8 COMBUSTION UNDER PRESSURE - University of Bradford · 8 COMBUSTION UNDER PRESSURE ... 8.3.7.1 Bed...

1 Jul_y '19T Section 8

Section Inde

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8 COMBUSTION UNDER PRESSURE

8.1 The Advantages of Pressurised Combustion

8.2 The Use of a Gas Turbine with a Fluidised Combustor

8.2.1 Feasibility

8.2.2 Potential applications8.2.2.1 Compact boiler8.2.2.2 Power generation

8.2.2.2.1 Pure gas turbine cycles8.2.2.2.2 Mixed cycles

8.2.2.3 District heating8.2.2.4 Endothermic chemical reactions

8.2.3 Design implications

8.3 Effects of Pressure on Bed Performance

8.3.1 Fluidisation quality

8.3.2 Combustion efficiency

8.3.3 Elutriation

8.3.4 In-bed heat transfer

8.3.5 Sulphur retention

8.3.6 NOx emissions

8.3.7 Turndown8.3.7.1 Bed circulation8.3.7.2 Airslide transfer

8.4 References

1 July 1979 Section 8

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8 COMBUSTION UNDER PRESSURE

In this Section the design of fluidised combustors operating at above

atmospheric pressures (pressurised combustion) is discussed. Pressurised

combustion implies operation of a gas turbine in conjunction with a fluidised

combustor so that the pressure energy in the off-gases can be recovered. Once

a choice of a gas turbine that is compatible with fluidised combustors has been

made the requirements of the chosen turbine will dominate the selection of many

of the parameters of the fluidised combustor design and operation. For this

reason the various effects that an increase of operating pressure can have on

fluidised combustor design and performance are described only in general terms

in this Section. The effects of operating pressure are also described in

individual Sections of this Manual as appropriate.

A brief discussion of the advantages and likely applications of

pressurised fluidised combustion is included to put the subject into perspective.

Howeyer, it must be emphasised that each application will need a detailed design

study involving close collaboration between the turbine manufacturer, the

combustor manufacturer, and CSL and its partners.

8.1 The Advantages of Pressurised Operation

Pressurised fluidised combustion offers two main advantages over

atmospheric pressure fluidised combustion. Firstly, it is possible to reduce

* the size of the combustor for an equivalent heat output. Secondly, at least a

part of the energy of combustion must be recovered via a gas turbine which opens

up possibilities for the use of more efficient themodynamic cycles for power

generation than those employed with conventional conbustors.

The smaller size of pressurised combustors arises in the following way.

As the pressure increases the air volume decreases proportionately and the oxygen

partial pressure rises. If conditions of constant heat duty, fluidising

velocity and excess air are maintained the bed cross-sectional area must become

directly proportional to the reciprocal of the pressure, and the heat release

per unit bed area will increase directly with the pressure. This situation is

illustrated by Figure 8.1, which shows the calculated bed area needed per MW,

plotted as a function of pressure for various values of fluidising velocity.


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In order to keep the bed temperature constant with increasing operating pressure,

therefore, it is necessary to remove more heat by increasing the in-bed surfaces.

This can be achieved by reducing the in-bed tube spacing but a practical limit is

reached when the solids flow becomes restricted and the heat removal rates decrease

again. For further increases it is then necessary to increase the bed depth merely

to accommodate the required heat transfer surface. See also Sections 10.2 and

8.3.4. As an approximation, therefore, the bed, depth will increase in direct

proportion to the pressure. It follows then that the bed volume tends to remain

a constant, independent of the operating pressure. However, the bed is only one

part of a fluidised combustor; an entry zone for air distribution and a freeboard

zone are also required, and as their height does not vary with pressure their

combined volume will decrease with increasing pressure as the bed cross-sectional

0 area decreases. Since the air distribution zone and the freeboard together

account for about half of the total combustor height, it can be seen that the

overall combustor volume decreases significantly as the operating pressure

increases. A further consequence of this decrease in size, which is favourable

from the constructional point of view, is that the combustor heightjwidth ratio

increases with increasing pressure.

The reasoning outlined above is somewhat oversimplified. Factors such

as the size and location of steam headers and the arrangement of hot gas ducting

can also have an important effect on the overall size of the plant. Nevertheless,

considerable reduction in combustor size is anticipated through pressurised

operation.

.Several additional advantages can also be gained through pressurised

operation.

(a) With increasing pressure the fluidisation characteristics of a

fluidised bed tends to give smaller bubbles and better gas]solids

contacting. This, together with the deeper beds of tube banks

that are necessary results in higher combustion efficiencies.

Operation without recycle of elutriated material, even at the

higher efficiencies required in power station applications,

becomes possible. See 8.3 below.

(b) Pressurised operation can give a higher sulphur retention by

additives because of the deeper beds needed and also because the

Section 81 July 1979

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bt Pressure - k N/m2 abs Conv No.

0- 200 400 600 800 1000 1200 1400 1600

MW

I !i~.iiii,7, i: :;;; ' L w F 1111

0.1 - 0 .....=0- _'- ar....__.-Exces air 20%

i J '" ~, 11i |~ ~ |? - ~ ~ ;Overail cycle efficiency 3R8%riSi i Bed

· g i | 1! 11 1 91t il ! j | 1 1, ... I. .1 uiising vl ocity

.0 2 4 6 8 1 0il- 12 2 M 14 1,

Figure 8.1Effect of Pressure on Combustor Crss-sectioinal Area

!) .Effect of Pressure on Combustor Cross-serxional Area~~~~~~~~iI


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increase in the carbon dioxide partial pressure as the operating

pressure increases causes less additive decrepitation. See Section 11.

This effect is particularly noticeable when dolomites are used as

additives.

(c) NO, emissions are lower for pressurised operation than for atmospheric

pressure operation. See Section 12.

8.2 Use of a Gas Turbine with a Fluidised Combustor

8.2.1 Feasibility

Details of the experimental pressurised combustors used have alreadybeen outlined in Table 2.1 of Section 2. All these units employed diesel orelectrically driven air compressors for convenience and flexibility. Nofluidised combustor directly linked to a gas turbine driven air compressor hasyet been operated by CSL or its partners. However, a unit has been operated

in the United States (8.29) and has confirmed that such an operation is practicable.

The feasibility of using currently available gas turbines in conjunctionwith pressurised fluidised combustors will depend, primarily, on the levels of,

(a) dust emission, and

(b) alkali and trace element emission

in the exhaust gases from the fluidised combustor being sufficiently low to avoidany detrimental effects on turbine blading. Encouraging results have been foundon the experimental units from an extensive series of measurements of theseemissions and their effect on samples of static turbine blading. C8.a to 8.8, 8.27).

The gas cleaning equipment of the experimental units are primary, secondaryand tertiary cyclones and other inertia type separators. With this equipment it

has been found possible to reduce the dust concentration in the combustor off-gasto levels below those specified by some national environmental regulations.

(8.15). See Section 18. It is beyond the scope of this Manual to give a firmfigure for the maximum allowable dust loading of the combustor off-gas as this

will depend on the choice of gas turbine and on the requirements of the individualapplication. However, it appears likely that dust emission can be reduced to

K.


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acceptable limits although cleaning equipment that is more efficient than that of

the experimental units may well be required to meet both gas turbine and environ-

mental specifications.

