Issue 002
Section 3
Section Index
CODY NO.
3.
3.1
The Design Task
3.1.1
3.1.2
3.2
DESIGN PHILOSOPHY
Principal design considerations
Basic combustor concepts
Design Procedure
3.2.1
3.2.2
Preliminary considerations
Stepwise procedure
3.2.2.1
3.2.2.2
3.2.2.3
Calculation of combustor cross-sectional area
Bed maintenance and in-bed heat transfer
Feedstreams and control
3.3
Computer Programs
3.4
Minimum Fluidising and Terminal Velocities
3.5
Effects of the Principal Operating Parameters
3.5.1
3.5.2
3.5.3
3.5.4
Fluidising velocity
Bed temperature
Excess air level
Bed depth and splash height
3.5.4.1
3.5.4.2
3.5.4.3
3.5.5
Properties of bed solids
3.5.5.1
3.5.5.2
3.5.6
3.5.7
3.6
Static bed depth
Expanded bed depth
Splash height
Materials
Particle size
Coal feed particle size
Air and fuel distribution
Combustion and Unit Efficiences
3.6.1
3.6.2
Combustion efficiency
Overall thermal efficiency
3.7
Pattern of Heat Release
3.8
Heat Transfer from the Bed
3.9
Sulphur Retention
3.10
Turndown
3.10.1
3.10.2
3.11
General considerations
Choice of method
References
Issue 002
Section 3
Page 1 of 56
Cony No.
3.
3.1
DESIGN PHILOSOPHY
The Design Task
3.1.1
Principal desien considerations
The design of a fluidised bed combustor involves satisfying the
requirements of a minimum of three processes.
They are a combustion process to
generate the heat, a heat transfer process to remove the heat as it
is released
so that the bed is kept at a constant temperature and a fluidisation process to
5g
provide a suitable bed of particles so that the other processes can operate
steadily.
The existence of this third process is the main reason why the
design of fluidised bed combustors is
fired equipment.
different from that of conventionally
When sulphur retention is
desired the requirements-of the
sulphur retention process must also be satisfied.
The bed of a fluidised combustor is supported by the air distributor
and is fluidised by the combustion air.
place where the combustion occurs.
Its primary function is to provide the
The fuel is
fed to the bed at one or more
points so that the fluidisation process can disperse it quickly and evenly
throughout the bed.
In boiler applications the bed is also used for the
transfer of heat to cooling surfaces immersed in it.
In such applications it
is desirable that the heat of combustion is released as far as possible within
the bed as the rates of heat transfer from the bed to surfaces contacted by it
are up to ten times greater than those to surfaces in the freeboard above the
splash zone or from the gases downstream of the combustor.
sulphur retention is
desired,
form of solid calcium sulphate.
In addition,
when
the bed is used to trap the fuel sulphur in the
The sulphur is more easily disposed of in this
form than through the use of a flue gas desulphurisation stage.
The processes of combustion, heat transfer, fluidisation and sulphur
retention can interact on each other and values of the design parameters must
often be a compromise between conflicting requirements.
considerations are briefly as follows.
The principal design
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Copy No.t
The mean bed temperature is a compromise.
At low temperatures
combustion efficiencies may become unacceptably low because of reduced reaction
rates and increased elutriation of unburnt carbon.
On the other hand, the
temperature must never be too high because of the risk of sintering the bed
material.
The normal range of bed temperatures is 750 - 950C (1380 - 1740°F)
but a temperature around 850°C (1560°F) is an optimum for sulphur retention.
The bed depth is commonly in the range 0.15 - 1.0 m (6 - 40 in.)
static.
The depth must be sufficient to allow adequate dispersion of the
fuel and sufficient reaction time for the combustion process.
However, the
shallower the bed the less the power required to provide the air supply. Also
the rate of erosion of in-bed tube banks is decreased as the bed depth is
reduced.
Nevertheless the depth must be sufficient for the bed to cover any
in-bed tubing at MCR conditions and values at the upper end of the above range
may be necessary if
a high degree of sulphur retention is
required.
For furnace and incineration applications the excess air level is
dictated by the bed heat balance and is likely to be upwards of 100% depending
on the calorific value of the fuel.
For boiler applications the excess air
level is commonly in the range 20 - 50%. With coal firing the combustion
efficiency drops rapidly as the excess air level falls below about 20% while
the flue gas heat losses increase with increasing excess air level.
The fluidising velocity must be sufficient to
give adequate
fluidisation at all operating conditions and there is an economic incentive to
increase the fluidising velocity as this will lead to a reduction in the
combustor cross-sectional area.
On the other hand, the rates of solids
elutriation and of erosion of combustor and components both increase markedly
with increasing fluidising velocity while the residence times for combustion
and sulphur retention decrease.
The fluidising velocity is generally in the
range 1.0 - 2.5 m/s (3 - 8 ft/s) with values of 2.0 m/s (6.5 ft/s), or less,
recommended for obtaining long component life.
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Cony No.
The rate of elutriation of solids from the cornbustor is
increased with
increasing fluidising velocity and bed depth and with decreasing bed particle
size and freeboard height.
High elutriation rates may increase the proportion
of heat release in the freeboard and, of course,
increase the duty of the gas
clean-up system.
An initial bed of solids must be provided at start-up;
sand is commonly used.
graded silica
This start-up bed persists for the combustion of gas,
oil and coals that are washed, sized and relatively low in ash content.
During
the combustion of other solid fuels, and when sulphur retention is used, the
initial bed may be replaced by one of ash from the fuel and/or sulphur
retention sorbent.
The bed particle size range and mean size must be
appropriate for the bed material and fluidising velocity and for a given bed
material must be maintained constant within close limits so that fluidisation
conditions remain unchanged. The necessary bed maintenance measures range from
simple solids top-up, which replaces the fines lost by elutriation, to bed
cleaning procedures during the combustion of solid fuels for the removal and
separation of oversize material and the recycle of specification size solids.
In the heat transfer process the major part of the heat of combustion
is removed either as sensible heat in the combustion gases or by transfer to
in-bed cooling surfaces.
In the absence of in-bed cooling the furnace and
convection bank heat transfer designs will be similar to those for conventional
firing although the gas temperatures will be lower.
applications using in-bed cooling there is
However for boiler
an important interaction between
heat transfer and bed conditions because the heat transfer rate is a function
of the mean bed particle size as well as the in-bed tubing spacing and the tube
wall and bed temperatures.
It can be seen from the above that it is necessary to take a number of
inter-related variables into account to find suitable values for many of the
parameters of a fluidised bed combustor.
requirements
For some applications,
of the three processes of combustion,
the various
heat transfer
and
fluidisation may be best met by using combinations of fluidised beds each
operating under different conditions.
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op N
An examnple of this division would be the use of a burn-up bed to complete
the combustion of fine carbonaceous particles elutriated from
the main
combustion bed. See Section 4.5.2, Issue 002.
3.1.2
Basic combustor concepts
Fluidised bed combustor designs can generally be classified into
either of two basic concepts termed "deep bed` and "shallow bed"
combustor
designs. These terms are used to describe a complete design concept
as well as
an approximate bed depth. Shallow bed combustors generally have
static bed
depths in the range 0.1 - 0.5 m (4 - 20 in.) and use the above bed
feeding
method for solid fuels; see Section 4.2.2, Issue 002. Turndown
is usually
provided by the bed expansion method (see Section 3.10 below) in boiler
applications but can be supplemented by on/off operation or
the use of
compartmented beds for greater flexibility. The scale of operation
is
relatively low from about 30MW (1 x 108 Btu/h) downwards. Deep bed
combustors
generally have static bed depths above 0.5 m (20 in.) and use
the in-bed
feeding method for solid fuels; see Section 4.2.2, Issue 002. Turndown
is
generally provided by the use of compartmented beds although other methods
are
possible.
This
concept
is
considered suited to
the larger and more
sophisticated applications greater than 30MW (1 x 108 Btu/h).
Historically, deep bed combustors were developed first for the
combustion of unwashed coals in power generation applications. Since
much of
the mineral matter in unwashed coals is present as discrete ash
of a size
similar to that of the combustible coal it
was necessary to crush the coal so
that the ash particles formed on combustion were sized correctly to
act as the
bed solids. At the same time the in-bed feeding method (Section 4.2 and
15.1.2)
was developed to handle such crushed coals. Relatively deep beds,
0.5 - 1.0 m
(20 - 40 in.) static depth, were used to provide a relatively long
in-bed
contact time to give the maximum once through combustion efficiency.
An
empirical mathematical model of the combustion process in both the bed
and the
freeboard has been developed for deep bed combustors with in-bed
feeding; a
computer programme of the method is available. See Section 4.3, Issue
002 and
reference (3.1). The model is based on data obtained from a 1 m square
pilot
combustor at CRE. The deep bed concept is intended.primarily for large
scale
applications and an 18MW (60 x 106 Btu/h) cross type steam boiler
has been
converted successfully to fltiidised had rnmbustitnn using that concept (3.15).
*
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CoPy No.
-.-
The shallow bed concept has been developed more recently following the
discovery that relatively large particles of coal, up to 50mm (4 in.), could
be
dropped onto the surface of a comparatively shallow bed and could be dispersed
sufficiently rapidly throughout the bed to burn with high combustion
efficiencies similar to those obtained with the in-bed feeding method.
The
shallow bed concept uses beds with static depths in the range 0.1 0.5 m
(4 - 20 in.) with overbed feeding of solid fuels. See Section 4.4, Issue
002.
Turndown is obtained, in furnaces by varying the fluidising velocity
at
constant excess air level or, in boilers by, using the bed expansion
method
augmented for high degrees of turndown by on/off operation or the use
of
compartmented beds.
Since the bed pressure drop is reduced in the shallow bed
system and the overbed feeding of soid fuels is a simpler method than
the
in-bed feeding method, the shallow bed concept has been found to be better
suited for the simpler and smaller scale applications than the deep
bed
concept. Currently, the shallow bed concept has been used successfully
for
several hundred industrial installations in the 1 - 30MW (3 - 100
x 106
Btu/h) range. No mathematical model of the combustion process has
been
developed for the shallow bed concept with above bed feeding but an empirical
correlation for combustion efficiency is available based on extensive
industrial scale data (see Section 4.4.3, Issue 002). An empirical model
for
predicting the output from a boiler when the bed expansion method of turndown
is used is also available based on extensive industrial scale experience (3.3).
