Fuel Processing Technology 126 (2014) 366–382

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Fuel Processing Technology

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A CFD study of biomass pyrolysis in a downer reactor equipped with anovel gas–solid separator — I: Hydrodynamic performance

Xi Yu a, Yassir Makkawi a,⁎, Raffaella Ocone b, Martin Huard c, Cedric Briens c, Franco Berruti c

a European Bioenergy Research Institute (EBRI), School of Engineering and Applied Science, Aston University, Birmingham B4 7ET, UKb Chemical Engineering, Heriot-Watt University, Edinburgh EH14 4AS, UKc The Institute for Chemicals and Fuels from Alternative Resources (ICFAR), Western University, London, Canada

⁎ Corresponding author. Tel.: +44 121 204 3398.E-mail address: y.makkawi@aston.ac.uk (Y. Makkawi)

http://dx.doi.org/10.1016/j.fuproc.2014.05.0200378-3820/Crown Copyright © 2014 Published by Elsevie

a b s t r a c t

a r t i c l e i n f o

Article history:Received 6 March 2014Received in revised form 12 May 2014Accepted 19 May 2014Available online 9 June 2014

Keywords:Biomass pyrolysisDowner reactorGas–solid separationCFD modelingHydrodynamics

This studypresents thefirst part of a CFD study on the performance of a downer reactor for biomass pyrolysis. Thereactor was equipped with a novel gas–solid separation method, developed by the co-authors from the ICFAR(Canada). The separator, which was designed to allow for fast separation of clean pyrolysis gas, consisted of acone deflector and a gas exit pipe installed inside the downer reactor. A multi-fluid model (Eulerian–Eulerian)with constitutive relations adopted from the kinetic theory of granular flowwas used to simulate themultiphaseflow. The effects of the various parameters including operation conditions, separator geometry and particle prop-erties on the overall hydrodynamics and separation efficiency were investigated. The model prediction of theseparator efficiencywas comparedwith experimentalmeasurements. The results revealed distinct hydrodynam-ic features around the cone separator, allowing for up to 100% separation efficiency. The developedmodel provid-ed a platform for the second part of the study, where the biomass pyrolysis is simulated and the product qualityas a function of operating conditions is analyzed.

Crown Copyright © 2014 Published by Elsevier B.V. All rights reserved.

1. Introduction

The escalating global concern over the exhaustion of non-renewableenergy sources leads to the recent development of a range of novel tech-nologies for the use of renewable energy resources, such as biomass, solarand wind. Among these resources and technologies, biomass pyrolysishas emerged as a very promising renewable alternative for bio-oil pro-duction. In a large commercial scale, this could be carried out in a dual flu-idized bed (DFB) system with various optional arrangements. Theschematics in Fig. 1 demonstrate examples of these arrangements. Inthis study, we are interested in the downer-riser type of a dual fluidizedbed, shown in Fig. 1-b, where the biomass pyrolysis takes place in thedowner side of the reactor, while the riser side is used for combustion,thus providing the heat required for the pyrolysis through the circulatinginert heat carrier solid (such as sand). This arrangement has the followingspecific advantages for bio-oil production through fast pyrolysis:

i. The downer pyrolysis reactor can be operated with very low carriergas (e.g. nitrogen) flow rates, which is desirable in some cases to re-duce up-stream pre-heating and downstream processing.

ii. Reducing the gas and solid back-mixing [1–3] thus, limiting the spreadof the gas/solid residence time distribution, i.e. near to plug flow.

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iii. Relatively low cost, simple operation/control and high energyefficiency.

iv. The char combustion in the second reactor will guarantee sustain-able operation and better control of the pyrolysis temperature inthe first reactor.

However, in order to achieve high conversion efficiency (more than70% bio-oil yield) in a downer reactor there remains twomain technicalchallenges:

i. Control of the pyrolysis gas residence timewithin the hot zone of thereactor (ideally 1–2 s). Longer residence time of the pyrolysis gas athigh temperature initiates a range of undesirable side reactions,which could adversely affect the quality of the product bio-oil [4,5].

ii. Control of the downstream contact between the pyrolysis gas andbio-char. The bio-char, formed during pyrolysis, acts as a vaporcracking catalyst, therefore should be separated as soon as the pyrol-ysis vapor is released [6].