The situation as regards the emission of alkali and trace elements is

also favourable because of the comparatively low bed temperatures used in fluidised

combustion compared with conventional power generation plant. The bed temperatures

currently recommended lie in the range 750 - 950 °C (1380 - 1750 OF). For coal-

fired combustors at these temperatures the alkali components of the ash, which

are known to play a part in fouling and corrosion of heat transfer surfaces and

turbine blading, have vapour pressures that are several orders of magnitude lower

than at the temperatures in conventional combustion systems. As a result it is

* found that there is a marked reduction in the alkali concentrations in the off-gas

from fluidised combustors. For oil-fired fluidised combustors similar reductions

in the off-gas concentrations of alkali compounds, compounds of heavy metals like

vanadium, and sulphur compounds (provided a sorbent is added) are found. For

coal firing typical experimentally measured values of the off-gas concentration of

sodium lie in the range 0.2 - 2.0 ppm v/v; similarly, values for potassium lie

in the range 0.1 - 1.0 ppm v/v. Although these alkali concentrations are approx-

imately a power of ten lower than those found in conventional combustion systems,

they are still above values postulated as being safe for the avoidance of corrosion

of turbine blading. However, such safe values are based on marine turbine test

conditions. Alkalies derived from coal or oil combustion appear to be in a less

aggressive form and the results from the corrosion tests using static turbine

blading (8.2 to 8.5) appear encouraging. See also Section 14.

Current experience suggests that corrosion of turbine blading should not

be a problem for turbine gas inlet temperatures below 820 °C (91510 OF), and alloys

are currently being developed that will allow this temperature to be increased

further to 900 OC (1650 OF).

8.2.2 Potential applications

In this subsection the most promising potential applications and

thermodynamic cycles for pressurised fluidised combustion are outlined. The

treatment is illustrative and is not intended to be in any way exhaustive; it

is included so that the implications of the choice of cycle on combustor design

* may be appreciated.


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8.2.2.1 Compact boiler

In this application, which is outlined in Figure 8.2, the combustor

is not used to generate electrical power but only to raise steam or to provide

hot gases. The combustor off-gas is therefore cooled so that the gas turbine

generates only sufficient power to drive the air compressor. Such an application

is somewhat specialised but could be used to supply steam to a factory or for

district heating.

8.2.2.2 Power generation

Large scale generation of electrical power is thought td be an

.9 application of prime importance for pressurised fluidised combustion. No plant

is yet operational but various studies made both by CSL and in the general

literature suggest that significant savings in capital and operating costs should

be obtainable using pressurised fluidised combustion. Also, of course,

fluidised combustion can readily handle poor quality and/or high sulphur fuels in

an environmentally acceptable manner. See also Sections 11 and 12.

Thermodynamic cycles for power generation may be classified broadly

into two categories:

(a) where the working fluid is air and/or combustion products

(b) where there are two or more working fluids; e.g. water/steam and

combustion products; or helium and air/combustion products.

Furthermore the gas turbine working fluid may pass once through the system,

which is then described as an "open cycle", or it may be recirculated in a

"closed cycle". The subject is vast and only a few of those cycles that can be

based on fluidised combustion are butlined below.

8.2.2.2.1 Pure gas turbine cycles

Pressurised fluidised combustion potentially converts "dirty fuels" to

acceptably "clean" fuels. Therefore, it has attractions for coal or residual

oil firing of pure gas turbine cycles which at present need clean fuels, nuclear

energy or large costly air heaters.


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Conv No. ·

steamfluidized StT ea

combustor

water

flue gas air

Figure 8.2

_Compact Boiler.

fluidised combustor

fuel

flue gas air

Figure 8.3

Simple Open Cycle


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I>.*'~~~~~~~~~ _ _ _~~~~~~CO NO .-- "

fluidized combustionair heater

fuel

9.9-

gas air

Figure 8.4

Open "Airheater" Cycle

fluidized rJcombustionair heater l

air < ~~~~~~~~~~flue gas

0~~~~~~~~~~~~~~~~~~~~~~~~~~~~

airthermal

. < uel _> output

2-stage air _compressor / X

with intercooler

_9 < thermaloutput

Figure 8.5

Airheater Cycle with Heat Recovery for District Heating

l July 1979 Section 8

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* l. Simple open cycle. This cycle is shown in Figure 8.3 and is the simplest

possible. The working fluid is the fluidising air which is heated to

the gas turbine inlet temperature as it passes through the bed. Although

this cycle is the ultimate in simplicity, it does not utilise one of the

main advantages of fluidised combustion - the high rates of heat transfer

to in-bed surfaces - and operation must be at high excess air levels to

remove the heat of combustion. It will also be necessary to clean the

whole of the working fluid before it enters the turbine.

2. Open cycle with in-bed air heating. This cycle is shown in Figure 8.4.

The air from the compressor is split into two streams, one of which

fluidises the bed, while the other (up to about two thirds) passes through

tubes immersed in the bed, where it is heated to a temperature approaching

that of the bed before being mixed with the gases cleaned after leaving the

bed. This cycle requires a smaller combustor than the open cycle (1. above)

and it is necessary to clean only the combustion gas which is the smaller

part of the gas flow to the turbine. Combustion will also take place with

a much smaller amount of excess air than with the simple open cycle.

Actual plant designs will incorporate more sophisticated provisions

for heat recovery than are shown on Figure 8.4. As an illustration,

Figure 8.5 shows diagrammatically an air heater scheme that has been

prepared (8.10) based on the use of an existing oil-fired gas turbine.

As well as power generation this scheme incorporates heat exchangers for

recovering the sensible heat remaining in the turbine exhaust gases in the

form of a conventional waste heat boiler. Heat is also recovered via an

air compressor intercooler. Such a scheme would be suitable for an

application needing a combination of power generation and district heating.

However,the thermal output could equally well be via a steam boiler for a

conventional steam turbine power generation system. The overall thermal

efficiency of air heater cycles with sophisticated heat recovery provisions

is comparable with that of mixed cycles (8.9). See Section 8.2.2.2.2 below.

Air heater cycles will normally operate at excess air levels that will

give excellent combustion with a minimum number of fuel distribution points.

They are more simple in operation than mixed cycles. (See Section 8.2.2.2.2

below). Start-up is quicker and the provisions for turndown are less


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complicated. A supply of water is not necessary, which may well be an

advantage for some applications. However, air cooled in-bed surfaces

normally operate at temperatures closer to the bed temperature than do

water cooled surfaces. As a result air cooling requires a larger cooling

surface area for the same heat duty and the selection of appropriate heat-

resisting alloys is more restricted. See Section 15, Table 15.2,

page 15-45-, and also Section 14.

3. Closed cycles. This cycle is illustrated in Figure 8.6. In this cycle the

combustor air and the working fluid for heat removal via in-bed surfaces ate

separate and the latter, being in a closed cycle, heed not be air. Helium,

for example, has a better combination of physical properties than air and

0 has been used (8.11). Another feature is that the closed cycle turbine will

generally operate at a lower pressure ratio than conventional gas turbines,

but at a significantly higher pressure level. This leads to substantial

increases in the tube side heat transfer coefficient and hence to a reduction

in the heat exchanger surface required.

Semi-closed cycles have also been proposed for use with pressurised

fluidised combustion (8.12). Air is the working fluid and the fluidised

combustor is either in the closed or the open part of the cycle. Such

cycles are stated to have good start-up and part load characteristics and

are thus well suited for peak load power generation.

4. Air storage systems. In a normal gas turbine generator plant only about 1/3

of the turbine output is converted into electricity, the other 2/3 being

required to drive the air compressor. With compressed air stored in an

underground cavern, the whole of a turbine output could be used for power

generation at peak load periods. Such a system is illustrated in Figure 8.7.

An "air heater" cycle is used to drive a turbine which, instead of generating

electricity compresses air for storage underground. The stored air is

subsequently fed back into the system at a controlled rate, heated in the

fluidised combustor and then passed to the turbine where all the power output

is used for electrical generation. (8.14). Alternatively, air may be

compressed using off-peak electrical power. An important application using

this principle is an air storage scheme integrated with a power generating

system with a large percentage of nuclear power operating at a steady load.