Both the shallow and deep bed combustion concepts are suitable for use
with fuels in any physical form. At similar operating conditions similar
once
through combustion efficiencies are obtained with either concept. For
coal
firing the once through efficiencies are generally sufficiently high for
the
smaller scale applications; see Section 3.6 below. The combustion efficiency
can be further increased by grit refiring or by the use of a separate
carbon
burn-up bed; see Section 4.5.
Although defined by the static bed depth the distinction between the
two concepts is essentially one of solid fuel feeding method and the degree
of
sophistication. The shallow bed concept offers the following advantages
in
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2
~~~~~~~~~~~~~~~~C
Section 3
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'
comparison with the deep bed concept:
a) Lower power consumption.
b) Little or no pretreatment of
coal fuels.
c) Can be accomodated in relatively
low height
combustion chambers
without excessive elutriation of
bed solids.
d) Has a similar combustion efficiency
but using a simpler solids
feeding system.
e) Has a reduced rate of erosion
of combustor walls and in-bed tube
banks.
However,
the shallow bed concept has the
following disadvantages:
i) It may be necessary to set a limit
on the large stone content of a
solid fuel or to use bed cleaning
systems and/or sloping plate or
sparge pipe air distributors.
ii)
A higher Ca/S ratio for a specified
sorbent and specified sulphur
retention may be required.
Generally, the shallow bed concept
is recommended for the smaller
scale or simpler application while
the deep bed concept is considered
more
suited to the larger scale and
more sophisticated type of application.
In this Section a procedure is outlined
by which a complete fluidised
bed combustor design may be built
up step by step for either design
concept.
Subsequent Sections give detailed
methods for individual design steps.
A bed
heat balance is calculated first
to establish values for the
basic operating
parameters. For a given fuel flow
and bed temperature the bed cross-sectional
area is then determined by the fluidising
velocity and excess air level while
the combustor height is dependent
on the bed depth and freeboard
height.
3.2
DesignProcedure
3.2.1
Preliminaryconsiderations
stages.
Process design calculations for a
fluidised combustor usually have
two
The first stage is a preliminary design
and costing study to select a
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Cony No::_
design concept.
In the second stage the detailed mechanical designs of the
components of the selected concept are considered. It should
be noted that the
information in this Manual relates only to the fluidised bed
combustor design
although in selecting a suitable design it will be necessary
to take into
account the performance of associated equipment.
is
For the purposes of the preliminary design a fluidised bed
combustor
assumed to comprise,
(a)
Air plenum, distributor and, for oil-firing,
(b)
Fluidised bed and freeboard containment
(c)
Immersed cooling surfaces (walls and tube banks) contacted by
the bed.
Bed management systems
(d)
the oil nozzle design
The main results of the procedure are a heat and mass
balance over the
combustor and an estimate of the following principal parameters.
(a)
The combustor cross-sectional area and the height needed
for the
expanded bed, splash zone and freeboard.
(b)
(c)
The immersed cooling heat transfer surface area.
The sorbent flowrate if sulphur retention is required.
(d)
The duty of the bed cleaning equipment.
(e)
The solids elutriation rate for the combustor and the duty
of gas
clean-up equipment.
(f)
The general arrangement of the air distributor and, for oil-firing
applications, the design of the oil feeding nozzles.
The subsequent physical arrangement of the combustor
and its
integration with any downstream equipment for convective heat
transfer or other
utilisation of the hot freeboard gases will be similar in principle
to that for
conventional combustion and is outside the scope of this
Manual.
The procedure should be used initially to establish the
design
parameters at MCR conditions and then again to check that
the values
Issue 002
Section 3
Page 8 of 56
Cony
No.l
established will allow operation at the desired turndown
conditions.
For the smaller, or more simple, application the parameter
values
determined using the procedure will be sufficiently accurate
for use as a basis
for the detailed mechanical designs of the second stage.
For the larger and
more complicated applications the resulting values may
be used as the first
approximation of an iterative process.
Further iterations may be needed to
give the required accuracy because of the interactions
of the fluidisation,
combustion, heat transfer and sulphur retention processes
on each other. For
all applications, however, it is desirable to select a
method of turndown and
to check that the combustor will operate correctly at
the maximum turndown
anticipated. This check is particularly important for oil
or multi-fuel fired
combustors.
3.2.2
StePwise procedure
3.2.2.1 Calculation of combustor cross-sectional area
The suggested procedure is shown diagrammatically in Figures
3.1 to
3.3. Further amplification and explanation of the procedure
at some of the
steps is given below. It is suggested that the procedure
should be used first
for a design at MCR conditions and then to check that
satisfactory operation
can be obtained at the conditions of maximum turndown.
A number of
computational aids to this stepwise procedure are available;
see Section 3.3
below.
In step 1, at the outset of the design, it is assumed
that the
following items of preliminary data are available.
1.
Combustor heat output, or for incineration applications,
the fuel
flowrate.
2.
Fuel specification;
calorific value, chemical analysis, relevant
physical properties. Note that the use of a high ash fuel
may affect
the design of the air distributor system and/or the
bed regrading
system.
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Paze 9 of 56
Cony No.
V
3.
The type of application.
This item assumes a general idea of how the
heat of combustion will be removed together with temperatures and
pressures on the coolant side of heat transfer surfaces.
If the
combustion air is preheated an estimate of the air inlet temperature
is needed.
If
the fuel is oil then 20% of the combustion air should
be regarded as
requirements.
an initial
estimate
of the oil transport
air
A gas exit temperature may be needed for the estimation
of stack gas heat losses.
4.
Ambient conditions for the proposed installation.
air and cooling water temperatures.
5.
Mean and maximum
Mean barometric pressure.
Environmental regulations at the proposed site for the emission of
sulphur dioxide, grit, nitrogen oxides and any other particular
compounds present in, or generated by the combustion of, the fuel.
This item will include the desired sulphur retention.
6.
The
requirements,
in
terms
of
allowable
contamination
and
compatibility of operating conditions, of equipment immediately
downstream, of the combustor.
7.
Desired range and frequency of load variation.
8.
Minimum acceptable combustion efficiency.
9.
For some applications it may be necessary to specify the required
combustor pressure drop.
10.
Fuel feeding method.
11.
The proposed start-up system. The combustor design may be affected by
the type and availability of the start-up fuel as well as by the
reactivity of the main fuel.
For some applications experimental tests (See section 17) may be
advisable to check that the proposed fuel does not introduce any features that
might make a fluidised bed unworkable.
Possible causes might be agglomeration
or sintering of the bed particles; the escape of gross amounts of fuel from the
bed unburnt,
or the generation of toxic materials during combustion.
The
efficiency of the additive for sulphur dioxide retention may also require
determination experimentally.
Issue 002
Section 3
Page 10 of 56
Copy No.
1.
Specify application
_J_
2.
'
o
parameters
Assume operating
paramelers
Yee
ei
s
| 3acombustion
Calcuale.
No_
13. Calculate or estimate combustionj
efficiency and heat release
j
1 in4abed and freeboard
4LEslimate proportion of in-bed
heat release
5. Calculate or estimate stoichiometric
air. Calculate air and flue gas
flows and gas composition
retention
Yes
No
aremeters. Calculate sorbent flow
7 Calculate bed heat balance
per unit moss of main fuel
Aulolhernmol,
Yes
\
No
So. Adjust votue of bed temperalure or excess air or air preheat
No
8b.Calcutlae auxiliary luel flow for
autothermal conditions
djusim ents
or balance
Bailer
9 Calculate immersed surface heat
I
transfer load to
gire balance
J
f
Furnace
Incineralor
or
YesYes
balance
iain fue
NooesNo
10. Specify heat output and overall
thermal efficiency.or calculate
heat loss)Calculate main fue flowJ
Jr1.
Calculate air and flue gas flows,
for whole unit
12.
Calculate bed cross-sectional
area
Figure 3.1
Design Procedure Part I: Bed Cross-Sectional Area
No
Issue 002
Section 3
Page 11 of 56
Copy No!>:.
2.
From Part I
Calculate Bed
Cros-Sectionolt orea
I3. Specity turndown method i
/4.
Estimate static bed depth
Calculate exPonded deph
15. Culc late splash heighlht
ecI
nc mcudna mecn
No
psizend
bed persists
iGa
gt7_
7I.Calculate oed,Utn
shandlin
p.|
es P
eqapument
Fuel
c17buCalculale
versize ash rate,
b d me
__r_a'te
Solid
bC'lute bed make-up rte
ach etutri,tion rate
I.I
and makeup roles
uc. e
Iclancoing equipmen
Ret r-
ombto
wall type
tactry
Water
cooted
P20.Calculate h.t. of.,surfac
|
area and heal load
c; .r~n prolee]ion
Assume PitCh
3. Calculate hi. coefl. tube
surtface and lenh
erosoon Protection
ube
~~TrYes <
_
___No
25.Select and design start-up i
equipmen
Figure 3.2
Design Procedure Part II: Combustor Height and Cooling
Issue 002
Section 3
Page 12 of 56
Cony No.
From Part H
25. Select and design startup equipment
26. Estimate distributor Ap
Design air nozzles
Gas
e
Fue
|
Oil
Sid
|29a Design gas distribution
grid
S27a.pecity feeding method I
Estimate No.o1 feed points
29.
Design feed syste
30.
I
Specify oil nozzle type
Calculate nozzle density
|29~~~~~~~~~~~c'
Yes
b.
Design
UNo
oil nozzles
N
Design sorbent teed
~~systte
31. Specify control philosophy
and instrumentation
Figure 3.3
Design Procedure Part El: Feedstreams, Start-up, Control
Issue 002
Section 3
Page 13 of 56
Cony No..
In step 2 it is necessary to select
probable values for the bed
temperature, fluidising velocity, bed
excess air level and the static bed
depth. Guidance on the selection of
appropriate values is given in later
sub-sections below.
An estimate of the air inlet temperature
(or the air preheat
temperature) is also required in step
2.
Step 3 and 4.
