SORBENT DISPERSION TECHNIQUES IN FLUE GAS FOR REMOVAL OF TOXICS

J. Environ. Eng. Manage., 18(5), 319-338 (2008)

SORBENT DISPERSION TECHNIQUES IN FLUE GAS FOR REMOVAL OF TOXICS

Jaikrishnan R. Kadambi,1,* Amirthaganesh Subramanian,1 David Su-Ik Jang,1 Caner U.

Yurteri,1 Vaibhav P. Kadaba,1 Mark P. Wernet 2 and Sudeep Sastry1

1Department of Mechanical and Aerospace Engineering Case Western Reserve University

Cleveland, OH 44106, USA 2National Aeronautics and Space Administration

Glenn Research Center Cleveland, OH 44135, USA

Key Words: Flue gas, laser Doppler anemometry (LDA), particle laden jet, particle image velocimetry (PIV), removal of toxics, sorbent dispersion, two phase flow.

ABSTRACT

Mercury removal from flue gases resulting from coal fired power plants is a prime concern. The goal of the Ohio Coal Research Consortium is to promote the use of Ohio coal by addressing the issues related to the control of trace substance emission, especially mercury, as well as issues relating to sulfur emission control. One mercury removal technology is injection of activated carbon in flue gas. The activated carbon process requires 3 × 10-3 to 10-4 weight ratio of carbon to mercury. Proper contacting i.e. interaction, between carbon and the flue gas is important for the removal of mercury. The uniform distribution of a small amount of sorbent (e.g. for 90,720 kg h-1 of coal, about 90.7 kg h-1 of carbon are required) in the flue gas ducts of cross section as large as 3.0 × 3.6 m is a challenging task. Carbon must be injected into humidified flue gas without collecting on the duct walls; it must mix thoroughly and be retained in the flow for sufficient amount of time for mercury removal. There is a need to identify and evaluate effective contacting methods for carbon/flue gas for the removal of mercury. Since the cost of control of mercury with carbon injection mainly depends upon the amount and type of activated carbon needed, it is important to determine all factors that could minimize the amount of activated carbon needed. The cost of carbon injection can vary by an order of magnitude (from $ 100,000 to > $ 1 million per year in carbon cost for a 500 MW plant not including equipment and installation costs) over the range of type of carbon and rate of carbon injection used. This paper addresses the issues involved in optimal contacting of activated carbon and flue gases so as to obtain the minimum necessary rate of carbon injection, which is an important consideration for making the process economically viable.

*Corresponding author Email: [email protected]

INTRODUCTION

1. General Aspect The removal of mercury from flue gases in coal

fired power plants is a prime concern. Crowell et al. report that the average, minimum and maximum val-ues of mercury in ppm for Ohio coal are 0.204, 0.003 and 1.1 respectively [1]. The average value for mer-cury was the second highest of eastern United States coals surveyed with Pennsylvania coals having the highest average value of 0.276 ppm. US Environ-mental Protection Agency’s scientific advisory board

is expected to come up with the recommendations re-garding control of mercury emissions resulting from flue gases in the near future. Therefore mercury con-trol in flue gas emissions is relevant in the use of Ohio coal. Technology to reliably achieve specific emission mercury reductions is not available. One promising mercury removal technology is activated carbon injec-tion at reduced flue gas temperatures. The activated carbon process requires 3×10-3 to 10-4 weight ratio of carbon to mercury. Proper contacting i.e. interaction, between carbon and the flue gas is important for the removal of mercury with minimal amount of carbon injection so as to make the process economically vi-

320 J. Environ. Eng. Manage., 18(5), 319-338 (2008)

able. The uniform distribution of a small amount of sorbent (e.g. for 90,720 kg h-1 of coal, about 90.7 kg h-1 of carbon are required) in the flue gas ducts of cross section as large as 3.0 × 3.6 m is a challenging task. Carbon must be injected into the humidified flue gas without collecting on the duct walls. Carbon must mix thoroughly and be retained in the flow for sufficient amount of time to remove mercury. There is a need to identify and evaluate effective contacting methods for carbon/flue gas for the removal of mercury. This study involves the development and evaluation of the method(s) of sorbent (activated carbon) injection to obtain uniform dispersion and good mixing with the flue gases for effective removal of mercury and thus ensure minimal requirements of carbon.

A review of the literature undertaken earlier [2] indicated a substantial interest in the use of activated carbon for mercury removal. The range of carbon par-ticle size varied from 17-25 µm. Since the cost of con-trol of mercury with carbon injection mainly depends upon the amount and type of activated carbon needed, it is important to determine all factors that could minimize the amount of activated carbon needed and could enable the use of cheaper activated carbon types. The optimal contacting of the carbon and flue gases (uniform dispersion, mixing and longer retention in the duct) to obtain the minimum necessary rate of car-bon injection is an important consideration for making the process economically viable.

Initially, the sorbent i.e. carbon particles, have to be mixed properly with air so that the air carbon parti-cle mixture can be conveyed into the flue gas duct through a pipe. This is taken care of by making sure that sufficient pipe length (40 diameters) as well as the atomizing air velocity is high enough to have fully de-veloped homogeneous flow prior to injection into the flue gas. The carbon injection system requirements can be stated as: (1) The carbon must be injected into the humidified flue gas without collecting on the duct walls; (2) Fast mixing, long particle residence time in mixing section, with high level of particle scattering and uniform particle distribution in the duct; (3) A simple mechanism, which requires no or very little additional power for operation, ideally uses the energy of the incoming flue gas stream, to excite the mixing layer; and (4) Relatively larger spreading angle and higher entrainment.

Though the activated carbon process requires 3 × 10-3 to 10-4 weight ratio of carbon to mercury the amount of carbon required is still relatively small (amount of mercury can vary from 1-10 μg m-3 in the flue gas). Based upon the amount of carbon to be in-jected and a jet velocity of about 30 m s-1, the mass loading of the nozzle (ratio of mass of carbon in-jected/mass of nozzle air flow) can vary from 0.28 to 1 depending upon the number of nozzles used. How-ever, these numbers depend upon the atomizing air ve-locity as well as the number of nozzles used. In gen-

eral the nozzle mass loading will be light.

2. Two-Phase Flows and Particle Laden Turbulent Jets

Dilute (solid occupying less than 5% of the total

mix volume) two-phase flow consisting of solid parti-cles carried by a turbulent gas flow falls under three different regimes as described by Pai [3]. When parti-cle density and size are large, the carrier has little or no effect on the particles. As the velocity of the carrier is increased, the particles begin to follow the flow but not exactly. In this second phase, the particles have an effect on the flow structure and vice-versa as a result of exchanges in energy and momentum. As the parti-cle size becomes sufficiently small, they follow the streamlines of the carrier phase and can be considered a pseudo-fluid. In this third phase, the particles have a negligible effect on the flow of the carrier. The second regime is of primary interest in the study of particle-laden turbulent jets. It is an inherently complex prob-lem due to the existence of a discrete phase that inter-acts with and in turn affected by the carrier. For either case, the carrier or particles dominate the flow behav-ior, but in between, the flow structure of each phase is unique yet coupled to each other. As opposed to a steady mean flow where sudden changes do not occur, the fluctuating nature of turbulence tends to accentu-ate the separate behavior of each phase.