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fluidized combustionair heater

open closed

Figure 8.6

_Closed Cycle Plant

air

fluidisedcombustor

air _ >

gas

Figure 8.7

Air Storage Scheme-.

.July 1979 Section 8

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EEm ~ ~ ~ ~ ~ ~ A

ZS~Z

Figure 8.8

Supercharged Boiler Combined Cycle Diagram


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8.2.2.2.2 Mixed cycles

The supercharged boiler cycle shown in Figure 8.8 is that cycle

involving both air and water as working fluids which is considered most likely to

provide worthwhile improvements in efficiency and capital costs when fired by a

pressurised fluidised combustor (8.9, 8.12). Because the combustion air is

heated by compression to over 250 0C (500 OF) alternatives to exhaust/air heatexchange have to be found for recovering heat from the steam turbine exhaust.In the arrangement shown in Figure 8.8 some of the feed heaters of the standardsteam cycle have been replaced by a low level economiser.

Calculations made for a coal fired combustor show. that the thermalefficiency of an arrangement like Figure 8.8 would be about 40% (based on thegross calorific value of the fuel), which is about 6 percentage points betterthan for a conventional steam plant incorporating stack gas scrubbing. (8.9, 8.30).Operation over a wide load range should be feasible, a feature of particularadvantage for sites not served by power 'listribution networks. An advantage ofmixed cycles over pure gas turbine cycles when using coal-fired combustors is thatany fall off in turbine power through deposition or erosion has a much smallereffect on the output of the whole plant.

In conventional steam generation systems there is every incentive tooperate with low values of excess air. This is no longer the case in combinedcycles where the cycle efficiency increases as the excess air increases, up toabout 100% excess air. For the cycle shown in Figure 8.8 the efficiency wouldincrease from 40% at 20% excess air to 41% at 100% excess air (8.13). Thisincrease arises basically because more of the heat is being incorporated into thecycle at a high temperature in the gas turbine. Operation at high excess air isbeneficial to combustion, corrosion and turndown, improves controllability,especially for rapid load increases, but increases the pressure shell and gascleaning costs.

It has been suggested that the thermal efficiency would be increased byusing potassium for `topping` a conventional steam cycle by combining the functionsof a potassium condenser with those of a steam boiler, superheater and reheater.

(8.12). A pressurised fluidised combustor would provide a favourable heatingsystem for generating the potassium vapour, since fireside corrosion would be


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less than with other combustion systems. A thermal efficiency of 48% ispredicted, but this increase might not compensate for the higher capital and

operating costs likely to be incurred through the use of potassium.

8.2.2.3 District heating

Any of the schemes shown in Figures 8 .2 to 8.6 and 8.8 would be suitablefor district heating applications. The fluidised combustor can be used equally

to provide only for the district heating needs or to generate power with thedistrict heating as a heat economy measure. A scheme for the latter alternative

using an air heater cycle is discussed in-reference (8.10).

8.2.2.4 Endothermic chemical reactions

Pressurised fluidised combustion offers potential economies forchemical reactions that are both endothermic and are carried cut at elevated

pressures. The reforming of hydrocarbons to produce towns gas, ethylene orammonia is a typical example, requiring temperatures of 650 - 1000 °C(1200 - 1830 OF) and pressures up to 4000 kN/m 2 (40 atm). Under these

conditions metallurgical considerations are of major importance in the design

of the plant because of tube metal creep. By operating a fluidised combustion

furnace under pressure, the pressure difference across the tube walls could be

reduced or eliminated. Also the heat transfer flux in a fluidised bed is more

uniform than in a conventional futnace. Economies in both capital and

maintenance costs could thereby be obtained.

8.2.3 Design Implications

When determining values of design parameters for pressurised fluidised

combustors various factors additional to those necessary for atmospheric pressure

fluidisation must be considered. The majority of these stem from the use of a

gas turbine.

1. Operating pressure. The operating pressure will be mainly determined by

the choice of gas turbine for the particular application. Currentlyavailable gas turbines operate in the pressure range 400 - 1600 kN/m 2

(4 - 16 atm). with a preference for the range 800 - 1200 IG/m 2 (8 - 12 atm).


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2. Gas cleaning. The off-gases from a pressurised fluidised combustor will

always pass through a gas turbine although they may first be diluted with

cooler air. See 3. and 4. below. Provision for solids removal at bed

temperature and pressure is needed. Cyclones are suitable for the bulk

of the cleaning. They should be refractory lined and be designed to

facilitate the removal of collected solids since under some adverse

operating conditions the latter may contain a sufficiently high proportion

of carbon to be combustible. However, it is recommended that the final

stage of the gas cleaning equipment and all downstream ducting should be

made of metal. Any possibility would thereby be avoided of any portions

of refractory lining being carried through to the gas turbines.

3. Gas turbine control. The gas turbine may be either in single shaft

operation, directly linked both to the air compressor and an alternator,

or in two, or three, shaft operation, with the compressor and alternator

driven by separate turbines. While two or three shaft operation gives

greater flexibility of operation the singlc shaft system gives a longer

time interval for corrective action in the event of an alternator trip,

because of the stabilising influence of the directly connected air

compressor.

The two or three shaft system allows the turbine-compressor set to be

run at variable speed if needed to alter the air flow. The single shaft

system must be run at synchronous speed and the only control variable for

load variation becomes the turbine inlet temperature. A by-pass air

stream around the combustor is therefore required to vary the turbine

inlet temperature independently of the bed temperature. Either system

may be used with any of the schemes shown in Figures 8.2 to 8.8.

4. Load control. This subject has already been discussed in general in

Section 3.6 of Section 3. A detailed study will be required for each

individual application. It is somewhat easier to arrange for adequate

turndown ratios using gas turbine cycles because of the greater influence

of bed temperature on heat removal. However, the turndown obtainable

by reduction of bed temperature alone is always limited. For most

applications the use of compartmented beds and the by-passing of air

around the combustor to reduce the turbine inlet temperature will also

be required.


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Load control is most difficult in mixed cycles operating at minimal

excess air levels. Combustors operating under these conditions will

require deep beds in which it may be difficult to reactivate individual

compartments.' Various methods are currently under examination that

might be used as an alternative to the use of compartments with bed

slumping for varying the heat transfer to in-bed surfaces. In one

alternative, bed material is transferred to a refractory storage vessel.

This exposes the upper tubes of the tube bank and reduces the heat transfer

coefficient there. See Section 10.4. A second alternative method uses

a patented (8.16) recirculating bed system in which the combustion and

heat transfer functions of the bed are physically separated. See also

Section 8.3.7.

5. Combustor start-up. Conventional oil-fired or steam driven turbo-

alternators can be speeded up and synchronised in only a few minutes.

A much longer time is required for fluidised dombustors. A start-up

burner using a premium fuel will be required in a line supplying the

gas turbine and also a source of heat for the combustor bed. It will

be necessary to bring the combustor up to its operating temperature

one compartment at a time. Alternatively, the bed depth can be

progressively increased to submerge more in-bed tubing. Again, a

detailed study will be required for each individual application.

6. Excess air level. In pressurised fluidised combustion the excess air

level is not determined mainly from combustion efficiency considerations,

as in atmospheric pressure combustion, but by cycle type and from

considerations of turndown and gas turbine control. As the excess air

level increases; for a given heat input, so also does the rate of heat

removal as sensible heat from the bed, but the in-bed heat transfer

surface that is needed will decrease. Also the combustor cross-section

will increase in direct proportion to the excess air level. A final

choice of excess air level, therefore, will require a detailed economic

study.