Char, as well as other entrained fine particles, can primarily beseparated from the pyrolysis gas by using conventional cyclones(reverse and co-current flow types). However, this carries the risk of in-creasing the contact time between the gas and char inside the cyclone.In addition, the cyclone inlet is commonly placed external to the reactoror away from the pyrolysis zone, thus, causing extra contact time be-tween the solid and gas. The extensive review conducted by Huardet al. [7] and Cheng et al. [8] on downer reactors and rapid gas–solid

Fig. 1. Demonstration of biomass pyrolysis in various dual fluidized bed reactorconfigurations.

Fig. 2. Schematic diagram of the experimental apparatus.

Fig. 3. Illustration of experimental (a) coldmodel downer setup and (b) gas–solid separa-tion zone.

367X. Yu et al. / Fuel Processing Technology 126 (2014) 366–382

separation techniques revealed that there are limited attempts onimplementing new design methods for rapid gas–solid separation inthese reactors.

Recent research at the Institute for Chemicals and Fuels from Alter-native Resources (ICFAR) has led to the development of a novel gas–solid separation device for a downer pyrolysis reactor. The device fea-tures a cone-shaped solid deflector positioned above a gas outlet pipe,both positioned concentrically in the downer pipe (see Fig. 2). Thiswas designed to achieve primary solid–gas separation and gas removal

Fig. 4. Particle size distribution of silica sand used in the experiment.

Table 1Dimensions of the downer reactor and the cone separator.

Parameters Value Parameters Value

L1 [cm] 98.6 Dc [cm] 7L2 [cm] 34.9 Ls [cm] 0, 3.5, 7a [cm] 0.35 Dgo [cm] 0.95α [degree] 60, 90, 120

368 X. Yu et al. / Fuel Processing Technology 126 (2014) 366–382

within the same device [9]. The separator allows for better control of thepyrolysis vapor residence time, therefore, reducing the severity of vaporover-cracking compared to other fast separation methods. Experimen-tal work by Huard et al. [10] has shown that this separator can achievevery high solid–gas separator efficiency above 99.99% when usingspherical silica sand particles of 200 μm diameter.

Computational Fluid Dynamics (CFD)modeling is one of the power-ful tools to analyze gas–solid flow behavior, including that involves in-tense heat transfer and chemical reactions. The co-authors from the

Table 2Constitutive relations for the gas–solid flow.

Solid pressure

Psi ¼ αsiρsiΘsi þ ∑3

j¼1; j≠i

dsi þ ds j� �3

4d3sig0;sis jρsiαsiαs jΘsi 1þ esi s j

� � (T1 ‐ 1)

Solid shear viscosityμsi ¼ μsi ;col þ μsi ;kin þ μsi ;fr (T1 ‐ 2)Collisional viscosity [21]

μsi ;col ¼ 45αsiρsi dsi g0;si si esisi þ 1

� � Θsiπ

� �1=2 (T1 ‐ 3)

Kinetic viscosity [22]

μsi ;kin ¼ αsiρsi

dsiffiffiffiffiffiffiffiΘsi

πp

6 3−esi sið Þ 1þ 25 esi si þ 1� �

3esisi−1� �

αsi g0;sisih i (T1 ‐ 4)

Friction viscosity [23]

μsi ;fr ¼Psi

sinϕ

2ffiffiffiffiffiI2D

p (T1 ‐ 5)

Bulk viscosity [24]

λsi ¼ 43αsiρsi dsi g0;sisi esisi þ 1

� � Θsiπ

� �1=2 (T1 ‐ 6)

Radial distribution function

g0;sisi ¼ 1− αsαs;max

� �1=3� �−1

þ 12 dsi∑

3

j¼1

αs j

ds j

(T1 ‐ 7a)

g0;sis j ¼ds j g0;si si þdsi g0;s j s j

dsi þds j

(T1 ‐ 7b)

αs ¼ ∑3

j¼1αs j

Gas viscosity

μg ¼ μ l;g þ μt;g ; μ t;g ¼ Cμαgρgk2gεg

(T1 ‐ 8)

Gas–solid drag coefficient [25]

βgsi ¼3ρgαsi

αg

4u2r;si

dsiCD

Resivr;si

� �u*g−u

*si

(T1 ‐ 9)

vr;si ¼ 0:5 A−0:06Resi þffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi0:06Resi� �2 þ 0:12Resi 2B−Að Þ þ A2

q� �(T1 ‐ 10)

A ¼ α4:14g ;

B ¼ 0:8α1:28g αg≤0:85

� �B ¼ α2:65

g αg N0:85� �

(

CD ¼ 0:63þ 4:8ffiffiffiffiffiffiffiffiffiffiffiffiffiffiResi =vr;si

p� �2 (T1 ‐ 11)

