Oxidation rates were proportional to total sulfite concentration
(effectively bisulfite ion concentration). Oxidation of calcium sulfite
always resulted in a decrease in solution pH. The rate of oxidation of a
0.2% calcium sulfite slurry at pH 7 was much higher than the saturated
solution and equivalent to that of a calcium sulfite saturated solution at
pH 4. Thus, the rate of sulfite dissolution is a significant contributing
factor to the rate of oxidation under normal operating conditions.
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Saturated solutions of calcium carbonate/calcium sulfite oxidize at a slightly
higher rate than calcium sulfite solution. Above pH 6, oxidation causes an
increase in solution pH; below pH 6 oxidation causes a decrease in solution
pH. It appears that above pH 6 there is sufficient carbonate present to
neutralize the sulfate formed whereas below pH 6 sulfite is in excess.
Due to the dissolution effect noted earlier, mixed carbonate-sulfite
slurries oxidize more rapidly than the corresponding solutions. Neither
copper sulfate nor Calnox 214 DN appreciably affects oxidation rates. Since
oxidation rate is strongly dependent on bisulfite ion concentration in
calcium solutions, other cationic impurities such as sodium or magnesium,
which can dramatically increase bisulfite concentration in solution in the
range of pH 5-6, may have a strong effect on oxidation rate. Although the
effect of oxygen partial pressure was not studied with calcium solutions, the
effects noted with sodium sulfite suggest that the rate of oxidation might
be higher in the hold tank than in the scrubbing tower due to the higher
oxygen partial pressure over the hold tank.
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REFERENCES FOR CHAPTER V
1. Reith, T., Physical Aspects of Bubble Dispersion in Liquids, Thesis,
Delft Technical University (1968).
2. Astarita, G., Marucci, 6., and Gioia, F., Proc. 3rd Europ. Symp. on
Chemical Reaction Engineering, 195 Pergamon Press, London (1964).
3. de Waal, K.J.A., and Okeson, J.C., Chem. Eng. Sci. 21. 559 (1956).
4. Srivastava, R.D., McMillan, A.F., and Harris, J.J., Can. J. Chem.
Eng. 46_, 181 (1968).
5. Rand, M.C. and Gale, S.B. "Principles and Applications of Water
Chemistry," Ed. Faust S.D. & Hunter, J.V., Wiley (1967) .
6. Cooper, C.M., Femstrom, G.A., and Mille, S.A., Ind. Eng. Chem. 36,
504 (1944).
7. Fuller, B.C. & Crist, R. H., JACS 63, 1644 (1941).
8. Westerterp, K.R., Van Dierendonek & de Kraa, J.A., Chem. Eng. Sci.
18, 157 (1963).
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VI. MATHEMATICAL MODEL OF PACKED BED SCRUBBER
A. Material Balances
A flow diagram of the bench scale scrubber system is shown in Figure
VI-1. Since heat transfer in a packed column absorber is known to be very
rapid, the system can be assumed to operate isothermally; that is, the
variations in the liquid and gas temperatures throughout the system can be
ignored for all practical purposes when the heat of reaction is negligibly
small .
Overall Material Balances
The conservation of sulfur, carbon and calcium within the overall
system shown in Figure VI-1 can be written as
Sulfur
MS02.<) + MS02.1 + MS02»2+ M[S1M + ^S.M ' ^S.P ' M[S]P = AS
Carbon
m n + Mrn 1 + Mrn 7 + Mte]M + ^r M ~ ^r T. ' MtClp = Ar
C02,U CO2>J- C02»^ M C,M C,P PC
Calcium
M[Ca]M - MSCa,P ' M[Ca^p = ACa
Where M is the absorption rate of S02 in the i— piece of equipment,
2' (gmole/sec)
C02,i is the absorption rate of CO- in the i— piece of equipment,
(gmole/sec),
S. , is the total concentration of species k as solid in the
K I 4>l*
' liH stream, (gmole/liter),
[k]. is the total liquid phase concentration of species k in
the 1-& stream, (gmole/liter) ,
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GAS
IN
t
GAS OUT
PACKED
COLUMN
HOLD
TANK
MAKE UP
PURGE
MAKE UP
TANK
RECYCLE
PUMP
Figure VI-1. Line Diagram of Experimental Scrubbing Apparatus
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A, is the accumulation of species k as solid within the overall
system, (gmole/sec),
and M is the make-up rate, (liter/sec).
When i = 0,1,2 the subscript refers to the scrubber, recharge tank and
holding tank, respectively,
k = S,C,Ca the subscript refers to sulfur, carbon and calcium,
respectively,
and 1 = M,P the subscript refers to the make-up and purge streams, respectively.
The above equations apply to steady state operation of the scrubber
system (that is, steady state with respect to the liquid concentrations)
and a constant rate of scale formation as indicated through the use of the
accumulation terms (A. ).
Steady liquid concentrations may be inconsistent with the steady build-
up of solids in the scrubber system. For example, it is known that the
accumulation of scale in the scrubber causes an increase in the S02 removal
which in turn would affect the liquid compositions. However, this process
normally takes place slowly and it can be assumed that Equations VI-1
through VI-3 can be applied at each instant of time. This is equivalent to
a quasi-steady state assumption.