7. Combustor configuration. The use of a tapered bed construction is

recommended for pressurised combustors. With a tapered bed a smaller

volume of air is needed to commence fluidisation of the bed than with

I-July 1979 Section 8

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Copy No. c

a parallel-sided bed. During the start up of pressurised combustors

this facility allows the pressure in tapered bed combustors to be raised

to normal operating values more quickly and smoothly. An additional

advantage of tapered beds is that the vigorous fluidisation conditions

existing towards the base of the bed makes any difficulties through

segregation of bed particles even less likely. The bed can be tapered

in one or both planes and can be made in a modular construction. See

also Section 15.

8.3 Effects of Pressure on Bed Performance

8.3.1 Fluidisation quality

The bed of a fluidised combustor must be maintained in a well fluidised

state at all times. Section 9.8.1 contains a description of what is meant by

"good fluidisation". Briefly, the term implies that a correct degree of bed

solids circulation is maintained so that elutriation is not excessive yet bed-to-

surface heat transfer and bed combustion occur at optimum rates. Attention to

fluidisation quality is particularly important in pressurised combustion because

a reduction in bed temperature, pressure or fluidising velocity shifts the

operating conditions towards defluidisation and there is less leeway than in

atmospheric pressure fluidisation.

Laboratory studies (8.17 to 8.20) have shown that as the pressure is

increased, firstly, the bubbles in a fluidised bed become smaller and secondly,

the proportion of gas in the bubble phase decreases. These effects would be

expected to lead to better solids/gas contacting, and hence to improved heat

transfer and combustion characteristics during pressurised fluidised combustion.

It is not clear, however, to what extent these findings are applicable to

large scale operation.

A graphical representation of fluidisation conditions that is

recommended for monitoring combustor operation is shown in Figure 8.9. In

this Figure a Velocity number, Vn, is plotted as a function of a Fluidisation

number, Fn. Vn and Fn are defined by equations 8.1 and 8.2 below and are

chosen so that the fluidising velocity occurs only in Vn and the particle size

only in Fn. Details of the derivation of Vn and Fn are given in Section 9.8.2.

-1JulIy 19-79 Section 8

Page 18 of 42

CoPY No. iA

The defining equations are,

Vn = Uf r - Pg)J ... ... ... ... 8.1[g p (PS Pg)

Fn = 0 d g )] ... . ... 8.2

Lines representing minimum fluidising conditions (subscript mf,

d = dp, Uf = UJf) and terminal free-fall velocity conditions (subscript t,d = dx, Uf = Ut) are shown as plots of Vn versus Fn on Figure 8.9. In

calculating Vnmf , Umf was evaluated using equation 3.1 of Section 3 with

* = 0.8 and ef = 0.5. In calculating Vnt , Ut was evaluated using equations3.3 and 3.4 of Section 3 with 4 =0.8. However, for large particles(dx L0 > 1000 pm) the correlation of reference (8.21) for cubes has been usedinstead of equation 3.5 for spheres. These choices are thought to correspond

best with particles typical of those in a fluidised combustor bed.

Two further lines, labelled good fluidisation and minimum safe

fluidisation are also marked on Figure 8.9. These lines have been derived

empirically to fit data thought to represent operation under conditions of good

fluidisation and operation just above segregation conditions respectively.

A series of arrowed lines on Figure 8.9 shows the direction of movement

of a point on the figure caused by a change in one of the operating parameters.The scales marked on each vector arrow are ratios of the new value to the

original value. On the temperature vectors the changes are marked both as

ratios of the absolute temperatures and also in °C. When considering changes

in general it is important to distinguish whether the gas mass flow or the gas

volume flow (proportional to the fluidising velocity) is being held constant.

Since velocity, pressure, mass flow and temperature are not independent, thereare a number of possible vectors. Thus the vector ?P varied, Uf constant,

Gg changes" infers that P is changed by the ratio 2, 4, 6, etc at a constant

velocity Uf, with a corresponding change in mass flow, whilst "P varied,

G constant, Uf changes" infers that P is changed by the ratio 2, 4, 6, etc

with a corresponding change in velocity to maintain a constant mass flow.


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These vectors apply over the whole range of the graph and it will beseen that their effect can differ at different values of Vn or Fn because theeffect of pressure on physical properties is such that points representingoperation at atmospheric pressure generally occur to the left hand ends of thecurves of Figure 8.9, where the slope of the lines is greatest, while pointsrepresenting pressurised operation generally occur in the middle and towardsthe right hand (as the pressure rises) of the Figure.

The use of Figure 8.9 may be illustrated by considering the operationof a hypothetical pressurised fluidised combustor. Normal operation might berepresented by the point B on the "good fluidisation" line of Figure 8.9. Areduction of the bed temperature from 950 to 750 °C (1740 to 1380 OF) wouldmove the operating point from B to C. A subsequent reduction of the fluidisingvelocity by 25% would then shift the operating point to D , where the qualityof fluidisation is only just satisfactory. Further reductions in any of theparameters, bed temperature, fluidising velocity and air mass flow, or anincrease in the pressure, might them cause unsatisfactory particle flow and thepossibility of segregation of bed particles. Good fluidising conditions couldbe restored, however, by reducing the pressure at constant mass flow, therebymoving the operating conditions to point E. Sequences of changes such as thesecan occur during turndown, shut-down and start-up procedures. The aboveillustration emphasises that the effects on fluidisation quality of changes inoperating parameters must be considered carefully in pressurised fluidisedcombustion. In particular, it can be seen from Figure 8.9 that it is notpossible to fluidise the bed of a cold pressurised combustor at start-up unlessthe pressure is considerably below the normal operating value.

It must be emphasised, however, that in pressurised combustionthe interactions between the gas turbine, air compressor and fluidised combustorare very complex and each application will require individual analysis.

Analysis of the data supporting the "good fluidisation" line ofFigure 8.9 has shown that the ratio of maximum particle size, dMX, to meanparticle size, d , varies from 2.3 to 5.6 with a mean of 3.3. In this

panalysis dMx was defined as the 99% cut point on a Rosin-Rammler plot; i.e.less than 1% w/w of the sample was greater than size dMx. If the conservativeassumption is made for coal firing that the particle size distribution of the

1 July 1979 ~~~~~~~~~~~~Section 8

PLage 20 of 42

(e Cony No.~~~~~~OW14

-ro ~ ~ ~ ~ +L

t~tjJ14 h:Q. FIL; 1141 L4 12.6 M pi p M p-WZ ~ ~ QII' 1

9~~~~~~~~~~~~~~~~~~~~~~~~~~

a, ,orn"~~~~~~~~~~~~~~~~~~~~~~~~~~~tttT+ St A t-+-~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~

0~~~~~~~~~~~~~~

2 ~~>2

.. . .. ...... i..r..8.9


Page 21 of 42

Copy No. ' .'

ash forming the bed is the same as that of the feed coal, then the coal feed top

size necessary for obtaining "good fluidisation" can be related to the mean bed

particle size. Thus, to ensure "good fluidisation" when coal firing using

in-bed feeding techniques. it is recommended that the coal feed top size (defined

in the same way as dmx) should not exceed 3.3 times the mean particle size of

the bed under normal operating conditions.

8.3.2 Combustion efficiency

In pressurised combustion combustion efficiencies in excess of 98%

will normally be required and very little combustion can be tolerated in the

freeboard. There are several reasons for avoiding freeboard combustion.

_ft Firstly, the power generated by the gas turbine depends on the gas inlet

temperature which is of the same order as the combustor bed temperature. Thus

in contrast to atmospheric pressure combustion, freeboard cooling is not normally

provided in pressurised combustion and there is no significant design temperature

change in the off gases as they pass through the freeboard and gas cleaning

equipment. Freeboard combustion could, therefore, cause overheating of the

gas clean-up equipment. Also high combustion temperatures may sinter the

relatively soft particles elutriated from the bed to give harder more abrasive

particles that might erode turbine blading. Furthermore, the occurrence of

freeboard combustion can initiate an undesirable cycle of events that requires

special provisions to control. Corrective action to reduce. high off-gas

temperatures by reduction of bed temperature or by increasing gas velocity will

only tend to reduce combustion efficiency and aggravate freeboard combustion.