In multi-fuel applications the procedure
should be used
first for the main fuel and the resulting
design parameter values checked for
compatability with those for the alternative
fuel. A new section of the Design
Manual (Section 20) is currently being
compiled on Multi-fuel Firing and
reference should be made to the design
approach outlined there. A complete
model of the combustion process is only
available for coal firing with in-bed
feeding (Steps 3a and 4a).
Empirical correlations for the combustion
efficiency during coal firing with above
bed feeding are available; see Section
4.4.3 Issue 002. For oil and gas firing
the overall combustion efficiency is
generally greater than 99%.
The proportion of freeboard combustion
is normally significantly less
during oil-only or gas firing than
during coal firing. In multi-fuel
applications it is desirable to match
as far as possible the bed/freeboard heat
release ratio for oil firing to that
for the combustion of the main fuel.
A
suitable match can be obtained through
appropriate reductions in the oil
feeding nozzle density. See
Section 5 and also Section 3.7
below.
In Step 5 note that in the absence of
a complete fuel analysis the
minimum information needed to use the
procedure is the fuel gross calorific
value and an estimate of the stoichiometric
air ratio. An estimate of this
ratio may be read from Figure 3.4 with
sufficient accuracy for initial design
purposes. When the fuel is coal and
the stoichiometric air ratio is estimated
it will also be necessary to assume
a value for the combustion efficiency
as
the correlations used in Step 3 for the
latter require a knowledge of the coal
oxygen content.
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Section 3
Page 14 of 56
Cony No
Calculation of the sulphur sorbent flow in Step 6 will
need values for
the sulphur retention required, for the sorbent top
size and activity, for the
carbon dioxide concentration in the flue gases and
for the sulphur content of
the fuel. Guidance on estimating suitable values
is given in Section 3.9
below.
In Step 7 the bed heat balance is calculated. As
the bed must always
remain in heat balance during steady operation
the subsequent steps in the
procedure depend on the nature of both the,fuel
and the application. For
premium fuels and furnaces the excess air leviel required
to give a balance is
calculated (Step 8a) while for premium fuels and
boilers the area of cooling
surfaces in contact with the bed that is necessary
to provide a balance is
calculated (Step 9). For lower grade fuels that
are just autothermal it will
generally be possible to obtain a balance through adjustments
of the excess air
or bed temperature values (Step 8a). Non-autothermal
fuels can be burnt if
additional heat is provided by preheating the
air (Step 8a) or by the
combustion of an auxiliary premium fuel (Step
8b).
In the procedure of Figure 3.1 the heat balance
is calculated on the
basis of the unit mass of main fuel/unit time. This
method has the advantage
that any changes needed in the operating parameters
to give a heat balance can
be finalised before the need to calculate or estimate
the overall thermal
efficiency, which itself is dependent on the
operating parameters.
The bed heat balance includes the following items.
Heat INPUT
(a)
in-bed heat release (i.e. fuel calorific value
x combustion
efficiency x proportion of heat release occurring
within the
bed)*.
(b)
sensible heat in the inlet air.
(c) sensible heat in all the solids inlet streams.
sorbent plus fines recycle if employed).
*
(fuel plus
The form of item a) arises because the combustion
efficiency is normally
estimated for the combined bed and freeboard.
Issue 002
Section 3
Page 15 of 56
4j~~~~~~~~~~~~~~
~~~~~Cony
No ..
40400
u---'
Btu/lb
20000
40000
60000 (as fired basis)
H2
30
Stoichiometric
air ratio
SR
20
kg air/kg fuel
or
lb air/lb fuel
(as fuel basis)
Ct
10
.Note. for bituminous coals
SR 'r
x 10 - 4 x EH
For SI Units
For British Units
0
EH
-
50000
Gross
e = 3.25
8 = 7.56
Fuel type
·
x
$olid
liquid
i
gas
100000
150000
calorific value (as fired basis)
Figure 3.4
Variation of Stoichiometric Air Ratio with Calorific Value
9,.
Issue 002
Section 3
Page 16 of 56
Copy Nott,
Heat OUTPUT
(i) heat transfer to in-bed and containment wall cooling surfaces.
(ii) sensible plus latent heat in the bed off-gas.**
(iii) sensible heat needed to raise the temperature of recycled fines
to the bed temperature.
(iv) heat of reaction of sorbent.
(v) sensible heat of ash, sorbent, or other inert solids discharged
from the bed.
(vi) sensible heat of solids elutriated from the bed.
(vii) radiation to and from the bed surface.
(viii) heat transfer to and from the freeboard by solids splashing from
the bed surface and falling back.
Only items (i) and (ii)
are of major significance in removing the heat
release and the division of the heat load between them will depend primarily
on
the type of application,
the bed temperature,
the excess air level and the air
inlet (preheat) temperature. Table 3.1 shows an approximate split of the
heat
output for typical values of these operating parameters. With no air preheat
and an excess air level of 30% the proportion of the heat release that must
be
transferred to immersed cooling surface for a heat balance is in the range
of
40-45% according to the bed temperature and the extent of freeboard combustion.
For the less complicated applications and for design studies it will
only be necessary to include the major items in the bed heat balance. These
are
items a), b),
i) and ii)
above.
If
sulphur retention is required item v) may
also become significant.
For the larger and more complicated application it
will be necessary to include all the items in later iterations of calculation.
**
Note that item ii) includes the heat to vapourise any water introduced
with the fuel or formed during combustion.
Issue 002
Section 3
Page 17 of 56
Cony No..
Step 10.
Guidance on probable values of the overall thermal efficiency
is given in sub-section 3.6 below.
Step 12.
step.
The total bed cross-sectional area needed is given in this
The area may be arranged in a circular or rectangular containment and
sub-divided if required to suit the method of turndown and the needs of the
application.
In oil-firing applications when a tapered bed construction is used
(see Section 5) the cross-sectional area calcutated in this step is that at the
bed surface.
3.2.2.2 Bed maintenance and in-bed heat transfer
Steps 13-15.
Methods of turndown are outlined in section 3.10 and
guidance on the selection of the static bed depth is given in Section 9.2.3,
Issue 002 and of the height of the splash region in Section 3.5.4 or Section 10.
Step 16.
The characteristics needed in the bed solids of a fluidised
bed combustor are outlined in Section 7 and the properties of suitable solids
are given in Table 9.1.4 of Section 9, Issue 002.
Steps 17-19.
These steps in the procedure are concerned with the
provision and maintenance of a suitable bed of solids and identification of the
bed mean particle size.
The matter is discussed in Section 9.
Briefly, three
situations are possible;
(i) the start-up bed may remain and require only periodic make-up.
(ii) the start-up bed may remain but require periodic or continuous
cleaning to remove oversize material together with recycle and top-up
of fresh material.
(iii) the start-up bed may be replaced by one of coal ash or sulphur
retention sorbent.
Issue 002
Section 3
Page 18 of 56
Copy No.Y
Table 3.1
Approximate Fractions of the In-bed Heat Release Removed as Sensible Heat
in the Gases and Transfered to Immersed Cooling Surfaces
Bed Temperature 900 'C (1550 'F)
Excess air, %
0
30
50
1 100
165
Air preheat temperature 20 °C (68 OF)
Fraction of heat to gases
Fraction to immersed surfaces
0.44
0.56
0.55
0.45
0.61
0.39
0.78
0.22
1
0
Air preheat temperature 400 °C (750 OF)
raction of heat to gases
Fraction to immersed surfaces
Conditions
0.30
0.70
Fluidising velocity
Static bed depth
Fuel calorific value
Combustion efficiency
Freeboard combustion
0.36
0.64
0.40
0.60
2 m/s
0.15 m
29 000 kJ/kg
97 %
5 %
0.50
0.50
0.63
0.37
Issue 002
Section 3
Page 19 of 56
Cony No''>'
The measures necessary in each situation for maintaining a constant
bed
depth are outlined in Section 9.7, Issue 002.
In situation iii) the bed
composition and hence the mean particle size will alter during
the initial
period of operation until new equilibrium values are established.
The addition
of fresh solids to the bed will generally be required for maintaining
the design
bed depth but could also be needed to maintain a suitable bed
particle size
inventory.
Steps 20-24 are only required when cooling surfaces in contact with
the
bed and splash solids are used.
The heat transfer coefficient between bed
solids and surfaces is significantly less if the surfaces
are vertical walls as
opposed to near horizontal tubes and in each case the coefficients
reduce in the
splash zone above the bed surfaces to values characteristic for
convective heat
transfer from a dusty gas. Methods of allowing for these effects
are outlined
in section 3.8 below and section 10.
The design procedure suggested is to calculate the heat transfer
rates
from the bed and splash zone region to the water walls (if present)
first. For
some combustors the resultant heat transfer may be sufficient to
give a bed heat
balance.
In general, however, additional bed cooling will be required,
in the
form of a bank of tubes immersed in the bed, to provide the
necessary heat
balance. The bed to tube surface heat transfer coefficient is a
function of the
bed mean particle size,
adjacent tubes.
the tube surface temperature and the gap between
Approximate values for these latter two parameters may be
assumed initially and modified later. When selecting a tube
bank geometry
Section 19 should be consulted for the measures recommended
for minimising
erosion on immersed tubes.
The same section also contains recommendations for
minimising erosion on the water walls.
3.2.2.3 Feedstreams and control
Step 25.
The start-up method adopted may have a marked influence on
the type of air distributor that will be required and should,
therefore, be
specified before the air distributor design is undertaken in
step 26. Details
p
Issue 002
Section 3
Page 20 of 56
Cony No.
of start-up methods and their characteristics are given in Section 13,
Issue
002. Recommendations for the value of the air distributor pressure
drop
(needed in step 26) are given in Section 15, Issue 003 and the size and
number
of the holes in the standpipe nozzles should be chosen to give the required
pressure drop at MCR conditions. A method of calculating the nozzle pressure
drop is also included in Section 15, Issue 003.
Steps 27-30 are concerned with the design of the fuel and sorbent feed
systems. Recommendations are given in Sections 4,5, 6 and 7 as appropriate
for
the fuel used.
For oil firing a special procedure is given in Section 5.3 for
the calculation of the number and dimensions of the oil feed nozzles.
Step 31.
The control requirements of a fluidised bed combustor are
outlined in Section 13.6, Issue 002, together with a description of the control
philosophies available.