Subtracting Equation VI-3 from the sum of Equations VI-1 and VI-2 gives
I (M + M ) = M[A - A.J (VI-4)
1=0 S02»i CQ2' p
Where ^ = [S]^ + [C]^ - [Ca]^
Here the fact that the solid calcium salts have a one-to-one correspondence
of calcium to sulfur or carbon has been used to eliminate the solid concen-
trations appearing in Equations VI-1 through VI-3. It can be noted that
Equation VI-4 does not contain any solid compositions.
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A further simplification of Equation VI-4 is possible if it is noted
that A^= 0. This follows from the fact that the make-up slurry is formu-
lated by mixing solid calcium salts of sulfur and/or carbon with water; then
by stoichiometry AM = 0. Thus
Z (M + M ,) = MA (VI-5)
1=0 S02»i C02,i p
In the bench scale scrubber very little SC^ evolution occurs in the
hold or make-up tanks (except at very low pH) . This is concluded for
several reasons: (1) the tanks are for all practical purposes closed
containers, (2) very little surface area is available for mass transfer
between the liquid in the tank and the gas space above the liquid, and
(3) if the partial pressure of S(>2 builds up to a significant level in the
liquid phase, it will be preferentially discharged in the scrubber since
in the scrubber conditions are conducive to mass transfer whereas in the
hold or make-up tank they are not.
Under the assumption of no SC>2 evolution from the hold or make-up tank,
Equation VI-5 becomes
M + I M = MA (VI-6)
b°2»° 1=0 C02'i p
As defined above, the absorption rates Mon _ and M - . are positive for
" CC/O , i
absorption. They are negative in sign when material is evolved from the
system.
Without further assumptions Equation VI-6 does not appear to be useful.
However, with the assumption that the liquid in the holding tank maintains
equilibrium, the Delta concentration of the purge stream, A , and Equation
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VI-6 take on unusual significance and utility as discussed below. (It
is assumed that the slurry in the holding tank and the purge slurry has
the same composition.)
Analysis of Holding Tank under the Equilibrium Assumption
The idea that the liquid in the holding tank maintains equilibrium
(8)
has been used by Wen and Uchida to successfully analyze recycle scrubber
systems using limestone slurry and stirred holding tanks with residence
times between 10 and 40 minutes. The minimum time in which equilibrium in
a stirred tank is reached for limestone slurry systems is not well known.
(3)
Borgwardt reported a kinetic expression which could predict the volumetric
rate of sulfite disappearance from limestone scrubbing slurry fairly
accurately for stirred tank reactors indicating 4 to 6 minutes minimum
residence time are required for the solution to reach equilibrium.
Under the assumption of equilibrium of the holding tank, the delta
concentration, A , can be related to the liquid phase concentrations of the
other species present through equilibrium calculations. Figure VI-2 presents
a typical equilibrium graph of pH versus concentration. However, the reverse
is not true. It can be seen from Figure VI-2 that at certain delta concen-
trations two or three pH's are possible. Thus if A is calculated from
Equation VI-6, some mechanism must be provided to resolve the pH ambiguity
which might occur. This can be done by following the relationships between
pH and the amount of SO2 absorbed and CC>2 evolved from the system through
the use of a "titration" diagram.
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30
20
o
o>
1 10
o
§
c
E
3
Jotal Sulfur, [S]
Solution Saturated with CaSO-j
' C00~ '
Total Carbon [C]
A = [S] + [C] - [Ca]
pH of the Solution
FIGURE VI-2 EQUILIBRIUM CONCENTRATIONS OF TOTAL SULFUR, CARBON AND CALCIUM
AND THE DELTA CONCENTRATION AS A FUNCTION OF pH. SOLUTION SATUR-
ATED WITH CaSO3, TEMPERATURE IS 125°F AND PARTIAL PRESSURE OF CO2
RESTRICTED TO LESS THAN 0.2 ATM.
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"Titration" Diagrams
The scrubbing operation can be viewed as a process in which a volume
of make-up slurry is allowed to absorb a certain number of moles of 862
and C02 is allowed to evolve from the volume of make-up slurry. The C02
is evolved in accordance with the following rules.
1. No C02 evolution occurs until the partial pressure of
C02 in equilibrium with the make-up slurry to which S02
has been added reaches a certain maximum partial pressure
XT A Y
2. Once the partial pressure of CO^ reaches PCQ , only
enough C02 is evolved from the solution so as to maintain
MA V
the partial pressure above the solution at PCQ .
The process described above is similar to an ordinary titration
experiment. The progress of the titration can be followed on a titration
diagram which can be prepared using equilibrium curves such as the one
shown in Figure VI-2 and Equation VI-6 rearranged in the following form.
The. terms on the right hand side of this equation represent the moles of
862 and CC>2 added to the system per volume of make-up slurry, respectively.
In what is to follow, the construction of the titration diagram for
the following conditions will be described:
1) Make-up slurry is 0.5 weight per cent
2) The liquid temperature is 125 °F.
3) The limiting partial pressure of C02, P, is 0.2 atm.
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4) Only CaCO* and CaSC>3 can be present in solid form.
and 5) Sulfate ions are not present.