Resi ¼dsi ρg u

*g−u

*sij j

μg

(T1 ‐ 12)

Solid–solid drag coefficient [19]

βsi s j ¼3 1þesi s j

� �π2þCfr;si s j

π28

� �αsi

ρsαs jρs dsi þds j

� �2

g0;si s j

2πρs d3si þd3s j

� � u*

si−u*

s j

(T1 ‐ 9)

Diffusion coefficient of granular energy [21]

κΘsi¼ 150ρsdsi πΘsið Þ12

384 esi si þ1ð Þg0;si si 1þ 65αsi g0;si si esisi þ 1

� �h i2þ 2αsi

2ρsdsi g0;sisi esi si þ 1� � Θsi

π

� �12

(T1 ‐ 13)

Collisional energy dissipation [24],

γΘsi¼

12 1−e2si si

� �g0;si si

dsi π1=2 α2

siρsΘsi3=2

(T1 ‐ 14)

Fig. 5. Computational domain and meshing in a cross-sectional view.

ICFAR have previously used an Eulerian–Lagrangianmodeling approachto investigate the effect of the particle elasticity on the separator effi-ciency in the same novel separator investigated in this study [9].While this approach revealed important details of the particle–wall col-lision and its effects on the separator efficiency and mechanism, thesimulation domain was limited to the separator zone only and thetotal solid volume fraction was limited to a maximum of 4 × 10−5.The Eulerian–Eulerian (also referred to as two-fluid) is another model-ing approach that has the advantage of being robust and realistic incomputational time, especially when considering a large number ofparticles or large simulation domain. Unlike the Eulerian–Lagrangianapproach, which treats each single particle as a dispersed phase in

Table 3Particle size distribution employed in the CFD model.

Solid phase I Solid phase II Solid phase III

Particle size (μm) X10 = 124 X50 = 206 X90 = 328Volume fraction (%) 20% 60% 20%

Table 4Gas and solid phase boundary/operating conditions used in the CFD model.

Boundary/operating condition Experiment Model

Gas mass flow rate,mg [kg/s] 0.0039, 0.0239, 0.0439 0.0039a, 0.0239, 0.0439Gas density, ρg [kg/m3] 1.2 1.2a

Gas dynamic viscosity[kg/(m·s)]

1.8 × 10−5 1.8 × 10−5a

Gas outlet pressure, Pgo [Pag] 0 (ambient) 0a (ambient)Particle density, ρP [kg/m3] 2650 2650a

Solid mass flow rate,ms [kg/s] 0.017, 0.083 0.004a, 0.02, 0.04, 0.08Mean particle size, ds [μm] 188 Mixturea (124, 206, 328)Angle of cone deflector [degree] 60, 90, 120 60a, 90, 120Separation length, Ls [m] 0 0a, 0.035, 0.07

a Default simulation conditions, unless otherwise specified.

Fig. 6. Example of the gas–solidflow structure in the overall downer column and around the separation device. (a) Gas phase velocity vectors, with the color code restricted to amaximumof 5 m/s to allow visualization. (b) Solid phase volume fraction in a section, with the color code restricted to a maximum of 4 × 10−4 to allow visualization.

369X. Yu et al. / Fuel Processing Technology 126 (2014) 366–382

the continuum fluid flow, the Eulerian–Eulerian approach treats boththe fluid and solid phases as an interpenetrating continuum. Studieson the Eulerian–Eulerian simulation of solid–gas hydrodynamics in adowner reactor have been previously reported by Ropelato et al. [11],Kim et al. [12] and Samruamphianskun [13]. This modeling approachwas found to be especially useful in predicting the effects of inletdesign and flow conditions on the solid distribution and dispersionbehavior. This CFD modeling approach has also been used by differ-ent researchers to study the phenomena of solid–gas separation incyclones [14–16].

In this study, the main objectives are:

i. to develop a valid Eulerian–Eulerian (multi-fluid) CFD model capa-bly of predicting the detailed hydrodynamic behavior in a downerreactor equipped with a novel gas–solid separation device;

ii. to use the developed model in investigating the effect of the op-erating conditions and various separator design parameters onthe overall hydrodynamics, with particular focus on the separatorefficiency;

iii. to provide a platform for the development of a predictive modelof the pyrolysis reactions and yield in the downer reactorequipped with the novel gas–solid separator.