Choice of the most appropriate system will depend on
the requirements of the particular application. The specification of
any
control hardware or instrumentation will be similar to that for conventional
combustion systems and is outside the scope of this manual.
3.3
Computer Programs
A number of computational aids are available on topics concerned with
the whole fluidised bed combustion design procedure or with specific design
aspects.
For the complete procedure a program called "DESIGN" has been written
at CRE in APLJ language using an IBM 3070 mainframe computer (3. 1). The program
models the in-bed feeding of crushed coals in relatively deep beds in fluidised
bed combustors.
The combustion, bed maintenance,
retention processes are modelled.
Heat, mass and size fraction balances are
carried out over the bed and freeboard.
can be included if
desired.
heat transfer and sulphur
Allowance for cooling in the freeboard
Issue 002
Section 3
Page 21 of 56
Cony Non
A program called "FBCDES" has been written in IBM BASIC for desktop
microcomputers to carry out the preliminary design calculations for a fluidised
bed conbustor (3.2). The program first calculates a heat balance, as described
in Figure 3.1, for a given fuel to determine if
it
would burn autothermally.
If the bed is not in heat balance the operating parameters can be adjusted or
an amount of auxiliary fuel can be added to ensure a balance. If excess heat
is generated in the bed the necessary bed cooling may be provided either by
innersed surfaces or by an increased air flow.
The sulphur retention process
is also modelled.
The output data include a bed heat balance and estimates of
the bed cross-sectional area, the imnersed surface heat transfer area, the air,
fuel and sorbent flows and the exhaust gas composition.
Particular aspects of the design procedure are modelled in the
following programs which are all written in IBM BASIC language for desktop
microcomputers.
Program "BEDEXP7".
This program (3.3) carries out a heat balance for
a fluidised bed boiler and can be used to predict the steady state operating
parameters and heat output characteristics of different tube layouts when the
bed expansion method of turndown is
used.
A knowledge of the variation of
immersed surface heat transfer area and tube bank volume with bed depth is
required as a subroutine in program BEDEXP7.
This information can be
calculated for any given tube bank configuration using program "SUBTUBE" (3.4).
A program "HOTSTART" (3.5) has been developed to calculate the rating
of the auxiliary burner used with the hot gas start-up system. The program
uses the method of program BEDEXP7 to carry out a bed heat balance and estimate
the fuel requirements during the start-up process.
Program "PLENUW" (3.6).
The heat losses from the plenum chamber of a
fluidised bed combustor can be estimated using this program.
On input of the
plenum geometry and temperature, steady state conduction, convection and
radiation heat transfer calculations are performed to estimate the heat loss.
Issue 002
Section 3
Page 22 of 56
Cony No>
Program
"STANDPIP" (3.7).
This program can be used to calculate the
pressure drop characteristics of the standpipe nozzles in the fluidised bed
combustor air distributor.
The correlations used in the model are based on
experimental work which is also described in reference (3.7).
The calculation procedure for sulphur retention sorbents given in
Section 11 of this Manual has been programmed separately;
called "SUIRET".
the program is
This program can be used either to estimate the sorbent flow
required for a given sulphur retention or the retention obtainable for a given
sorbent flow. (3.8).
The above programs on aspects of the design procedure and program
FBCDES are all available on a disk file through Combustion Systems Ltd. (3.6).
Additional programs concerning specific aspects of fluidised bed
combustor design are program OHBURNER for the dynamic simulation of temperature
increases in a fluidised bed during start-up using an oil or gas fired overhead
burner, a program for the prediction of the behaviour of a fluidised bed
furnace/drier system in which a proportion of the gas leaving the unit is
recycled through the bed (3.9) and a program for the economic ranking of
alternative technical solutions of erosion on in-bed tubing (3.10).
These
three programs are written in APL or IBM Basic languages.
3.4
Minimum Fluidising and Terminal Velocities
The minimum fluidising velocity is
the parameter that determines the
onset of fluidisation in a bed of particles; it has a value characteristic of
the mean particle size of the bed.
The terminal velocity is that velocity at
which a single particle is prevented from falling downwards and is carried away
by the rising gas stream.
velocity is
Unlike the minimum fluidising velocity, the terminal
a function of the diameter of an individual particle.
Thus,
although the terminal velocity determines the upper limit to bubbling
Issue 002
Section 3
Page 23 of 56
Copy No.'
fluidisation conditions there is no sharp cut-off point.
Instead the
elutriation rate rises with increasing fluidising velocity until eventually
even the largest particles can no longer remain in the bed.
Several equations
are available for calculating the minimum
fluidising velocity, Umf, and the
terminal velocity, Ut.*
For Umf
equation 3.1, or the equivalent form, equation 3.2, is recommended.
These
equations are based on the Ergun pressure;drop relation. (3.11).
Good
agreement between calculated and experimental data has been found and these
equations represent the best method of extrapolation.
Armf
where
-
150 Remf (1-Emf)/E3mf + 1.75 Re2 mf/E3 mf
Armf -[ g(4 dp)3 pg (fpsg)1
/j 2 and Remf
..
[
...
...
.....
3.1
dp, Umf tg] /
An equivalent form is,
U2 mf + Umf[150 (1-Emf)J]/[1.75 pgdp]-[g(ps-pg) E3mfap]/[1.75 pg]
=
0
...
Note that in equations 3.1 and 3.2, dp is in metres if SI Units are
used and in feet when British Units are used.
For calculating the terminal velocity of spherical particles
equations 3.3 to 3.5 may be used.
Again note that as these equations are
dimensionless dp must be in metres if SI Units are used and in feet if
British Units are used.
* Symbols are defined in Section 1 of this Manual
3.2
Issue 002
Section 3
Page 24 of 56
Copy No.
Laminar (Stokes) region, where Ar t < 3.6
Art - 18 Ret
or U
Transition region,
where 3.6 < Ar t < 105
Art - 18 Re t + 2.7 Retl
68 7
... 3.3
pg)/(2
g
...
......
...
3.4
Turbulent (Newton) region, where Ar t >>105
Art -
(Ret) 2 /3
or U ttt a1
[ 3 g(ps-pg)/pg]e,~2/3 dx.3
~[lp'p.p,r/p,]*P
or13.5
where Art - [ £Pg(Ps1g)
dx3] /I2
...
5...
and Ret - (dx Ut pg)/
Equation 3.4 is not convenient for rapid calculations and can be
replaced by an approximate linear form.
Ut
[4
(pspg) 2
g 2 )/(225
p)
]
dx
(3.12).
.....
3.4a
The bed particles in fluidised combustors are only approximately
spherical.
Beds of coal ash, for example, have a sphericity of about 0.8. The
terminal velocities of non-spherical particles are always less than those for
spheres and the reduction is greatest in the turbulent region. For calculating
the U t of non-spherical particles equations 3.2 and 3.3 may be used with
4d
x replacing
dx .
For the turbulent region equation 3.5a, taken from
reference (3.13), is suggested.
Ut
1.33 dx(pspg)g/(pg (5.31-4.88
Figures
3.5 and 3.6
)
for Art >,5 x 104.. 3.5a
illustrate the variation of Umf and U t with
pressure, temperature and particle size for typical ash particles of density
2100 kg/m 3 (130 lb/ft 3), Emf = 0.5 and a sphericity of 0.8, fluidised
Si
Issue 002
Section 3
Page 25 of 56
0
~~~~~~~~~~~~~~~~Copy
No.(.
with air.
3.5a.
The values have been calculated using equations 3.1 and 3.3, 3.4,
In figure 3.6 smoothed lines have been drawn at the junctions between
regions.
3.5
Effects of the Principal Operating Parameters
3.5.1
Fluidising velocity
The fluidising velocity is the key parameter to the design of
fluidised combustors.
If the fluidising velocity is increased in a
well-fluidised bed the following effects will occur.
(i) An increase in the proportion of gas bubbles in the bed.
(ii) Increased movement of bed solids and increased rates of
diffusion of fuel particles.
(iii) Increased rates of erosion of both the combustor walls and
in-bed tube banks.
(iv) An increased elutriation from the bed
(v) A greater rate of attrition and, for a fixed system geometry,
(vi) An increase in the oxygen supply.
(vii) Increased distributor and gas cleaning equipment pressure drop.
The operating value of the fluidising velocity must be a compromise in
a range between the minimum fluidising velocity, and the terminal free-fall
velocity corresponding to the mean bed particle size. Operation at minimum
fluidising condition is not feasible as there is
for the avoidance of hot spots.
insufficient solids movement
At all times the solids movement must. be
sufficient to disperse the fuel and obtain an even bed temperature and to carry
any oversize solids to bed drain ports at a sufficient rate.
describes the term "good" fluidising conditions.
This description
A method of checking whether
given operating conditions will result in "good" fluidisation is
Section 9.6, Issue 002.
given in
In general it is recommended that the fluidising
velocity should always be at least three times the minimum fluidising
velocity.
It follows that this recommendation will set the minimum allowable
Section 3
Issue 002
Page 26 of 56
Copy No!~
~ ~~0.04
0.02
0
0.08
F 0C
900
0
655
____
-
In,2
-6r
m
1MW
0
5P~~~~~~~~L
>0
,jf
-¶
r- I-F~
4F
E
20
1
6
o:,z~~~~
0
0.24
100 KN/m2 -1 atmUm
P
250C -750F
0
0.20
0.16
0.12
0
0
0
10
~
~
~
~~~~~~~~~~~
~
it
It
~~~~_____
30
00
Mean Particle
-
Size
-~~~~~n
10If,e~-m-1
0
00
40
60
pm~~~~~~~~~~~~~~~r
Effet o Presur and TeprtrEnMnmmFudsn
Veocty
Section 3
Issue 002
Page 27 of 56
Copy No>,
in.
dix
~~0.02
0
0.12
0.08
0.04
0.16
5
Er~~~~~~4
0.20
0.24
KN/m2 - 6jatmP~i
610
10
0~~~~
tlPj
16 KNjiIrnt16
5 &JBSliI4LSt4.
z. 20 ~~~~~~~~~~~~~~~~~~~~0
0~
~ ~ ~ ~~~~~~~~~~~~~~~~~~~~7
20
4~~~~~~~~~~~~~~~
'
-,
H15
10~~~~~~~~~~~~~~~~~~~~~~~~4
*
jjj
-
j
Patcl tize4
610 KN/m2
6atm
r
Figure 3 6~~~~~~~~~1
Velcit: Efectof Pessre ad Tmpertur
TermnalFreeFal
ft
30
Issue 002
Section 3
Paze 28 of 56
Copy No.