Figure VI-2 is the equilibrium diagram which is applicable at these
conditions. In this figure solid CaSO- is present at every pH and CaCO^
is present as solid above a pH of 6.4 and abeent below this pH. The pH
of 6.4 corresponds to the point at which the partial pressure of CC>2
reaches P__ and is referred to as the invariant point. At pH's below
MAY
the invariant pH the partial pressure of CO- is maintained at P in
accordance with the (X^ evolution rules established above.
At the start of the titration only CaCO^ is present in the slurry
so the initial pH cannot be found on Figure VI-2; however, this pH does
not have to be known precisely since in all our experiments enough SC^
is added to the slurry so that the final chemical composition of the
slurry is far removed from this point. As S02 is added to the slurry the
pH will fall and at the pH of 7.6 the calcium in solution will equal the
total carbon in solution. It is at this point that CaSOs begins to pre-
cipitate from solution. This point also is the end of Region I and the
beginning of Region II as shown in the titration graph in Figure VI-3.
In Region I the amount of S02 absorbed per liter of make-up is equal
to the liquid phase concentration of total sulfur, [S]. Thus over this
region the pH of the slurry as a function S02 absorbed per liter of slurry
can be followed easily using the equilibrium diagram (Figure VI-2).
Region II of the titration (as shown in Figure VI-3) extends from
the point where CaS03 first precipitates out of solution to the invariant
point.
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60
n Max
Make-Up 0.5% CaCOo, 50°C, P = 0.2 atm
J C02
Region V
40
01
C
o
» 20
3
3
CJ
Cumulative Amount of
CaSOo Precipitated
Solid CaCOo Remaining
in Slurry
Cumulative Amount
of C02 Evolved from
Slurry
I I
10
20
30
40
50
60
70
Cumulative Amount of SC^ Added to the Make-Up
Slurry, (mgmol SC^/Nter slurry)
FIGURE VI-3 TITRATION DIAGRAM FOR THE CaCO3~SO2 SYSTEM
c
o
Q.
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Since the pH range of Region II lies above the invariant pH, C02 does
not evolve from the solution in this region as discussed previously. Thus
in Region II the liquid phase concentration of total carbon, [C], is an
accurate measure of the CaSO~ precipitated. The pH at any point in Region II
can be calculated from the total amount of SC^ added per liter of make-up
slurry using Equation VI-7. Thus the other concentrations necessary to
calculate the amount of CaC03 dissolved or CaCO^ precipitated as described
above can be obtained from the equilibrium diagram.
Region III in the titration is the invariant region. This region
starts at the point at which the partial pressure of CCL reaches PCQ .
It is characteristic of this region that all the solid CaC03 must finish
dissolving before the system leaves this pH and this region and the liquid
concentrations remain constant here. Thus by Equation VI-7 the moles of
SC>2 absorbed in this region must be equal to the moles of C02 evolved and
the moles of CaCCK dissolved must equal the moles of CaS03 precipitated
in the invariant region. Region III appears in Figure VI-3 as the region
of constant pH.
Region IV of the titration lies between the invariant pH and the pH
at which the minimum in the total calcium curve occurs. In this region
no solid CaCO_ is present and the amount of SO- added per liter of slurry
in this region equals the increase in the total sulfur, [S], in solution
plus the decrease in the total calcium [Ca] in solution.
Region V is the region which has a pH less than the pH at which the
minimum in the total calcium curve occurs. In this region solid CaS03
dissolves as more SC>2 is added to the slurry and the added SO, per liter
of slurry equals the increase in the total sulfur concentration minus the
increase in total calcium in solution.
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The overall titration diagram is shown, in Figure VI-3. Here the pH and
accumulative amounts of C02 evolved, CaC03 dissolved and CaS03 precipitated
are shown as a function of the accumulative amounts of SC>2 added per liter
of make-up slurry.
The most important curve on the titration diagrams for the purpose of
simulating the ADL scrubber system is the pH versus S02 absorbed curve.
This curve can be used in conjunction with the pH-SC^ absorption relation-
ship which characterizes the packed bed scrubber. This latter relationship
is discussed in Section B. The pH and 802 absorption at which both the
titration and scrubber characteristics are satisfied is the operating
point of the scrubber recycle system. This will be discussed in more
detail in Section C.
The calcium utilization for the scrubber-recycle system can be
calculated from the titration diagram as long as there is no accumulation
of CaCC>3 in the system. Suppose, for example, that the titration and
scrubber characteristics are both satisfied at 40 mgmol per liter slurry.
Since the initial amount of CaC03 was 50 mgmol per liter of make-up slurry,
the calcium utilization is 84%.
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B. Absorption Parameters
In the analysis and modeling of scrubber systems probably the most
important quantities which must be known are the physical mass transfer
coefficients for the scrubber.
Correlations for both the liquid and gas side physical mass transfer
coefficients in packed beds are plentiful in the literature ' ' and
can predict these quantities fairly accurately if wall and end effects are
negligible. For small diameter packed beds such as the one used here,
where wall and end effects are apparently significant, corrections must
be made in these mass transfer coefficient correlations.
Although the mechanism of SC>2 absorption into limestone slurry is
not fully understood and the number of reactions which occur upon 862
absorption are too numerous to develop a complete analytical treatment of
S07 absorption into a limestone slurry; SC>2 absorption can, for the present
time, be handled empirically through an enhancement factor.
