Section 11
21 July 1978
Section Index (i
Copy No.
11
SULPHUR RETENTION
General Considerations
11.1
11.1.1
The behaviour of sulphur in combustion
11.1.2
The retention of sulphur
11.1.3
The chemistry of sulphur retention during fluidised
combustion
Effects of Operating Parameters on Sorbent Performance
11.2
11.2.1
Calcium:sulphur mol ratio
11.2.2
Bed depth and fluidising velocity
11.2.3
Sorbent particle size
11.2.4
Temperature, pressure and excess air
11.2.5
Sorbent type
11.2.6
Effects of recycle
Design Method
11.3
11.3.1
11.3.1.1
11.3.1.2
11.3.1.3
11.3.1.4
-
Calculation of sulphur retention required
Calculation of mol fraction of sulphur oxides
Calculation of weight of sulphur oxides per unit
volume of gas
Calculation of weight of sulphur oxides per unit
heat input
Calculation of retention required
11.3.2
Choice of sorbent
11.3.3
11.3.3.1
11.3.3.2
11.3.3.3
Calculation of sorbent requirements
Outline of procedure
Stepwise calculation
Maintenance of bed depth
11.3.4
/
11.3.5
11.3.5.1
11.3.5.2
11.4
Reliability of the design method
Compositions and flows of spent sorbent streams
Outline of approach
Calculation procedure
Thermal Considerations of Sulphur Retention
11.4.1
Thermochemistry of sorbent reactions
11.4.2
Sensible heat changes
11.4.3
Calculation procedure
11.5
Sorbent Reactivity
11.5.1
Variation with composition
11.5.2
Determination of Reactivity Index
,
Section 11
21 July 1978
Section Index (ii)
Copy No.
11.5.3
Sorbent characterisation test
11.5.4
Improving sorbent reactivity
11.5.5
Regeneration of spent sorbent
11.6
Sulphur Retention During Unsteady State Operation
11.6.1
Introduction
11.6.2
The kinetics of sulphur dioxide absorption
11.6.3
Calculation procedure
11.6.4
Considerations during start-up
11.6.5
Recommended operating plan
11.7
Sulphur Trioxide and the Acid Dewpoint
11.7.1
Introduction
11.7.2,
Calculation procedure
11.8
References
f,
21 July 1978
Section 11
Page 1 of 60
Copy No.
11
11.1
SULPHUR RETENTION
General Considerations
11.1.1
The behaviour of sulphur in combustion
Coal and oil contain sulphur and when these fuels are burnt the
sulphur is released as sulphur dioxide and sulphur trioxide.
are air pollutants.
Both of these
Sulphur trioxide can also react with water to form
sulphuric acid which can condense from the combustion gases at relatively
high temperatures, thus:creating corrosion hazards in the cooler regions of
the plant.
The ash of coal often contains free calcium oxide, and this can
"fix" some of the sulphur as calcium sulphate during the combustion process.
However, the quantity of sulphur retained by the ash is usually only a small
proportion of the total, and there are very few high-sulphur coals that contain
sufficient free lime in the ash to reduce the sulphur oxides emission to
acceptable levels.
The ash content of oil fuels is usually very low and also
has a neglegible calcium content so no material "fixing" of sulphur by ash
is possible during the combustion of oil.
One of the principal advantages of
fluidised combustion, however, is that irrespective of the fuel it is possible
to feed additives such as limestone or dolomite to the combustor where they
act as sorbents and augment the "fixing" of sulphur.
Sulphur is present in coal in two distinct forms:
as organic
sulphur compounds forming part of the coal substance, and as inorganic
sulphur compounds (mainly pyrites) forming part of the mineral matter
associated with the coal.
It has been found (11.1) during both devolatilisation
in a reducing atmosphere and combustion (in a fluidised bed) in an oxidising
atmosphere, that the pyritic sulphur is released more rapidly than the organic
sulphur.
Hence under conditions leading to high unburnt carbon loss (see
Section 4) some of the organic sulphur in the coal can be retained as part of
the unburnt fuel loss.
It can be stated, therefore, that the carbon/sulphur
ratio will be higher in the unburnt fuel loss than in the parent coal, but it is
recommended, nevertheless, that in all calculations it should be assumed
Section 11
Page 2 of 60
21 July 1978
_
Copy No.
that none of the fuel sulphur remains unburnt.
Some (usually < 5%) of the
sulphur appears in the flue gas as sulphur trioxide. The formation of sulphur
trioixide is favoured by high sulphur dioxide and oxygen partial pressures. It
is therefore favoured by operation at high total pressure and high excess air
levels and is suppressed by the retention of sulphur.
The sulphur in oil fuels is mainly organically bound and is released
along with the volatile hydrocarbons during combustion. It is generally accepted
that atomic oxygen species play a major role in determining the rate of formation
of sulphur dioxide and its subsequent conversion to sulphur trioxide.
11.1.2
(9*The
The retention of sulphur
air pollution legislation of many countries imposes national
limits on the emission of sulphur oxides from combustion plant. Many states,
provinces, counties and cities throughout the world supplement their national
limits by local limits which are often expressed in a different way. Section 18
gives a summary of the principal regulations regarding air pollution.
Some limits are expressed as the maximum permissible volumetric
concentration of sulphur oxides. Others are given as the maximum weight of
equivalent sulphur dioxide, sulphur trioxide or sulphuric acid per unit volume
of flue gas. Some countries, including notably the United States, impose a
limit on the weight of sulphur or sulphur dioxide emitted per unit gross
Some standards relate either the weight or the volume
of sulphur oxides to the stack height that gives an acceptable ground level
concentration, whilst others impose a limit on the total weight of sulphur
oxides that can be emitted from any one source. Finally, a limit that already
potential heat input.
exists in some regions of the United States and is now proposed as a national
standard for new combustion sources, is that 9 0 %t of the sulphur content of
the fuel must be removed before, during or after combustion.
t The 90% sulphur reduction requirement remains linked to weight per unit
heat input standards. If 90% removal still gives more than 0.52 kg of SO2
emission per giga joule heat input (1.2 lb SO2 per million Btu) then the removal
must be increased to limit the emission to that level. Alternatively, for low
sulphur fuels the emission need not be reduced to less than 0.087 kg SO2 per GJ
40)
(0.2 lb 50211o6 Btu).
Section.11
Page 3 of 60
21 July 1978
Copy No.
I
{
_
The relevant definition of air pollution standards are needed
for the initial step in the design of a sulphur retention system which is
to calculate the percentage retention of sulphur that is required.
The
Sulphur retention can be achieved by a variety of methods.
sulphur may be removed from the fuel before combustion, e.g. hydrotreating
of oils, or after combustion, e.g. wet gas scrubbing, or during combustion
by employing fluidised combustion with the addition of sorbents for sulphur
retention.
The choice of method for any particular application will be
determined by economics and cannot be discussed further here.
In this
Section it is assumed that sulphur retention is carried out during fluidised
combustion by adding sorbents such as naturally occurring dolomite or limestone.
The mechanism of the method is discussed and a design procedure is suggested
for calculating the amount of sorbent to be fed.
11.1.3
The chemistry of sulphur retention during fluidised combustion
The retention of sulphur by limestone or dolomite in fluidised
beds occurs through the following calcium oxide sulphation reaction.
CaO
+
SO2
+
°2
=
...
CaSO4
...
...
...
...
11.1
The calcium oxide results from calcination of calcium carbonate, which is a
major constituent of limestone and dolomite:
CaCO 3
=
CaO
+
CO2
...
...
...
...
...
...
11.2
Calcination is a reversible reaction, governed by chemical
equilibrium, which will occur only if the temperature exceeds a value that
is related to the partial pressure of carbon dioxide in the gas.
The
kinetics of calcination are also dependent on the temperature and carbon
dioxide partial pressure, and are of great importance in determining the
behaviour of sorbents.
-The
chemistry is slightly more complex when dolomite is used.
The calcination can be regarded as taking place in the following steps:
Decomposition :
CaCO 3 .MgCo 3
Half-calcination :
(CaCO3 + MgCO3)
=
(CaCO 3 + MgCO 3)
=
...
...
(CaCO 3 + MgO) + C02...
...
11.3
...
11.4
Section 11
21 July 1978
Page 4 of 60
Copy No.
(CaO + MgO) + C02
Ca calcination :
(CaCO 3 + MgO)
Ca sulphation
(CaO + MgO) + S02 + I 02
=
=
...
...
(CaSO4 + MgO)
K()
11.5
11.6
Reactions 11.3 and 11.4 can occur at temperatures below those
normally used for fluidised combustion.
Reactions 11.5 and 11.6 are
essentially the same as reactions 11.2 and 11.1 respectively, but occur in
the presence of magnesia which does not participate in the reactions.
At temperatures below the calcination temperatures, calcium carbonate
in limestone or "half-calcined" dolomite can react with sulphur dioxide as
follows:
3
2
& 02
CaS
+02
...
...
Reaction 11.7 occurs readily with half-calcined dolomite because the
inert magnesia acts as a skeleton and allows the reacting gases to penetrate
the proous particle.
Uncalcined limestone, however, has a low porosity, and
calcium sulphate, which has a greater molecular volume than calcium carbonate,
can render the outer surface layer of the particle impervious.
therefore, a high conversion to sulphate is difficult to obtain.
With limestone,
For this
reason dolomite is particularly suitable for operation under pressurised
conditions that prevent the calcination of the calcium carbonate.
11.2
11.2.1
Effect of Operating Parameters on Sorbent Performance
Calcium : sulphur mol ratio
The calcium : sulphur mol ratio (Ca/S ratio, Y,) is the most
significant of the operating parameters from the viewpoint of sulphur retention.
As would be expected, the sulphur retention under steady-state conditions increases
with increase in Ca/S ratio.
The effect of a change in Ca/S ratio on the fraction
of sulphur retained, R, is proportional to the fraction of sulphur emitted and
therefore diminishes with increase in Y. This can be expressed as:
t
C
Symbols are defined in Section 1 of this Manual.
S
Section 11
21 Jluly 1978
Page 5 of 60
Copy No.
dR/dY
=
k
=
1
(1 - R)
...
...
V;
...
...
...
11.8
...
...
...
11.9
which integrates to
R
where k
-exp(-k
Y)
...
...
is the proportionality constant in equation 11.8.
The value of k
depends on the sorbent used and on the operating conditions.
11.2.2
Bed depth and fluidising velocity
It is found, that as would be expected, sulphur retention is improved
by increase in bed depth and decrease in fluidisation velocity.
The ratio of
these parameters has the dimensions of time, and can be defined as a superficial
gas residence time, given by,
t
=
(seconds)
L b/U
...
It is found that the value of k
...
...
...
...
is proportional to t .
...
11.10
It is
tempting to infer from this observation that the rate of absorption of sulphur
dioxide is in fact a function of gas residence time.
There is ample evidence,
however, to show that it is more related to the inventory of unsulphated calcium
in the bed, which is a function of weight of bed (and hence bed depth), total
calcium concentration (hence sorbent feed rate) and fraction of calcium sulphated
(which depends on the residence time of sorbent in the bed).
For a given level
of excess air and Ca/S ratio, the sorbent feed rate is proportional to the
fluidising velocity, and for a given sorbent inventory in the bed, the residence
time is inversely proportional to the sorbent feed rate.
High fluidising velocity
can also reduce sorbent residence time by giving high sorbent elutriation losses.
The significance of these factors is demonstrated by the slow response to a
sudden change in sorbent feed rate;
under some circumstances it can take several
hours for sulphur retention to stabilise at a new level.