(
air flow under turndown conditions; the MCR flow may be two to three times
higher.
There is an economic incentive to increase the MCR value of the
fluidising velocity as much as possible as the combustor cross-sectional area
will be proportionally reduced.
However,
the effects (i) to (viii) listed
above are not all advantageous when the fluidising velocity is
final choice will be an economic compromise.
increasing fluidising velocity are as follows.
combustible material from the bed.
increased so the
The undesirable results of an
Firstly, the increasing loss of
The combustibles loss from the combustor as
a whole may remain unaffected, however,
if the freeboard is
to accommodate the increased combustion in it.
sufficiently large
Secondly, there will be an
increasing need for greater static bed depths to give the residence time
necessary both for combustion and sulphur retention.
result in an increased bed pressure drop.
This, in turn, will
Thirdly, the increased attrition and
elutriation may result in higher bed solids make-up requirements.
Finally, the
rates of erosion of both walls and tube banks will be increased.
Current recommendations for the fluidising velocity are:
for boilers - values in the range 1.0 - 2.5 m/s (3 - 8 ft/s) and
preferably below 2.0 m/s (6.6 ft/s).
The limiting factor is the need
to minimise erosion.
for furnaces - values up to 3 m/s (10 ft/s).
The limiting factor is
the need to avoid excessive entrainment.
The above recommendations are made on the assumption that the size of
the bed solids and the fuel can be chosen to match the fluidising velocity and
avoid excessive elutriation.
However, for one class of.fuels it will be
necessary to select the fluidising velocity to suit the fuel size.
In this
class belong those solid fuels that are only available as fine powders, e.g.
spent cracking catalyst.
If the fluidising velocity is not appropriate to the
particle size of such fuels, or if they are fed to an inert fluidised bed of
Issue 002
Section 3
Page 29 of 56
Cony Nob
larger sized particles, then the combustion efficiency in the bed may be low as
the material will be substantially elutriated from the bed before combustion is
completed. These remarks apply to dense phase fluidised combustors, the design
of which is
the subject of this Manual.
It is
possible to operate fluidised
beds under such conditions that the entire bed is elutriated, separated from
the gas stream and recycled (3.14).
Such equipment is referred to variously as
"dilute phase fluidised bed combustors",
"solids transport combustors" or "fast
fluidised bed combustors" and its design is
3.5.2
qutside the scope of this Manual.
Bed temDerature
The experimental rigs and prototype plant have been operated with mean
bed temperatures up to 9500C (1750°F).
Occasionally bed temperatures of
1000°C
(1830°F) have been
reached.
The
normal range of
temperatures is considered to be 750 - 9500C (1380 - 1750°F).
operating
The lowest
temperature for self ignition and sustained combustion is in the range
400-800°C (750-1470°F) depending on the fuel. See Table 13.2 of Section
13,
Issue 002.
At
efficiency is poor.
temperatures below 7500C (1380°F) the combustion
Significant quantities of carbon monoxide and volatiles
may escape from the bed unburnt and elutriated solids are comparatively rich in
carbon compared with those for conventional firing. The upper limit to the bed
temperature of 9500C (1750°F) is set for coal firing by the danger of
sintering the ash (possibly plus sulphur retention sorbent) in the bed. In
this connection it should be remembered that individual burning particles may
be up to 2000C (360°F) hotter than the bulk bed temperature. For purely
oil firing applications the upper limit to the bed temperature may be higher
than 9500C (1750°F) and determined by the choice of bed solids. However,
for multi-fuel firing applications the criteria for coal-firing will apply.
The retention of sulphur by sorbents is markedly affected by the bed
temperature during atmospheric pressure operation (see Figure 3.8 and Section
11) and the optimum temperature for both limestone and dolomite sorbents is in
the range 800 - 8700C (1470 - 1600°F).
Issue 002
Section 3
Page 30 of 56
Copy Not
The bed temperature will very often be varied over the normal range to
obtain load variation and designs must be checked at the extreme temperatures as
well as at the nominal mean value.
3.5.3
Excess air level
The operating excess air level will be determined by the type of
application. For furnaces and applications without immersed surface bed cooling
the excess air level will be that value necessary to maintain a bed heat
balance, which will be of the order of 150 - 200% (see Table 3.1) when premium
grade fuels are burnt.
When immersed cooling surfaces are present, as in boiler applications,
the bed heat balance is maintained by adjustment of the immersed surface area.
There is an economic incentive to reduce the excess air level as far as possible
to minimise the heat losses with the stack gases. The lower limit to the excess
air is governed by the combustion requirements.
During the combustion of solid
fuels the combustion efficiency begins to drop rapidly at excess air levels
below 20% (See Section 3.6) and an operating value of about 30% is recommended
for normal operation with solid fuels. However, if staged combustion is used
for the reduction of NOx emission levels then the bed itself is operated under
slightly sub-stoichiometric conditions and additional secondary air is admitted
to the freeboard to give an overall excess air level of around 30%.
concept is the subject of ongoing research.
This
For the combustion of oil or gaseous fuels in boilers the excess air
level can be reduced to 10% without any significant adverse effect on. the
combustion efficiency.
However, when these fuels are an alternative in
multi-fuel applications it is generally most convenient to design for the
combustion of all the alternative fuels at the higher excess air levels needed
for the solid fuel combustion.
Issue 002
Section 3
Page 31 of 56
Copy No.<
3.5.4
Bed depth and splash heifht
3.5.4.1 Static bed depth
There is a considerable variation in the static bed depth recommended
depending on the fuel burnt,
application.
the fuel feeding method and the type of
There is an economic incentive to use the minimum possible static
bed depth as the bed pressure drop, and hence the power for providing the air
supply, are directly proportional to the bed depth.
Also the rate of erosion
of any tube banks inmersed in the bed is reduced as the bed depth is decreased;
see Section 19,
For
Issue 001.
gaseous
fuels,
provided
they
are
fed in
a premixed or
pseudo-premixed form (see Section 6), low bed depths of the order of 50 mm (2
in.) are sufficient to obtain satisfactory combustion.
Additional depth may be
required to accommodate in-bed tube banks.
For the over bed feeding of washed coals and similar solid fuels
static bed depths in the range 0.15 - 0.2 m (6 - 8 in.) have been found
sufficient to give high combustion efficiencies.
These depths are also
generally sufficient to accommodate the in-bed tube banks needed when the bed
expansion method of turndown is used in boiler applications.
When crushed coals and similar fuels are fed using the in-bed feeding
method the static bed depth should be at least 0.5m (20 in.)
to achieve the
rates of fuel dispersion needed for high in-bed combustion efficiencies.
For the combustion of oil fuels in a purely oil-firing mode a minimum
static bed depth of 0.5m (20 in.) is recomnended to minimise both the number of
oil feed nozzles required and the proportion of freeboard combustion.
In
multi-fuel applications when a higher proportion of the heat release is needed
in the freeboard to match the characteristics of coal-firing the static bed
depth may be reduced down to 0.3 m (12 in.). If the static bed depth is less
than 0.56m (22in.) an adjustment to the number of oil nozzles needed (the oil
Issue 002
Section 3
Page 32 of 56
Copy No.
nozzle density) will also be required: see Section 5, Issue 002.
-
Oil firing
with static bed depths less than 0.3 m (12 in.) is not recommended.
When sulphur retention is desired any of the values of static bed
depth recommended above may need to be revised upwards according to the
reactivity of the sorbent and the degree of sulphur retention required.
For
example, Figure 3.10 shows that the Ca/S molar ratio needed to give a 50%
sulphur retention using a sorbent of average activity (Reactivity index of 40)
is reduced from 2.5 at a static bed depth of 0.15 m (6 in.) to 1.5
at a static
bed depth of 0.4 m (16 in.).
3.5.4.2 Exuanded bed deuth
Expanded bed depths depend on the type and mean particle size of the
bed solids and on the fluidising velocity.
Values of the expanded bed depth
can be calculated using the correlations given in Section 9, Issue 002 or they
may be estimated using Figure 9.2 of that section.
As an example; sand beds at
a fluidising velocity of 2 m/s (6.5 ft/s) and with mean particle sizes of 800
and 1200 u m (0.031 and 0.047 in.) have expanded bed depths 1.7 and 1.5 times
their respective static bed depths.
3.5.4.3 Snlash heig-ht
A splash zone exists above the notional expanded bed depth where the
bed solids concentration reduces from that typical of the expanded bed to that
typical of a dusty gas.
The top of the splash zone - the splash height
-
is
usually estimated as 1.5 times the expanded bed height.
3.5.5
Properties of bed solids
3.5.5.1 Materials
The properties of materials that are suitable for forming the initial
bed of a fluidised bed combustor are given in Table 9.1 of Section 9,
and Table 13.1 of Section 13, Issue 002.
Issue 002
Issue 002
Section 3
AW
Page 33 of 56
Cony No.
>
For gas and oil firing applications and for the combustion of washed
coals with less than 10% ash the initial bed will persist and become the bed
for normal operation.
When sorbents are added for sulphur retention and for
unwashed coals and many solid fuels the initial bed becomes replaced by one
composed of sorbent and/or fuel ash.
Details of such beds are also given in
Table 9.1.
The ability of coal ash to form a suitable bed depends on the amount
and properties of the ash formed.
No hard and fast rules can be given because
of the wide variations between different coals and Section 9.7, Issue 002
should be consulted for further details.
3.5.5.2 Particle size.
The bed of inert particles will normally be comprised of a range of
sizes.
Indeed a bed of uniformly sized particles does not fluidise well and
some fines are needed for good fluidisation.
A bed of particles of varying sizes can be characterised by a mean
particle size.
The mean most suited for fluidisation purposes is the surface
mean diameter,
*
dp, defined by
~~~~~~i=n
;
1
e
%
dp
~xii
...
e>z
*s
3.6
(di j)
i-1
where xij is the weight fraction between the size limits di and d.