In this section the correlations for the mass transfer coefficients
for physical absorption and enhancement factor for the A. D. Little scrubber
will be discussed.
Mass Transfer Coefficients for Physical Absorption
(2)
S02 absorption experiments using NaOH solution as the scrubbing
liquor has been carried out in the A. D. Little packed column. At the
values of pH (8.92 in equilibrium with the scrubbing liquor is nil;
and therefore, the transfer of S02 to the NaOH solution can be considered,
for all practical purposes, gas film controlled.
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Under the assumption of gas film controlled absorption of 862, the
governing equation for the transfer of SO- in the scrubber is
a d
PT dz -8-so2
A 2
Where G is the molar gas flow rate, (gmole/cm sec),
Pqn is the partial pressure of S02 in the bulk gas phase, (atm),
PT is the total pressure, (atm),
z is the height of the column measured from the bottom, (cm),
k is the gas side mass transfer coefficient for physical
absorption, (gmole/cm atm sec),
and a is the specific interfacial area available to mass transfer,
(cm'1) .
Integration of Equation VI-8 over the height of the packed column yields
p in
k a - ^ In S°2 (VI-9)
8 T out
PS°2
Where Z is the total height of the packed column, (cm),
and PCA , PCA are the inlet and outlet partial pressure of S09, (atm)
S02 S02 2.
The gas side mass transfer coefficient for physical absorption, k a,
O
for ADL's packed column can be calculated using Equation VI-9 and the
(2)
experimental data for S02 scrubbing with NaOH solutions . The calculated
coefficients are shown in Table VI-1.
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TABLE VI-1
Run //
3
4
5
Gas Side Mass
for Physical Absorption
^ gmole
G 2
cm sec
1.98 x 10"3
1.92 x 10~3
2.60 x 10"3
Transfer Coefficients
for A.D. Little's
L g
L 2
cm sec
0.312
0.203
0.203
Packed Column
gmole
8 3
e cm atm.sec
2.39 x 10~4
2.07 x 10~4
1.96 x 10~4
With reference to Table VI-1, Run.Nos. 1, 2 and 6 have been excluded.
For Run Nos. 1 and 2, the percent of S0_ removal was reported to be
greater than 97.5. The exact percent removal must be known at these high
removal percentages because the calculated k a is extremely sensitive to
O
the S02 removal in this range. Run No. 6 was discarded because it appeared
inconsistant with the remainder of the runs.
Not enough S02~NaOH scrubbing data have been analyzed to establish
the functional dependence of the gas side mass transfer coefficient on flow
rates. However, if the functional dependence of k a on the gas and liquid
o
flow rates, which is reported in the literature , is assumed to be valid
for the k a in A. D. Little's packed column, then a relationship between
O
the gas side mass transfer coefficient for physical absorption and the
liquid and gas flow rates can be established.
The correlation for the gas side mass transfer coefficients for a
column packed with 1/2" Berl saddles operating under gas and liquid flow
rate similar to the flow rates used in the experiments is given by:
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ka=a£°-24 G°'7
g
Where a = 0.00792
If the same functional form of k a is assumed for the A. D. Little
g
data shown in Table VI-1, the value of a which minimizes the sum of the
squares of the difference between the calculated and experimental gas side
mass transfer coefficients for the packed column is
aADL = °-°222
The correlation of the gas side mass transfer coefficient for physical
absorption is then
*0 ")L *ft 7
k a = 0.0222 LU'^ GU
-------
Sherwood's correlation for 1/2" Berl saddles is
°-72
Where k. is the liquid side mass transfer coefficient for physical
absorption, (cm/sec)
and B = 0.0270
If the constant g is affected by the interfacial surface area in similar
manner as the value of a in the k a correlations then the correlations
8
for k^a (liquid side mass transfer coefficient for physical absorption) in
the A. D. Little packed column is
v a = * (S) £°'72 = 0.0758 £°'72 (71-13)
L d
Mass Transfer with Chemical Reaction in the Liquid Phase
The molar flux of 80^ across the gas-liquid interface in a wet scrubber
can be written in terms of the gas or liquid side resistance as
N_. - k (P_. - P* ) - k. $ (CA ~ CA> (VI-14)
S0« g SOo ^Oo LI AJ A
Where NSQ is moles of SOo absorbed per unit time per unit interfacial
2 area, (mole/cnrsec),
fcn is the partial pressure of S09 at the gas-liquid interface,
Ov/O t \ ^
2 (atm),
<|> is the enhancement factor for mass transfer in the liquid
film due to chemical reaction, (dimensionless),
C. is the 112803 concentration at the gas-liquid interface,
•*• (gmole/cm^),
o
and C. is the 112803 concentration in the bulk liquid phase, (gmol/cm ).
The enhancement factor, $, takes into account the reaction of the
diffusing H-S03 with components found in the liquid phase.
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The concentration of ^803 at the gas-liquid interface, CA , can be
related to the partial pressure of 802 at the interface, Pgi , by Henry's
law
pi = HC. (VI-15)
SO 2 A£
Where H is the Henry's law constant, (atm cnr/gmol).