Combustion efficiency in the bed depends mainly on gas residence time,

defined as (combustor bed volume)/(volumetric air flow), and on bed temperature,

excess air level and the number of fuel feed points. The dependence, of course,

also varies according to the fuel fired.

8.3.2.1 Coal firing

In coal-fired combustors the fuel particles burn in the bed unless they

are elutriated, or ejected by splashing through the bursting of bubbles. The

effects of operating variables dn combustion efficiency are discussed in Section 4

for atmospheric pressure operation. Fluidised combustion under pressure is


Page 22 of 42

Copy No. .-..

generally similar but differs in two important respects. In the first place, as

the operating pressure increases the gas bubble size decreases with the consequence

that gas/solids contacting is improved. Secondly, pressurised combustors are

generally designed with relatively deep beds and closely packed tube banks. Both

of these factors lead to high combustion efficiencies and the virtual absence of

freeboard combustion. In view of these differences from atmospheric pressure

combustion it is recommended that the method for calculating combustion efficiency

given in Section 4 should not be used for pressurised combustion. Instead, curves

based on experimental results are presented in this Section and should be used for

.estimating the combustion efficiency.

For good combustion it is recommended, currently, that bed areas up to

0.93 m2 (10 ft2) may be fed per feed nozzle for operating pressures up to

610 kN/m 2 (6 atm) absolute using in-bed feeding; no experience is available with

overbed feeding. Normally, it is possible to select operating conditions to

give adequate combustion efficiencies in once-through operation and thereby avoid

the complications of dust recycle or of a separate carbon burn-up bed. See also

Section 4.4.

For a given system and fuel the operating parameters that may affect the

combustion efficiency are the bed temperature, excess air level, bed depth and

fluidising velocity. These last two variables can be combined into a single one -

a gas residence time, tg , defined as,

bed volumetg volumetric gas flow

or, for parallel sided beds, as,

Lbtg =Uf

The general effect of bed temperature on combustion efficiency is shown

in Figure 8.10. While there are only marginal improvements to the efficiency in

the 850 - 950 OC (1560 - 1740 OF) range, reduction to 800 °C (1470 OF) causes

a significant drop in the efficiency.

*Symbols are defined in Section 1 of this Manual.


Page 23 of 4a

Copy No.

The effect of excess air level is shown in Figure 8.11. Only

marginal increases in combustion efficiency are seen once the excess air is

greater than 25%, but a marked reduction is caused if the excess air level drops

below 20%.

The effect of fluidising velocity and bed depth combined into the single

variable, gas residence time, is shown in Figure 8.12. It is not always

possible in experiments to vary the bed depth over a wide range because of the

need to provide adequate depth for in-bed cooling surfaces. If any effects of

bed depth itself are ignored then Figure 8.13 is obtained where the combustion

efficiency is plotted directly as a function of the fluidising velocity. The

justification for this method is that the combustion loss is almost wholly due

to entrained carbon, and the latter would be expected to be some function of

the fluidising velocity at the surface of the bed, provided that there is

adequate residence time for combustion in the bed itself. The data so far

available do not allow a choice of correlation to be made and for this reason

both.figures.have been included. Apart from experimental error it is also

suspected that some scatter of the data points is due to a second order effect

of coal type and fines content.

The general conclusion, however, from Figures 8.10 to 8.13 is that a

combustion efficiency of greater than 98.5% for a bed temperature range of

750 - 950 °C (1380 - 1740 OF) will require a residence time of 2 seconds or

greater, preferably using a fluidising velocity less than 1.5 m/s (5 ft/s).

9 Fortunately, for most pressurised applications, the values of operating

parameters determined by considerations other than combustion are also compatible

with the above requirements for obtaining good combustion. The operating

pressure and maximum bed temperature are dictated by the choice of gas turbine,

but are likely to be in the ranges 610 - 1600 kN/m2 (6 - 16 atm) and

800 - 900 0C (1470 - 1650 OF) respectively. Excess air levels are dictated

mainly by the choice of cycle and application type. For Group 2a applications

the excess air is likely to be in the range 100 - 300 % and for Group 2b

applications in the range 25 - 100 %. Fluidising velocities of 1.5 mi/s (5 ft/s)

do not result in unduly large comnbustor cross-sections and, in Group 2b

applications, the requirements for accommodation of the necessary in-bed heat

transfer surfaces call for deep beds so that gas residence times in the range

2 - 4 s are normally needed. See also Section 8.3.3 below.


Page 24 of 42

Bed temperature - OFCony No

1300 1400 1500 1600 1700 1800

99

98

mm-m0.25 .Gas residence time - s97

96

Overallcombustion efficiency

94

i :1::: f iEF

F~~~~~~~~5~~ =, Excess airi i 2%

700 800 900 1000

Bed temperature - C

Figure 8.10

Effect of Bed Temperature on Combustion Efficiency


Page 25 of 42

* Coov No. 'if'

100

':i '

.. -r L :- r S~~ . . -. _ r _+_r_*r;._......... = === = =t = ==E _= == == ====

::; '_,r : ,,, U ~: Fi -I1T - X_ _: ' I 8edmmpere I residence [W -=:_- - l : ~ ~- --- A 8 1600

96| ! M ; HE

[ X X~~~~~. i. 7 X ;|iF. .

94~~~~~~~~~~~~~~~~ ~_: :::::= =-: =3 -:::L a --]._

::lt:::: i l ~ ~ =-- ~ :_ L 3 _-.'edtemperature Gas residence e93::O:: i -- s

4 tX +1:- _._ _ m m m t | ~~~~~~~~A |870 |1600 | 3 l

t 1: : 9~~ W m W i|i .=-= | ~~B , 8 0 0 1470 |2

coJmbustion :: :: : - -=-= _ =

0 20 40 60 80 100 120 140 160

Excess air %

Figure 8.1 1

Effect of Excess Air Level on Combustion Efficiency


Page 26 of 42

Copy No.

850°C9 =1 c15800F

ff _}m2e X X W 1 1470°F > t fficiency

F8 I I I i gi ue 8 .'11380OF

Efet o Ga Reic Ti o bst E f f ie

97~ ;i;l~i'/]!!~ I :1: ; -: :i::Exces

9 1 'iefficiency ..// d - %...........

O 1 2 3

Effect of Gas Residence Time on Combustion Efficiency

,~, [II ........... W1101 ;: |gW = LWL=0 0 m it~I,=, T M"

| t S S W' 96 t- ::i ,i i? " i' : ! : i 5i i : i l r: i ! ; ! i i; ii ! ! [! !

0 , 1. 2 3,, , , , , , I ! , i , , !

•~~~~~~a esdnetm ,~ , l; /!:!I , t,.l=b ¢!h h J , ;=:= =~~' i gu r e , 8.12 l r:, :; ;, !, I i, i!!¢, :¢ i

1: :1' 1 1 fec of Ga Reiec ='iTrTimeI on: Co bsto Efficiency,i]: '!,: :l ''1;' ?........... . . . .

_N~~~~I ~ t

_1 July 1979 Section 8

Page 27 of 42

Fluidising velocity at bed surface m/sopy No.

0 1 2 3 4700

98

840- 870168000 g E

1470°F

95'* 96 g

5 g t t i i g 900 W

94

Overall

efficiency

q-%

0 2 4 6 8 10 12 14 16

Fluidising velocity at bed surface ft/s

Figure 8.13

Effect of Fluidising Velocity on Combustion Efficiency

$


Page 28 of 42

Copy No.