11.2.3
Sorbent particle size
The value of k
sorbent particle size.
in equations 11.8 and 11.9 falls with increase in
This is because the rate of sulphation of sorbent is
directly related to the specific surface area of the particle which, for a
spherical particle, is inversely proportional to the diameter.
21 July 1978
Section 11
Page 6 of 60
Copy No.
(
>
It has been found possible to relate sorbent performance to the
maximum particle size, dM , of the sorbent material;
inversely proportional to d
the value of k
is
*. This is, of course, an empirical simplification
for a material that does not depart radically from a normal type of particle
size distribution. There are no experimental data available on the effect
of using a narrow particle size range on the value of k for a given value
of dMx.
Operating conditions should be selected such that the size grading
of the sorbent is sufficiently coarse to avoid excessive elutriation (which
would effectively reduce the value of k ) but not so coarse as to lead to
S
segregation.
The increase in the value of k
with decrease in dM
is illustrated
by the curve between points A.and B in Fig. 11.1, which shows data, reported
in reference (11.6), obtained in a series of tests with progressively diminishing
particle size of the sorbent (a low-reactivity limestone).
The tests were at
about 645 C (11550 F), with a bed depth of 0.25 m (10 inch) and a fluidising
velocity of 4 mls (13 ft/s). At B, the sorbent fed contained a high proportion
of material which, having a free-fall velocity less than the fluidising velocity,
could therefore be elutriated from the bed.
a short residence time in the bed.
This material consequently had only
At C all of the sorbent fed could be elutriated
from the bed.
Between B and C, therefore, the diminishing residence time of
sorbent resulted in a sharp fall in the value of k . Once point C was reached,
further reduction in the particle size of the sorbent fed had little effect on
the particle residence time, which was not much greater than that of the gas.
Consequently the benefits of increasing specific surface area with diminishing
particle size again manifested themselves, and the value of k
by the curve between C and D.
rose, as shown
It can be seen from Fig. 11.1 that in some circumstances there may be
an advantage in employing finely-ground sorbent to improve the sulphur-retention
capacity of a sorbent.
11.2.4
Temperature, pressure and excess air
The rate of reaction between calcium oxide and sulphur dioxide changes
only slightly with reaction temperature. The apparent activation energy for
the reaction, deduced from the results of experiments performed in the pilot
Section 11
21 July 1978
Page 7 of 60
Copy No.
dMx inch
0.05
0.01
0.005
0.5
DO
I
0.4
03
ks
0.2
B
·0i
30
50
100
500
1000
2000
30
50
100
500
1000
2000
dMx /um
Figure 11.1
Variation of ks with dMx
I!
21 July 1978
Section 11
Page 8 of 60
Copy No.
plants (11.1),
and (11.14), is only about 29 MJ/kg mole (12 600 Btu/lb mole or
7 k cal/g mole).
This value is close to published values for the reaction
(11.3, 11.4).
As well as altering the reaction rate, the combustor bed temperature,
in conjunction with the pressure and excess air level, causes other very
important effects which can be summarised briefly as follows:(a) At some temperature, which is usually up to about 5000 (100IF)
above calcination temperature, the value of k
maximum.
passes through a
This "optimum temperature" therefore increases with
pressure and falls with increase in excess air.
(k-(b)
The peak in the value of k
is usually more pronounced with
limestones than with dolomites.
(c) The fall in the value of k
with increase in temperature becomes
slightly less marked with increase in the maximum particle size of the
sorbent.
(d) Pressurised operation brings about a major improvement in the
value of k
for dolomites at all temperatures.
All these phenomena can be quantitatively explained by reference
to the calcination behaviour of sorbents.
To understand this, it is important
to appreciate that the rate at which a sorbent particle reacts is initially high,
but that as it becomes sulphated the reaction rate falls by several orders of
magnitude, so that under typical operating conditions in a fluidised combustor,
a particle residence time of several hours is needed to achieve a reasonably
high fractional sulphation.
Hence the sulphur retention obtainable with any
given set of operating conditions will be influenced adversely by factors tending
to curtail the residence time of the sorbent particle in the bed.
The two factors that control the particle residence time are the
rates of withdrawal of bed material and the elutriation of fines.
Fines are
present in the fresh sorbent feed, and are also generated by attrition and
decrepitation. Decrepitation probably occurs only during heating up and
calcination of the sorbent whereas attrition occurs during the whole of the
time a particle is in the bed.
Because decrepitation occurs to an extent that
depends on operating conditions, it brings about reductions in sulphur retentin-
Section 11
Page 9 of 60
21 July 1978
Copy No.
2
that are also related to operating conditions. Decrepitation can therefore
be a more significant factor than the fines content of the sorbent fed or its
rate of attrition.
The available data indicate that decrepitation is related to the
volumetric rate of evolution of carbon dioxide that occurs when sorbent
undergoes calcination in the bed. This is probably because the sorbent particles
are disrupted by the gas evolved from within the particle.
Studies of calcination kinetics (12.2) indicate that the rate of
calcination of calcium carbonate is inversely proportional to particle size
and is strongly influenced by temperature, with an activation energy of about
170 MJ/kg mole (40.6 k cal/g mole or 73 100 Btu/lb mole). It is also dependent
on the proximity to equilibrium, being zero when the partial pressure of carbon
dioxide exceeds the equilibrium value. The equilibrium partial pressure is also
temperature-dependent, with an activation energy of 159 MJ/kg mole (38 k cal/g
mole or 68 400 Btu/lb mole). Hence the rate of calcination is favoured by high
temperature, low pressure and by high excess air (which gives a-low mol fraction
of carbon dioxide in the gas).
Dolomite decomposes and "half-calcines" at all conditions prevailing
in fluidised combustors including those conditions that prohibit calcination
of calcium carbonate. The kinetics of the 'half-calcination' of magnesium
However, since operating conditions are so remote
from the relevant equilibrium, complete calcination of magnesium carbonate is
probably rapid, and the effective rate is independent of temperature and of
carbonate are not known.
particle size.
The volumetric rate of carbon dioxide evolution is proportional to
the rate of calcination and the absolute temperature, and.inversely proportional
to the absolute pressure. Conditions that increase the volumetric rate of gas
evolution increase the severity of decrepitation and reduce the sulphur retention.
This is illustrated by the four diagrams in Figure 11.2, which show the variation
of k
with bed temperature for a number of operating conditions.
In each diagram the broken line shows the value of k S that would
be obtained if no decrepitation occurred.
of calcium carbonate.
Circles show calcination temperatures
21 July 1978
Section.11
Page 10 of 60
Copy No.
The first diagram, Figure 11.2a, presents a curve for limestone,
at atmospheric pressure. Below the calcination temperature, the value of
k
is low because the limestone particle lacks porosity and high sulphation
is not possible. Close to the calcination temperature, the value of k rises
to about the value expected with no decrepitation.
At higher temperatures,
decrepitation results in a divergence between expected and observed values
of kI, with the observed value peaking at or just above the calcination
temperature.
At operating pressures above atmospheric, Figure 11.2b, the
calcination temperature is raised to an extent that depends on the excess
air level; at high excess air levels there is a lower partial pressure of
carbon dioxide and hence a lower calcination temperature (shown by the full
circle) than at low excess air levels (calcination temperature shown by the
open circle).
At any temperature above the calcination temperature decrepitation
has a smaller effect on k for pressurised operation mainly because the higher
pressure reduces the volume of gas evolved.
At atmospheric pressure, dolomite, Figure 11.2c, suffers severe
decrepitation at all likely combustor temperatures because of the halfcalcination of the magnesium carbonate. It is commonly observed that the
sorbent content of bed material when feeding dolomite is lower than when
feeding limestone. For this reason dolomite has a low apparent reactivity,
although its inherent reactivity,
that of the limestone.
revealed by the broken cutve,
is higher than
Since half-calcination develops adequate internal
porosity, there is no abrupt increase in the value of k at the calcination
temperature, and the peak in the curve, resulting from calcination of calcium
carbonate and consequent increased decrepitation, is therefore less pronounced
than with limestone.
Figure 11.2d illustrates how the performance of dolomite improves
during pressurised operation. The decrepitation from both half-calcination
and full calcination is reduced because of the effect of pressure on the volume
of gas evolved, and is is observed that the retention of dolomite by the beds of
pressurised combustors is high. On a molar basis most dolomites are superior to
limestone in pressurised combustors at all conditions unfavourable to calcination
Section 11
21 July 1978
Page 11 of 60
Copy No.
*iS~~~
~
Bed Temperature Tb OF
1400
Bed Temperature Tb OF
1600
1400
1600
I!
-
I- .
Excess Air:
ks
Low
High
A'
..
-~...
W~~~
(b) Limestone at high pressure
(a) Limestone at atmospheric pressure
950
850
750
1400
950
850
750
1600
1400
1800
Excess
(d) Dolomite at high pressure
(c) Dolomite at atmospheric pressure
750
850
Air:
-
A
950
750
BedTemperature Tb oC
850
Bed Temperature Tb °C
Figure 11.2
Effects of Temperature, Pressure and Excess Air ons k
pl
\
950
21 July 1978
Section 11
Page 12 of 60
Copy No.
' ,A-'
of calcium carbonates, but under `calcination" conditions this superiority can
be lost.
Other possible explanations for the temperature effects have been
proposed.
They include:
(a) The suggestion that water present on the surface of the sorbent
is necessary for the reactions between sulphur dioxide, oxygen
and calcium oxide (11.1).
Increase in temperature might result
in accelerated desorption of water, a process that would be
favoured by reduction in pressure.
(b) Sorbent particles circulate between oxidising and reducing
zones, resulting in decomposition of calcium sulphate by
sorbent regeneration reactions (Section 11.5.5) at a rate
that is dependent on temperature (11.20).
The equilibria
involved are favoured by decrease in pressure.
A similar
suggestion is made in Reference (11.11).
(c) "Ageing' or sintering of the calcium oxide crystallites after
calcination, leading to reduction in internal reaction surface
area (11.14).
Sintering is a process that is accelerated by
temperature, but since it occurs after calcination the onset
of the process is favoured by low pressure.
(d) The suggestion that the first step in the reaction sequence is
the oxidation of sulphur dioxide to sulphur trioxide, and that
it is the latter that reacts with calcium oxide (11.16).
The oxidation
reaction equilibrium is favoured by high pressure and low temperature,
whereas the reaction kinetics are favoured by high temperature and
high pressure.
Whilst some of the above mechanisms may contribute to the observed
effects of temperature pressure and excess air, it has been found that the
decrepitation mechanism can account for them quantitatively and it has therefore
been taken as the basis for the calculation procedure.
.21 Jul.y 1978
Section 11
Page 13 of 60
Copy No.
<'
Decrepitation would not be expected to be a factor when using
finely ground limestone, since this material is in any case completely
elutriated.
Data on the effect of temperature with finely ground limestone
is sparse. Data from two references (11.1), (11.5) each cover only
very
restricted ranges of data and show no temperature effect. Data from
a third
reference (11.6) covers the approximate range 8500 C to 1000C (15500
F to
1800 F) and shows that increase in temperature reduces the reactivity
of the
finely ground material, though to a smaller extent than with coarse
material.
The general operating conditions and combustor performance are not
given
precisely in these references, however, and the apparent falling
off in sulphur
retention efficiency may have a number of explanations. For example,
with low
excess air and consequent high temperatures the equilibrium becomes
favourable
for the regeneration reaction, i.e. unfavourable for sulphur retention.