This equation holds for sufficiently dilute solutions.
The Henry's law constant, H, as a function of temperature has been
given by Vivian as
InH = 17.360 - g'* (VI-16)
Where T is the liquid temperature in degrees Kelvin.
An expression for the interfacial concentration of ^803, CA , can be
obtained by substituting Equation VI-15 into VI-14.
H 2 (VI-17)
ic
8 H
Where Pg* is the partial pressure of S02 which could be maintained in
2 equilibrium with the bulk liquid phase, (atm).
Substitution of this equation in Equation VI-14 gives
'SO Ik ' k.4.1 (PSO,
*• L_ 8 " J *•
The rate of S02 absorption in a differential height of the scrubber,
dz, can be written as
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A dpsn i
_G S02 = f 1 H "I1 ( . (VI-19)
PT dz Ik a (jikjal k S02 S02
In Equation VI-19 both and P * are functions of position in the
bU2
column and both depend on the mechanism of SO- adsorption into recycle
limestone slurries. For convenience, the enhancement actor can be defined
such that Equation VI-19 can be written with P * . Thus
- dpso
G 2
Pso2
It must be noted that this equation is approximately correct and no
assumption about >PS02 and PS02 can be iS11016*1 relative to Pg() , the partial pressure
of S02 in the bulk gas phase.
Equation VI-20 can be integrated over the height of the packed column
to give
PI H T*
Where K^a = 1 -r=— + -£— I and is called the overall gas side mass
I k 3 y*^r ^ 1
L 2 LI _I
•* o ^*
G |^kga ^aj transfer coefficient, (gmol/cm3atm.sec).
Applying the mean value theorem to Equation VI-21 yields
~in
2
Where | and K a are the "average" enhancement factor and the "average"
overall mass transfer coefficient, respectfully.
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The "average" enhancement factor, $ was calculated from Equation VI-22
F21
and our experiment data . As can be seen in Figure (VI-4), the enhancement
factor is a strong function of pH. The enhancement factor was found to be
nearly independent of the liquid recycle rate and temperature.
Large values of the enhancement factor correspond to the gas film
being the dominating resistance to the transfer of SO- and similarly small
values of the enhancement factor indicate the liquid film resistance to be
important.
If Equation VI-9 were used to calculate k a for the CaCO, data, i.e.,
assuming zero ihterfacial partial pressure of SO-, the calculated k a would
be a strong function of pH because the liquid film resistances had been
ignored. However, the true gas film mass transfer coefficient must be
greater than the k a value calculated from Equation (VI-9) using the CaCO,
data. The maximum k a calculated using Equation (VI-9) is 37.4 and occurs
O
at pH of about 6.75. The k a predicted by Equation (VI-11) for the same
O
flow rates of gas and liquid is 37.8. This agreement is very good and
tends to validate the assumptions made in obtaining the correlation for the
gas side mass transfer coefficient (Equation (VI-11)) from the NaOH scrubbing
data for the packed bed unit.
The correlations for the gas and liquid film mass transfer coefficients
for physical absorption in the packed bed given by Equations (VI-11) and
(VI-13), respectively, along with the empirical correlation for the enhance-
ment factor, <|>, given in Figure VI-4 can be used for predicting SO. removals
in the A. D. Little scrubber. However, the value of the enhancement factor,
should only be used within the limits of the data from which it was obtained
(Table VI-2).
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100
50
20
Versus pH Used in Simulation
of Experiments
-r 7.53
6 =
v 6.85-pH
6.0 < pH < 6.75
0 =Exp [-3.054 + 0.869 pH]
4.5 < pH < 6.0
pH of the Holding Tank
FIGURE VI-4 THE "AVERAGE" ENHANCEMENT FACTOR FOR MASS TRANSFER
IN THE LIQUID FILM IN PACKED BED AS A FUNCTION OF THE pH
OF THE HOLDING TANK IN SCRUBBER RECYCLE SYSTEM
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oo
D
cr
r-r
r^
{L
R
Make-up Rate
(gpm)
Recycle Rate
Gas Flow Rate
(scfa)
% 02 in Gas
Inlet SC>2
(ppm)
Liquid
Temperature
Stoichiometry
TABLE VI-2
Range of-Data Used in the Construction of Figure VI-4
Scrubbing Slurry
1.3 to 5.2
0.25, 0.50, 1.0
0%
2000
100, 125
0.4 to 2.8
0.5
-------
A comparison between the calculated and experimental SO. removals is shown
in Figure VI-5. Here the percent of SO- removal was calculated using Equation
(VI-22) and the value of the enhancement factor given by the solid line in
Figure VI-4. The experimentally observed pH was used to evaluate the enhance-
ment factor in Figure VI-4. It can be seen from Figure VI-5 that most of the
data for SO. removals for Shawnee limestone and CaCO_ are within the ten percent
error band; however, the CaSO- data for SO- removals are quite scattered and is
due to the scatter of the enhancement factor data and error in drawing the line
through the enhancement factor data in Figure VI-4.
C. Simulation of the A. D. Little Scrubber System
In Section A the titration diagram, which relates the amount of S0~
absorbed in the scrubber per volume of make-up slurry to the pH of the holding
tank, was described.
