Measuring the viscous flow of molten coal ash

Fuel Vol. 75, No. 12, pp. 1476-1479, 1996 Pn: SOOM-2361(96)00129-9

ELSEVIER

Short Measuring

Brian

the viscous

R. Stanmore

flow

and Stuart

Copyright 0 1996 Elsevier Science Ltd Printed in Great Britain. All rights reserved 0016-2361/96 %15.00+0.00

Communication of molten

coal ash

Budd

Department of Chemical Engineering, University of Queensland, 0 4072, Australia (Received 23 June 1995; revised 24 April 1996)

The condition which determines whether ash will cause fouling on heat transfer surfaces in p.c.-fired boilers is that the viscosity must lie between 0.1 and 10 MPa s. Ash under these conditions is stiff, exists in two phases and therefore exhibits a yield stress. A new type of rheometer consisting of a standard thermomechanical analyzer modified to act in the squeeze film mode is proposed to measure ash fluidity. Flat pellets of sample heated to operating temperature under a stream of nitrogen were subjected to a constant compressive load. Analysis of the sample thickness at any point and its rate of deformation provided the Bingham viscosity and yield stress. An ASTM glass at 750°C gave a viscosity near the standard value, while a brown coal fly ash exhibited a yield stress of ~1 kPa at 1190°C. Copyright 0 1996 Elsevier Science Ltd. (Keywords:coal ash, viscousflow; rheometry) In boilers which fire pulverized coal, fouling of heat transfer surfaces is one of the major causes of unplanned outage’. Extensive research into the nature and progress of fouling has identified the flow properties of the ash as a crucial variable during deposition*. To form a fouling deposit, the ash must be sticky enough to adhere to a surface and be incorporated into the ash layer. If the existing ash deposit and the ash particle striking it are ‘solid’, the particle will tend to bounce off. On the other hand, if the particle or the target deposit is fully liquid, the particle will be captured but have sufficient fluidity to flow from the surface as slag. The fluidity or ease of flow of fluids is generally characterized by the viscosity. The viscosities of interest for fouling are reported* as lying between lo5 and 1O’Pas (O.l-10MPas). These values were determined for synthetic glasses that are Newtonian in character. It is likely that ashes nominally in the O.l10 MPa s (Newtonian) range, which are at temperatures above lOOo”C, are in fact two- or multi-phase mixtures. These fluids are not Newtonian, although they have been characterized as such. In particular, a yield stress probably exists, so that the Bingham model would be more appropriate than the Newtonian. That is, the stress generated by shearing is the sum of a plastic component (the yield stress) and a viscous stress component proportional to the shear rate as in Newtonian flow. A Bingham fluid does not flow until a stress exceeding the yield stress is applied:

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where r = shear stress (Pa), r. = yield stress (Pa), 77~= Bingham viscosity (Pas) and y = dw/dx = shear rate (s-l). The differences in behaviour between these two classes of fluid may account for some of the apparent anomalies in the melt/flow behaviour of ashes as demonstrated by standard ash fusion tests. There are two parameters that are sensitive to temperature, rather than one, i.e. yield stress and Bingham viscosity. Differences in their temperature dependence may cause the standard flow temperature to be close to the initial deformation temperature with one coal but to be well removed from it with another. A simple Newtonian description does not allow for such differences. The presence of a yield stress means that some parts of an ash deposit are unsheared, because in those parts the shear stress is not sufficient to overcome the yield stress and so initiate flow. MEASUREMENT VISCOSITY

OF ASH

Two types of rheometer are currently in use to measure ash viscosity, based on a shear field generated either by rotation or by penetration. The rotational instruments are generally standard heads, e.g. Haake, fitted with special rotating elements operated in a furnace. The ‘cup and bob’ is preferred because of simple operation and analysis. By varying the conditions in a rotational viscometer it is possible to fit the Bingham model. For the present application, rotational rheometers suffer from the fact that they are suitable only for relatively mobile fluids. Their maximum capability appears