The mean bed particle size and size range must suit the desired
fluidising velocity and particle density.
Suggested values of particle size
are given in Table 13.1 of Section 13, Issue 002 for the initial beds.
Issue 002
Section 3
Page 34 of 56
Cony No:'57
In normal operation the bed material tends to an equilibrium state
provided all feed and take-off streams remain constant. Fines in excess of the
equilibrium requirements will be quickly elutriated while the larger size
fractions remain until either they are gradually reduced in size by abrasion
and elutriated, or they are removed in the bed overflow stream. Variations in
feed streams are therefore only self correcting if excess under-sized material
is fed. All oversized solids that are fed must be removed otherwise they will
collect at the bottom of the bed and eventually cause defluidisation. When the
bed overflow stream (see Figure 9.1 of Section 9, Issue 002) is substantial,
usually during the combustion of high ash solid fuels or when sulphur retention
sorbents are added, the bed solids offtake ports should be sited to remove the
lower layers of the bed solids and prevent any build-up of oversize material.
When oversize materials
enter, but
the overflow stream
is small or
non-existent, then bed cleaning equipment will be necessary. The terml bed
cleaning means the removal, periodically or continuously, of a portion of the
bed solids followed by their separation into an oversize fraction, which is
rejected, and a within-specification fraction, which is returned to the bed.
Bed cleaning equipment is described in Section 9, Issue 002 along with methods
for calculating the size of the overflow stream that is necessary for effective
bed cleaning.
Thus the match between mean solids particle size and the fluidising
velocity is maintained during operation either through the attainment of a
suitable equilibrium particle size distribution or through the use of suitable
bed cleaning techniques. Fresh bed solids can also be added to maintain the
desired particle size distribution. A method of checking that the quality of
fluidisation of a given bed is suitable is given in Section 9, Issue 002. That
Section also shows how the separate effects of those factors that may alter the
bed size composition may be calculated so that the equilibrium particle size
distribution and mean particle size may be estimated.
Issue 002
Section 3
Page 35 of 56
~~~~~~~~~~~~~~~Copy
No. /
*
3.5.6
Coal feed narticle size
It may be necessary to regulate the top size of the coal feed during
coal-firing so that the inert materials formed during combustion can
be handled
by the overflow stream and/or the bed cleaning equipment.
The in-bed feeding method was developed for the combustion of unwashed
coals using a flat plate nozzle standpipe type, of distributor (see Section
15,
Issue 003). Unwashed coals require crushing If the combustor is
operating at
fluidising velocities in the recommended range and also so that they
can be
transported pneumatically.
Table 3.2 gives suggested values for the coal top
size at various fluidising velocities.
Table 3.2
Suggested Values of Unwashed Coal Feed Top Size
for Various Fluidising Velocities
Coal top size
in
mm
1/4
3/16
1/8
1/16
6.34
4.74
3.17
1.58
Fluidising velocity
pm
6340
4740
3170
1580
ft/s
9
6
4
2
-
14
9
6
4
m/s
2.7
1.8
1.2
0.6
-
4.3
2.7
1.8
1.2
Unwashed coals may contain inert materials of the same top size
as
the coal itself. For this reason the use of material with a top
size larger
than those quoted in Table 3.2 would require specialised designs
of
distributors and arrangements for the removal of ash in the overflow
stream.
See also Section 4 and 15, Issue 002.
Washed coals are generally available in the UK in a graded form
"singles", sized 25 - 12 mm or "doubles" sized 50 - 25 mm - or in an
-
Issue 002
Section 3
Page 36 of 56
Cony No.<
-
ungraded form "smalls", sized 25 - 0 mm.
All these forms are suitable for
firing without further treatment using the flat plate nozzle standpipe type of
distributor and the technique of above bed feeding to the surface of the bed.
By using this technique it
is possible to "float" comparatively large particles
of coal on the fluidised bed.
As the coals have been washed the ash formed
during combustion is very fine and is almost wholly elutriated.
3.5.7
Air and fuel distribution
The performance of a fluidised combustor can be profoundly affected
by the manner of distribution of air and fuel into the bed and it
extremely important to use appropriate methods.
is therefore
Any performance figures and
design recommendations given in the Manual are only valid if the distribution
methods recommended in Section 15, Issue 003, are adhered to.
3.5.7.1
Air distribution
The preferred distributor design is the nozzle standpipe type with
the holes or slots in the nozzles being horizontal to prevent the back flow of
solids from the bed.
The hole size and number of nozzles should be calculated
according to the recommendations of Section 15, Issue 003.
3.5.7.2
Solid fuel distribution
Alternative feeding methods are available.
For the method of in-bed
feeding the feed is crushed to a top size, see Table 3.2 above, and is
delivered by pneumatic transport to nozzles feeding to the base of the bed
through the air distribution plate.
Nozzle spacing must be as uniform as
possible over the combustor cross section and areas of up to 1.8 m2 (20
ft
2
) may be fed satisfactorily from each nozzle. (3.15).
In general the
greater the number of fuel nozzles, the higher the combustion efficiency.
In-
bed feeding is also reconmended if the fuel is in the form of a fine powder as
it will result in the highest proportion of in-bed combustion.
.
Issue 002
Section 3
Page 37 of 56
Copy No..,
The alternative method is termed overbed feeding where the fuel is
fed directly onto the bed surface via a chute or spreading device.
has been used extensively for feeding washed coals in
fines content.
This method
lump form with a low
The combustion efficiencies obtained using overbed feeding are
similar to those found with the in-bed feeding method.
3.5.7.3
Oil distribution
Special arrangements are required to overcome feeding difficulties
such as cracking of the oil on entry to the bed.
The oil is
fed at an
appropriate rate and temperature to injection nozzles situated near the base of
the bed at points where the bed solids movement is most rapid.
Methods for
calculating the number of oil feed nozzles and their detailed design are given
in Section 5, Issue 002.
3.5.7.4
Gas distribution
It is generally desirable that the heat release of the fuel occurs as
far as possible within the fluidised bed so that advantage can be taken of the
higher rates
of heat transfer
from the bed solids compared with the
conventional rates of convective heat transfer.
To obtain combustion within
the bed with gaseous fuels it is necessary that they are fed in a premixed or
pseudo-premixed form with air as described in Section 6.
3.6
3.6.1
Combustion and Unit Efficiencies
Combustion efficiency
The term combustion efficiency is used to describe the extent of
combustion in the bed and freeboard acting together as one unit.
efficiency of combustion may be affected by the following factors.
(i)
(ii)
(iii)
excess air
bed temperature
fluidising velocity
The
Issue 002
Section 3
Page 38 of 56
Copy No.;
(iv)
(v)
(vi)
bed depth
method of fuel feeding
air distribution
(vii)
properties of fuel
(viii)
operating pressure
The factors which have the largest effect are the excess air level,
the fuel properties, which include the reactivity of solid fuels, and the bed
temperature.
The fluidising velocity and static bed depth have very little
effect provided their values are in the ranges recommended in the previous
sections.
For all fuels it is important that the air distribution is
uniform
over the entire combustor cross-section; otherwise some localised regions may
operate at a low excess air level with adverse effects on the combustion
efficiency.
The combustion efficiency of gaseous fuels is virtually 100% when
they are fed with air in a premixed or pseudo-premixed form at excess air
levels above 10% and at bed temperatures in the normal operating range of
750-9500C (1380-17500 F).
The combustion efficiency of oil fuels is generally greater than 99%
provided the excess air level is above 10%.
The combustion losses are mainly
in the form of carbon monoxide and increase rapidly as the excess air level is
reduced below 10%; see Section 5, Issue 002.
S
The combustion efficiency of solid fuels is more affected by the
excess air level than is that of oil or gaseous fuels.
As an illustration the
variation in combustion efficiency of a bituminous coal is shown as a function
of excess air level in Figure 3.7 for both the in-bed and above bed feeding
methods.
At high excess air levels and a given bed temperature similar
combustion efficiencies are given by either feeding method but the above bed
method is somewhat more sensitive to changes in the excess air level.
Combustion losses through unburnt volatiles or carbon monoxide are
normally negligible.
The main loss occurs through the elutriation of fine
Issue 002
Section 3
Page 39 of 56
Copy No,
c
E<~~
o
o
-4~~~
oE~a
0 °
am I
a
-
Cn
I
O
I
(
m._
i
o0
o
c
o
T-
a1
-W
30
C-
9~~~~~~~~~~~~.
(Bituminous
Oxygen
cols,
10%
content
W/w dmmf)
.
Issue 002
Section 3
Page 40 of 56
Si
Copy No.
carbonaceous material from the freeboard.
No significant correlation has been
found between the combustion loss and either the proportion of fines or the
mineral content in the feed. However a limited amount of evidence suggests
that above average losses may occur with coals having both a high mineral
content and a high proportion of fines in the feed.
The overall combustion efficiency for once through operation can be
improved by the recycle of fines back to the main bed. To obtain very high
overall efficiencies for atmospheric pressure operation, however, it will
probable be necessary to refire the carbon fines in a separate burn-up bed.
For the in-bed feeding method, and providing the excess air level is
above 15%, the changes in the combustion efficiency caused by any one variable
are only a few percentage points in the areas of interest. The effect of bed
temperature is similar to that shown for above bed feeding in Figure 3.7. The
effect of bed depth is very slight in the range 0.6-1.4 m (2-4 ft) static
depth.
The effect of fluidising velocity is also sligit.
At a bed temperature
of 850 C (1560°F) and an excess air level of 13% the combustion efficiency
was found to increase from 93 to 96% as the fluidising velocity decreased from
0
2 to 0.5 m/s (6.5 to 1.6 ft/s) (3.16).
Table 3.3 suggests values of overall combustion efficiency that
*
should be attainable by various modes of operation at atmospheric pressure.
Table 3.3
Suggested Values of Overall Combustion Efficiency
In-bed coal feeding.
Mode of Operation
Atmospheric pressure operation.
Overall Combustion Efficiency
Once through
92 - 94
Once through + recycle
94 - 97
Once through + recycle +
carbon burn-up bed
> 97
Issue 002
Section 3
PaEe 41 of 56
ConY NoA..t
No experimental data is available for the influence of coal
reactivity on combustion efficiency with in-bed feeding but trends similar to
those found with above bed feeding would be anticipated.