In Section B the relationship between the amount of S02 absorbed in the
scrubber, the gas and liquid flow rates through the scrubber, and the pH of
the holding tank was given by Equation (VI-22) as
pin
S0
- v (VI-22>
so2
The amount of SO- absorbed in the scrubber per liter of make-up slurry is
given by
M
S02'° . GS_ in out (VI_23)
M MPT ^S02 *S02} ^
2
Where S is the cross sectional area of the packed column, (cm )
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100
80
O
a)
E
o
0)
QC
CM
o
V)
0)
o
I.
V
O_
0)
«-*
JO
J
(3
60
40
20
Symbol Make-Up Flue Gas
O Limestone
• Limestone 0-
D CaSO3
• CaS03 02
0 CaC03
I I
20 40 60
Percent of SO-> Removal from the Flue Gas
80
100
FIGURE VI-5 COMPARISON OF THE PERCENT SO2 REMOVAL FROM THE FLUE GAS WITH
THE PREDICTED PERCENT SO2 REMOVAL CALCULATED USING THE OBSER
pH OF THE HOLDING TANK TO EVALUATE THE ENHANCEMENT FACTOR
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Substitution of Equation VI-22 into this equation gives
MSO ,0 in
m-24)
Once the flow rates, total pressure and the inlet partial pressure of 802
are fixed, the right hand side of Equation VI-24 becomes only a function
of the pH of the holding tank. Equation VI-24 can be thought of as an
equation describing the scrubber characteristics in terms of the pH of
the liquor and the amount of SO 2 absorbed in the packed bed.
Two relationships between the amount of S0» absorbed in the column
per volume of make-up slurry and the pH of the holding tank have been
established. One is given by the titration diagram as shown in Figure
VI-3 and the other is the scrubber characteristic which is given by Equation
VI-24. At an operating point of the scrubber-recycle system both of these
relationships must be satisfied.
MAY
In Figure VI-6 the titration diagram for P = 0.2 and 1.0 atm. ,
C°2
and the scrubber characteristic for several make-up rates and a recycle
rate of 0.5 gal/min. have been superimposed. The intersection of the
titration line and the scrubber characteristic can be interpreted as an
operating point of the scrubber-recycle system.
The procedure described above for finding the operating point of the
scrubber-reyclce system has been used to simulate various experiments
carried out by A. D. Little. In these . simulations the maximum partial
pressure of CO- was taken to be 0.2 atm. This value of the maximum
partial pressure of C02 gave the best agreement between the observed and
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8 -
_
V)
0)
o
I
7_
6 _
5 _
Titration Curve; 50°C, Make-Up 0.5% CaCOg
Scrubber Characteristic; 50°C, Pm = 2000 ppm
SOn
Recycle Rate 0.5 gpm. Gas Flow Rate 5 scfm
M is the Make-Up Rate
M = 5.0 gph
M = 3.9
t
— Intersection is the
Operating Point
M = 1.3
20
40
60
80
mgmol SC^ Added per Liter of Slurry
FIGURE VI-6 GRAPHICAL DETERMINATION OF THE OPERATING
POINT OF THE SCRUBBER-RECYCLE SYSTEM
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calculated pH of the holding tank. This agreement is shown in Figure VI-7
and it can be seen that in the majority of cases the observed and calculated
pH's are within one-half of a pH unit of each other.
The choice of the maximum partial pressure of CO- equal to 0.2 results
in the calculated percent S02 removal being slightly less than the observed
percent S02 removal. However, as can be seen in Figure VI-8, the agreement
between these two quantities is still fairly good. The agreement could
be made better by choosing a slightly higher value for the maximum partial
pressure of CO-; however, this would increase the error between the observed
and calculated pH of the holding tank. The fact that the C02 partial
pressure in the holding could vary sometimes considerably, P _ =0.2 atm.
CAJrt
on the average, seems reasonable.
Also, agreement could possibly be improved between the observed and
calculated pH of the holding tank and percent SO- removal if a more accurate
value of total calcium concentration (liquid plus solid) in the make-up
slurry had been used in the simulations. For example, in Run 22 it was
reported that the make-up consisted of a 0.5 percent (by weight) CaCO^ in
the slurry and this was the value used in the simulations. However, chemical
analysis of the make-up slurry revealed that the total calcium concentration
of 50 mgmol/liter. It is not known why this anomaly occured; however, the
total calcium concentration in the make-up feed is very important in
constructing the titration curve and should be known accurately.
Sulfite oxidation in the scrubber-recycle system does not appear to
pose any particular problem in simulating the system by the method described
above. Figure VI-9 shows the pH titration curves for a make-up slurry of
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o
c
OJ
c
jo
o
0)
.c
a
3
(3
12]
Make-Up
CaC03
CaC03
CaC03
CaC03
Limestone
Flue Gas
S02, N2
S02, N2
SO2, N2
S02, N2
S02, N2
I
Limestone S02, N2> 02
I ...