Fuel 1996 Volume 75 Number 12

to be -0.1 MPa s, which is at the lower limit of viscosities of interest for fouling studies. The common correlations for calculating viscosity, e.g. Watt and Fereday4, and Urbain et al.’ give reasonable predictions at lower viscosities6-* but tend to perform badly as the fluid ash becomes stiffer and crystals form. Watt and Fereday stipulate that their correlation is valid only when crystals are absent, and Watt found no correlation between slag composition and crystallization temperatureg. There is a need for a better instrument to measure ash flow behaviour under the conditions where fouling is imminent. The equipment should be such that it can be adopted as a tool for standard ash testing. It should be a standard laboratory instrument, require only a small sample size, operate at low shear rates on high-viscosity fluids and allow analysis with non-Newtonian models. The squeeze film rheometer fulfils these requirements. The sample to be tested is retained between two horizontal plates and is compressed axially by driving the plates together as depicted in Figure 1. From the magnitude of the compressive force, the sample size and the rate of approach of the plates, the rheometric arameters for the fluid can be deduced PO . The velocity field is nonuniform, with the highest velocities at the outside edge of the plate surfaces. In a fluid with a yield stress, there is a nonsheared plug at the centre. Squeeze film (or parallel plate) rheometers have been used for the rheometry of a range of fluids, including partly cured rubber”, high-polymer viscoelastic fluids12~13,masticated brown coal14 and mudcakei5.

Measuring the viscous flow of molten coal ash: B. R. Stanmore and S. Budd

/ / I I I I I I I

Graphite / plunger

( Graphite /’

C”P

where F = imposed force (Pa), hL = equilibrium plate separation (m) and R = sample radius (m). For the constant-volume case, the value R is substituted from V = rR*h to give

(3)

Sample

I+-/ I ! I , I I 1 Force

Figure 1 Sample arrangement in the squeeze film rheometer

analyser The thermomechanical (TMA) has been applied to rheometry in penetrometer configuration, and is marketed in this form16. It has also been used to measure the viscosity of glasses in a bending-beam configuration17. However, there appears to be no record in the literature of its use as a squeeze film rheometer for coal ash. This paper presents preliminary results which demonstrate that a squeeze film rheometer for molten coal ashes can be based on a commercial TMA.

ANALYSIS OF RHEOMETER RESULTS An analytical interpretation of the performance of squeeze film rheometers is available from a previous paper”. To make the analysis tractable it is assumed that the film is thin: h/R < 0.1, where h is plate separation and R is plate (sample) radius. Then force and material balances give an analytical solution to the equation of motion. By matching the movement under load of the plates in the rheometer to the theoretical solution, values can be assigned to the rheological parameters. The rheometer can be operated as either a constant-radius or a constantvolume instrument. In constant-radius operation the sample is extruded out of the region between the plates, whereas the sample remains between them in the constant-volume situation. Solutions are available for both versions”. Application of the approach to the rheometry of brown coal is described elsewhere14. For any model which includes a yield stress, the value of 7. can be established from a static situation. If the system is allowed to come to equilibrium, in the constant-radius case the yield stress is derived from the limiting height hL by

(2)

These results are not often of practical use because of the long time necessary for equilibrium to be achieved and the fact that any crystals present at small plate separations may obstruct movement. Dynamic results offer a more useful result, and for molten ash the constantvolume situation is more manageable. The form of the solution depends on the ratio of viscous to plastic forces as defined by a dimensionless plasticity number S,: s

=

R(-dh/dh.s h2r,,

”

70 v 3’2 + 2 m( -dh/dt) F=-------

2

V2

nh5

(5) Note that when there is no movement, (-dh/dt) is zero but the expression does not reduce to the equilibrium equation for yield stress. This is because the solution is valid only for large S,, which is not the case for no movement, when S,, is in fact zero. Note also that the above substitution of V in place of R at the limiting-height condition does not strictly apply to flow situations. Experimental deformation data can then be analysed by rearranging Equation (5):

v3’2

.rr’f2F

(-dh/dt) (6)

With Newtonian fluids there is no yield stress and the first term disappears. S, is

(-dh,&)

Since the rate of plate approach is proportional to the fifth power of the separation, there will be a rapid initial closure and a slow final approach to equilibrium. Care must be taken to obtain a good estimate of plate separation throughout. Equation (7) can be integrated with the boundary condition h = hoat time t = 0, to give 1 --h4

1 8rFt hi = w

If the bulk of the flow resistance comes from plastic forces, S,, will be low. This situation is more likely during a squeeze film test. When S, < 0.5, the forcedeformation relation is more complicated than for large S,,:

(4)

where h = plate separation (m) and t = time(s). The value of S, varies with conditions and is zero when there is no movement, as it contains the rate of approach of the plates dh/dt. When it is large (here >OS), the resistance to flow is predominantly viscous, and when it is small the resistance is predominantly plastic. When most of the resistance to flow is due to viscous rather than plastic stresses, the plasticity number is large. If S,, is >0.5, the relation between force and displacement is given by

70

Fh’ = $

&I2 =

= V”2(-dh/dt)rjB 2$/2h5& 0

,1/2h5/2

always infinity for Newtonian fluids, so that the equations apply. Then

2ro V312+ 4(2qB~o)1’2V714 779 37+/2 x

h5/2 1 (9)

(-dh/dt) [

‘I2

Note that 70 appears in both terms, and that the above equations hold strictly only for thin pellets, i.e. h/R < 0.1. Similar correlations can be derived for the constant-radius condition. When analysing TMA data, the following equations were used for manipulating the gradients (denoted as grt) and intercepts (icp) of the required plots: Newtonian,

h5 vs. (-dh/dt):

Bingham (large S,), h5/* vs. (-dh/dt): 6’B=

7-o=

27rF(grt) 3~2

r1’2F(icp) ~312

(11) (12)

Bingham (small S,,), h512 vs. [(-dh/dt)/ h5/2]I/* : 49.rr3J2F2(grt)2 VB =

32ro VII2

(13)

(14) The mean shear rates for the various configurations are given by the following18: Newtonian:

-ym=

--$ [

1

“2(-dh,dt)

Fuel 1996 Volume 75 Number 12

(15)

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Measuring

the viscous flow of molten coal ash: 6. R. Stanmore and S. Budd

Bingham (small S,):

8&’ __@*+...

1(16)

For the values of S, operative in the following tests, the terms in S, were negligible, i.e.

3in =

[

1

V 11.2 (-dh/dt) -9,&’

Figure 2

s (pm s-‘)

Newtonian analysis of NIST 717 glass at 750°C

(17)

The measured force-time-thickness traces from the rheometer were fitted with the analytical solution to obtain a yield stress and a Bingham viscosity.

EXPERIMENTAL The Rigaku TMA was run as standard, with sample deformation measured by a linear displacement transducer (LDVT) located in the frame of the furnace. Discs of graphite were machined to act as retaining plates for the sample as shown in Figure 1. The deformation was applied electromagnetically at 5 and log force, but the value could not be changed during operation. During ash sample preparation, the powder was compacted by slightly wetting with water to assist cohesion and then pressing in a specially constructed mould 4mm in diameter at 20MPa for several minutes to give cylindrical pellets. A previous study into ash fusion behaviour has shown that water or dextrin solution binders have little effect on fusion behaviourlg. After pressing, the pellets were removed from the mould, dried and weighed, and their height was measured with a micrometer. They were stored in airtight containers until use. Most bituminous coal ashes were difficult to press, but brown coal ashes produced strong pellets. The performance of the system as a rheometer was checked using the standard 7 17 borosilicate glass obtained from NIST in the USA. A 3Smm diameter cylinder was cut from the glass block with a rotary saw and then parted to give a number of pellets whose ends were ground to give parallel surfaces. They were then measured for thickness with a micrometer. For a test, the pellet was placed in the retaining system and carefully mounted in the TMA. Loads of 5 or log force were applied with a flow of high-purity nitrogen passing over the sample. The furnace temperature was increased at lOKmin_’ to the test temperature.