11.2.5
Sorbent Type
Many of the differences between sorbent types have already been
discussed in Sections 11.1.3 and 11.2.4. It has been shown that after
correcting for the effects of decrepitation, the rate of reaction of
all
sorbents is increased by increase in temperature with an apparent
energy
of activation of about 29 MJ/kg mol, (12 600 Btu/lb mol), or (7 k cal/g
mol).
If the value of k
0
is corrected for the effect of decrepitation; to
the value that would be obtained with a superficial gas residence time
of 1
second; and with a maximum particle size of 1 mm (0.04 in.), it is found
that the resulting datum value, denoted ko, is given by the Arrhenius
equation:
ko
where
A
exp{
E
/T
...
...
...
...
...
As
has a value expressing the effective reactivity and is
called the sorbent Reactivity Index.
E
has the value of 3530, equivalent to an activation energy of
29 MJ/kg mol, (12 600 Btu/lb mol), (7k cal/g mole).
In equation 11.1 e equals 1 for Tb in
K or 1.8 for Tb in
R.
11.11
Section 11
21 July 1978
Page 14 of 60
Copy No. ,
The sorbent Reactivity Index, A , has a range of values between
about 8 for an uncalcined pure limestone to about 46 for pure dolomite.
value of A
The
is related to the porosity characteristics of the sorbent, and
probably also to other physical properties, e.g. rate of attrition in the
bed.
It is high when the sorbent has both a high porosity and a good balance
between large "transport pores` facilitating deep penetration of the particle
by reacting gases and small pores that endow the particle with a high internal
surface for reaction.
Most dolomites, dolomitic limestones and magnesian
limestonies have values of A
close to 46.
However, for limestones it is not
possible to give any rules for predicting the value of A .
of A
S
The highest value
= 42 has been obtained with a stone of low (about 82%) purity, and the
lowest value of A
= 23 with a pure stone (over 99% purity), but there is some
evidence indicating that some reactive stones have a high purity and some impure
stones have only a low reactivity.
This is to be expected, since there is a
wide range of types of limestone, and there is no obvious reason for the porosity
characteristics of a calcined stone to be related in any predictable way to the
purity.
of A
Until relevant test data are available, therefore, a value for limestone
= 30 is recommended for preliminary assessments.
When limestone is finely ground, there is much less variation in
the value of A .
This is because the finie grinding gives great accessibility
to the reactants and a high specific surface area for reaction without relying
on favourable porosity characteristics.
time reduces the effective value of A
S
However, the short particle residence
to about 8 for both the most reactive
and the least reactive of the limestones that have been tested in work described
in references (11.1),
11.2.6
(11.5)
and (11.6).
Effects of Recycle
Recycling of elutriated fines is sometimes carried out to improve
combustion efficiency, and when limestone or dolomite are being used to retain
sulphur, partly sulphated material becomes recycled and improves the<sulphur
retention.
There appears to be no improvement in sulphur retention from fines
recycle when the fresh sorbent is fed in a finely ground form.
21 July 1978
Section 11
-Page 15 of 60
Copy No.
11.3
Design Method
The rate of sorbent to be fed in any particular application depends on
the design and operating conditions and on the sorbent characteristics as well
as on the sulphur content of the fuel and the percentage sulphur retention needed.
Calculation of the sorbent rate involves the following steps.
1.
Calculate the sulphur retention required.
2.
Selection of a sorbent and determination of its reactivity index.
See Section 11.3.2.
3.
Calculation of the Ca/S molar ratio for the given sorbent and the
sorbent flow rate. See Section 11.3.3.
4.
Calculation of the compositions and flows of spent sorbent streams.
See Section 11.3.5.
5.
Calculation of the thermal effects of sulphur retention.
See Section 11.4.
6.
The sorbent rate found in step 3 above may affect the following
combustor parameters,
mean bed particle size
elutriation rate
combustion efficiency
'
See Section 11.3.1.
-
Section 9
-
Section 4, 5, 6 or 7 as
appropriate
J
heat transfer to in-bed surfaces,
(Group 2 Applications only) - Section 10.
Changes in the above parameters may necessitate changes in the bed depth,
fluidising velocity or sorbent particle size which, in turn, may alter the
sorbent flow.
A detailed calculation is therefore iterative and should repeat
from step 3 until the required degree of convergence is obtained, although for
preliminary estimates a single iteration may be sufficient.
The calculation procedure given in the following sections are, in
themselves, suitable for solution by a pocket calculator. However, the
iterative nature of a detailed calculation makes the use of a computer advisable.
A suitable program is available and is described in reference (11.7).
For any particular application it will probably be necessary to
calculate the required sorbent flow rate at several different combinations
of operating conditions which represent the maximum allowable variations
Section 11
21 July ]978
Page 16 of 60
[f >(
Copy No.
to be expected during both normal and part load operation.
See Section 3.6
for a general discussion on the requirements for turndown.
The effects of
unsteady state operation during start-up and load changing or fluctuations in
fuel flow must also be taken into account before a final choice of sorbent
flow rate is made.
11.3.1
(See Section 11.6).
Calculation of the sulphur retention required
A range of calculation procedures is required to conform with the
range of ways in which the limits are expressed, as discussed in Section 11.1.2.
In most cases it is necessary to calculate the emission of sulphur oxides that
would be obtained if all of the sulphur content of the fuel appeared in the
combustion gases, and then calculate the retention required.
Sometimes it
is necessary to check that a given percentage retention gives an emission
falling within prescribed limits while for some of the design calculations it
is necessary to know the.partial pressure of sulphur oxides, so methods for
estimating the mol fraction of sulphur oxides, will also be described.
11.3.1.1
Calculation of mol fraction of sulphur oxides
The procedure is as follows:-
1.
Calculate the moles of sulphur oxides produced by unit weight of
fuel, w.
W$ = S'/32.07
...
... ....
...
...
...
...
....
11.12
where S' is the weight fraction of sulphur in the fuel as fired.
2.
Calculate the stoichiometric moles of air, wa, required for combustion
of unit weight of fuel.
(a) If the composition of the fuel is known, then
wa
(C'/12.01 + H'14.03 + S'132.07 - 0'/32)/0.21
.
..
where C', H', S' and 0' are respectively the weight fractions of
kS
carbon, hydrogen, sulphur and oxygen in the fuel.
11.13
21 July 1978
Section 11
Page 17 of 60
Copy No.
(b) If only the gross calorific value of the fuel, Ea. is known,
calculate w
from:
wa
EH
=
...
...
...
...
where e = 1.11 x 10J5 kg moles/kJ when E
...
...
11.14
is in kJ/kg
2.59 x 105 lb moles/Btu when Eh is in Btu/lb.
0
3.
...
Calculate the moles of wet gas, w , and dry gas, Wd, from combustion
of unit weight of the fuel with X% excess air at a percentage combustion
efficiency, n, from:
w
=
(wa/100)(0.03
n
+ 100 + x)
Wd
=
(waI/OO)[ 100
+
X-
-
(-
)]
...
...
...
...
11.15
...
...
...
...
11.16
where the fuel type parameter, 4, is 0.96 for all solid fuels and 0.94
for all liquid fuels.
4.
Calculate the mole fraction of sulphur oxides in wet gas, m , and
in dry gas, md, when all of the fuel sulphur appears as sulphur
oxides in the gas.
m=
md =
11.3.1.2
w
...
w
s /wd
...
...
..
...
...
...
...
11.17
..
..-.
*..
...
11.18
Calculation of weight of sulphur oxide per unit volume of gas
The procedure involves converting the moles of sulphur compound
to the appropriate weight, and the moles of gas to a volume.
1.
Multiply w
by the appropriate molar volume to give V , the volume
of wet gas from combustion of unit weight of fuel.
2.
3
3
Calculate the weight of sulphur compounds (kg/m , g/l, lb/ft3)
that would be obtained in unit volume of gas if all of the fuel sulphur
appeared as that compound in the gas from the appropriate ratio below:
Section 11
21 July 1978
Page 18 of 60
Copy No.
32.07 w4/V
'7
for weight of sulphur
W
64.07 w S/V
for weight of sulphur dioxide
80.07 w /V
for weight of sulphur trioxide
96.09 w /V for weight of sulphuric acid
Calculation of weight of SO2 per unit heat input
11.3.1.3
The weight of sulphur dioxide per unit gross potential heat
input that would be obtained if all of the fuel sulphur appeared in the
gas as sulphur dioxide is 64.07 w /EH (or 2 S'/E).
11.3.1.4
1.
Calculation of retention required
Let e
0
be the sulphur oxide emission when all of the fuel sulphur
appears in the gas, expressed in the units required by the relevant
Thus
legislation.
e
-
m
mole fraction wet gas
e
=
md
mole fraction dry gas
e
=
'
64.07 m /V
e
=
64.07 w /EH =
weight SO2 per unit volume
2 S'/EH weight SO2 per unit heat input
If e is the permissible sulphur oxide emission in the same units
as e, then the total fractional sulphur retention needed, R, is
given by:
R = 1 - e/e
2.
0
....
...
...
...
...
...
...
...
11.19
Some legislation specifies the value of R and it is necessary to
check that this gives an emission within certain limits using the rearranged
form:
e = e
(1 - R)
...
...
...
...
...
11.20
21 July 1978
Section 11
Page 19 of 60
Copy No.
V%
3.
If the ash content of the fuel, and the ash analysis are
known, estimate the sulphur retention by ash, R, from:
P
R
c
4.
(0.4 x wt. fraction StO in ash) x Cwt.fraction ash in coal) ...
(wt.fraction S in coal)
11.21
Calculate, Ra' the fractional reduction in sulphur oxides
emission by sorbent, from:
R = 1
a
If R
(I R)
(1- R)
........
...
.
...
...
...
...
11.22
cannot be estimated assume that it is zero and R = R.
a
c
11.3.2
Choice of Sorbent
The choice of sorbent will usually be dominated by economic consideration:
and in particular by the cost of transport from the quarry to the plant. Transport
costs for limestone or dolomite can outweigh the ex-quarry cost of the stone by
a substantial factor except for quite modest distances, e.g. less than about
30 km (20 miles). It may often be more economic to employ an inferior but locallyavailable sorbent than a reactive stone from a more remote quarry. The economics
of crushing, grinding and other forms of pretreatment also have to be considered.
The designer therefore starts with a set of options regarding the
source and pretreatment of sorbent, and calculates for each option:
(a) the cost of supplying, preparing and feeding the sorbent to
the combustor.
(b) the cost of removing spent sorbent from the combustor and
exhaust gases and disposing of it in an environmentally
acceptable manner.
(c) the cost of handling any consequential effluents, such as
water for quenching spent sorbent, or the products of sulphur
recovery from regeneration processes.
The calculations have to be performed in detail.
For example, in
choosing the source of sorbent, the designer should not be misled by the
higher Reactivity Index of most dolomites compared with limestones. The low
calcium content of dolomite - 22% - puts it at a disadvantage compared with
limestone - calcium content 40%.
Although under "non-calcining" conditions
Section 11
21 JulY 1978
Page 20 of 60
Copy No. L 4>-
(e.g. pressurised combustors) dolomites are superior to limestones on a Ca/S
mole ratio basis, there is often little difference on a weight basis.
Under
"calcining" conditions (e.g. atmospheric pressure combustor) dolomites suffer
more severely than limestones from decrepitation, and are often inferior to
limestone even on a mole basis.
The next choice to be made is of the particle size distribution
of the sorbent.