Marble was one of the earliest researcher to ap-ply modern fluid mechanics to solid-gas flows [4]. He introduced the concept of temperature and diameter of solid particles in the analysis of low solid particle vol-ume flow. Murray [5] was the first to systematically derive the fundamental equations based on continuity of a particle-gas mixture. Soo [6] stressed the signifi-cance of particle size distribution in the fundamental equations of particle-gas mixtures and Vasiliev [7] discussed some results pertaining to turbulent flow of a particle-liquid mixture.

Early experimental work on particle-laden air jets using hot-wire anemometers was carried out by many investigators [8-11] that indicates the centerline velocity decay for the two-phase jet is slower than without particles. Also, the narrower jet profile width was observed as compared to single-phase gas jets.

A key parameter characterizing particle disper-sion in turbulent jets is the Stokes number (St), de-fined as a ratio of the particle response time (tp) to the appropriate time scale of turbulent flow (tf) [12].

f

p

t

tSt =

The particle response time assuming Stokes drag law is defined as:

μρ18

2pp

p

dt =

Kadambi et al.: Carbon Dispersion into Flue Gas 321

where ρp is the particle density, dp is the particle di-ameter, and μ is the fluid viscosity. The proper selec-tion of turbulent flow time is less exact due to the ex-istence of several time scales in a turbulent jet. Flow time should be derived from global measurements of the velocity and length of large-scale structures in the near field. An example would be the comparison of eddy length to its velocity scale. It is also commonly taken as a comparison of jet diameter to jet velocity. Aggarwal suggested that using the dominant fre-quency of the shear flow can lead to the proper flow time scale [12].

Shuen et al. [13] compared three different mod-els for the prediction of the structure of turbulent par-ticle-laden round jets. Models considered were a lo-cally homogenous flow (LHF), deterministic sepa-rated flow (DSF), and stochastic separated flow (SSF). LHF neglects slip between the phases while DSF takes slip into account but neglects the effect of gas fluctuations on particle velocities while SSF considers both. The models were evaluated for particle sizes ranging from 215-1060 μm for mass loadings (ratio of the particle mass flux to the jet mass flux) on the order of 1.0. They found that LHF and DSF over and under-estimated the flow development rates respectively, while the SSF model produced good agreement to previously existing measurements except at high mass loadings. Shuen et al. [14] further concluded that specification of particle properties plays an important role in the prediction of the jet, while gas-phase prop-erties have a secondary role.

Modarress et al. [15] used LDA (Laser Doppler Anemometry) to measure velocity and turbulence quantities of a co-flowing circular jet laden with 50 μm glass spheres at mass loadings of 0.32 and 0.85. It was found that the presence of particles reduces the gas velocity fluctuation as well as Reynolds shear stresses. The level of damping was proportional to the mass loading.

Hishida and Maeda studied the effect of 62 and 132 μm particles on the turbulent structure of a free round jet [16]. Mass loading was held constant at 0.3 and a monochromatic 3-beam 2-D LDA was used to obtain velocities and Reynolds shear stresses. It was observed that the centerline velocity decay for the par-ticle-laden jet was reduced along with the jet spread-ing rate. Smaller particles intensified this effect. Cen-terline particle velocities decreased slower than the re-spective gas velocity, and more so with larger parti-cles. The presence of the particles reduced the Rey-nolds shear stresses in the two-phase jet. The particle velocity shear stress values were found to be an order smaller than the gas-phase.

Park and Chen conducted an investigation using LDA on a confined turbulent jet with and without 40 μm spherical glass beads [17,18]. Low mass loadings of 0.065 and 0.13 were studied, and mean and fluctu-ating velocities were measured for both the solid and

the gas. As found in previous work, the presence of the particles reduced the rate of velocity decay, but it was also found that the recirculation zone was length-ened. They also observed that the fluctuations in the velocity of the particle-laden jet were greatest in the near field and decreased downstream. The increase in mass loading of the particles increased the effect that the particles had on the flow. Also, the gas phase ve-locity was greatest at exit and decreased downstream.

Mostafa et al. [19] conducted an experimental and theoretical study of a free jet laden with 105 μm glass spheres at mass loadings of 0.2 and 1.0. They found that the jet turbulence and width were decreased with the higher mass loading. The introduction of par-ticles created a flatter mean velocity profile at the jet exit, and the gas velocity downstream was found to be greater than the particle velocity.

Barlow and Morrison performed velocity meas-urements in a dense (mass loading of 6.5 and 54), low Re (780), gravity augmented free jet laden with 163 μm spherical glass spheres [20]. The high mass load-ing resulted in reduced slip velocities between the phases, a reduction in fluctuation in the velocities of both phases and decreased jet spreading.

Chinnapalaniandi and Kadambi studied particle-laden jet cocurrent flow interactions. A particle-laden jet facility was designed to specifically study the transport process of sorbent particle injection for in-duct flue gas desulfurization [21]. Phase Doppler in-terferometry and LDA were utilized for measuring particle velocity, size, and concentration. The particle size ranged between 2 and 100 μm, and mass loading was varied between 0 and 10. It was found that pres-ence of the particles decreased the rate of decay of the centerline mean axial gas velocities of the jet. The particle and gas phase axial and radial root mean square fluctuating velocities decreased with increasing mass loading. This resulted in the reduction of the spreading rate in the particle-laden jet. Also, larger particles tended to concentrate around the centerline of the jet.

Sakakibara et al. [22] made velocity measure-ments in the near field of a round jet laden with 55 and 86 μm glass beads. They found that the larger par-ticles more faithfully followed the fluctuation of the fluid. This was attributed to the fact that the mean ve-locity of the larger particles was closer to the convec-tion velocity of the vortex rings in the jet.

Yuu et al. [23] used direct numerical simulation (utilizing a finite difference method) to describe the mechanism of turbulence in gas-particle turbulent jets. Eulerian air velocities and Lagrangian particle trajec-tories were calculated simultaneously and found to be in good agreement with experimental results. Experi-ments were conducted using 58 μm mean diameter glass beads mixed into an air jet (Re = 1700, mass loading = 0.6) produced by an 8 mm diameter nozzle and confined in a 295 × 295 × 1000 mm vessel. LDA


was used to measure both the particle velocities in the particle-laden jet and the gas velocities of the single phase jet. The gas phase was seeded with 0.8 μm NH4Cl particles. It was found that the existence of the beads caused a decrease in the momentum spread and an attenuation of turbulence, evident by a narrower width of the axial velocity jet profile for the particle laden jet. The turbulent intensities of the particle laden jet were 20-50% lower than for the single-phase air jet.