8.3.2.2 Oil firing

Various tests have been made burning residual fuel oil and heavy fuel

oil in pressurised combustors (8.13, 8.22, 8.31 - 8.34) and it is considered

that this could be carried out on the large scale successfully if the oil

distribution system is designed according to the guidelines given in Section

15.1.3.1. The design of distribution systems for feeding oil fuels to pressurised

fluidised combustors is also a current area of ongoing development (8.34). See

also Section 5. Because of the need for distributing oils at a relatively large

number of inlet points in a manner that ensures rapid dispersal throughout the

bed in order to minimise freeboard combustion, the overall combustion efficiency

is normally 99.5% or greater. A possible advantage of oil firing in some

applications is that these combustion efficiencies. can be maintained at excess

air values as low as 15%.

8.3.2.3 Gas firing

Very high local combustion intensities are caused in gas firing which

calls for a very rapid circulation of bed material to avoid the possibility of

fusion of bed particles. For this reason gas is not recommended as a fuel for

pressurised fluidised combustion except at start-up. When used for start-up

it is recommended that the change to the main fuel should be made before the

operating pressure reaches 200 - 300 kN/m2 (2 - 3 atm).

8.3.2.4 Other fuels

It is considered that pressurised combustion requires a guaranteed

source of fuel since start-up and shut-down procedures are more laborious than

for atmospheric pressure fluidised combustion. For this reason the fuels

discussed in Section 7 of this Manual are not considered suitable for pressurised

fluidised combustion.

8.3.3 Elutriation

In almost all pressurised combustion applications the dust content of

the hot combustor off-gas must be reduced before the gases can be passed to the

gas turbine. Detailed design of suitable hot gas clean-up equipment is not


Page 29 of 42

Copy No. &

considered as within the province of this Manual. The subject is however, a

current area of ongoing development.

The performance required of the hot gas clean-up system will be largely

determined by the particle concentration and particle sizes that can be tolerated

by the gas turbine but the avoidance of environmental pollution by the stack gases

should not be overlobked. See Section 18. The load on the clean-up system will

be determined by the gas volume flowrate and the particle elutriation rate from the

freeboard. The rate and particle size range of the elutriated solids can be

calculated to a first approximation using the procedures outlined in Section 9.

However, more accurate estimates of the kind needed for a final design of a

clean-up system can only be obtained at the present time from tests on an

0 experimental combustor using fuel, additives and operating conditions matching as

closely as possible those chosen for the full scale combustor. See Section 17

for testing facilities.

Experimental elutriation rates are being analysed currently and a

correlation is being developed. The rates vary widely according to the operating

conditions and the type and specification of both the fuel and the sulphur

retention additive used. The fluidising velocity is the dominant operating

variable and Table 8.1 illustrates its effect for a typical coal fired combustor

operated with dolomite as a sulphur retention sorbent.

Table 8.1

9 Effect of Fluidising Velocity on the Elutriation of Mineral Matter

Fluidising velocity Elutriation - % of input Total elutriationmn/s ft/s Ash Dolomite Total kg/MWh lb/106 Btu

0.77 2.5 58 29 42 12.2 7.91.37 4.5 85 35 56 14.9 9.62.13 7 100 87 97 28.7 18.5

For oil firing lower particulate emissions from the bed have been noted

amounting to about 10% of the dolomite input and consisting almost entirely of

partially sulphated dolomite with negligible refractory bed material. (8.22).

The fluidising velocity was 1 m/s (3.1 ft/s) in these tests.

I July 19°79 Section 8

Page 30 of_~2,

Copy No.

8.3.4 In-bed heat transfer

The correlations and design methods outlined in Section 10 for

calculating heat transfer rates are unaffected by the operating pressure and

may be used equally for pressurised combustion or for atmospheric pressure

operation. As the combustion intensities are generally greater it becomes

necessary in pressurised combustion with in-bed heat removal (Group 2b applications)

to pack in-bed tubing as closely as possible. However, it is recommended that

not more than 17% of the bed volume should be tubing.

The bed-to-surface convective heat transfer coefficient, which is

correlated by equation 10.9 C10.9a) of Section 10, decreases as the gap between

the tubes is reduced. The relative change in the coefficient is small and is

illustrated in Table 8.2 for various tube diameters and tube gaps. Now the

recommendation of a 17% maaximum for the volume occupied by tubes implies a

minimum gap, z, for any given tube diameter, Do. For an equilateral

triangular tube pitch, for example, the relationships are,

r 1~~2tube volume j | Do 1 ... . .. ... ... ... 8.3bed volume 2 ;[3 Do + z

and when the value of the left hand side of equation 8.1 is 0.17 the minimum

recommended tube gap is given by

z = 1.31 Do0 ... ... ... ... ... ... ... 8.4

Table 8.2 also shows, for illustration, values of the tube volume

and tube surface area per unit bed volume and a relative heat flux per unit bed

volume. The Table illustrates that the relative convective heat flux per unit

bed volume increases as the tubes are packed closer together, in spite of the

reduced convective heat transfer coefficient, and also that the heat flux

increases as the tube diameter is decreased. Thus, there is clearly an

incentive to use the. minimum tube gap with small diameter tubes. However, a

minimum tube diameter of 25 mm (1 in.) is recommended. Other limitations on

pitch, etc., and on operating conditions are given in Table 10.1 of Section 10.


Page 31 of 42

Copy No. -

Table 8.2

Relative Effects of Tube Gap on In-bed Heat Transfer

Tube dia. Do in. 1 1 1 2 2

Pitch, Ph in. 2.31 3 5 4.61 6(equilateral)

Tube gap, z in. 1.31(min.) 2 4 2.61(min.) 4

Tube vol/bed vol % 17 10 3.6 17 10

Tube surface/bed vol ft-1 8.16 4.84 1.74 4.o8 2.42

Relative hlc 0.975 1.0 1.07 1.02 1.07

Relative flux 7.9 4.84 1.86 4.i6 2.59(tube surface/bed vol) x h.

It will be found necessary in most pressurised fluidised combustor

applications to use greater bed depths than those employed in atmospheric pressure

applications merely to accommodate the required in-bed heat transfer tubing,

even at the closest recommended spacing. Experimental units have been operated

successfully with beds up to 2.5 m (8 ft) in depth. With such deep beds it is

necessary to allow for the effect of bed depth on the convective bed-to-surface

heat transfer coefficient as described in Section 10.2.4.2.

The in-bed tubing may be arranged in any convenient geometry but it is

recommended that no clear vertical passage for bed particles and gas bubbles

should be available. Thus, referring to Figure 8.14, arrangements (a) are

recommended while arrangements (b) are not. However, under some operating

conditions it has already been shown that vertical finned exchangers can produce

acceptable combustion conditions (8.6) but the in-bed heat transfer coefficient

may be somewhat lower than for equivalent plain horizontal tubing.

8.3.5 Sulphur retention

The effect of pressure on the performance of sorbents for sulphur

retention added to the bed of the combustor are described in detail in Section 11.2.

During pressurised combustion the performance of dolomites is found, in general,

to be superior to that of limestones. This trend is the reverse of that

experienced during atmospheric pressure fluidised combustion. The explanation

is thought to lie in the effects of pressure on the decrepitation and mechanical

breakdown of sorbent particles during calcination. A certain amount of particle

_ 1 July 1979 Section 8

Page 32 of 4j

Covpy No.

a). Ib).