5 6
pH of the Holding Tank as Measured by A.D. Little
FIGURE VI-7 COMPARISON OF THE CALCULATED AND THE MEASURED HOLDING
TANK pH FOR THE A.D. LITTLE SCRUBBER-RECYCLE SYSTEM
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0)
D
O
re
S
cT
CO
•*-•
§
£
<3
Limestone S02, N
Limestone
60 80
Percent of S02 Removal from the Flue Gas
FIGURE VI-8 COMPARISON OF THE CALCULATED AND EXPERIMENTALLY OBSERVED
PERCENT S02 REMOVAL IN EXPERIMENTAL SCRUBBER-RECYCLE SYSTEM
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0.6% CaCO- with, in one case, the make-up being saturated with respect to
CaSO^ and, in the other, no sulfate ions being present. For both these
curves it is assumed that none of the sulfur dioxide added to the slurry
is oxidized. Also shown in Figure VI-9 is the pH titration curve for 20%
oxidation of sulfite to sulfate. For this curve the initial slurry was
assumed to contain only CaCO-j. By 20% oxidation, it is meant that 20% of
the SO- goes into the slurry as sulfate ions and the remainder is added as
sulfite.
It can be seen from Figure VI-9 that there is little difference between
the two oxidation curves and that the 20% oxidation curve lies between
them throughout the invariant region. After the invariant region the oxida-
tion curves decrease in pH faster than the no oxidation curves. In general,
the rate of decrease in the pH titration curve after the invariant region
is greater for larger percent of oxidations. The pH titration curves
corresponding to no sulfate-no oxidation and sulfate saturation-no oxidation
bound the pH titration curves for the case of CaCOg make-up with oxidation
of sulfite (at least through the invariant region). Hence, these curves
can be used to estimate an upper and lower bounds on the pH of the holding
tank and the SC^ removal efficiency in the case of CaCO-j make-up with
stoichiometry greater than one. Since these bounds are close to each other,
it appears that it is not necessary to know the exact percent oxidation
of sulfite to sulfate in order to calculate the pH of the holding tank and
S0? removal efficiency of the scrubber-recycle system.
After the invariant region the sulfate saturated-no oxygen curve can
be used to give an upper found on the pH and 862 removal efficiency.
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9 r
to
O
I
- Titration Curve; 50°C, Make-Up 0.6% Shawnee Limestone (assumed 100% CaCOg),
pMax = o.2 atm; No Oxidation
^\Jn
Scrubber Characteristic; 50°C, P'" = 2000 ppm
Gas Flow Rate 5.2 scfm 2
7!7~.~ 777 Titration Curve; 50°C, Make-Up 0.6% CaCOg, PMax = 0.2 atm; Oxidation
20% Oxidation, CO2
80% Oxidation Sulfite to Sulfate /
/ Recycle Rate 0.25 gpm - Make-Up Rate
/ 5.3 gph
/ /Recycle Rate 0.5 gpm -
No Oxidation / / Make-Up Rate 5.3 gph
Saturated with Respect'
CaSO,,
No Oxidation
Recycle Rate 0.25 gpm
Make-Up Rate 2.9 gpm
Recycle Rate 0.5 gpm''
Make-Up Rate 2.9 gph
I L
10
20 30 40 50
mgmol S02 Added per Liters of Slurry
60
70
FIGURE VI-9 GRAPHICAL DETERMINATION OF THE OPERATING POINT OF THE
SCRUBBER-RECYCLE SYSTEM WITH A COMPARISON OF THE EFFECTS
OF EITHER HAVING NO SULFATE OR SULFATE SATURATED SLURRIES
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2
Table VI-3 shows the results of simulating Run 14 using the methods
described above.
TABLE VI-3
Simulation of Run 14
Inlet S02 2000 ppm, Gas Rate 5.2 scfm (4% 02),
Liquid Temperature 125°F, Make-up 0.6% Shawnee Limestone
Calculated
% S02
Calculated pH
Run
No.
14-1 '
2
3
4
5
6
7
8
9
10
Observed
PH
5.93
5.45
5.3
5.6
5.9
5.9
5.6
5.4
5.7
6.1
no_
so4
6.5
5.9
6.35
6.5
6.5
6.5
5.9
6.35
6.5
6.5
sat .
so^
6.2
5.8
6.2
6.2
6.2
6.2
5.8
6.2
6.2
6.2
Observed
% S02
Removal
81
70
64
74
86
85
74
63
73
87
Removal
no_
S04
92
83
82
84
92
84
83
82
84
92
sat.
504
89
83
79
79
89
79
83
79
79
89
In this table the pH.of the holding tank and S02 removal efficiency was
calculated using both the no sulfate-no oxygen and saturated sulfate-no
oxygen titration curves shown in Figure VI-9. It can be seen from Table
VI-3 that the choice of titration curves makes little difference in the
calculated pH or SO,, removal efficiency. However, the titration curve for
the saturated sulfate-no oxygen case appears to give better agreement in
the calculated and the observed pH's and removal efficiencies.
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A comparison between the calculated and observed pH of the holding
tank for Run 14 is shown in Figure VI-7 and the comparison of the removal
efficiencies is shown in Figure VI-8. While the agreement is not spectacu-
lar, it is still good in view of the simplicity of the calculation. It
should be noted that the deviation between the observed and calculated
S02 removal efficiencies is due to the error in predicting the pH of the
holding tank. It has already been demonstrated in Figure VI-5 that the
scrubber model can predict SOg removals with fair precision if the pH of
the holding tank is accurately known.