Prevost et al. [24] performed phase Doppler anemometry measurements of fluid/particle correlated motion in the far field of a round jet. A fully devel-oped vertical jet was produced with a Re number 13,100 corresponding to a exit velocity of 20 m s-1. The jet was laden with poly-dispersed glass beads with 60% of the beads ranging in diameter from 10 to 50 μm. Tracer particles were used to simultaneously measure gas phase and particle velocities. Particle ve-locities were divided into four size classes: 10-15, 20-25, 30-35 and 40-45 μm in order to investigate the ef-fect of size on particle motion. The Stokes number was calculated using the Stokesian particle relaxation time as in Aggarwal, and the characteristic time of the jet tj was defined as:

',cf

j uLt =

where L is the large eddy length scale and u’f,c is the standard deviation of the centerline axial fluid velocity fluctuation [12]. As a first approximation beyond 10 nozzle diameters downstream, L was taken as the half width of the jet, L = 0.1x as in Dimotakis et al. [25]. Noticeable relative velocity between the particle and fluid velocities was found at the nozzle exit due to flow-particle interaction in the tube and the presence of rebounds of the particles along the wall of the tube. Planar visualization of the nozzle exit flow supported this observation. This effect decreased the axial flow velocity of the particles (relative to the gas flow) while increasing the fluctuation in the velocity. Mean axial velocity for all particle classes was accelerated by the gas reaching a maximum at various down-stream locations depending on the size class and then decreased reaching the gas velocity near 40 diameters downstream. The same behavior was shown for the fluctuation of the axial velocity. This showed that for each size class, there was a correlation with the value of the Stokes number.

Yurteri undertook an investigation of transport processes involved in flue gas desulfurization [26]. The work was continuation of the investigation by Chinnapalaniandi and Kadambi [21]. LDA and phase Doppler anemometry were used to measure particle size and velocity. A particle-laden jet was generated by a 7.3 mm (inner diameter, d0) nozzle using 50 μm glass beads at a mass loading of 1.66 and 2.5. Concur-rent flow average velocity in a 88.6 × 88.6 mm square

duct was varied between 1.66 and 15 m s-1. It was found that the larger mass loading results in a reduc-tion of the jet spread angle. Jet spread turbulence in-tensity levels were increased with decreasing concur-rent flow velocity. The axial and radial turbulent in-tensity levels were reduced by the presence of parti-cles when compared to the single-phase jet. Recircula-tion was observed for low concurrent velocities.

3. Jet Dispersion and Mixing Techniques

Many different techniques offer mechanisms for

the enhancement of turbulence that lead to an increase in mixing levels. Mixing is characterized by jet growth and entrainment rates, which can be related to velocity profiles and turbulence. Studies have shown that under certain conditions, particles will tend to concentrate in areas of high strain, and generation or suppression of turbulence in the fluid has a profound effect on the mixing effectiveness and particle disper-sion of a particle laden flow [27].

Yuu et al. [28] investigated the turbulent diffu-sion mechanism of a round air jet laden with 20 μm fly ash. Higher Stokes number (based on nozzle di-ameter and exit velocity) resulted in a decrease in the particle diffusivity. It was found that, in general, the particle diffusivity was smaller than for the fluid par-ticle. It was concluded that Stokes number and large eddies play an important role in particle transport properties.

Subramanian and Ganesh used a double concen-tric jet to enhance entrainment [29]. Particles were fed into the outer annulus jet while the central jet re-mained single-phase. Entrainment rates were to be found greater than without particles. This was ex-plained by the larger spanwise coverage of the con-centric jet. In previous work, they found that particles introduced into the central flow decreased the en-trainment rate.

Crowe et al. [30] reviews important issues in-volving particle mixing in free shear flow. Based upon a thorough examination of previous experimental and numerical work, it was concluded that particle disper-sion is controlled by the motion of large-scale vortex structures. Brown and Roshko first identified large-scale vortex structures in characterizing free shear flows [31]. Crowe et al. [30] reported (based on nu-merical modeling) that for small Stokes numbers, the particle dispersion follows the same rate as the fluid, but for larger particles (0.1 < St < 1.0), the dispersion can be greater than the fluid. In addition, particles with large Stokes number remain unaffected by the turbulence. They found that in a majority of experi-ments, the data for particle diffusivity were reported in terms of the particle Schmidt number:

p

ep D

Scν=


where νe is the effective kinematic viscosity of the fluid, and Dp is the particle diffusion coefficient.

Launder and Spalding reported that the Scp for tracer particles in a round jet is 0.70 and 0.50 for a square jet [32]. Values obtained less than these would indicate that larger particles would have dispersion levels greater than the fluid particle. Crowe et al. [30] found that lower values were obtained in other studies [29,33,34] when compared to those of tracer particles. A conceptual model for this occurrence has been de-scribed by Singamsetti [35]. Particles could be en-trapped in large-scale structures until centrifuged be-yond the structure. This would cause larger particles to be dispersed more readily than smaller ones. If this were the case, higher concentrations of larger particles would be located off the jet axis while smaller ones would remain in the core. Kobayashi et al. [36] gath-ered data where dense particles exhibited a pro-nounced tendency to have velocities directed away from the jet axis opposing the turbulent entrainment.

Quinn investigated turbulent free jets produced by rectangular shaped slots [37]. It was found that near-field mixing is increased and the greatest mixing occurred when the potential core was the shortest. Longmire and Eaton used open-loop forcing to control the structure of a particle-laden round jet [38]. By do-ing so, they were able to vary the vortex length scale and change particle dispersion. Jet (Re = 1.9 × 104) with 55 μm size particles. Flow visualization was im-plemented through glycerin smoke and a copper vapor laser. images were taken by a 35 mm camera and digi-tized into a computer for number density analysis. Mean velocity components were measured using both LDA and hot-wire. It was noted that particles were ejected beyond the turbulent structures often never to return. This was the dominant mechanism for disper-sion for the natural jet. When forcing was applied (via an acoustic speaker between the range of Strouhal number StD = 0.3 to 1.1), the nozzle exit distribution of particles was altered by an increase in radial spread. A particle moment function Dx was used to quantify radial dispersion:

∫=2

1

r

rx crdrD

where c is the local number density, and r is the radius. The analysis held the lower limit r1 at 0.5 rather than 0 in an attempt to exclude the particles clustered near the jet axis and r2 was chosen as the outer limit of the jet. The dispersion of the particles with the application of forcing greatly exceeded the case without. This was attributed to the early generation of vortices larger than those of the natural jet. Mean axial velocity pro-files for both cases possessed similar shapes. The forced jet showed only a slight decrease in velocity near the nozzle exit. Radial velocity for the forced jet showed a significant increase in outward direction be-yond one nozzle diameter downstream. Near nozzle

exit, both cases exhibited similar profiles with inward radial velocity. When a second sub-harmonic fre-quency was superimposed onto the fundamental, con-trol of vortex pairing was achieved. This allowed the formation of denser particle clusters in the jet as well as a delay of large dispersion to further downstream locations.

Raman used unsteady fluid mass addition to en-hance the mixing of a single-phase rectangular jet [39]. The frequency, magnitude and phase of opposing transverse jets were controlled in an attempt to excite the large scale rectangular jet and control shear flow. A mass addition of 12% resulted in a 35% reduction in potential core length and a 60% increase in normal-ized mass flux.