The size distribution should be chosen to avoid the
possibility of segregation of the coarser fractions under any likely operating
conditions.
For this reason the maximum particle size is generally taken to
be not more than that of the coal to be fed for coal firing, or not more than
that of the inert bed material for the firing of other fuels.
For coal firing
there are some advantages in using a sorbent with a somewhat finer size
distribution than the coal;
although a low value of dM
accelerates
decrepitation this is more than offset by a higher specific surface.
Removal
of elutriable fines obviously reduces the amount of sorbent elutriated
slightly.
However, the size distribution of sorbent is also governed by
economics:
i.e. whether the costs of extra crushing, or of sieving and
rejection of fines, are justified by savings in sorbent fed or in removal of
fines from the exhaust gases.
Use of finely-ground sorbents can give an improvement in sulphur
retention efficiency by increasing the availability of particle surface.
The
greatest benefits would accrue with sorbents of low reactivity, or with
operation under "non-calcining" conditions, particularly with low fluidising
velocities.
The benefits increase with fineness of grinding, as do the
penalties, which are the high cost of fine-grinding, and the need for
efficient removal of fines from the exhaust gas.
Finally, the choice has to be made whether to use one of the
techniques described in Section 11.5.4 for improving sorbent reactivity,
or in Section 11.5.5 for regenerating spent sorbent.
The penalties of
these techniques have to be weighed against the savings in fresh sorbent
required.
Recycling elutriated fines also improves the utilisation of
sorbent, but this is a technique that would generally be recommended only
to obtain improved combustion efficiency;
from the sulphur retention
21 July 1978
Section 11
Page 21 of 60
Copy No.
ajX.'{
standpoint alone, the savings would generally not justify the added complexity,
thermal loading on the combustor bed, or increased particulates emission that
would result.
For preliminary estimates and as a starting value for more detailed
calculations it is recommended that a sorbent should be chosen with a top size
equal to that of the coal feed or inert bed material for coal or other fuel
firing respectively.
11.3.3
Calculation of sorbent requirements
11.3.3.1
Outline of procedure
The calculation procedure for a coarsely-crushed sorbent is to
calculate a datum value, ko, of the proportionality constant in equation
11.8 and then to correct k for the effects of decrepitation, sorbent
0
particle size, and superficial gas residence time to obtain a value of
kS.
For a finely ground sorbent k
is calculated directly.
The effects of
recycle are then incorporated and the resulting kS value is used to calculate
the Ca/S molar ratio and the sorbent flow.
11.3.3.2
Stepwise calculation
For a coarsely-crushed sorbent begin at step 1. For a finely
(
ground limestone sorbent begin at step 10 and ignore steps I - 9 and step 11.
1.
Calculate a datum value, k
in equation 11.8.
,
of the proportionality constant
This datum value is the value of k
the maximum sorbent particle size, dM
when
is 1000 pm (0.0394 inch),
when the superficial gas residence time, t , is I second,
when the effect of decrepitation is negligible, and when there
is no fines recycle.
It is obtained from equation 11.11 which
is repeated below for convenience.
kwhere 0
ASexp(-3530 e/T1)
1 for Tb in
bhb
...
K or 1.8 for Tb in
...
0R.
...
...
...
11.11
Section 11
Page 22 of 60
21 July 1978
Copy No.
<K
depends on the sorbent type, as explained in
The value of A
For dolomites, and for dolomitic and magnesian limestones,
a value of 46 can be taken. Most limestones will have values of A 5 in the
Section 11.2.5.
range 20 to 40 and unless test data are available, a value of 30 should be
Curves of k /A plotted against temperature
assumed for preliminary assessments.
are shown in Figure 11.3.
2.
Calculate the partial pressure of carbon dioxide, P , in the
wet gas from
Pc = P i/(203 + X - 100 Cb
where the fuel type parameter, i,
3.
.
=
...
...
...
...
11.23
17.60 for coal or 14.3 for oil.
Calculate the equilibrium carbon dioxide partial pressures,
P ,from
ce,
e exp(- 19 140 e
= 1.2 x 10
P
...
XTb)
...
...
11.24
2
where, 6= 1 when the units of P and Pce are kN/m and
0.01 when P and Pce are in atmospheres
0
e 1=
I for Tb in
Rt.
plotted against temperature are shown in Figure 11.4.
Curves of P
4.
K and 1.8 for Tb in
Check that calcium carbonate calcination is possible by
evaluating the ratio P /P
.
If this ratio is less than
unity calcination of calcium carbonate is possible.
If calcination is possible proceed to step 5;
otherwise go
to step 7.
5.
Calculate the decrepitation parameter, r, in the following way.
First calculate, r
,
the effective calcium carbonate calcination
rate at zero carbon dioxide partial pressure, from:
r
(*.~~~~~~~~~~~~xb
L
1.17
x 10
/Ce
d )j exp(-20 440 el/Tb)
...
.
...
11.25
Section 11
21 July 1978
Page 23 of 60
Coov No.
i
;'
0.06
0.05
ko/A s
0.04
0.06>
0.02
1000
S00
800
700
Bed Temperature Tb OC
0.06
___
_i--:X_
mg g S E S S X i ,
0.05
k0IA5
0.04
0.03
0.03
-
X
(British
gg; | Units)
t mX E
g
0.02
1200
1600
1400
Bed Temperature Tb OF
Figure 11.3
with Bed Temperature
Variation of ko/A
s
1800
2000
Section 11
21 July 1978
Page 24 of 60
CopV No. ¾iL
70
700
..
60
6004
O
kNlmZ
-3
kN/m,
d
30
300
20
200
10
:.
~~~~0
650
700
750
-800
__~-,-,--.~,,'
Xt WW X X gWW
850
900
Bed Temperature Tb OC
Figure 11.4 (SI Units)
Variation of P0c with Bed Temperature
950
gE~~~~~~
1000
1050
Section 11
21 July 1978
Page 25 of 60
Copry No.
g #
g } gW
Fg WW
0.9 g g W
i
g9
g M X~Ft
8
0.8 ~=T~;:·;F~;iF~:T~-T;Ff;FPsF~rRAr=~~Crt~
~:~t
ce
CO
PCO
~r-lt·~f~i·FhiW~i~t~~;
~~~0.7..7
0
043
0.2
T
S
S
-.-
t
f
At
4
1
0.1
0
0
1200
1400
1600
Bed Temperature Tb F
Figure 11.4 (British Units)
Variation of Pce with Bed Temperature
1800
2000
21 July 1.978
Section 11
Page 26 of 60
Copy No.
i)
1*
where e = 1 when dMx is in pm and 25 400 when dMx
is in inches,
and e0 = 1 when T is in K and 1.8 when Tb is in
R.
bb
A curve of (dM½r ) plotted against temperature is
shown in
Figure 11.5.
6.
Calculate, r , the effective calcium carbonate calcination
rate
at the prevailing carbon dioxide partial pressure
from:
c
=c
ro (1 _pc
~~P )
ce
........
.
.
. ...
...
11.26
Go to step 8.
7.
If calcination is not possible then r
8.
If the sorbent to be used is a dolomite or a dolomitic
or
magnesian limestone, calculate the effective magnesium
carbonate
calcination rate, r , (whether or not r is zero),
from
m
=
6.1
x 10j4 z
...
...
is, of course, zero.
...
...
...
...
11.27
where, Z, is a sorbent composition parameter representing
the mol
ratio of magnesium carbonate to calcium carbonate
in the sorbent.
It has a value of 1 for pure dolomites and can be
evaluated from the
composition of the sorbent as:
1.391 x weight ratio of MgO/CaO in stone
1.648 x weight ratio of Mg/Ca in stone
(Mol fraction CO - Mol fraction Ca)/(Mol fraction
Ca)
1.274 x Wt.fraction CO2 - Wt.fraction CaO)/(Wt.fraction
CaO)
9.
Calculate the decrepitation parameter, r, which is
proportional
to the volumetric carbon dioxide evolution rate,
from:
r=
100 e T (r
b
c
+ r )/(P
m
e
)
...
...
...
...
11.28
Section 11
21 July 1978
Page 27 of 60
Copy No.
~~~40
~~~~~~~~~~~~~4
dMx
dMx ro
ro
3.
_
...
.
_
.
....
..
.
_
30
.
20
---- --
2
'
,
-,
=
1
650
700
750
800
850
Bed Temperature Tb °C
Figure 11.5 (SI Units)
Variation of dMx ro with Tb
900
950
1000
1050
21 July 1978
Section 11
Page 28 of 60
Copv No.
105, x dMax r
O
-H
-~X:
-:---_
_.
d
Mx
H
t
ariation
.
104
ro
80
E
6-
.-
40
o M r wit
,,,
,,W
,=-frV
1200
1400
1600
bR
1800
2000
Section 11
21 July 1978
Page 29 of 60
Copy No.
and 0.01 for P in atm
where e = 1 for P in kN/m
and
10.
KU)
01 = 1 for Tb in K and 1.8 for T in R.
~~~~~b
b
Calculate tg from,
t
expanded bed volume
volumetric gas flow
tg
For beds of constant cross-section t
may be obtained
from equation 11.10.
t
11.
Lb/Uf
=
(seconds)
...
...
...
...
11.10
..
...
...
...
11.29
Calculate the value of k S from
r
ko
li
lr)l[0dm
where 6= 1 for dMx in pm and 25 400 for dMx in inches.
Note that if r
0 (i.e. the sorbent is a limestone and
conditions are below those for lime calcination) the value of
k
is reduced to about 40% of the value given by equation 11.29.
Go to step 13.
12.
Correction for finely ground limestone.
When the sorbent
maximum particle size, dMx, is such that the particles are
elutriated almost immediately at the prevailing operating
conditions then the sorbent is termed "finely ground".
Section 3.5.1 should be consulted for equations for particle
terminal velocity and Section 9.5.2 for determination of the
size of particle assumed to be elutriated immediately.
For finely ground limestone sorbent ignore equation 11.29
and determine the value of k
k
0
[ tg/( d
]
from:
...
..
where e =0.001 for dMx in pm and 25.4 for dM
Go to step 14.
....................
in inches.
..
11.30
21 JuIY 1978
Section,l
Page 30 of 60
et
Copy No.
13.
Correction for fines recycle.
t s)
When the fresh sorbent is
coarsely crushed, the recycling of fines offsets the effects
of decrepitation. The magnitude of the effect of recycling
probably depends on the recycle ratio, but there are insufficient
data available to elucidate this point. The effect is also greater
for dolomite than for limestone.
Replace the denominator in equation 11.29 by the term (1 + r - r )
where rr, the sorbent recycle parameter, has the average value
0.35 for limestones and 1.25 for dolomites.
Recycling does not bring about any improvement when the fresh
sorbent is finely ground material.
14.
Calculate the Ca/S mol ratio, Y, needed to obtain the desired
fractional reduction, Ra' in sulphur oxides emission from:
Y =
- (1/k
s
I n (1 - Ra)
a
..............
.
.
.
...
If Y is less than R
11.31
this would represent the physically impossible
situation of more than complete sulphation of the calcium. In these
circumstances assume sulphation is just complete by putting Y = R
If sorbent is to be supplied at more than the minimum rate needed
to achieve the desired reduction in sulphur oxides emission
(e.g. to maintain bed depth, Section 11.3.3.3, or to ensure
adequate inventory of sorbent for load changing, Section 11.6.4),
the value of R
that will be obtained can be calculated from
equation 11.31 in its alternative form, equation 11.9.
15.
Calculate the sorbent mass flow.