Possible sources of variation between the observed and calculated pH
values for Run 14 in Figure VI-7 include:
(1) The assumption that Shawnee limestone is 100% CaCO~.
The amount of CaCO-j in the make-up slurry has important
consequences in the construction of the titration
diagram as discussed previously.
(2) The presence of other elements such as magnesium.
The magnesium level in the solution, because of the
high solubility of the magnesium salts, will greatly
affect the equilibrium compositions which are
important in constructing the titration diagram.
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REFERENCES
1. Bennett, C. 0. and Myers, J. E. Momentum. Heat and Mass Transfer.
McGraw-Hill Book Company, Inc., New York, N.Y. (1962).
2. Berkowitz, J. B., Ketteringham, J. M., and Shooter, D. "Evaluation
of Problems Related to Scaling in Limestone Wet Scrubbing," r.eport
prepared for the EPA (April, 1973).
3. Borgwardt, R. H., EPA Progress Report No. 9 (April, 1973).
4. Leva, M., Tower Packings and Packed Tower Design, The United States
Stoneware Company, Akron, Ohio, 2nd Edition (1953).
5. McCabe, W. L. and Smith, J. C. Unit Operations of Chemical Engineering, '
1 McGraw-Hill Book Company, Inc., New York, N.Y., 2nd Edition (1967).
6. Treybal, R. E. Mass-Transfer Operations. McGraw-Hill Book Company, Inc.,
New York, N.Y., 2nd Edition (1968).
7. Vivian, J. E.,, The Absorption of S02 into Lime Slurries: An Investi-
gation of Absorption Rates and Kinetics, report prepared for the HEW
Department (September 1973).
8. Wen, C. Y. and Uchida, S. Absorption of SO? by Alkaline Solutions in
Venturi Scrubber Systems. report prepared for the EPA (July 1973).
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TECHNICAL REPORT DATA
(Please read Instructions on the reverse before completing)
1. REPORT NO.
EPA-650/2-75-031
2.
3. RECIPIENT'S ACCESSION-NO.
4. TITLE AND SUBTITLE
Scale Control in Limestone Wet Scrubbing Systems
5. REPORT DATE
April 1975
6. PERFORMING ORGANIZATION CODE
. B. Berkowitz. u. onooter. d. M.
L.N.Davidson, K.M.Wiig (A.D.Little Inc.); C.Y.Wen,
W.J.McMichael, R.D.Nelsen Jr. (U. of W. Va.)
8. PERFORMING ORGANIZATION REPORT NO
C-75092
9. PERFORMING ORG tNIZATION NAME AND ADDRESS
Arthur D. Little, Inc.
Cambridge, Massachusetts 02140
10. PROGRAM ELEMENT NO.
1AB013; ROAP 21ACY-038
11. CONTRACT/GRANT NO.
68-02-1013
12. SPONSORING AGENCY NAME AND ADDRESS
EPA, Office of Research and Development
NERC-RTP, Control Systems Laboratory
Research Triangle Park, NC 27711
13. TYPE OF REPORT AND PERIOD COVERED
Final: 12/72-11/73
14. SPONSORING AGENCY CODE
15. SUPPLEMENTARY NOTES
16. ABSTRACTTne repOrfgiVes results of tests of a number of phosphate and polymeric
additives—which have proven effective in controlling scale in some commercially
encountered calcium-containing systems—for scale control potential in limestone wet
scrubbers. Additives selected were Lime Treet; Acrysol A-l, A-3, A-5; Calnox 214
DN; Darex 40; Dequest 2000; PD-8; sodium hexametaphosphate; Calgon CL-14; sodium
pyrophosphate; Versenex 80; Quadrol; and sodium tripolyphosphate. Calnox 214 DN
and Calgon CL-14 were found to be particularly effective in controlling sulfate scaling
in the bench scale scrubber used for testing: both reduced the rate of scaling by 75%
under conditions previously shown to lead to catastrophic sulfate scaling. The kinetics
of oxidation of calcium sulfite in calcium carbonate/sulfite slurries was studied arid
compared with the oxidation of sodium sulfite solutions. Rates of oxidation in the cal-
cium system, found to be proportional to bisulfite ion concentration, increased in the
presence of solid calcium sulfite. Therefore the rate of sulfite dissolution is a con-
tributing factor to the oxidation under normal operating conditions. Cationic impur-
ities, such as sodium or magnesium, which can increase bisulfite concentration in
solution in the 5-6 pH range, are expected to accelerate oxidation.
KEY WORDS AND DOCUMENT ANALYSIS
DESCRIPTORS
b.lDENTIFIERS/OPEN ENDED TERMS
c. COSATI Field/Group
Air Polluation
Scrubbers
Limestone
Additives
Scale (Corrosion)
Fouling
Oxidation
Gypsum
Air Pollution Control
Stationary Sources
Phosphates
13B
07A
11G
07B, 07C
8. DISTRIBUTION STATEMENT
Unlimited
19. SECURITY CLASS (ThisReport)
Unclassified
21. NO. OF PAGES
97
20. SECURITY CLASS (This page)
Unclassified
22. PRICE
EPA Form 2220-1 (9-73)
91
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