Fan et al. [40] measured particle concentration and size distributions of particle-laden turbulent coax-ial jets using laser diffraction and a tomography data transform technique. Gas phase dispersion was greater than particle dispersion, and particle dispersion de-creased with increasing particle size. An increase of velocity ratio served to increase the gas dispersion. Mean particle size at the jet edge was found 15-20% smaller than in the jet centerline.

Swirl flows have been given a considerable amount of attention due to their common occurrence among many engineering applications. Some exam-ples are vortex shedding from aircraft wings, cyclone separators, and primary and secondary flows in tur-bomachinery. Limited information is available on the effects of a swirl velocity on a particle-laden jet for the enhancement of mixing. Chigier et al. [41] showed that the empirical constants describing the decay of velocity and the spread of the jet were functions of the degree of swirl. Trabold et al. [42] showed that there is a tendency for a higher entrainment rate when the length of a tube type nozzle was reasonably small. Also, in the nozzle exit region, entrainment was sensi-tive to the geometry of the exit. Mehta et al. [43] stud-ied the effect of different rates of swirl on the stream-wise development of the mixing layer. Increased swirl showed an increase in Reynolds shear stress levels. Farokhi et al. [44,45] found that the introduction of swirl leads to an increase in the decay of the mean centerline velocity (Udc), an indication of enhanced mixing.

De la Rosa et al. [46] performed an experimental study on the effect of swirl on the velocity and turbu-lence of a liquid spray. Recirculation in the core of the swirling field was found for high swirl number and not found for medium and low swirl number. Plots of kinetic energy, normal Reynolds stresses, and Rey-nolds shear stresses showed double-peak radial distri-butions indicating regions of high energy or shear. The double peak distributions decayed downstream. Normal Reynolds stresses were found to be signifi-cantly greater than the Reynolds shear stresses requir-ing the consideration of all three components of nor-


mal stress in any turbulent swirling spray. Park et al. [47] investigated the rate of entrain-

ment of swirling jets. Weak swirling jets showed little increase in entrainment while strong swirling jets pro-duced increases of up to 400%. Cutler et al. [48] found that turbulence generated from vortex break-down provides a second mechanism for the increase of mixing.

Park et al. [49] performed numerical simulations of droplet laden swirling flows. For a single-phase swirling jet at Re = 800, it was found that the dynam-ics of large-scale structures were strongly affected by the degree of swirl. Low and intermediate levels of swirl produced vortex rings to roll up closer to the nozzle exit. Frequency and pairing interactions be-came progressively stronger as swirl number was in-creased. Strong swirl increased the jet spread angle while weaker swirl enhanced the effects of vortex structures and shear layer growth and entrainment.

The effect of tabs on high speed jets has been re-searched intensively by Zaman et al. as a means of vortex generators [50]. Tabs that protrude into the jet at the nozzle exit produce distortion of the jet. Tabs can be different sizes, shapes, and at various angles of incidence. They can also be placed in different con-figurations. The streamwise vorticity created by tabs can significantly increase mixing. It was also found that for most combinations of tabs on a circular nozzle, the mass fluxes downstream were much higher than that for asymmetric nozzles. The most effective tab was the triangular-shaped delta tab. Four such tabs equi-spaced on the nozzle exit resulted in significant downstream jet mixing. Two sources of streamwise vorticity were postulated. One was due to the uphill pressure rise generated by the tab, and the other was the vortex filaments shed from the sides of the tabs. The two effects seem to add in the case of the triangu-lar tab.

Zaman has also investigated the effect of delta tabs on asymmetric nozzles for an under expanded jet at a Mach number of 1.63 [51]. The asymmetric noz-zle was found to produce a jet with greater mixing (indicated by greater downstream mass fluxes) than for a circular nozzle. However, it was also found that a circular nozzle with four delta tabs produced greater mass fluxes when compared to the asymmetric noz-zles with and without delta tabs. Delta tabs combined with asymmetric nozzles proved to only provide a slight increase in mixing. Only for the case when two large delta tabs placed on the narrow sides of a 3:1 ra-tio rectangular nozzle were higher mixing levels ob-tained. Under this case, the jet width fanned out at a large angle after undergoing one axis switch. It was also noted that the axis switching phenomenon could be affected by the delta tabs.

Zaman et al. [52] compared jet spreading and en-trainment characteristics experimentally for several asymmetric nozzles as well as for a asymmetric noz-

zle and axisymmetric nozzle with delta tabs for a jet Mach number range of 0.3-1.8. At subsonic conditions, the jet spreading for the asymmetric nozzle was found slightly greater than the axisymmetric one. At super-sonic conditions, the onset of screech generated much faster spreading for the asymmetric nozzle. Jet spread-ing for the nozzles with tabs was found much greater at both subsonic and supersonic conditions than for any other nozzle geometry.

EXPERIMENTAL METHODS

1. Objectives

In this paper, methods of activated carbon injec-

tion to obtain uniform dispersion, good mixing with the flue gases to enhance removal of mercury are dis-cussed. The tests were conducted in the particle injec-tion test facility in Case Western Reserve University (CWRU). The flow conditions encountered in the flue gas ducts in power plants can be simulated in this fa-cility. Non-intrusive laser based Particle Image Ve-locimetry (PIV) and LDV were used for obtaining the characteristics of the nozzles. The nozzle designs are chosen to enhance the dispersion and mixing charac-teristics of a particle-laden jet to the concurrent flow. Nozzles with innovative design to obtain uniform car-bon dispersion and improved sorbent (carbon)-flue gas mixing will be developed and evaluated.

2. The Particle-Laden Jet Facility

An experimental study of turbulent particle-

laden jets in concurrent flow was conducted using the Particle Laden Jet Facility located in the Laser Flow Diagnostics Laboratory, CWRU. Equipment used for included a Dantec Inc. 60x FiberFlow Series Phase Doppler Anemometry system, an in-house Nd:YAG pulsed laser PIV system, and a concurrent flow parti-cle-laden jet facility. The test set-up was described in Jang [53].

The study of a particle-laden jet concurrent flow requires the conditions of uniformity in both the parti-cle-air jet and the concurrent duct flow and fully de-veloped turbulent flows upon interaction. These re-quirements were satisfied by the particle-laden jet fa-cility (Fig. 1). The setup allowed the measurement of both single phase and particle-laden jets and takes into account the geometric, kinematic, and dynamic simi-larity relationships of a full-scale flue duct system [20,26]. The setup has undergone modification for specific experiments with the current setup described for this experiment. The main flow was directed through a clear Plexiglas duct of an overall length of 341 cm and a square cross-section of 8.86 cm. The test section was located 243 cm from the entry of the duct and has a length of 38 cm. It was machined from 12.7 mm thick Plexiglas sheets with a removable 4.37


1. REGULATOR2. PRESSURE GAUGE3. FLOWMETER4. VENTURI INDUCTOR5. PARTICLE RESERVOIR6. CONTROL VALVE

10

9 8 7

6

5

4

2 3

1

1

2

2

2

1

111

7. TEST NOZZLE 8. TEST SECTION 9. SQUARE DUCT10. CYCLON SEPARATOR11. SEEDING GENERATOR

TO COMPRESSORS

Fig. 1. Schematic of particle-laden jet facility.