For a pure sorbent, limestone or dolomite, the mass
flow of sorbent required, M , is:
For limestone
3.12 Y times the mass flow of sulphur
For dolomite
5.75 Y times the mass flow of sulphur
For an impure sorbent, these figures need to be multiplied
by a factor calculated from one of the alternatives given
in Table 11.1.
1
21 July 1978
Section .11
Page 31 of 60
Copy No.
K!
Table 11.1
Mass Flow Factors for Impure Sorbents
Factor for limestones
11.3.3.3
Factor for dolomites
0.40 /(Wt.fraction of Ca)
0.217 /(Wt.fraction of Ca)
0.56 /(Wt.fraction of CaO)
0.304 /(Wt.fraction of CaO)
0.44 /(Wt.fraction of CO2)
0.477 /(Wt.fraction of CO2 )
Maintenance of bed depth
In some circumstances the use of a sorbent helps to maintain the
bed depth required to obtain good combustion efficiency and to submerge in-bed
heat exchange surfaces.
In these circumstances the sorbent may be fed at a
higher rate than that required to reduce sulphur oxide emission to the legal
limits.
This technique of maintaining the desired bed depth is somewhat
wasteful because of the losses of sorbent resulting from decrepitation and
elutriation, but it avoids the complexity of separate feeding of yet another
solid, e.g. sand, which would have suffered less from size degradation.
In
these circumstances the sulphur retention that will be obtained can be easily
estimated by solving Equation 11.31 for Ra for the known value of Y and k S
11.3.4
Reliability of the design method
The reliability of the above design procedure has been assessed by
comparing experimental and calculated values of retention.
The experimental
values, which were also used to derive many of the constants (e.g. A ) in
the procedure, were taken from an extensive series of tests using the equipment
and operating conditions shown in Table 11.2.
This method of assessment was chosen because the alternative of
comparing calcium/sulphur molar ratios is very sensitive to small errors in
the measurement of sulphur retention at high values of sulphur retention;
e.g. when R
> 0.97.
21 July 1978
Section .11
Page 32 of 60
Copy No.
>.
It was found that 95% of the 164 calculated values of R were
a
within ± 0.17 of the observed values.
It was noted that the predicted values of Ra may be higher (e.g.
by about 0.3) than the experimental values when the freeboard height is small
having regard to the fluidising velocity.
This is because excessive loss of
sorbent by elutriation can reduce the effectiveness of the sorbent in the bed.
Limited data suggest that this situation occurs when the freeboard gas residence
time, tf, becomes less than 0.8 seconds.
Data obtained under these conditions
were not used in the derivation of constants for the design method and have
been excluded from the above assessment.
-
The reliability of the design method was further assessed by
comparisQns made with results from the Renfrew combustor, bed area 9.6 m
(100 ft ),
which were not used in the derivation of the constants of the
design method.
The 95% confidence limits for this comparison, ± 0.19 based
on 6 points, is comparable with the limits obtained with the first assessment
of 164 values.
11.3.5
Compositions and flows of spent sorbent streams
11.3.5.1
Outline of approach
Sorbent reacts with sulphur oxides at a rate that depends on the
partial pressure of sulphur oxides;
been sulphated;
on the temperature;
on the extent to which the sorbent has
and on sorbent particle size.
Fine
material, either originally present in the sorbent feed or subsequently formed
by decrepitation or abrasion, can react relatively more quickly than coarse
material.
However this faster sulphation is partly offset by the fact that
fines have a.shorter residence time in the bed because they are rapidly
elutriated.
In general, therefore, the extent of sulphation of sorbent in
the two output streams (elutriated sorbent, and sorbent extracted with bed material)
are different.
S~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~1
Section 11
21 July 1978
Page 33 of 60
Copy No.
Table 11.2
Operating Conditions for Equipment used in Sulphur Retention Tests
CURL
CRE
Location
Bed
Size
m
ft
0.9 x 0.45
3 x 1.5
Slumped
bed depth
m
ft
0.8 - 1.4
2.7 - 4.3
Operating
pressure
kN/m 2
abs.atm
100
I
Fluidising
velocity
0.15 dia
0.5 dia
0.82 dia
2.25 dia
0.4
1;3
0.2 - 0.8
0.7 - 2.7
0.5
1.6
100
I
100
1
100
0.3 x 0.3
I x 1
I
m/s
up to 2.3
0.9
0.6 - 0.9
ft/s
up to 7.3
3.0
2.0.- 3.0
coal
coal
coal
Fuel
1.8 - 3.3
- 11
6
coal
Sorbents studied
I
1
I
UK limestone
UK dolomite
US limestone
US dolomite
1
1
2
-
-
Location
CURL
Bed
Size
m
ft
1.3 dia
1.2 dia
4.0 dia
3.75 dia
shell boiler
Slumped
bed depth
m
ft
0.7
2
Operating
pressure
kN/m2
abs.atm
Fluidising
velocity
m/s
ft/s
0.3 dia
1.0 dia
1.3 x 0.6
4 x 2
0.9 x 0.6
3x 2
0.7
2
0.5 to 1
1.5 to 3
0.7
2.5
up to 2.1
up to 6.5
100
I
100
I
up to 600
up to 6
up to 600
up to 6
up to 600
up to 6
2.6 - 4
8 - 12
1.6 - 3.2
5 - 10
up to 1.8
up to 6
0.8
2.5
up to 3
up to 10
oil
oil
Fuel
Sorbents studied
UK
UK
US
US
limestone
dolomite
limestone
dolomite
1
-
I
_
coal
coal, oil
coal
.
-
I
2
I
1
I
2
21 July 1978
Section 13
Page 34 of 60
Copy No.
'
The approach is to calculate the average fraction of sorbent
sulphated; to estimate the fraction of input calcium that is elutriated;
and then to estimate the fractional sulphation of the elutriated sorbent.
The compositions and mass flows of the two spent sorbent streams can then
be calculated.
11.3.5.2
1.
Calculation procedure
Calculate the average fraction of sorbent sulphated, k
a
2.
from:
a
Estimate the fraction of input calcium elutriated (or, when fines
are being recycled, the fraction lost from the system as elutriated
fines not recycled), f
,
from the empirical equation:
f
0.19 r - k r + k (6 U)
...
...
...
...
11.33
r r
e
f
where r is the sorbent decrepitation parameter from equation 11.28
and r is the sorbent recycle parameter (see step 13, Section 11.3.3.2)
r
e
which of course is zero when the fines are not recycled; kr
is a constant having the value I for limestone and 0.5 for dolomite;
and 8 has the value 1 for Uf in m/s and 0.3 for U in ft/s.
The calcium elutriation constant k has a value of about 0.5 for
e
limestones and 0.7 for dolomites. These values of ke, which
surprisingly appear to be independent of the value of d,x$ neverthe less are affected by the fines content of the sorbent fed.
Limited data show that if fines have been removed the value of
k is reduced by one third. When finely-ground sorbent is fed,
e
the value of f appears to be independent of fluidising velocity
or fines recycle, and a value of f = 0.97 can be assumed.
e
Equation 11.33 has been derived from data for combustors having a
deep freeboard and in which the superficial gas velocity remains
sensibly constant throughout the bed and freeboard.
21 July 1978
Section 11
Page 35 of 60
Copy No.
t
i
At high values of U
and/or r, the value of f calculated by
f
~~~~~~~e
Equation 11.33 can exceed 1, which is clearly impossible for
steady-state operation.
In these circumstances it is necessary
to make an assumption (e.g. that f
ments.
= 1) for preliminary assess-
An alternative assessment can be made by calculating
the weight of calcium in the bed by the procedure given in Section
11.6.3.
Then from a knowledge of the bed composition and the
weight of ash being extracted from the bed (Section 9) the value
of f
can be deduced from a mass balance.
In many instances,
however, the quantity of sorbent being fed may be just the quantity
needed to maintain the bed, and the value of f = 1 will be
e
appropriate.
3.
Calculate the fraction of calcium remaining in the bed, i.e.
not elutriated, fb' as the complement of f :
fb
b
4.
Estimate
=
K
1 - f
e
,
.......
...
...
...
...
...
...
11.38
the fractional sulphation of elutriated calcium, from
the empirical equation:
XKe
5.
IKa + fbife
*--
...
...
...
...
...
11.35
Calculate Kb, the fractional sulphation of calcium remaining in the
bed, from:
b
6.
a
efe)/ b
...
...
..
.
...
Calculate the mass flows of spent sorbent material elutriated, M
and retained by the bed, M b. as follows.
11.36
se
In the equations
below, the subscript x denotes either subscript e or b as
appropriate.
The calculations also enable the compositions
of the two spent sorbent streams to be estimated.
(a) Calculate M1 ,
M
=
3.40 M
the mass flow of calcium sulphate, from:
fXK
Ca
...
...
...
...
...
11.37
Section. ]
21 July 1978
Page 36 of 60
Copy No.
(ii
(b) Calculate M2x, the mass flow of magnesium oxide resulting
from decomposition and half-calcination of the dolomite content
of the sorbent, from:
M2x
=
1.66 M
f
Mg
...
...
...
...
...
...
11.38
(c) If conditions permit calcination of calcium carbonate in the
sorbent (See Section 11.3.2.2 for the test that calcination is
possible) assume that calcination is complete and calculate M3x,
the mass flow of calcium oxide, from:
M3x
= 1.40 M Sx f
X~~
Ca (I - KX)
...
...
...
11.39
(d) If calcination is complete, then M4x, the mass flow of calcium
carbonate, is zero.
If calcination is not possible then M3
is zero
and M4x is calculated from:
M4 =
2.50 M
f
Ca (1 - K )
...
...
...
...
11.40
(e) Calculate M5 , the mass flow of impurities in each sorbent
stream resulting from inert or stony constitutents of the
sorbent fed:
Msx
=
(I - 2.50 Ca -3.47 Mg)
M f
...
...
...
...
11.41
(f) Calculate Msx from:
sxC
Msx
=
11.42
Mix + M2x + M3x + M4x + Msx *....
The weight fractions of CaS0 4, MgO, CaO, CaCO 3 and impurities
can be calculated as the ratios M ]X/Ms,
M4x /Msx and Msx/Msx respectively.
M 2 /MSX, M 3 /Msx,
21 July 1978
Section I1
Page 37 of 60
Copy No.
(g) Calculate the mass flow of carbon dioxide evolved from
sorbent M6, from:
M6 = M
(1.10 K
Ca + 1.81 Mg) + 0.785 (M3
...
...
+ M3 b)
11.43
(h) Calculate M 7, the mass flow of sulphur oxides and oxygen
absorbed from the gas:
M7 .
11.4
0.588 (Mle + Mlb)..
...
...
...
...
11.44
Thermal Considerations of Sulphur Retention
11.4.1
Thermochemistry of sorbent reactions
The sorbent reactions listed in Section 11.1.3 have been rewritten
to include the heats of reaction and are shown below with heats of reaction in
kJ/g mole.
The reaction heats have been calculated from heats of formation of
the compounds involved at 25 C (77 F) listed in (11.8).
To convert the data to
the more familiar thermochemical units of kCal/g mole divide by 4.1868;
to
convert kJ/g mole to Btu/lb mole divide by 2.326 x 10-3.
CaO + S022 +
CaCO3
02
=
CaS04 + 486.1 kJ/g mole
=CaO + CO2 -183.0
CaCO 3 MgCO 3
=
kJ/g mole
(CaCO3 + MgCO 3)
...
...
...
...
11.1
......
11.2
...