10 0 0 0 0 0 0 o 00 0 0 0 0 o o 0

1 000000 0 0 0 000 0loo 0o o000010000000 0000

0000

000 000 000 000

l Il ~000 000 000

I 000 000 000 000

01 0 0 L 0 00 00 00I

I I INot recommended Recommended

Figure 8.14

in-bed Tubing Arrangements

D~


Page 33 of 42

Copy No. <<

breakdown is needed to give access to the interior of the particles. If the

particles are broken down too finely, however, they are rapidly elutriated from

the bed before much reaction has occurred. In dolomites the magnesium carbonate

portion calcines at all conditions used in both atmospheric pressure and pressurised

fluidised combustors. The calcium carbonate portion also calcines at conditions

used in atmospheric pressure operation but its calcination rate is reduced as the

pressure rises, with the result that particle breakdown for dolomites becomes

closer to the optimum compromise during pressurised fluidised combustion.

Final choice of a sorbent will, of course, depend not only on the

sorbent efficiency but also on the local availability and economics of sorbent

supply. The disposal of the bed material should also be considered as magnesium

oxide hydrates more slowly than calcium oxide. See also Section 16.

The procedure for calculating the amount of sorbent needed to achieve a

required sulphur retention during pressurised operation is exactly as described

in Section 11.3; the effects of pressure are incorporated into the decrepitation

parameters. Operating pressures higher than atmospheric pressure cause the bed

temperature for.optimum sorbent performance to occur at higher values and also

the performance - bed temperature curve becomes flatter with increasing pressure.

The deeper bed depths, and longer residence times, commonly used.for pressurised

combustion are advantageous as regards sulphur retention. In many pressurised

combustion applications a 90% sulphur retention can be obtained with a Ca/S

molar ratio of 2.0.

8.3.6 emissions

NOx emissions from pressurised fluidised combustors have been found to

be less than those from combustors operating at atmospheric pressure. Figure 8.15

illustrates this effect for a coal-fired combustor (8.23). Typical values of

NOx concentrations in the combustion gases from coal-fired pressurised fluidised

combustors lie in the range 80 - 180 ppm v/v. NO2 concentrations for similar

operating conditions are typically about 5 ppm v/v. During tests on an oil-fired

pressurised combustor (8.22) values of NOx concentrations in the combustion gases

were around 100 ppm v/v when using an oil fuel containing 2230 ppm of nitrogen.


Page 34 of 42

Copy No.

No predictive models are yet available to the designer for calculating

the No. concentrations emitted under given conditions but the basic mechanism of

the process is now sufficiently well understood for qualitative predictions to be

made (8.24, 8.25, 8.26). A fuller discussion is given in Section 12. The

formation and decomposition of nitrogen oxides during fluidised combustion is a

complex process. The main compound formed is nitric oxide, NO. It is formed

during combustion from the nitrogen compounds in the fuel and also from

atmospheric nitrogen. In fluidised combustion NO formation from atmospheric

nitrogen accounts for about only 10% of the total NOx emissions (8.25). This

has been demonstrated by feeding helium-oxygen mixtures in place of air (8.26).

Formation of NO from fuel nitrogen is favoured by high oxygen concentrations

and hence occurs preferentially low down in the bed near the distributor and

under high excess air conditions. There are two significant NO decomposition

reactions active in coal-fired fluidised combustors. One is a homogeneous

reaction between NO and gaseous nitrogen compounds (e.g. NH3) from coal volatiles.

The other is a heterogeneous reaction between NO and carbon (8.25). The latter

reaction becomes significant at bed temperatures above 800 °C (1470 OF) (8.25).

Both decomposition reactions occur throughout the bed and not predominantly at

the bottom; their significance for pressurised fluidised combustion lies in the

fact that they require good contacting between gas and bed particles for their

effectiveness. Indreasing operating pressure gives smaller bubbles and better

gaslsolids contacting (See Section 8.3.4), and this, along with the deep beds

usually used in pressurised combustion, are thought to be the main reasons why

increasing operating pressure reduces NOx emissions as shown in Figure 8.15.

Additional support is given to this hypothesis by the observation that momentary

fluctuations in NOx emissions also decrease with increasing pressure (8.23).

This would be expected if the bubbles become smaller and more numerous at high

pressures.

The overall effect of the NO formation and decomposition reactions

coupled with the degree of bed/solids contacting is that NOx emissions from

pressurised fluidised combustors are increased both by operation at high excess

air and at high bed temperatures, although a fall in the rate of increase is

experienced at bed temperatures over 800 0C (1470 OF).

For fuels that do not contain nitrogen it would be anticipated that

Nox emissions would be around 10% of the values quoted above for coal firing.,

i July 1979 Section 8

Page 35 of 42

Pressure- kN/m2Copy No.

0 100 200 300 400 600

800 I 4 iFF3rP R NOxemission

700 lb106 Btu

ncentra tion iCobs1.0

600

ppm - v/v

(assuming 20%500 excess air)

400

0.5300

200

100

2 3 4 5 , '

Pressure -atm.abs.

Figure 8.15

Effect of Pressure on NOx Concentration in Combustor Off-gas

S

1 Jufly 1979 Section 8

Page 36 of 42

Copy No. V '

This is confirmed by tests using propane (at start-up) when NO, concentrations

around 14 ppm v/v were observed (8.2).

8.3.7 Turndown

Turndown methods are discussed in general in Section 3.6.2. In

Section 8.2.3 the importance of turndown in pressurised combustor design is

outlined and it is explained that selection of the most appropriate method for

a pressurised fluidised combustor application must be the subject of a detailed

study. Two promising methods of turndown that are particularly applicable to

pressurised fluidised combustion are outlined below.

O 8.3.7.1 Bed Circulation

In this method separate portions of the bed (or separate beds) are used

for the combustion and heat transfer functions of a fluidised combustor. (8.16).

In cold model work it has been demonstrated that circulation rates between beds

of the order required to give a wide load change can be maintained stably (8.28).

The equipment, shown diagrammatically in Figure 8.16, consisted of two beds one

above the other and separated by a "one way screen" which allowed bed particles

to move upwards preferentially. A suitable design made from cooling tubes is

shown in the insert of Figure 8.16. A downcomer occupying 11% of the total bed

cross-sectional area was situated at one side. When fluidised with circulation

air it allowed bed material to return from the upper bed to the lower one.

In an operating combustor the combustion would occur in the lower bed

and heat transfer would 'occur in the upper bed via in-bed cooling surfaces immersed

in it. The higher the circulation rate of solids between the beds the more nearly

will the temperature of the upper bed approach that of the combustion bed. As

the solids circulation rate is reduced the temperature of the upper bed will fall,

and with it the rate of in-bed heat removal.

8.3.7.2 Airslide transfer

In this method of turndown control a part of the combustor bed is

transferred by an "airslide" out of the combustion/in-bed heat transfer zone to

an adjacent storage zone. The solids level in the combustor bed is thereby


Page 37 of 42

Co'py No.

lowered to expose some in-bed heat transfer surface to the freeboard with a

consequent decrease in heat transfer coefficient and heat transfer rate.

See Section 10.4.

Cold model studies (8.35, 8.36, 8.1) using the apparatus shown in

Figure 8.17 have demonstrated that bed material can be readily transferred from

the side container to the combustor bed with a reasonable pressure drop (3.5 kN/m 2)

(0.5 psi) in the experiments). Bed material could also be moved from the bed

to the side container using an external pneumatic elevator without difficulty.

A conceptual design for a large power generation application showed that this

turndown method is a practicable proposition (8.1) and the concept has been

patented (8.37).