Fig. 2. Test Nozzle Configurations. mm thick optically clear front window.

Particles were introduced into the nozzle through a venture inductor and a screw feeder. A pitot-static tube was located in the center of the duct 105 cm downstream of the duct entrance to measure the Udc of the duct. The relationship between Udc and mean ve-locity (Ud) for a fully developed turbulent flow in a smooth square duct was:

dcd UU =20.1

The duct flow temperature is approximately 300 K but can vary slightly depending on the temperature of the laboratory and outside environment. Cyclone seeding generators were used to seed the main flow as well as the jet flow. The main flow seeder was located 130 cm upstream of the test section. Talcum powder (particle size 1-3 μm) was used as the seeding agent.

The two component of LDA consisted of a 4 Watt Lexel model 95 Argon ion laser, Dantec 60 × 40 series fiber-flow transmitter, 60 × 11 2-D fiber optic probe, 57 × 10 backscatter mode receiving optics, 55 × 80 photomultiplier tubes, 58 × 10 covariance processor, Pentium 200 MMX computer. The fiber optic probe was mounted on a 3-D traverse with a resolution of 0.0254 mm. Average of 3000 velocity measurements was used at each measurement location.

A PIV system developed in house at Laser Diag-nostics Laboratory of CWRU consisted of New Wave Research Co. MiniLase III-15 50 mJ (5 ns)-1 pulse double pulsing Nd:YAG laser, which generated a 3.5 mm diameter beam with a wavelength of 532 nm. The laser beam was formed into a 0.093 mm thick sheet of light using a combination effect of cylindrical and spherical lens (focal length 500 mm) arrangement, and

Table 1. Specifications for Delta Tab Nozzle (DTN)

No. Number of tabs for

equally spaced Tab Angle φ

(º) Size (mm)

1 3 135 2 × 2 × 2 2 4 135 2 × 2 × 2 3 4 135 4 × 4 × 4 4 4 90 2 × 2 × 2

a planar section of the seeded flow was illuminated. Double exposure images of the particles were re-corded using a Kodak Megaplus 1.4 CCD camera and EPIX 4-MEG frame grabber board and stored in a PC. The CCD camera has a resolution of 1320 × 1040 pix-els. Each pixel is 6.8 μm square with 100% fill factor. The camera and illumination system were mounted on a traverse (0.0254 mm resolution). Two-dimensional instantaneous velocity fields were obtained from the PIV system. The acquired double exposure images were processed by the auto-correlation data reduction technique, which was enhanced with a fuzzy logic al-gorithm in order to optimize the quality of the proc-essed velocity vector maps. The auto-correlation soft-ware performs the correlation of the images in the fre-quency domain.

Various nozzle geometries were selected along with a simple base nozzle based upon the work of Ka-dambi et al. [54]. The six nozzles investigated are de-scribed as follows. The base or simple nozzle (SPN, Fig. 2) consists of a round copper tube with d0 of 7.3 mm. The nozzle consisted of one 90° bend and then a straight 450 mm section prior to exit. The remaining five nozzles were modifications of the simple nozzle and their basic geometry is shown in the figure. The asymmetric jet was generated by an asymmetric noz-zle with an elliptical exit. The circumference of the jet was the same as for the simple jet, and the exit aspect ratio is 5:9. Four different delta tab nozzles (DTN) consist of triangular tabs in various configurations at the nozzle exit (Table 1). The main parameters of the DTN are the tab size, number and angle. In order to determine the bulk effect of these parameters, each one was varied independently. Each nozzle consists of a parameter that is different, yet has something in common with another. Thus, a comparison is possible while minimizing the number of nozzles necessary. The delta tabs were hand sheared from metal sheets and located by the assistance of jigs on the nozzle exit area.

3. Test Conditions and Preliminary Tests

The test conditions are provided in Table 2. Av-

φ


Table 2. Particle-laden jet test conditions

Test items Symbol Conditions Duct Overall Length 3410 mm Duct Length to Test Section 2430 mm

Duct Size Cross Section 88.6 mm × 88.6 mm square

Duct Average Flow Velocity Ud 15 m s-1

Duct Flow Reynolds Number Red 88,600

Duct Flow Temperature < 300 K

Nozzle Inner Diameter do 7.3 mm

Nozzle straight section length 450 mm

Nozzle Average Gas Exit Velocity Uj 30 m s-1 Jet Reynolds Number Rej 14,600 Jet air mass flux 1.47 × 10-3 kg s-1 Minimum nozzle straignt section gas or particle residence time 0.015 s Particle Type Engineering Class Glass Beads Particle Manufacturer Cataphote Inc.

Mean Particle Size dp 49.9 μm

Particle Size Standard Deviation 13.6 μm Particle Density ρp 2450 kg m-3 Index of Refraction n 1.51

Mean Mass Loading m 1.7 Particle mass flux at jet exit 2.49 × 10-3 kg s-1 Particle concentration at jet exit

Volumetric Number/volume

0.0807 CV% 38.8 particles mm-3

Stokesian Particle Time constant tp 0.01890 s

erage duct velocity and average jet velocity were held constant for all tests at 15 and 30 m s-1 respectively. For the single phase experiments, the main duct flow and the jet were seeded with tracer particles (talcum powder). For experiments with the particle-laden jet, no artificial tracer particles are introduced. Only the particle dynamics were measured. Experiments with particle injection were performed with a mass loading of 1.7. Spherical glass beads were used for particle in-jection, and Microtrac test results for the particles re-port a mean size of 49.9 μm and a standard deviation of 13.6 μm.

Preliminary measurements were conducted using the parameters listed in Table 2 for the simple nozzle using LDA in order to verify the proper operation of the facility. Talcum powder was used to seed both the main flow and the jet flow. No particles were intro-duced into the jet. The jet was traversed across its horizontal span at an axial location, 10 nozzle diame-ters downstream of the nozzle exit. The resulting ve-locity profile is given in Fig. 3. The z-position indi-cated the off-axis span wise direction and the jet was normalized with respect to Uc. The velocity was sym-metric about the center axis. Away from the centerline of the jet, the velocity approaches the duct flow veloc-ity.

Preliminary measurements of the round turbulent jet produced by a simple nozzle verified the operation of the facility. The flow conditions met the desired

1.2

1

0.8

0.6

0.4

0.2

0-2 -1.5 -1 -0.5 0 0.5 1 1.5 2

U/U

c

Z-Position (z/do ) Fig. 3. Mean axial fluid velocity profile for the simple

nozzle at 10 nozzle diameters downstream. behavior for the study. The interaction of particles with the free stream flow was studied by the introduc-tion of particles to the jet using various nozzle designs.