11.3
...
11.4
- 31.8 kJ/g mole
(CaCO3 + MgCO 3) = (CaCO 3 + MgO) + C02 - 99.6 kJ/g mole
(CaCO3 + MgO) = (Ca
+ MgO) + CO2 - 183.0 kJ/g mole
(CaO + MgO) + SO 2 + 1 02 1 (CaS04 + MgO) +.486.1
......
kJ/g mole ...
CaCO 3 + SO2 + 1 02 = CaSO 4 + CO 2 + 303.1 kJ/g mole
...
...
11.5
11.6
11.7
As mentioned in Section 11.1.3, Reactions 11.5 and 11.6 are essentially
the same as Reactiois 11.2 and 11.1 respectively, and their reaction heats are
unchanged.
Reaction 11.7 is a combination of Reactions 11.1 and 11.2, and the
reaction heat of 11.7 is therefore the algebraic sum of the heats of the two
reactions 11.1 and 11.2.
The sulphation reactions 11.1, 11.6 and 11.7 are exothermic.
The remaining reactions 11.2 and 11.5 are endothermic.
The overall net heat
21 July. 1978
Section 11
Page 38 of 60
Copy No.
of reaction, therefore, depends on the sorbent type (limestone or dolomite);
on whether the calcium carbonate is calcined;
the sorbent is sulphated.
and on the extent to which
At one extreme, a sorbent used under calcining
conditions with a low sulphur retention and a high Ca/S ratio will absorb
considerably more heat than is furnished by the sulphation reactions, and this
heat has to be supplied by combustion of fuel, thus reducing the thermal
efficiency of the combustor.
At the other extreme, a sorbent that becomes
highly sulphated, particularly if used under non-calcining conditions, evolves
a significant amount of heat and reduces the fuel consumption.
11.4.2
Sensible heat changes
The heats of reaction listed above refer to reaction at 25 C
(770F).
As is commonly done, it is convenient to assume that reaction
occurs at this temperature, and. that the reaction products are heated to
the bed temperature, Tb.
The overall thermal effect of the use of SO2
sorbents therefore also includes the sensible heat contehts of the spent
sorbent streams leaving the combustor.
In instances where this heat
constitutes a significant loss it may be worth taking steps to recover
a part of the sensible heat of spent sorbent.
11.4.3
Calculation procedure
The calculation procedure involves calculations of the net
reactions heat at 25 0 C (77 F) and of the net sensible heat losses in
spent sorbent and gas.
t.
Calculate Hr, the net heat of reaction, from the individual
reaction heats found from the following equations.
Note
that in equations 11.45 to 11.49, e has the value I for M
in kg/s and H
in kW, and 2.326 for M
in lb/h and Hr in Btu/h
(a) Calculate H 1' the heat of sulphation of calcium carbonate
from
H
=
rl ~
M
K Ca (7560/0)
s a
...
...
...
...
(b) Calculate H 2' the heat of decomposition and halfcalcination of the dolomite content of sorbent, from
...
11.45
21 July 1978
Section.1!
Page 39 of 60
Copy No.
Hr
=m Smg (5410/a)
...
...
...
...
...
ii.46
(c) If conditions permit calcination of the calcium
carbonate content of the sorbent (see Section 11.3.3.2,
step 4, for the test that calcination is possible), assume
that calcination is complete and calculate the reaction heat,
Hr3
from:
Hr
if
MSCa
m
(1
Ka)
(4570/e)
...
...
...
...
...
11.47
..
114
calcination is not possible, Hr3is, of course zero.
(d) Calculate H
Hr= Hri 2.
-
Calculate Hse
from:
r2 Hsb
....
r3..
and H1
the net sensible heat losses in the
spent sorbent in the elutriated stream, in spent sorbent extracted
with bed material and in the carbon dioxide respectively, and H7,
the net sensible heat saving arising from the absorption of sulphur
oxides plus oxygen from the gas stream, as follows:
(a) For each sorbent stream and for the gas leaving the<system
interpolate the appropriate value of sensible heat capacity At
the relevant temperature above ambient temperature (15 C or 590F)
0
~~~~in kJ/kg or Btu/lb from Figure 11.6 as follows:
H
se
Hs
H
6
H
Spent sorbent in the elutriated stream
Spent sorbent in the overflow bed material
Carbon dioxide
Sulphur oxides and oxygen
7
(b) Calculate the net sensible heat loss from the system, HS
from:
HS
(Hsemse + HsbMsb + H6M
6
-H 7
M7 )
/e
...
...
11.49
Sectioni 11
21 July 1978
9f ~60
fraC.4.e
Covy No.
ZZ-~~~~~~~~~~~~~.
JE
JA=~~~~~~~~~~~~
QT-~~~~~~~~~~~~~~~C
M -W
~~
i 901~~~~~~~~~~~~~~~~~~~~~~~~0
9M E-~~~~~~~~~~~~~~~~~~~~
Figure 11.6
Variaion
(SI
Units)~~~~~~~~~~~~~~~~~~~~~C
f SenibleHeatswithTempeatuH
~
~
~
~
Section 11
21 July 1978
Page 41 of 60
). Copv No.
:=:
o0
0
a
0
G,
_
,
-7-
c
of SesibieHeat withTempeatur
r~~~~~aito
:~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~,L
4*tX
=
W1
=;°1;3
2f
~S
S 010 W
Figure 11 .6 (British Units)
Variation of Sensible Heats with Temperature
o
21 July 1978
Section 11
Page 42 of 60
Copy No.
3.
(2V
The overall thermal effect of retaining sulphur can now be
evaluated. The net heat loss from the system, HL, is the
sensible heat loss, H , less the algebraic value of the heat
of reaction, Hr
HL
'
HS
s
Hr
r
.. ...
...
...
...
...
...
11.50
In general HL will have a positive value, i.e. there will be
a net heat loss from the system resulting from the use of a
sorbent.
11.5
11.5.1
Sorbent Reactivity
Variation with composition
It has been seen in Sections 11.2.5 and 11.3.3.2 that the
comparative reactivities of sorbents are expressed by the values of the
pre-exponential term A in equation 11.11, defined as the Reactivity
Index of the sorbent.
It has also been seen that the value of A
S
can
vary over wide limits depending on the porosity characteristics of the
sorbent. Dolomites generally have a value of A in the neighbourhood
-
of 46.
Most limestones have values of A
in the range 20 to 40.
The value of A
is generally (though not invariably high with
sorbents of low calcium content and vice versa. This is illustrated by
the high Reactivity Index of most dolomites which have a calcium weight
fraction of about 0.2, and by the low Reactivity Index (around 20) of many
of the purest limestones with calcium weight fractions of about 0.4.
Similar observations are reported in (11.9).
There are so many exceptions
to the general rule outlined above, however, that it would be misleading
to formalise it as a sound means for predicting Reactivity Index.
11.5.2
Determination of Reactivity Index
In many instances it will be possible to evaluate the Reactivity
Index of a sorbent from available performance data. If these are not available,
and the cost can be justified, the best procedure is to conduct a test on a
21 July 1978
Section 11
Page 43 of 60
Copy No.
small pilot plant having a combustor of diameter of at least 150 mm (6 inch).
The ideal would be to operate with fuel from the same source, and with
conditions as close as those relevant to the plant being designed.
Operation
should preferably be with two or more Ca/S mol ratios, each for sufficiently
long duration to achieve steady-state conditions and accurate mass balances.
The calculation of A
from previously available performance data
or from the test results would then be as follows.
1.
For a known value of Y and Ra , solve Equation 11.31 for k
2.
For known values of t, dM
and r, solve Equation 11.29
for k.
0
3.
For the known value of Tb, solve Equation 11.11 for A S
If the cost of combustion tests is not justified, e.g. if a choice
is to be made from a number of candidate sorbents, an alternative procedure
may be employed.
This is a laboratory-scale "Sorbent Characterisation Test"
that has yielded encouraging results, although at the present time it gives
values of A
that are less reliable than those obtained from the small pilotscale plants as described above. The test is discussed in Section 11.5.3 below.
11.5.3
Sorbent characterisation test
The sorbent reactivity test rig and test procedure is described
in Section 17.5 and reference (11.10). The rig is used under standardised
operating conditions, and the extent of sulphation, KA, of the sample of
sorbent is measured at 845 C (1553 F) over the period, tAhours, during
which the SO2 concentration in the dry exhaust gas rises to within 100 vpm
of the concentration (1800 vpmt) obtained without the sorbent sample. The
duration of this period is also measured.
For an advancing-interface gas-solids reaction the conversion,
K, of a spherical particle of diameter, d, reached after a time, t, is
given by:
d {1
-
(1 -
6
}
=
XX) t
..
...
...
...
...
11.51
21 July 1978
Section 11
Page 44 of 60
Copy No.
/
-
where x is a rate constant related to temperature and to particle
reactivity.
The left hand side of equation 11.51 represents the
thickness of the layer of reaction product.
When reaction products seriously impede the access of gaseous
reactants to the reaction interface, as in the sulphation of lime, the
effective rate constant falls in proportion to the thickness of the
reacted layer.
This is confirmed by data presented in (11.10).
Equation
11.51 has to be modified to the form:
(d/t) {I - (I -
X - X' d {I - (I - Kr}
K)3}=
...
11.52
where X' is a constant.
Since equation 11.52 involves two constants, X and X', a
simpler form is needed for present purposes, and it is found that over
a wide range of values of K, a reasonable approximation is given by:
[
-
(1
St
K)J]
K
=
constant
in which the constant is proportional to k
...
...
...
...
11.53'
in equation 11.11, and
therefore, for a given temperature, to the value of A
S
Hence, having evaluated KA and tA from the sorbent reactivity test,
S
calculate A
from
r~~~~~~~
=1
(2.22 x 10
/tA)(dA)
exp(3530
/T)]
-
(1
-
...
where e = i for dA in pm and 25 400 for d Ain inches.
e, = I for TA in °K and 1.8 for TA in OR.
For the operating conditions of the test as described in Section 17.5 and
reference (11.10), d is 600 pm (0.024 inch) and T' is 1118 K (2012 0 R)
A
A
845 C(1553 F) . The term in square brackets then reduces to (1880/t
A
and equation 11.54 becomes,
e~~~~
A=
1880/t
[I
-
(1
-
KA)41
11.54
21 JulY 1978
Section 11
Page 45 of 60
Copy No.
11.5.4
',
Improving sorbent reactivity
The use of sodium chloride as an additive for improving the
reactivity of limestone or dolomite sorbents has been reported. (11.9, 11.11).
It has been shown (11.9) that other chlorides, e.g. calcium chloride, are also
effective additives for increasing the reactivity of sorbents, but that excessive
quantities of additive reduce reactivity.
Sorbents that are highly reactive
generally suffer from some deactivation when treated with additives.
There may be a similar effect from chlorine in the coal.
It
was observed (11.1) that the sulphur retention with a particular British
coal was higher than that with other British and United States coals.
The
coal in question had a high chlorine content (0.6% compared with 0.1 to 0.2%
for the other coals studied).
The explanation of these phenomena appears to be that chloride
forms liquid phases with carbonate or sulphate, and that this promotes
crystal growth leading to increased pore size and improved access of
gaseous reactants to the centre of the particle.
Excessive use of additive
reduces the internal reaction surface, and this explains their harmful effect,
particularly on reactive sorbents.
It has been found (11.9) that precalcination and heat treatment
also improves sulphur retention, the improvement increasing with the
period of pretreatment.
Several effects are involved.