See Figure
3.8 and
Section 4, Issue 002.
For the above bed feeding method Figure 3.7 shows that it is
desirable to operate at an excess air level over 30% (or even above 40% at low
bed temperatures) to obtain high values of combustion efficiency. Reduction of
the bed temperature through the normal operating range decreases the combustion
efficiency by about two percentage points!. (Figure 3.7).
The fuel reactivity is the other principal parameter affecting the
combustion efficiency during above bed feeding. For coals the oxygen content
has been found to characterise their reactivity and Figure 3.8 shows how the
combustion efficiency is reduced as the coal becomes less reactive, other
conditions being kept constant.
Provided the fluidising velocity is below 3 m/s (9.8 ft/s) the
combustion efficiency is not significantly affected by variations of the
fluidising velocity. A gain of one percentage point may be obtained if the
fluidising velocity is reduced over the range 3.0-1.7 m/s (9.8-5.6 ft/s).
However, as the fluidising velocity is increased above 3 m/s (9.8 ft/s) the
elutriation rates of solids increase sharply and the combustion efficiency can
be reduced significantly (3.17). The extent of the reduction will depend on
the freeboard design.
Variations in the bed depth have little effect on the combustion
efficiency provided the beds meet the recommendations for minimum depth given
in Section 3.5.4 above.
Correlations for predicting the combustion efficiency for above bed
feeding are available; see Section 4.4.3, Issue 002.
Issue 002
Section 3
Page 42 of 56
Copy No
-
100
Anthrocites
Bituminous
Combustion
efficiency
ed temperature
Excess air
9000 C (1650 0 F)
30%
Static bed depth
150mm(6On.)
Fluidising
80
0
2
Coaol
4
oxygen
6
8
content % w/w
Figure
10
velocity
2m/s (6.5ft/s)
I
12
dmmf basis
3.8
Effect of Coal Reactivity on Combustion Efficierncy
for Overbed Feeding
9
Issue 002
Section 3
Page 43 of 56
Cop
3.6.2
No.
Overall thermal efficiency
No firm values can be quoted for the overall thermal efficiency as it
depends on the design of the heat recovery system downstream of the combustor,
which is outside the scope of this Manual.
For fluidised bed boilers of similar design the overall thermal
efficiency will be mainly influenced by the combustion efficiency.
Results
from a number of industrial installations less than 10 MW (3.4 x 107 Btu/h)
in heat output suggest that the following equation may be used for estimating
purposes.
not
-
0.84
x1n
...
...
...
...
...
3.7
A slightly larger constant term in equation 3.7 may be appropriate
for larger scale applications.
3.7
Pattern of Heat Release
A knowledge of the proportion of the fuel heat release that occurs in
the bed is required for calculating the bed heat balance.
For the combustion
of crushed coals using in-bed feeding methods a model of the combustion process
is available that allows the heat release in the bed and freeboard to be
calculated separately.
See Section 4, Issue 002.
For other fuels and feeding
methods estimates of the proportion of freeboard heat release are available
based on industrial operating experience.
For the combustion of washed, graded coals (UK singles or doubles)
using above bed feeding methods it is estimated that 5% of the heat release
occurs in the freeboard.
If such coals are ungraded and contain all the fines
(UK smalls) then up to 10% of the heat release may occur in the freeboard.
For oil fuels the proportion of the heat release occurring in the
freeboard depends on the oil grade, the excess air level and the oil feed
nozzle density.
With a nozzle density of 4 per square metre (0.37 per
Issue 002
Section 3
Page 44 of 56
Cony No.\
/
square foot), which is the minimum value recommended
for multi-fuel firing
applications with an untapered bed, the proportion
of heat release in the
freeboard can be estimated from Figure 5.23
of Section 5, Issue 002, for
various oil grades and excess air levels. The
freeboard heat release is
generally in the range 2-10% which is similar to
that for coal firing. As the
otk-nozzle-density-iwsiicreased1abive -this inimnum
value the proportion of heat
release in the freeboard will reduce but no precise
estimates of the reduction
are currently available.
For gas firing with the gas fed in a premixed or
pseudo-premixed form
(See Section 6) the proportion of heat release in
the freeboard is estimated to
be less than 5% of the total heat release.
3.8
Heat Transfer from the Bed
The only situation in fluidised bed combustors
where the heat
transfer rates differ from those found during conventional
combustion is where
surfaces are contacted by the bed solids.
In all other respects heat transfer
rates are similar to those for conventional convective
heat transfer and design
considerations for them are outside the scope
of the Manual.
The heat transfer coefficient to immersed surfaces
is made up of a
convective and a radiative component which are
considered to be additive. The
convective component is affected by the bed particle
size, the temperature of
the surface through which heat is being transferred
and the gaps between
adjacent surfaces allowing flow of the fluidised
solids. The convective
component is generally in the range
200-500 J/m2s
(35-90
5o
Btu/ft 2 hOF). The radiative component similarly
is in the range 100-200
J/m2SoC (17-35 Btu/ft2 hoF).
When the heat transferred to the in-bed surface
is removed by boiling
water and if the tube-to-water film heat transfer
coefficient is assumed to be
850 J/rn2sC (150 Btu/ft 2 hOF), and the conduction
through the tube wall
is neglected, a value of 250 J/m2 soC (45 Btu/ft 2
hOF) for the overall
Issue 002
Section 3
Page 45 of 56
!
~~~~~~~~~~~~~~~~~~~Cony
No. a'~
heat transfer coefficient is suggested for estimating purposes or as the
first
iteration of the design calculations.
For more detailed design Section 10 should be consulted. It is
necessary to differentiate between surfaces, such as tube banks, lying
wholly
or partially within the bed and those, such as containment walls, that
are
contacted only on one side by the bed.
The splash zone above the bed also
requires separate consideration.
Equations for calculating both heat transfer
components in various bed geometries are given in Section 10. The equations
for the convective component of heat transfer were developed from data obtained
with relatively deep beds of coal ash and corrections may be necessary
when
they are applied to shallow bed combustion (bed less than 400 mm (16
in)).
Further details are given in reference (3.3) and Section 10.
3.9
Sulphur Retention
The sulphur retention process involves the addition of a solid
sorbent which retains the sulphur evolved during combustion in the form
of a
solid sulphate for subsequent disposal.
The sorbent is a source of calcium
carbonate which decomposes to calcium oxide at the bed temperature.
Graded
limestones, or less commonly dolomites, are the usual sorbents for atmospheric
pressure operation.
The sorbent flow depends on the chosen calcium/sulphur molar ratio
which varies with the degree of sulphur retention required. Figure 3.9
shows
how the retention varies with the Ca/S molar ratio for a typical limestone
sorbent used in a boiler application.
Other factors which affect the molar
ratio needed for a given retention are the bed temperature, the sorbent
top
size , the sorbent activity and the bed depth and fluidising velocity
which
together determine the gas residence time in the bed. The manner in which
each
of these factors, when varied on its own with all other values kept constant,
is
illustrated in Figure 3.10 for a limestone sorbent.
The selected set of
conditions chosen is that for average operation and each individual
graph
illustrates the calculated change in Ca/S molar ratio for variation
of one
particular parameter.
Issue 002
Section 3
Pave 46 of 56
Copy No "
100
90
80
70
/60
~Typical
~
60 ~
Sulphur
retention
Limestone Sorbent
%
8500°C
2 ms
3000pm
Static
0.3 m
so50
Bed temperature
Fluidising velocity
Sorbent top size
bed depth
Sorbent
40
activity
Excess air
Flue gas C02
/
40
30%
13%
30
20
0'
1
2
3
Ca/S
4
5
molar ratio
6
7
8
Figure 3. 9
Variation of Sulphur Retention with Ca/S Molar Ratio
S~~~Vaito
Issue 002
Section 3
Page 47 of 56
Copy NNo.A
Typical Limestone
1400
1500
6
1600
1700
Performance
OF
0
5
10 ft/s
6
5
4
4
3
2
2
0
0
750
0
800
850
900
Bed temperature °C
950
0.04
0.16 in.
0.12
0.08
1
2
3
Fluidising velocity m/s
0
12
6
18
in.
4
0
-2
. 1
.
0
1000
2000
3000
4000
Sorbent top size microns
6
5000 0
Selected conditions
EBed temperature 850°C
Fluidising velocity 2 m/s
Sorbent top size
3000pm
Static bed depth
0.3m
Sorbent. activity
40
o Selected condition
for Limestone
sorbent
5
4
0.2
0.4.
0.6
Static bed depth m
3
Excess
air
30%
Flue gas C02
10
00X
13%
3
10
20
30
40
Sorbent
SO
60
70
activity
Figure 3.10
Effect of Operating
50%
m,
Parameters on Ca /S
Sulphur Retention
Ratio for
Issue 002
Section 3
Page 48 of 56
Copy No.W
_
When the bed temperature is in the middle of the normal operating
range about 850 0 C (1560°F) the next most important factors are the bed depth
and the sorbent reactivity.
found to be very variable.
The activity of limestones as sorbents has been
It depends on the stone pore size and pore size
distribution, on its hardness and on other factors which are not yet completely
understood.
A laboratory test rig is available for the determination of sorbent
reactivity.
See Section 17.
For estimating purposes a value of 40 is suggested
for the activity of an average limestone.
3.10
3.10.1
See also Sections 11 and 17.
Turndown
General considerations
Any plant of which a fluidised combustor is a component will almost
certainly be required to operate over a range of loads.
The requirements of
load regulation must be considered at the inception of any proposal because they
may have far reaching effects on the design.
For example, in some applications
low turndown ratios (< 1.5:1) may be easily obtainable with a single bed
combustor but high ratios (> 4:1) can only be obtained using compartmented or
multiple bed combustors.
The combustion rate in a fluidised combustor is reduced by feeding
less fuel, which may be carried out in four different ways.
(a) The fuel rate alone is reduced.
The air rate stays the same and
consequently the excess air increases.
(b) Both the fuel and air rates are reduced together to maintain a
constant excess air value.
(c) A portion of the total bed cross-sectional area is shut down and
normal operating conditions maintained in the remainder.
(d) The combustor is operated in an on/off mode with periodic bed
slumping.