RESULTS AND DISCUSSIONS

1. Particle Laden Jet Concurrent Flow Interaction

Test on the six nozzle designs (Table 3) were

conducted in the particle laden jet facility as per test conditions outlined in Table 2. LDA measurements were made at six different downstream locations (x/d0 = 0.17, 5, 10, 15, 20 and 25). At each location, the probe was traversed in the spanwise direction to pro-


Table 3: Summary of test nozzles and their corresponding symbols

No. Test nozzle Symbol Shap

1 SIMPLE NOZZLE SPN

2 ASYMMETRIC NOZZLE ASN

3 DELTA TAB NOZZLE 1 DTN1




vide a profile of the jet properties. Spatial resolution across the jet was 1.27 mm near the jet centerline and was increased up to 5.08 mm near the outskirts of the jet envelope. 3000 velocity samples with accuracy better than 1% were acquired at each point location. Mean values of axial (u), vertical/tangential and/or ra-dial (v) velocity were obtained with an accuracy better than 1%. Particle laden jet concurrent flow interac-tions (mixing properties) are also affected by the aver-age turbulent intensity, Reynolds shear stress, jet half-width, and axial centerline velocity decay rates.

The jet half-width is defined as the distance across the jet, where the jet velocity is half of the cen-terline velocity. The rate of centerline mean velocity decay is represented as the slope of (u0/uc)

2 where u0 is the centerline jet exit velocity and uc is the center-line jet velocity.

Turbulence makes the behavior of the flow diffi- cult to predict, and the existence of two separate phases introduces additional complications. The particles may have different velocity characteristics from the carrier gas. Smaller particles will more readily mimic the flow of the gas. One may also deduce that turbulence produces local accelerations in the flow and that particle inertia is important. Heavier particles need greater time to accelerate them to surrounding gas flow values. The Stokes number is an important parameter in the classification of particle behavior in particle laden turbulent jet. It is the ratio of tp to tf. The tp may be found assuming Stokes drag law [12,24]. The 50 μm mean diameter particles used in this study have a response time of 0.0189 s. Earlier investigations discussed in initial sections of this paper provide substantial evidence of the effect of large- scale turbulent structures in governing the behavior of particle dispersion. Therefore, the selection of a tf

should reflect this. The characteristic time is com-

Fig. 4. Mean axial velocity profiles for a turbulent round

jet in concurrent flow(a) fluid, (b) particle. monly taken as the ratio of jet diameter to jet velocity [12,24] compared jet half width to the standard devia-tion of the fluctuating component of the axial velocity. Stokes number was found to decrease as the particles moved further and further downstream for a turbulent round jet. The current study computes the turbulent time scale as such.

2. Turbulent Round Jet

LDA measurements were made in the simple

nozzle with and without particles. The Re number for the jet and the duct flow were 14,600 and 88,600 re-spectively. Spherical glass beads (50 μm mean diame-ter) were used for the particle-laden jet with a mass loading of 1.7. Tracer particles were used to measure the fluid velocities for the single-phase gas jet, and only the glass particle velocities were measured in the particle-laden jet. Velocity measurements were made across the horizontal span (z-direction) of the jet at six downstream locations, x/d0 = 0.35, 5, 10, 15, 20 and 25. Mean axial and vertical velocity components were sampled simultaneously and determined from 3000 data points. The mean axial fluid velocity profiles for the simple nozzle are shown in Fig. 4a. The velocity is normalized with respect to the maximum nozzle exit velocity (Uo). The velocity profile is narrowest near the nozzle exit. The velocity at the outer span of the jet approaches the concurrent flow velocity. The mean axial particle velocities are shown in Fig. 4b. The ve-locity is normalized by the maximum nozzle exit ve-locity of the gas phase jet. The pre-exit velocity of the gas phase for the particle-laden jet was maintained at the same value as for the single-phase jet. Even so, the particle exit velocity is less than 90% of the single-


Fig. 5. Fluid and Particle-phase mean axial centerline

velocity evolution for a turbulent round jet in concurrent flow.

phase jet velocity. The addition of particles tends to decrease the velocity of the gas phase due to momen-tum exchange. In addition, the particles flowing through a tube will interact with the walls which tend to decelerate the particles. The fluid then accelerates the particle until it interacts with the wall again. This effect tends to produce fluid velocities, which are greater than the particle velocities as observed by Prevost et al. [24]. Although the fluid and particle ve-locities in a particle laden jet were not measured, this effect can be observed in Fig. 5. The fluid velocity is decelerated while the particles are accelerated and reach a maximum near x/d0 = 5. This suggests that the fluid velocities for the particle-laden jet are greater than the particle velocities at exit conditions. The lower two curves show the square inverse of the first two curves. For a turbulent round jet, the curve tends to be linear downstream from the nozzle exit (> five nozzle diameters). The slope of the curve is a measure of the rate of decay of the centerline velocity. The centerline velocity decay for the particles is lower than the single phase jet near the nozzle exit, however, far downstream (x/d0 = 25) the slopes of both curves are similar.

The tendency of the particles to dampen the tur-bulence characteristics of the jet can be observed in Figs. 6a and 6b. The turbulence levels of the particles are lower than the values for the single phase jet. However, a decrease in turbulence is observed as the jet moves downstream for both cases. In addition, the particle turbulence levels near the nozzle exit are greater than the fluid levels only near the jet centerline. Near the nozzle boundary, the turbulence levels for the fluid increase dramatically and remain elevated downstream.

3. Particle Mean Velocity Profiles

Fig. 6. Mean axial velocity profiles for a turbulent round

jet in concurrent flow (a) fluid, (b) particle. Mean velocity vectors measured at various

points along the jet for each of the test nozzles are presented in 3-dimenstional space in Figs. 7-12. The flow direction is from right to left with the nozzle exit centered at the origin (0,0,0). All three axes are nor-malized with respect to the nozzle diameter. Only the velocity components in the x-y plane are presented. The figures illustrate the spatial evolution of the jet velocity.

Velocity vectors for the horizontal center plane for the simple nozzle are shown in Fig. 7. The velocity field near the edge of the jet becomes increasingly uniform conforming to the concurrent free stream. Figure 8 shows the velocity field generated by the asymmetric nozzle. The nozzle’s major axis is aligned along the y-direction. The flow in the vertical plane is directed away from the centerline of the jet and pos-sesses a wider spatial distribution of high velocity. The axial velocity of the particles is accelerating upon exiting the jet. The velocity field in the horizontal cen-ter plane is similar to that for the simple nozzle.

The flow corresponding to the DTN1 is shown in Fig. 9. The nozzle is oriented with one tab located on the positive y-axis. The tab has a dramatic effect on the velocity vectors near the nozzle exit especially near the tab. Similar effect is also observed for the DTN2 (Fig. 10), which has two delta tabs located on each x and y axis, and again, the tabs direct the flow away from the center of the jet near the nozzle exit. This suggests the particles may come in physical con-tact with the tabs and deflect through the jet. Figure 11 shows the results of DTN3 which has a larger tab size. The larger tabs produce more turning near the nozzle exit and particles are directed away from the jet.


Fig. 7. Particle-phase mean velocity for simple nozzle

(SPN).

Fig. 8. Particle-phase mean velocity for asymmetric

nozzle (ASN).

Fig. 9. Particle-phase mean velocity for delta tab nozzle

1 (DTN1).