(a) Precalcination ensures that the material fed has a
developed internal porosity so that the reactivity of a
limestone is high even for sulphation under "non calcining"
conditions.
(11.12)
(b) Precalcination obviates the occurrence of decrepitation
caused by calcination in the combustor.
(c) Heat treatment results in crystal growth within the
particle, and. this leads to the development of large pores
permitting access of gaseous reactants to the interior of the
particle.
As with the use of additives, excessive treatment
21 July 1978
Section .11
Page 46 of 60
Copy No.
K>
of a good sorbent can result in some loss of reactivity.
Both the use of additives and the techniques of precalcination
and heat treatment are in the early stages of investigation and have not
been developed to the state of commercial application.
There are as yet
inadequate data for predicting improvements obtainable in the value of the
Reactivity Index, or for optimising the sorbent pretreatment.
It is obvious
that the techniques will increase the cost of unit mass of sorbent, and may
increase the complexity and fuel consumption of the process.
The techniques
may also increase the corrosion of cooling surfaces in the bed, freeboard and
throughout the gas path to the stack, and of gas turbine blades in pressurised
fluidised combustion applications.
In some circumstances, however, these
penalties may be worthwhile.
11.5.5
Regeneration of spent sorbent
Spent sorbent can be treated to release sulphur dioxide from the
calcium sulphate, leading to the regeneration of calcium oxide, by heating
to a temperature of 1000 - 1100 C (1800 - 20000 F) or more under mildly
reducing conditions.
This process has not been studied by CSL or its partners
but is briefly reviewed here for the convenience of the designer.
The chemistry
of the process is as follows:
CaSO4 + CO
=
CaO + SO2 + CO
- 202.9 kJ/g mole
...
...
11.55
CaSO4 + H2
=
CaO + SO2 + H20 - 244.1 kJ/g mole
...
...
11.56
The reactions are endothermic.
The heats of reaction are calculated from the
heats of formation at 25TC (77 F) using reference (11.8).
The reactions
probably occur in stages, with the intermediate formation of calcium sulphide
which then reacts with sulphate.
CaSO4 + 4 CO
=
CaS + 4C0 2
3 CaSO4 + CaS = 4 CaO + 4 S02
...
...
...
...
...
11.57
...
...
...
...
...
11.58
Reaction 11.57 is favoured by low temperatures, e.g. about 800 C
(1450 F).
21 July 1978
Section .11
Page 47 of 60
Copy No.
An alternative regeneration process has been reported that
involves conversion of sulphate to sulphide by reaction 11.57, followed
by regeneration of the sulphide in a separate step with carbon dioxide
and water vapour at about 600 C (1100°F):
CaS + H20 + CO2
-
...
CaC03 + H2S
...
...
...
11.59
However it has been found that this reaction sequence leads to rapid
deactivation of the lime, probably because of the formation of eutectics
in the CaS-CaSO 4 system.
This alternative regeneration process, which
was originally proposed because it should be favoured by a high operating
pressure, is therefore not recommended.
Although early studies of regeneration were mainly concerned
with operation at atmospheric pressure, (11.13), much of the recent work
has been at high pressure (11.9),
(11.12).
The regeneration reactions are similar to those occurring in the
"Chemically Active Fluidised Bed" process, described in reference (11.19),
for gasifying a sulphur-containing fuel (usually oil) to produce a hot,
combustible, low sulphur gas for firing boilers.
with air at about 800 to 900 oc (1450 to 1650
which is the "chemically active` material.
The fuel is gasified
F) in a fluidised bed of lime
The lime absorbs sulphur to form
calcium sulphide, and this is regenerated to calcium oxide in a separate
bed at a temperature of about 1000 to 1100
less reducing atmosphere.
C (1800 to 2000
F) and with a
The gasifier and regenerator are the two units in
a system with continuous circulation of sorbent.
For fluidised combustion, regeneration can bring about significant
savings in the need for fresh sorbent, and can also reduce the quantity of
spent sorbent to be disposed of.
However, there are penalties to be
paid in plant complexity and capital cost;
in the need to consume additional
fuel to heat the spent sorbent to regeneration temperature and to supply the
beat of the strongly endothermic regeneration reactions;
and in recovery of
the sulphur dioxide evolved and its conversion to an acceptable form for sale
or for dumping (e.g. as elemental sulphur).
These penalties must be set
against the costs of fresh sorbent and the environmental or economic costs
21 July 1978
Section 11
Page 48 of 60
Copy No.
of dumping spent sorbent.
--
-'
Disposal of spent sorbent is in some respects less of
a problem than in most non-regenerative wet flue gas desulphurisation processes
using lime, where the effluent slurries tend to be excessively difficult to dewater.
However, as legislation imposes increasing demands upon the level of sulphur
retention required which in turn increases the amount of sorbent required to
meet these levels, then the overall economics of regeneration could become more
attractive.
11.6
Sulphur Retention During Unsteady State Operation
11.6.1
Introduction
All of the calculation procedures described in Section 11.3 are valid
for operation under steady-state operating conditions following a period in which
the bed weight and composition have been stabilised.
Such conditions are typical
of those in the test runs on the pilot plants during which the experimental data
were obtained.
Base-load commercial plant could also be expected to operate
largely under such steady-state conditions.
Most commercial plant, however, will
need to follow a fluctuating load, and operating conditions may therefore change
rapidly and cover wide ranges.
As can be expected, there are changes in sulphur oxides emissions when
operating conditions are changed.
Some
changes are rapid, and others are slow.
The designer has to make provision for these changes if the operation of the plant
is not to result in emissions that intermittently exceed the local limits.
In this Section a general method is given for predicting approximate
rate of response of the combustor to changes in operating parameters.
11.6.2
The kinetics of sulphur dioxide absorption
The problem of predicting response times hinges on the kinetics of
sulphur dioxide absorption which can be expressed by a rate equation of the form:
dK/dt
(P 5 02/d) (1
-
[X
X
X
d {1
-
(1
-
K)
.
...
11.60
21 July 1978
Section ll
Page 49 of 60
Copy No.
\t''
where dK/dt is the rate of sulphation of a particle of diameter, d, at a
sulphur dioxide partial pressure, Pso . The constants X and X' are the
same as those used in equation 11.52 which was an integrated and slightly
simplified form of equation 11.60.
At the conditions prevailing in fluidised combustors, the
reaction proceeds rapidly when K is close to zero, with (di/dt) values
in the range I to 10 h . However, it is unusual to operate with low values
of
The value of (dK/dt) falls rapidly with increase in K and for the sorbent
inventory of a typical bed, mean values, denoted (dK/dt) are in the range of
0.01 to 0.1 h . These mean values relate to mean residence times, t , of
several hours. The value of t is defined as the total calcium inventory of
K.
the bed divided by the hourly input of calcium.
It follows from these considerations that a change in operating
conditions can lead to an immediate change in sulphur oxides emission
which may be rectified only slowly by a change in Ca/S ratio.
The
composition of the bed has-to be changed to permit the new inventory of
calcium with a new extent of sulphation to be achieved. The process can
obviously be accelerated by initially feeding the sorbent at a higher Ca/S
ratio than that ultimately required; this not only changes the inventory
more quickly, but temporarily gives the sorbent present in the bed a lower
value of K and therefore a higher effective reactivity. However, there are
no experimental data to predict these temporary effects.
The calculation procedure given below enables the time spent in
reaching steady-state conditions and the weight and composition of the partly
sulphated sorbent inventory of the bed at the new steady state to be estimated
approximately.
11.6.3
Calculation procedure
The calculation procedure makes use of an empirical relationship
to estimate mean sorbent residence times in the bed. Equations 11.52 or 11.60
cannot be employed in these calculations because the sorbent inventory of the
bed is composed of particles
-
Section 11
21 July 1978
Page 50 of 60
Copy No.
KI<
(a) that have resided in the bed for a range of times and
are therefore reacted to different extents so that there is
not a unique value of
K
or of t for the bed;
and
(b) that are of different sizes and specific surface areas
(i.e. no unique value of d), and that therefore become
sulphated at different rates.
The calculation procedure is as follows:
1.
Estimate, t
(hours), the mean residence time needed by the sorbent
to achieve the required value of Ka from:
t
where e
8,
=0.013 Ka (8 dMx)(1 PS%0
...
k)
...
...
...
11.61
2
= I for dMx in pm and 25 400 for dM, in inches.
in kN/m2 and 100 for Pi in atm.
1 for P
2
The values of Ka
are for steady-state operation at the
PS0 , dmx and k
new conditions, with
and k
K
being obtained from equations 11.32 and
11.11 respectively as previously described.
2.
Estimate f for steady-state operation at the new conditions
b
from equations 11.33 and 11.34.
3.
Estimate t , the time taken to achieve steady-state operation
at the new conditions, from:
ts
4.
=
tm /fb
....
...
...
...
...
...
I I.62
Estimate fR' the fractional change to the new steady-state sulphur
retention at a time t,
fR
R
5.
..
=
(t/t
from:
s
)3
........
...
...
...
...
...
...
...
11.63
Calculate also Wca, the weight of calcium in the bed at the new steady-
state, from:
WCCa
=
tmm M SCa
s
...
...
...
..
...
...
...
11.64
By assuming that the bed has the same composition as the material withdrawn
from it (Section 11.3.5.2) the weight and composition of the partly-sulphated
sorbent inventory of the bed can be established.
21 July 1978
Section 11
Page 51 of 60
Copy No.
11.6.4
,
Considerations during start-up
Start-up of the combustor is a particular instance of unsteady-
state operation when the sulphur-retention efficiency of sorbents will be
severely limited by low temperatures.
To illustrate this point, the ratio
of the value of Y needed to that at 800 0 C (1472 0 F) is plotted in Fig. 11.7
over the range of combustor temperatures at which coal
be initiated during start-up.
purposes of illustration only.
or oil.:firing might
The curve of Figure 11.7 is intended for
Firm experimental data are not yet available
for the lower temperatures and the curve has been calculated by the procedures
described in Section 11.3.
Unless the local air pollution control legislation grants dispensations
for start-up periods, it is recommended that:
(a) The initiation of firing of the sulphur-containing fuel
be deferred until a bed temperature is reached at which
the sorbent can bring about the necessary reduction in
emission.
(b) The bed used for start-up be composed of fresh sorbent,
preferably previously calcined to minimise the effects
of decrepitation and (with limestone) the poor access
to internal surface with uncalcined stone.
11.6.5
Recommended operating plan
It is likely that with commercial plant operated on a loadfollowing basis it will be necessary to change load quickly, e.g. at the
rate of 5% per minute.
For a boiler operating at 2:1 turndown, restoration
of full load might be needed in as little as 10 minutes.
The sulphur input
during such a change would be at least doubled, and if the bed had been
stabilised to give the minimum acceptable sulphur retention at the low
load the emission is likely to peak at many times the limit.
Under all
operating conditions the input of fuel will fluctuate, even at nominally
steady-state operating conditions.
These fluctuations will be reflected
in exaggerated fluctuations in the sulphur oxides emission for the same
reasons as described above for turndown changes.
Section 11
21 July 1978
Page 52 of 60
Covy No.
Bed Temperature Tb OF
1400
1200
1000
1600
Y at T b
Y at 8000 C
I
0
500
600
700
800
Bed Temperature Tb OC
Figure 11.7
Effect of Bed Temperature on the Ca/S Ratio Needed
900
k
21 July 1978
Section 11
Page 53 of 60
Copy No.