Page 38 of 42

Copy No. ,t

Design of "One Way" Screen

r000 0> 000 `0 0 0 0ooo h Joe o directions of particle

preferential flow

upper bed

o o -- 0 _ I downcomer

O O o o. O O/

0 0 0 0 0 0_ _ 0 0O 0 o -O

"on0 0 0 0 0wa "one way"

screen

o o o o o o 0 0 0 O O O _i -- porous plastic

- distributor

lower bed -I "

fluidising solidsair circulation air

Figure 8.16

Recirculating Bed Apparatus

0


Page 39 of 42

Cony No. < '-

,~ 02 side container(slumped b eh)

I\I \ ll l transfer port fluidised

I \ \(3 xx C8) bed

\ \

\\

02-~ jjli~ side container i

, I~~~~~~~~~~~~~~~~~~Ar$ide d .etaIs

(slumped bed)

X,t , externalIj~~ -1 * \ \.r f'-pneumnatic elevator

transfer \

(fluidised bed)

airslide

distributor

Figure 8.17General Arrangement of Airslide Cold Model

I\


Page 40 of 42

Copy No. -

8.4 References

8.1 NCB, CURL. "Report on work carried out up to March 1971 on

pressurised fluid-bed combustion of coal". Nov (1971).

8.2 NCB, CURL. "Pressurised fluidised bed combustion - Report on

test run No. 5". Report to US ERDA under contract No.

E (49-18) - 1511, Sept (1975).

8.3 NCB, CURL. "Pressurised fluidised bed combustion - Quarterly

technical progress report for January to March 1977". Report

No. 39 to US, ERDA under contract No. Ex-76-C-01-1511,

April (1977).

8.4 NCB, CURL. "Pressurised fluidised bed combustion - Quarterly

technical progress report April to June 1977". Test 7 details.

Report to US, ERDA under contract No. EX-76-C-01-1511.

July (1977).

8.5 "'Feasibility study at BCURA Ltd in respect of certain test

conditions for fluidised combustion of coal". Interim summary

of tests 1-9. June 1975 - Feb 1976. May (1976).

8.6 NCB, CURL. "P.F.B. test on Curtiss-Wright heat exchanger at

Leatherhead". Report to US, ERDA under sub-contract FE 112003

for prime contract Ex-76-C-01-1726. October (1977).

8.7 NCB, CURL. "Pressurised fluidised bed combustion - Part 1:

Tests on a 12 inch diameter combustor". Report under contract

CC 4070. April (1978).

8.8 NCB, CURL. "Pressurised fluidised bed combustion - Part 2: Tests

on the 3 ft by 2 ft combustor". Report under contract CC 4070.

May (1978).

8.9 Davidson, B.J. & Moore, M.J. "Preliminary Assessments of the

Relative Thermal Efficiencies of Advanced Coal Burning Systems".

Symp. on "Power from Coal". Instn. Mech. Eng. London (10 April

1979), paper C20/79, pp. 53-65.

8.10 Woodall-Duckham. "Prospectus and phase I proposal for 66 MW gas

turbine power station with 116 MW district heating scheme

incorporating fluidised bed combustion". Reports Nos. PF 4101,

Mar (1974) & PF 4101/001 Jly (1974).

8.11 Keller, C. & Schmidt, D. ASME Paper No. 67-CT-10, (1967).


Page 41 of 42

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I8.12 Hoy, H.R. & Stantan, J.E. "Fluidised combustion under pressure".

Joint meeting Chemical Institute of Canada and ACS Division of

fuel chemistry. Toronto, 24th May (1970).

8.13 Roberts, A.G. et alii. "Fluidised combustion of coal and oil

under pressure". Fluidised Combustion. Institute of Fuel

Symposium Series No. 1, Vol 1, paper D4, September (1975).

8.14 Gramonti, A.J. et alii "Conceptual design of compressed air energy

storage electric power systems" Applied Energy 231, 4, (1978).

8.15 Hoy, H.R. & Roberts, A.G. "Further experiments on the pilot-scale

pressurised combustor at Leatherhead, England"'. Proc. Fluidised

Bed Combustion Technology Exchange Workshop, p. 43, April 13-15

(1977).

> ~ 8.16 British Patent No. 1 326 651, 15 Aug 1973, and the equivalent

US Patent No. 3 868 993.

8.17 Subswari, M.F., Clift, R. & Pyle, D.L. "Bubbling behaviour of

fluidised beds at elevated pressures". Fluidisation - Proc. 2nd

Engng. Foundation Conf., Cambridge, UK, p.50 (1978).

8.18 Varadi, T. & Grade, J.R. "High pressure fluidisation in a two-

dimensional bed". Fluidisation - Proc. 2nd Engng. Foundation

Conf., Cambridge, UK, p.55, (1978).

8.19 Guedes de Carvalho, J.R.F., King, D.F. & Harrison, D. "Fluidisation

of fine particles under pressure". Fluidisation - Proc. 2nd Engng.

Foundation Conf., Cambridge, UK, 'p.59, (1978).

8.20 Crowther, M.E. & Whitehead, J.C. "Fluidisation of fine particles

at elevated pressures". Fluidisation - Proc. 2nd Engng. Foundation

Conf., Cambridge, UK, p.6 5 (1978).

8.21 Pettyjohn, E.S. & Christiansen, E.B. Chem. Eng. Progr. 44, 157,

(1948).

8.22 BCURA Interim report. "Initial Experience when Burning Residual

Fuel Oil in the Pressurised Fluidised Bed Combustor at Leatherhead",

(April 1974).

8.23 CSL Report. "Feasibility Study at BCURA Ltd. in Respect of Certain

Test Conditions for Fluidised Combustion of Coal". (May 1976).

8.24 Shaw, T. "Reduction of Air Pollution by the Application of Fluidised

Bed Combustion under Pressure with Special Reference to Emissions of

NOx . Symp. on the Control of Gaseous Sulphur and Nitrogen Compound

Emission, Salford, UK, (April (1976). Instn. Chem. Engrs.

S


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Copy No. *-v

8.25 Pereira, F.T. & Beer, J.M. "A Mathematical Model of NO Formation

and Destruction in Fluidised Combustion of Coal". Fluidisation -

Proc. 2nd Engng. Foundation Conf., Cambridge, UK, p.401, (1978).

8.26 Furusawa, T. et alii. "NO Reduction in an Experimental Fluidised

Bed Coal Combustor"'. Fluidisation - Proc. 2nd Engng. Foundation

Conf., Cambridge, UK, p.314, (1978).

8.27 "Pressurised Fluidised Bed Combustion". US Office of Coal Research

R & D Report No. 85, Interim No. 1, (Sept 1973).

8.28 CSL Report. "A Two Stage Recirculating Fluidised Bed Combustor to

Operate over a Wide Load Range - Cold Model Work". (Jan 1979).

8.29 Combustion Power Co. Inc., "Energy Conversion from Coal Utilising

cpu-400 Technology". Final Report to US Energy Research and

Development Administration under Contract No. EX-76-C-01-1536,

Vol.1, (March 1977).

8.30 Freedman, S.I. & Eustis, J.N. "The Energy Conversion Alternatives

Study (ECAS) of Coal Fueled Utility Power Generation Systems".

US Energy and Research Administration. (1976).

8.31 BCURA Report to BP. "A Note on Oil Distribution in Fluidised Bed

Combustors - A Sumnary of Work carried out at BCURA and Sunbury",

(April 1975).

8.32 BCURA Summary report to BP. "Tests 3 and 4 on the 6 ft 2 Pressurised

Combustor", (Feb 1976).

8.33 NCB, CURL. "Heavy Fuel Oil Burning under Pressure". Report under

contract No. L-SY-101-1-1, (April 1978).

8.34 NCB, CURL. "Tests with an Oil-Atomising Nozzle in a 12 in. diameter

Fluidised Bed Operating at Pressure". Report to BP, (Nov 1978).

8.35 BCURA Report No. FCP 13, (May 1970).

8.36 BCURA Report No. FCP 21, (July 1971).

8.37 British Patent No. 1 397 80c, (18 June 1975). US equivalent 3 859 963.

(14 Jan 1975).

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