DTN4 nozzle (Fig. 12) uses the same small delta tabs as for the nozzles DTN1 and DTN2 (Figs. 9 and 10) but with a greater incidence angle. The effect is to de-crease the magnitude of the velocity of the outwardly directed particles when compared with Fig. 10. Near the nozzle exit, the physical deflection of the particles may serve as a significant mechanism for particle dis-


2 (DTN2).


3 (DTN3).


4 (DTN4).

persion.

4. Turbulent Intensity The variation of turbulent intensity for the axial

velocity along the horizontal span of the duct at vari-ous axial locations is plotted in Figs. 13a-13f for the six nozzles. An examination of the figures indicate that nozzle DTN3 (Fig. 13e) generates the highest


Fig. 13. Particle-phase axial velocity turbulent intensity

profiles across the horizontal center plane.

Fig. 14. Particle-phase axial velocity turbulent intensity

profiles across the vertical center plane. level of turbulence along the horizontal duct span. Based upon this figure, the DTNs and the asymmetric nozzles are of interest from the point of view of mix-ing.

Figures 14a-14e shows the variation in the turbu-lent intensity for the axial velocity along the central vertical span (plane) for the asymmetric nozzle and the four DTNs. Again as can be observed from this figure, DTN3 (Fig. 14d) generates the highest level of turbulence along the central vertical plane.

The vertical (tangential) velocity turbulent inten-sity profiles across the horizontal center plane are given in Fig. 15. The turbulence levels for the swirl and DTNs are greater than that for the simple and asymmetric nozzles. Also, near the nozzle exit, there are highly elevated turbulence levels. DTN3 shows turbulence levels that increase as one moves down-

Fig. 15. Particle-phase vertical (tangential) velocity

turbulent intensity profiles across the horizontal center plane

Fig. 16. Particle-phase radial velocity turbulent intensity

profiles across the vertical center plane stream (Fig. 15e). Overall, values are greater than the other DTNs.

Radial velocity turbulent intensity across the ver-tical center place is given in Figs. 16a-16e for the asymmetric and the four DTNs. Similar to the obser-vations for turbulence intensities for the axial and ver-tical velocities, the DTNs show the higher turbulence levels. Figures. 16b-16e show increasing turbulence with downstream location. Also, the turbulence in-creases away from the jet centerline. These nozzles also show highly elevated levels of turbulence near the nozzle exit. Among these nozzles, nozzle DTN3 has the highest turbulence levels in the radial velocity.

5. Reynolds Shear Stress

Reynolds shear stresses across the horizontal


Fig. 17. Particle-phase Reynolds shear stress profiles at 6 downstream locations across the jet horizontal center plane:

(a) simple nozzle (SPN), (b) asymmetric nozzle (ASN), (c) DTN1, (d) DTN2, (e) DTN3, (f) DTN4. center plane are presented in Fig. 17. The shear pro-files for the simple nozzle (Fig. 17a) are nearly zero except for near the nozzle exit. Figure 17b shows the Reynolds stresses for the asymmetric nozzle. The pro-file is not symmetric, but instead, the stresses are a maximum on one side of the jet. The shear stresses for the DTN1 are presented in Fig. 17c. The figure shows the profiles spanning between negative and positive shear stresses. This behavior DTN2 is similar. DTN3, the nozzle with the largest tabs, produces shear stresses that have the greatest spread. Figure 17f for DTN4 also shows the same trend as with the first two DTNs.

The Reynolds shear stresses across the vertical center plane are presented in Fig. 18. The DTN1 (Fig. 18b) shows shear stresses that rise near the centerline and fall on the extremes. Figure 17c shows a similar profile for the DTN2. The profiles for the DTN3 (Fig. 18d) show large positive values of stress on one side of the jet. The stresses dissipate downstream.

6. Digital Particle Image Velocimetry Measurements

The PIV technique was used to map the flow for

seven of the nozzles under the same test conditions as in the LDA measurements defined by in Table 2. PIV autocorrelation technique was used to obtain the ve-

locity vectors and concentration estimates. Double ex-posure images were acquired. A field of view of 26.9 × 21.1 mm was used. Images of the jet were acquired along the jet centerline at the nozzle exit and at 17.4 nozzle diameters downstream. Subregion size was se-lected as 128 pixels. The flow field spatial resolution was approximately 2.6 mm. Concentration levels were found by a simple particle count for each subregion. PIV setup parameters and conditions are given in Table 4. The expected image particle size (de) corre-sponds to the imaged particle size including the ef-fects of the point-spread function. Details of the PIV setup are provided in Jang [53].

7. PIV Results

The PIV results are discussed in this section.

Figure 19 compares the axial velocity profiles ob-tained using LDA and PIV for the simple nozzle 20 nozzle diameters downstream of the nozzle exit. The two profiles have approximately the same magnitude and shape. The results from PIV differ slightly be-cause the instantaneous velocity (rather than the mean) is obtained. This provided a calibration of the accu-racy of the PIV system.

Figures 20-24 show instantaneous velocity vec-tor and concentration maps for the jet flow for a single


Fig. 18. Particle-phase Reynolds shear stress profiles at 6 downstream locations across the jet vertical center plane(a)

asymmetric nozzle (ASN), (b) DTN1, (c) DTN2, (d) DTN3, (e) DTN4. Table 4. PIV Setup Parameters

Parameter Symbol Value CCD Array size 1320 × 1040 pixels

9.0 × 7.0 mm Pixel size 6.8 μm (square) 8-bit gray level 0-255 intensity grayscale values Exposure Time 100 ms Illumination Wavelength λ 0.532 μm Pulse Duration 5 ns Output beam Diameter (e-2 width) 3.5 mm Spherical lens focal length F 500 mm Minimum light sheet thickness 0.092 mm Energy Output 50 mJ pulse-1 Double Pulsing frequency 10 Hz Time Delay Between pulses 15-21 μs Magnification factor M 0.333 f-number f/# 8-220 Subregion size 128 pixels Flow field spatial resolution 2.6 mm Expected imaged particle size De 3.19-6.11 pixels Maximum velocity recoverable De 44-31 m s-1 Minimum velocity recoverable 16 m s-1 Expected nozzle exit particle concentration 3.6 particles mm-2 field of view for each nozzle at two locations. The flow direction is from left to right. The gray scale of the velocity vectors represents the concentration and

the length the magnitude of the velocity. Concentra-tion estimates are approximate and indicate the distri-bution of the particles qualitatively. This is due to the


Fig. 19. Particle-phase axial velocity profile of the

simple nozzle (SPN) at 20 nozzle diameters downstream measured using PIV and LDA.

Fig. 20. Particle-phase velocity and concentration map

for the simple nozzle (SPN): (a) beginning at nozzle exit, (b) beginning at 17.4 nozzle diameters downstream.

effect of thresholding of the image level resulting in a lower count of particles. However, the values of relative concentration are helpful in determining the mix-ing properties of the particle-laden jet.