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0.204 0

0.005

0.010 5

1

0.015

0.020

-2 -1

[(-dh/dt)flT ] * (mm 4 s 2, Figure 3 Bingham analysis of Loy Yang ash at 1190°C

RESULTS AND DISCUSSION A 905pm thick sample of glass loaded with a 5 g force performed satisfactorily, but only a small section of the trace was usable, as there was a delay while the temperature equilibrated. The plot of h5 against (-dh/dt) at 750°C is shown in Figure 2. The points are fitted with a straight line which passes through the origin, as required for a Newtonian fluid. From the gradient of the line, 455 mm4 s, a viscosity of 0.78 MPa s was calculated from Equation (lo), which compares with the ASTM standard value of 1.29 MPa s. The mean shear rates during the test as calculated from E uation (15) ranged from 1 5 to 2 7 x ~O’S-~ A typical analysis’for a 950 pi thick Loy Yang ash pellet at 1190°C under 5 g force is shown in Figure 3 in the form of a plot of h5’* against [-(dh/dt)/h5’*]“*, which applies to the constant-volume case. It was found that since there was limited axial deformation, the radius was little changed, so that both the constantradius and constant-volume analyses gave very similar results. The yield stress was calculated from Equation (14) to be 980Pa and the Bingham viscosity 0 41 MPa s from Equation 13). The mean’ shear rate was 9 x 10d s-l

Fuel 1996 Volume 75 Number 12

and the mean value of S,, was 0.06. The ash would present a fouling problem at 1190°C. Varying the heating rate and the applied load gave results for the Bingham parameters that could change by a factor of three or four at the same temperature. A test of the Loy Yang ash carried out at 1260°C resulted in the total collapse of the sample. For this ash the initial deformation temperature is 125o”C, the hemispherical temperature 1260°C and the spherical 1290°C. The presence of yield stresses has been demonstrated in all the thermograms of ash in a viscous condition. These appear in the thermograms as asymptotic values for long times and also from the data analysis. Better procedures need to be developed before a reliable comparison can be made between ashes. CONCLUSIONS The application of a Newtonian model to coal ashes in the viscosity range O.l10MPas (equivalent) does not give a description of flow behaviour that is useful for describing flow in a fouling deposit. Ashes exhibit plastic flow in this range and should be fitted with a model which includes a yield stress.

Measuring

Standard TMA instruments can be used to produce thermograms of coal ashes suitable for rheological analysis.

9 10

Watt, J. D. and Fereday, F. J. Inst. Fuel 1969,42,99 Urbain, G., Cambier, F., Deletter, M. and Anseau, M. R. Trans. J. Br. Ceram. Sot. 1981,80, 139 Hough, D. C., Sanyal, A., Annen, K. D., Gruninger, J. H. and Stewart, G. W. J. Inst. Energy, 1986, 59, 71 Jones, E. E. and Lindsey, J. S. Miner. Met. Process. 1987, (Feb.), 60 Jung, B. and Schobert, H. H. Energy Fuels 1992, 6, 387 Watt, J. D. J. Inst. Fuel 1969,42, 131 Covey, G. and Stanmore, B. R.

11

J. Non-Newtonian Fluid Mech. 1981, 8, 249 Scott, J. Trans. Inst. Rubber Ind. 1931,

4 5

6

REFERENCES 1

2

3

Pohl, J. H. In Proceedings, Australian Workshop on Ash Deposition, ACIRL, Brisbane, 1990, Paper 1 Srinivasachar, S., Helble, J. J. and Boni, A. A. In ‘Twenty-third Symposium (International) on Combustion’, The Combustion Institute, Pittsburgh, 1990, p. 1305 Raask, E. ‘Mineral Impurities in Combustion’, Hemisphere, Coal Washington, DC, 1985

the viscous flow of molten coal ash: B. R. Stanmore

7 8

12

7, 1935, 10,481 Brindley, G., Davis, G. M. and Walters, K. J. Non-Newtonian Fluid Mech. 1976, 1, 19

13

14 15

and S. Budd

Leider, P. J. and Bird, R. B. Znd. Eng. Chem. Fundam. 1974, 13, 336; Leider, P. ibid., 342 Covey, G. and Stanmore, B. R. Fuel 1980,59, 123 Sherwood, J. P., Meeten, G. H., Farrow, C. A. and Alderman, N. J. J. Non-Newtonian Fluid Mech. 1991, 39,311 ‘TAS 100 System’, Rigaku Thermal Analyser brochure Jewell, J. M. and Shelby, J. E. J. Am. Ceram. Sot. 1989. 72. 1265

Covey, G. Ph.D. Thesis, University of Melbourne, 1977 Doolan, K. J. ‘Coal Ash Characteristics Related to Ash Fusion TemperaEnd of Grant tures’, NERDDC Report, Project No. 1013, Department of Primary Industries and Energy, Australia, 1991

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