In (a) and (b) above the combustion intensity is reduced.
In c) it remains the
same in the active portion of the combustor providing the combustion efficiency
Issue 002
Section 3
Page 49 of 56
Cony No.
remains constant.
In d) the combustion intensity remains the same when
operating in the "on" mode.
Whatever method of load control is adopted the heat balance of the
fluidised bed for all applications must be such that the bed temperature is
maintained within the normal operating limits of 750-950°C (1380-1750°F).
Also the fluidising velocity must be such as to give "good" fluidising
conditions at all loads.
A definition of 'good" in this context is given in
Section 9.6, Issue 002.
The use of in-bed tubing in boiler applications imposes further
constraints on the fluidised combustor design.
The constraints arise because
heat is transferred predominantly by convection in fluidised combustion.
In
this respect fluidised combustors are different from conventionally-fired
equipment where heat is transferred predominantly by radiation and where the
transfer rate, being proportional to the fourth power of the absolute
temperature, may undergo large changes for quite small changes in combustion
temperature.
In fluidised combustion the effect of reducing the bed
temperature depends on the coolant cycle conditions.
For steam raising
applications the temperature difference between the bed and the boiling water
is large (of the order of 6000C (1080%F)).
Therefore only a limited change
in the rate of heat transfer to the in-bed surfaces can be achieved by
variation of the bed temperature.
Reduction of the bed temperature through the
entire operating range will give a load reduction of only 20-40%; the exact
value will depend on the proportion of heat removed by in-bed surfaces and that
as sensible heat in the off-gases.
For air heating cycles the coolant
temperatures are higher and the bed-to-coolant temperature differences
correspondingly lower.
Changes in the bed temperature therefore cause much
larger changes in the heat transfer rate. Nevertheless the effect of bed
temperature on in-bed heat transfer is always limited and must be carefully
considered during load reduction.
At any given bed temperature the balance
between the heat transferred to in-bed surfaces and that transferred in other
ways must remain unaltered at all combustion rates. If this balance is
disturbed the bed will either overheat or cool off.
Issue 002
Section 3
Page 50 of 56
Copy No.
Unless on/off operation is used it follows, therefore, that when
significant load reductions are required the heat transfer to the in-bed tubing
must be reduced by changing some other parameter than the bed temperature.
The
following alternative methods have been examined.
(a) Shut-down of a part of the combustor cross-section by using a
compartmented bed or multiple beds.
(b)
Exposure of a portion of the in-bed tubing to the splash zone,
where heat transfer rates are lower, by reducing the bed depth.
(c)
Separation of the combustion and heat transfer functions into
two separate beds with circulation of hot bed material between
them.
Shut down of a part of the combustor, method a), reduces the total
in-bed heat transfer proportionately.
It also reduces the air flow
proportionately and allows the fluidising velocity to be maintained in the
active part of the bed with a smaller total air flow.
One difficulty with this
method is that on shutting down one part of the bed, the bed temperature in the
remaining active part must be instantaneously increased from a minimum to a
maximum value to give a smooth variation of load between steps.
The choice
between a single compartmented bed or multiple beds will depend
largely on
economics and heat transfer surface layout. Compartmented beds if shallow (lm
(3 ft) or less) may require merely a zoned air supply to the distributor but a
physical separation in the bed between different compartments is always
recommended. See Section 13, Issue 002.
Method b) is termed the bed expansion method; the principle of
operation is illustrated in Figure 3.11.
The heat transfer rates to tubes are
a power of ten less in the freeboard than they are in the expanded bed; the
transition occurs gradually throughout the splash zone. If the tube bank is
situated just below the expanded bed surface at MCR conditions (as in
Figure 3.11 (a)) the natural contraction of the bed as the fluidising velocity,
and hence load, is reduced will cause the tube bank to be exposed to splash
zone conditions (as in Figure 3.11 (b)).
If the resulting reduction in
heat
transfer to the tube bank matches the change in load the bed will remain
Issue 002
Section 3
Page 51 of 56
Copy No. >.:
)'
2
16. 5)
Fluidising
velocity
m/s (ft/s)
1 13.3)
Bed
depth
m.
Bed
depth
in.
0.5
18
0.4
0.Figure
3.11
The Bed Expansion Method of Turndown -Bed
Variations for 14/25 Mesh Sand
0~~~~~~~~~iue31
Level
Issue 002
Section 3
Page 52 of 56
Copy No.,
in heat balance at constant bed temperature and excess air level.
The same
effect could also be achieved by the physical removal of bed solids.
The bed expansion method has been used extensively on industrial units
up to 30 NW
(1 x 108 Btu/h) heat output.
It
is
best suited to those
applications where the tube bank consists of only one or two rows of tubes.
Method (c) is more complicated than the other methods.
It has not been
developed although cold model studies (3.20) have demonstrated its technical
feasibility.
3.10.2
Choice of method
The choice of the method of load variation depends on the type of
application and the level of turndown required.
Furnaces.
combustors without immersed surfaces.
For coal-fired combustors it
fuel and air rates while still
is
usually possible to halve both the
maintaining "good" fluidising conditions. This
variation, in conjunction with a reduction of the bed temperature through the
normal operating range, can give a turndown of up to 2.5:1.
For oil-firing and
multi-fuel applications the need to maintain the oil transport air flow at MCR
values at all times restricts the degree of turndown obtainable through air
flow reductions.
Air flow and bed-temperature reductions together can give a
turndown of up to 2:1.
If a greater degree of turndown is required then the use of multiple
or compartmented beds should be considered.
Issue 002
Section 3
Page 53 of 56
Cp
No.
Boilers. combustors with immersed surfaces.
For both oil and coal-firing applications the bed expansion method
can generally be used to obtain a turndown of up to 2:1.
For higher degrees of
turndown multiple bed operation can be combined with the bed expansion
method.
A 6:1 turndown (2:1 bed expansion plus 3:1 multiple beds) has been
demonstrated on a 30 MW (1 X 108 Btu/h) unit (3.19).
Turndown ratios greater than 2:1 can; also be obtained through on/off
operation.
This method is probably best suited for hot water boilers and
smaller scale applications because of the temperature cycling involved and is
only applicable if periods of zero output can be tolerated.
*3~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~
4i
Issue 002
Section 3
Pave 54 of 56
to
Copy No.,
3.11
3.1
References
Collop, D.J. & Merrick, D.
"Design manual program user manual".
Brit. Coal, CRE Report No. PADB/79/15, (April 1979).
3.2
Ellis, J.E.
"Program FBC DESIGN for IBM PC: A Microcomputer program
for the design of an atmospheric fluidised bed combustor" Combustion
Systems Ltd., Report, (1988).
3.3
Moodie, J., Stockwell, D.H., Vickers, M.A. & Blenkin.P. "BEDEXP7 - A
computer program to aid the design of fluidised bed boilers using a
procedure to predict the effect of in-bed tubing arrangement on
boiler output characteristics", Brit. Coal, CRE Report No PH8,
1987).
3.4
(Dec
See also version with program listing CSL Report, (1988)
Stockwell, D.H. "SUBTUBE - A computer program to calculate the heat
transfer surfaces and volumes of immersed tubes in fluidised bed
boilers" Brit. Coal. CRE Report No. P373.
3.5
Moodie, J., Stockwell, D.H., Vickers, M.A. & Loynes, D. "Fluidised
bed combustors - a revised mathematical model of hot gas start-up.".
Brit. Coal, CRE Report No. PH 27/8, (Dec 1987).
3.6
Combustion Systems Ltd.
"IBM Basic computer programs." Disk file
containing 7 microcomputer programs.
3.7
(July 1987).
Iles, A.E., Moodie, J., Stockwell, D.H. & Vickers, M.A.
development of
"The
a mathematical model to predict the pressure drops
through standpipe air distributor".
Section Report No 271 (Dec 1985).
Brit Coal, CRE Combustion
Issue 002
Section 3
Page 55 of 56
Copy No.!
3.8
Ellis, J.E.
"
)
A computer program for the calculation of sulphur
retention in fluidised bed combustors".
Combustion Systems Ltd.,
Report (1988).
3.9
Atkinson, P.A.
"Fluidised bed driers - a mathematical model to
investigate the effect of recycling the exhaust gases".
Brit. Coal,
CRE Report No PI 40, (July 1986).
3.10
Davidson, J.E. & Waye, A.D.
"An economic comparison of technical
solutions to tube wastage in AFBCS".
Brit Coal, CRE Technical Note
87/972/2, (March 1987).
3.11
Davidson, J.F. and Harrison, D., ed.
"Fluidisation", Academic Press,
London, (1971).
3.12
Kunii, D. & Levenspiel, 0. "Fluidisation Engineering" pub J. Wiley
(1969), p.76 .
3.13
Pettyjohn, E.S. & Christianson, E.B. Chem. Eng. Progr. 44, 157,
(1948).
3.14
Engstrom, F., Ahlstrom, A. & Sahagain, J. "Circulating fluidised bed
technology", ed. P.Basu. Proc. 1st. Int. Conf. on CFB Technology,
Halifax, Nova Scotia, P.309-316,
3.15
"Fluidised Bed Combustion - Industrial Applications:
and 2".
3.16
(Nov 1985).
Report No 1/77/762A.
(9 March 1977).
NCB, CRE "Run reports for the 3ft combustor".
58, 64, 68, (1970).
Test series 1
Report Nos. 54, 57,
Issue 002
Section 3
Pave 56 of 56
Copy No.
3.17
Fisher, M.J.,
3.18
NCB, CRE "Run reports for the 3ft combustor".
Ford, N.W.J. & Robinson, A.W.
"Report of the
development of the high velocity Thermraiser fluidised bed boiler at
CRE". British Coal, CRE, Report ID 25, (Sept 1985).
Report Nos. 19, 21,
22, 24, 25, 31, (1969).
3.19
Fisher, M.J.
"The application of fluidised combustion of coal to a
low cost industrial water tube boilir integrated in a combined heat
and power system". British Coal, CRE. Contract No. EE 129/81
Consolidated Technical Report
for
1/11/81
to
31/3/86,
(Jan
1988).
3.20
BCURA Reports Nos. FCP 13, (May 1970) and FCP 21, (July 1971).