Figure 20a shows the nozzle exit conditions for the simple nozzle and the velocity is the largest near the centerline of the jet. Particle concentration is high along the jet center axis (3.6 particle mm-2). Disper-


for asymmetric nozzle (ASN): (a) beginning at nozzle exit, (b) beginning at 17.4 nozzle diameters downstream.

sion of the jet within the field of view is minimal. Figure 20b shows the flow field of the jet from 17.4 nozzle diameters downstream of the exit. The velocity levels as well as concentration have decreased along the centerline. Diffusion of the particles away from the jet axis because of turbulent mixing is shown by the spread in the concentration as well as the broader range of velocity. Near the jet edges for the asymmet-ric nozzle (Fig. 21a), there are large concentration gradients moving outward from the jet. The locations in the plot where vectors are absent are regions of ad-verse correlations resulting in incorrect vectors. At the downstream location (Fig. 21b) the jet has spread out to cover the entire range of the field of view and the concentration levels are lower and more uniform than at exit conditions.

The velocity vectors along the edges of the jet for DTN1 point away from the jet centerline (Fig. 22a). The trajectory of the particles directs them away from the jet and into the free stream flow at a velocity much lower than the jet velocity. The delta tabs are di-recting a portion of the particles away from the jet centerline. Within the jet core, the velocity field is di-rected uniformly downstream, and the effects of the tabs are not evident. For downstream profiles (Fig. 22b), the active dispersion of the particles results in

Y-Position (y/d0)



for DTN1: (a) beginning at nozzle exit, (b) beginning at 17.4 nozzle diameters downstream.

lower concentration values and a more uniform distri-bution of particles. The ejection of particles away from the jet axis by the DTN2 is also seen in Fig. 23a. Qualitatively, the vector plot is similar to the corre-sponding vector plot for DTN1. For downstream (Fig. 23b) the effect of the delta tabs is again evident by lowered concentration values along with greater uni-formity. The results for DTN3 (Fig. 24a) show the en-hanced dispersion of the particles as the jet velocity is fanning out into the free stream. It can be assumed that once the particles cross over into the free stream, and are no longer involved in the interaction of the jet and the free stream. At the downstream location (Fig. 24b) the concentration has greater uniformity and the velocity direction is uniform. The PIV measurements clearly show the nozzle exit behavior. This behavior is helpful in determining the dispersion of the particles due to the nozzle geometry. The tabs clearly direct the particles away from the jet centerline. Close to the nozzle exit, the particles are directed away from the jet at a large spread angle. Downstream, the angle nar-rows.

CONCLUSIONS

An experimental study of particle-laden jet con-



current flow interactions has been performed using both LDA and PIV. The dispersion and mixing char-acteristics of particle laden jet (50 μm glass beads as particles) for eight different nozzle geometries was studied. DTNs have been shown to exhibit elevated dispersion and turbulence characteristics.

The six different nozzles were; APN, ASN and four DTNs. The particle-phase mean velocity profiles, axial jet evolution, turbulent intensities, and Reynolds shear stress have been identified to study the disper-sion and mixing characteristics of 50 μm glass beads in a particle-laden turbulent jet. The nozzles with delta tabs have generally higher levels of Reynolds shear stress and wider jet width. PIV was used to provide the particle trajectory field and concentration values at the nozzle exit and at one downstream location. The velocity vector map for the exit velocity for the DTNs showed particles with velocities fanning out from the jet centerline. The particles are deflected by the tabs. Also, particle concentrations downstream of the DTN show greater uniformity than for the other nozzles. This results from high downstream turbulence levels. The larger the delta tab, the greater all of these effects, with DTN3 having the greatest effect. The asymmetric jet produced elevated turbulence levels when com-pared to the simple nozzle. Throughout the evolution of the jet, the aspect ratio of the jet remained oriented




in one direction.

LDA evaluation of the nozzles shows that char-acteristics of the DTN3 indicate the best mixing per-formance. The mean axial velocity profiles of the jet in the horizontal plane become uniform as one moves downstream. The mean vertical velocity at the nozzle exit vertical plane indicates that particles moving away from the jet centerline. Centerline axial velocity decay is the greatest between five and ten nozzle di-ameters downstream when compared to the other noz-zles. Stokes number beyond ten nozzle diameters downstream is slightly lower than the values of the simple nozzle. Axial velocity turbulent intensity levels around the jet centerline at the nozzle exit are elevated. Vertical velocity turbulent intensity levels are elevated at all downstream locations. Reynolds shear stress values show a large range of values across the hori-zontal width of the jet. Given the six nozzle designs, the DTN3 creates the best overall indication of en-hanced mixing performance.

PIV evaluation of the nozzles indicates that the DTN3 exhibits the greatest amount of dispersion when compared to the other nozzles. This is evident by the fanning of the velocity vectors away from the jet cen-terline at the nozzle exit. Also, at the downstream lo-cation, the concentration throughout the field of view is more uniform than at the nozzle exit. This is true for

all the delta tab nozzles, yet DTN3 shows the greatest dispersion at the nozzle exit due to an increased num-ber of fanning velocity vectors. Based upon the pre-ceding observations, it appears that the DTNs have superior mixing characteristics as compared to the other nozzles.

ACKNOWLEDGEMENTS

We wish to thank NASA Glenn Research Center

(GSRP award for D. Jang), the Department of Me-chanical and Aerospace Engineering, Case Western Reserve University, Cleveland, Ohio and the Ohio Coal Development Office, Columbus, Ohio for their support for this work. We also appreciate the help of Mr. D. Conger, Case Western Reserve University, Cleveland, Ohio.

NOMENCLATURE

m Fluid viscosity ρ Fluid density ψ Sphericity σ Standard deviation θ Swirl Angle φ Tab angle λ Wavelength σm Marginal standard deviation ρp Particle density de Effective e-2 particle diameter dg Geometric particle diameter D0 Nozzle diameter dp Particle diameter Dp Particle diffusion coefficient ds Diffraction-limited spot diameter Dx Particle moment function f Focal length f/# F-number (ratio of focal length to aperture

diameter) I Particle image intensity L Large eddy length scale Lz, Ly Jet half width M Geometric magnification factor m Mass loading (mass flux of solid: mass

flux of fluid) n Index of refraction P Gaussian distribution function Pj Jet driving pressure Q Flow rate r Radial direction Re Reynolds number RSS Reynolds shear stress Scp Schmidt number St Stokes number Std Strouhal number tf Turbulent flow time scale tj Jet characteristic time tp Particle response time constant


u, v Velocity Uθ Tangential velocity u’ Fluctuating component of velocity u’f,c Standard deviation of the centerline axial

fluid velocity fluctuation Uc Centerline mean axial jet velocity Ud Average duct flow velocity Udc Duct centerline mean flow velocity Uj Nozzle average jet velocity U0 Centerline jet exit mean velocity Ur Radial velocity V Mean vertical velocity ve Effective kinematic viscosity xs Equivalent surface diameter xv Equivalent volume diameter

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Discussions of this paper may appear in the discus-sion section of a future issue. All discussions should be submitted to the Editor-in-Chief within six months of publication.

Manuscript Received: February 11, 2008 Revision Received: June 20, 2008

and Accepted: July 18, 2008

SORBENT DISPERSION TECHNIQUES IN FLUE GAS FOR REMOVAL OF TOXICS

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