The solution to the problem will have to be to operate with
better than the minimum acceptable sulphur retention, particularly when
operating under low load conditions, to ensure an adequate sorbent
inventory for dealing with load changes.
The extent to which sulphur
needs to be "over-retained" at low load will depend on the likely loadchanging pattern of the particular plant.
11.7
11.7.1
Sulphur Trioxide and the Acid Dewpoint
Introduction
When fuels containing sulphur are burnt, most of the sulphur
appears in the exhaust gases as sulphur dioxide, but a proportion invariably
appears as sulphur trioxide, which forms sulphuric acid vapour.
Sulphuric
acid has a high dewpoint (around 150 C (300 F) depending on its concentration
and the water content), and this, together with its corrosive properties,
makes it necessary to predict whether condensation in the cooler regions of
the exhaust gas system is likely to occur.
Sulphur trioxide can be regarded as being formed by oxidation of
sulphur dioxide according to the equation:
02 + ~2
2 = So..
so2
1 °2
so3.......
.
. ... . ... .
11.65
At equilibrium the mol fraction of sulphur trioxide, is given by:
-Kp
=
m 50 /(m50
,p ~mso3
s2
Pi)
02
..*.
*..
...
...
...
in which the equilibrium constant, K , falls with increase in temperature.
p
Equation 11.66 shows that the equilibrium concentration of sulphur trioxide
is proportional to the concentration of dioxide, and to the square roots
of the oxygen concentration and of total pressure.
Hence the equilibrium
concentration of sulphur trioxide is reduced by retention of sulphur but
increased by pressure and excess air.
The reaction kinetics are complex, as explained in (11.17), and
are evidently not fully understood.
One reason for this is that it is
11.66
21 July 1978
Section 11
Page 54 of 60
Copy No.
l
difficult to measure sulphur trioxide concentrations accurately: great
care has to be taken in sampling to avoid reaction with or absorption on
the sampling probe.
Sulphur trioxide is also absorbed by the walls of
furnaces, exhaust gas ducts, etc., and becomes desorbed slowly when the
concentration in the gas falls or the temperature rises, giving nonequilibrium concentrations in the gas for prolonged periods (e.g. several
days). The reactions may be heterogeneous, catalysed by e.g. fuel ash,
and/or homogeneous, involving gaseous constituents such as carbon monoxide
and nitrogen oxides.
The subject has not been extensively studied in connection with
fluidised combustion. Such measurements as have been made on the pressurised
*
~~~~~~~~~~~~~2
pilot plant at 350 to 600 kN/m (34 to 6 atm) pressure, both with and without
sorbent addition, indicate that equilibrium has been approached by a fraction,
6 , defined as the ratio of the observed to the equilibrium value of k , in
the range 0 to 0.2, averaging about 0.1. (11.1, 11.18). Similar results
have been obtained on one of the atmospheric-pressure pilot plants. However,
measurements made at 900 to 970 kN/m2 (8.9 to 9.6 atm) over a wide range of
operating conditions (11.12) showed values of 6 in the range of 0 to 2.8,
averaging about 0.4.
No systematic variation with any operating parameters
could be detected, and these results illustrate the difficulties of studying
sulphur trioxide formation in fluidised combustors.
The calculation procedure given in Section 11.7.2 involves
assuming a value of 6 and using it, together with the equilibrium constant,
the concentrations of sulphur oxides and oxygen, and the total pressure, to
calculate the concentration of sulphur trioxide. The acid dewpoint can then
be estimated.
If this is too high for avoiding corrosion in the cooler regions
of the exhaust gas system, it may prove necessary to reduce the total sulphur
oxides concentration by increasing sulphur retention beyond the minimum needed
to meet local air pollution control requirements. Such an increase may be
particularly necessary if load variations are likely to produce the absorption/
desorption effects referred to above.
21 July 1978
Section 11
Page 55 of 60
Copy No.
11.7.2
1.
Calculation procedure
Calculate the oxygen.partial pressure in wet combustion gas, P0
from:
2
P02
2.
(0.21 P Wa/W w )
(1 - Cb + X/100)
...
...
11.67
Calculate Kp from:
where
%Kp=
(2.12 x 10
e
=
I for pressure in kN/m 2 and 10 for atmospheres.
=
I for Tb in OK and 1.8 for oR.
1
x 9) exp (l11 380 e1e/T~),
.
...
...
11.68
Figure 11.8 shows the variation of KP with bed temperature.
3.
Calculate the mol fractions of sulphur trioxide, mso , and sulphur
dioxide, ms0 , as follows:
(a) Assume a value for 6 .
A value of 6
of 0.2 is
recommended for operation in the pressure range 100 to
600 kN/m 2 (1 to 6 atm.) and a value of 0.4 for operation
at about 1000 kN/m
2
(b) Calculate mso
(10 atm.)
from:
3
(m Pe K 6 )I
=
(1 +
ms03
e0 2 p5s)
(c) Calculate mS0o from:
mso 2
=
me
m0 so3
P
2
K
P
)
..
...
.
-
..
...
...
11.69
.
11.70
In equations 11.69 and 11.70, m%, the mol fraction of sulphur oxides in
the wet flue gases is calculated by substituting m
in equation 11.20.
m
e
Thus
mW
w
(1 - R)
for e
and me for e
21 July 1978
Section 11
Page 56
f
60
CopyvNo. k~
4
~~~~0.50
-
Z
.7
0.4~~~~~~~~~-
0.40
0.35
-~
[iatm2]6-4-
---
0.25
E~~~~~~~4-
-
vt
0~~
650
700
~
~
750*-
~
t-
~~~f'
-0
-5
Be
~~~~~~~igre-4
~
-
~~FF
--
4Z'~
020~
-~~E
I-tI
0
Tepeatue4
nis
--
~~~~Vrato
ofK4ihTnprtr
T
5
O
_S
0015
Section 11
21 July 1978
Page 57 of 60
Copy No.
\ '
5.0 .
atm*5
4~0'N
'-::.
33-
;3.0
......
2.0
Figure 11 .8 (British Units)...
Variation of Kp with Bed Temperature
'--
.......
Section. 11
21 Ju.ly 1978
Page 58 of 60
Copy No.
Calculate mH2 0 ' the mol fraction of water vapour in the wet exhaust
4.
gas, from:"
m 2O
=
- Wd/W*-
*.-
................
...
11.71
2
Estimate the absolute temperature of the acid dew point, Td,
5.
from the empirical relationship below:
T' d
where
IfO
= .- 21 400 8/ {In(ms
0 =
I for Td in
)-+ 1.67 In (n20)
K and 1:.8 for, R.
- 37.1} ...
...
...
11.72
Section I1
.21 Ju-ly 1.978
Page 59 of 60
Copy No.
11.
&_.
i~
References
11.1
NCB "Reduction of Atmospheric Pollution".
Report No.
DHB 060971 to Environmental Protection Agency, (September 1971).
11.2
Ingraham, T.R. and Marier, P., Can. J. Chem. Engng. 170 - 173
(August 1963)
=1.1.3
Yang, R.T., Krishna, C.R., and Steinberg, M., Ind.Eng.Chem.
465-467, 16 (4),(1977)
11.4
Murthi, K.S., Harrison, D. and Chan, R.K., Environmental Science
& Technology, 776 - 781, 5 (9), (1971)
11.5
Anastasia, L.J., Carls, E.L., Jarry, R.L., Jonke, A.A., Vogel, G.J.,
(Argonne National Laboratory).
Paper to Second International
Conf. on Fluidized Bed-Combustion, Hueston Woods, Ohio.
(October 4-7 1970).
11.6
Glenn, R.D., and Robison, E.B., (Pope, Evans and Robbins) Paper
to Second International Conf. on Fluidized-Bed Combustion,
Hueston Woods, Ohio (October 4-7, 1970).
11.7
Merrick, D. and Collop, D.J. The Design of Fluidised Bed
Combustion Systems - Computer Program User Manual" NCB, PADB
Memorandum No.141, (December 1977).
11.8
Perry, R.H., Chilton C.H. and Kirkpatrick S.D. (Eds.), "Perry's
·Chemicai Engineers' Handbook", Fourth Edition,
New York:
McGraw-Hill Book Co.Inc., (1963).
11.9
Vogel, G.J.,.et al., (Argonne National Laboratory), "Supportive
studies in Fluidized Combustion".
Annual Report July 1976-June
1977 No.ANL/CEN/FE-77-3 to U.S.Energy Research & Development
Administration under Contract No.W-31-109-Eng-38.
11.10
Anderson, J.S. (NCB, CRE), "Characterisation of limestones using
a 0.15 m diameter fluidised bed test rig".
CRE Report No.MP
(November 1976).
11.11
Ehrlich, S., "A coal-fired fluidised-bed boiler".
Paper C4, Int.
Conf. on Fluidised Combustion, Imperial College, London,
September 1975.
Inst.Fuel Symposium Series No.l:
Combustion, Vol.1 (1975).
Sb
16-17
Fluidised
Section 11
21 July 1978
Page
60
:of 60
Copy No.
11.12
..
Hoke, R.C., Bertrand, R.R., Nutkis, M.S., Kinzler, D.D.,
Ruth, L.A.;, Gregory, M.W. and Magee, E.M. (Exxon Research and
Engineering Co.) "Studies of the Pressurized Fluidized-Bed
CoalsCombustion Process". Report No.EPA-600/7-77-107.to U.S,
Environmental Protection Agency, Office of-Research & Development
(September 1977).
11.13- Glenn, R.D. and Robison, E.B., "Characterization of Emissions
from Fluidized-Bed Combustion of Coal and Cpntrol of Sulfur
Paper to Second International Conf.
Emission with Limestone".
on Fluidized-Bed :Combustion, Hueston Woods, Ohio (4-7 October 1970)
11.14
NRDC "Pressurised fluidised bed combustion`' R & D Report No.85,
Interim No.1 to Office of Coal Research, U.S.Dept. of the
Interior, under Contract No.14-32-0001]1511, (1974)..
11.15
11.16
Roberts, A.G., Stantan, J.E., Wilkins, D..M., Beacham, B. and
Hoy, H.R. "Fluidised combustion of:coal and oil underpressure".
Inst.
Paper D4 to Int.Conf. on Fluidised Combustion, London.
(September 1975).
Fuel Symposium Series No.1.
Moss, G., "The mechanisms; of sulphur absorption in fluidised
Paper D2 to Int.Conf. on Fluidised Combustion,
Inst.Fuel-Symposium Series No.1. (16-17 September-1975).
beds of lime"'.
London.
11.17
Reid,W;.T., "External corrosion and deposits: boilers and gas
New York: American Elsevier Publishing Co. Inc.(1971).
turbines".
11.118 '.NRDC,"Pressurised Fluidised Bed Combustion - Complete Report of
Test: Run No.5" Report No.FE-1511-DRAFT to U.S.Energy'Research
and Development Administration, under Contract No. E(49-:18)-1511
(September 1975).
11.19
Craig, J.W.T., Johnes, G.L.,:Moss, G., Taylor, J.H. and Tisdall, D.E.
"Study of Chemically Active fluid Bed Gasifier for reduction of
sulphur oxide emissions:
Final report".
Report No.ER.1 KQA-F72
to U.S.Environmental Protection Agency under. OAP Contract
:CPA 70-46, (June 1972).
11.20
Jonke, A.A. et al (Argonne National Laboratory) "Reduction
of Atmospheric Pollution by the Application of Fluidized-Bed
Combustion". ANL Annual Report to Environmental Protection
Agency No. ANL/ES-CEN-1004,(1971).