The investigation was carried out theoretically and experimentallyin a cold flow reactor model equipped with the ICFAR novel gas–solidseparator and gas removal mechanism, as described in details in theExperiments and procedure section. The theoretical transient modelwas solved in three-dimensional coordinates using the Eulerian–Eulerian (two-fluid) approach, employing constitutive relations fromthe kinetic theory of granular flow (KTGF) [17]. In the second part of

this study, the developed hydrodynamic model will be extended to in-clude heat transfer and reaction kinetics to demonstrate the advantagesof the ICFAR separator in improving the performance and product qual-ity in a biomass downer pyrolysis reactor.

2. Experiments and procedure

The experimental work described here was carried out by the co-investigators at the ICFAR in Canada. The equipment consisted of acold flow gas–solid flow downer of 133.5 cm height and 7.0 cm diam-eter, equipped with the ICFAR novel gas–solid separator as shown inFigs. 2 and 3. This separator included a gas removal pipe and a conedeflector, where the bottom of the deflector and the tip of the pipewere located 98.6 cm below the downer inlet. A solid collectiontank of 20.4 cm diameter and 21.8 cm height was placed at the bot-tom of the downer column around 34.9 cm below the cone deflector.Compressed air at room temperature was supplied to the downerfrom a bank of calibrated sonic orifice nozzles. The Sauter mean di-ameter of the particle mixture used was 188 μm and its skeletal den-sity was 2650 kg/m3. The particle size distribution of the mixture isshown in Fig. 4.

The solid particles were delivered to the downer column from an airpressurized tank mounted above the downer main air inlet. The totalmass flow rate of air in the downer was mg = 0.0039 kg/s, whichcorresponded to a superficial gas velocity of Ug = 0.73 m/s. The solidmass flow rate was adjusted by changing the feed tank air pressureand this was varied betweenms = 0.017 kg/s and 0.083 kg/s, whichcorresponded to solids-to-gas loading ratios ofms=mg = 4.3 to 21. Thegas–solid mixture flowed co-currently in the downer before entering

Fig. 7. Example of the radial profiles of the (a) axial gas velocity and (b) solid volume frac-tion. The data was taken at the sample level of 3.9 cm below the gas exit tip.

370 X. Yu et al. / Fuel Processing Technology 126 (2014) 366–382

the gas–solid separation zone. Three different cone deflectors with var-ious internal angles of 60°, 90° and 120° were used. The downward fall-ing particles were collected in the tank at the bottom of the unit, whilethe gas stream, along with any entrained particles, exited the systemthrough the gas outlet pipe mounted in the center of the downer crosssection and below the cone deflector. A bag filter connected at the endof the gas exhaust line was used to collect the particles entrained inthe exiting gas stream.

At the start of each experiment, the total mass of solids fed into thesystem,min, was measured. The mass of the entrained solids collected inthe bag filter,mcollected, was thenmeasured at the end of each experiment.Thus, the experimental percentage total solid separation efficiency, η, wascalculated from the following expression:

η ¼ 1−mcollected

min

� �� 100%: ð1Þ

The mass flow rate of solids was determined by measuring the totalmass of solid collected in the filter bag and tank against the recordedtime.

3. Hydrodynamic model

The overall reactor hydrodynamics and gas–solid separationwas in-vestigated using the Eulerian–Eulerian (multi-fluid) model approachbased on the Kinetic Theory of Granular Flow (KTGF). The developedmodel was solved using the CFD software ANSYS FLUENT (Ver. 14).In order to mimic the wide size distribution of the solid mixture usedin the experiment, the simulation was carried out using a solid mixtureof three different particle sizes, as detailed in Section 3.3. Themain model equations for non-reacting isothermal gas–solid flow aregiven by:

Continuity equations:

∂ αgρg

� �∂t þ∇ αgρg u

*

g

� �¼ 0 ð2aÞ

∂ αsiρs

� �∂t þ∇ αsi

ρsu*

si

� �¼ 0 ð2bÞ

X3i¼1

αsiþ αg ¼ 1: ð2cÞ

Momentum equations:

∂ αgρg u*

g

� �∂t þ∇ αgρg u

*

g u*

g

� �¼ −αg∇P þ∇τg−

X3i¼1

βgsiu*

g−u*

si

� �þ αgρgg

ð3aÞ

∂ αsiρsu

*

si

� �∂t þ∇ αsi

ρsu*

siu*

si

� �¼ −αsi

∇P−∇Psiþ∇τsi þ βgsi

u*

g−u*

si

� �þ

X3j¼1; j≠i

βsis ju*

s j−u

*

si

� �þ αsi

ρsg

ð3bÞ

where:

τk ¼ λk−23μk

� �∇ � u*k

� �I þ 2μkSk ð4aÞ

Sk ¼12

∇u*

k þ ∇u*

k

� �T� �ð4bÞ

k represents solid or gas phase.To obtain the granular temperature, the FLUENT codewas optionally

set to use a partial differential equation (Pseudo Energy Equation) asfollows [18]:

32

∂ αsiρsΘsi

� �∂t þ∇ αsi

ρsΘsi

� �u*

si

24

35

¼ −PsiI þ τsi

� �: ∇u

*

siþ∇ κT∇Θsi

� �−γT þ

X3k¼1

ϕksi:

ð5Þ

The various closure and constitutive relations used in the model aregiven in Table 2. In order to take into consideration the solid–solid fric-tional stresses at the dense regions of the reactor, the friction equation

Fig. 8. Sectional view of the solid and gas axial velocity vectors demonstrating the various characteristic flow zones in the downer reactor and around the separation device. To allow vi-sualization of the low ranges, the magnitude of the velocity vector was restricted to a maximum of 5 m/s. The color code bar indicates the range of gas/solid velocity.

371X. Yu et al. / Fuel Processing Technology 126 (2014) 366–382

proposed by Schaeffer [19], as given in Eq. (T1-5), was used. Due to thehigh turbulence of the flow near the deflector zone the standard K–ep-silon turbulence and energy dissipation equations proposed by Launderand Spalding [20] were also incorporated in the model and these aregiven as follows: (See Table 1.)

Turbulence momentum equations:

∂ αgρgkg� �

∂t þ∇ αgρg u*

gkg� �

¼ αgGk;g þ∇ αg

μt;g

σkkg

� �−αgρgεg þ αgρgΠk;g :

ð6aÞ

Turbulent kinetic energy dissipation equation:

∂ αgρgεg� �

∂t þ∇ αgρg u*

gεg� �

¼ ∇ αg

μt;g

σεεg

� �þ αg

εgkg

C1εGk;g−C2ερgεg� �

þ αgρgΠε;g ð6bÞ

where:

Gk;g ¼ μt;g ∇u*

g þ ∇u*

g

� �T� �: ∇u

*

g

Cμ ¼ 0:09;C1ε ¼ 1:44;C2ε ¼ 1:92;σk ¼ 1;σε ¼ 1:3:

3.1. Computational domain and meshing

Fig. 5 shows the computational domain and the meshing used insolving the model. This was generated using a finite volume methodwith hybrid cells of structured and unstructured grids, giving a total of30,785 cells. In order to capture the steep hydrodynamic variationsaround the walls of the separation device (the conical deflector andthe gas exit pipe), the grid size was refined by setting the minimumand maximum grid sizes at 0.3 and 1.0 cm respectively. In the rest ofthe simulation domain the minimum and maximum grid sizes wereset at 1.0 and 5.0 cm respectively. The impact of the grid size on the so-lution accuracy was initially tested by setting three different meshingschemes and the grid size used in this study was found to give accept-able grid independent solution.

3.2. Computation procedure

The model equations were solved using the finite volume approach.First-order discretization schemeswere used for the solution of the con-vection terms in all governing equations. The relative error between any

Fig. 9. Illustration of the method used to determine the GDH demonstrated in typical re-sults obtained at the default model operating conditions (a) pressure gradient method(b) axial gas velocity method.

372 X. Yu et al. / Fuel Processing Technology 126 (2014) 366–382

two successive iterationswas specified by using a convergence criterionof 10−3 for each scaled residual component. The phase-coupled SIMPLE(PC-SIMPLE) algorithm [21], which is an extension of the SIMPLE algo-rithm to multiphase flows, was applied for the pressure–velocity cou-pling. The linearized equations for governing equations were solvedusing a block algebraic multigrid method. In order to ensure easy con-vergence of the various partial differential equations (PDEs) in themodel, the Courant–Friedrichs–Lewy (CFL) condition for three-dimensional PDE is followed:

C ¼ uxΔtΔx

þ uyΔtΔy

þ uzΔtΔz

≤Cmax ð7Þ

where Cmax is specified by the CFL condition to fall within the range of~1–5 [22]. In this study, a time step of 0.005 s was found to satisfy thiscondition.

3.3. Boundary and simulation conditions

The particle–wall restitution coefficient and the specularity coeffi-cient are two important parameters in determining the dynamics ofparticles at the wall region. The following wall boundary conditionswere employed in the model [23]:

usi ;w¼ −

6μsiαs;maxffiffiffiffiffiffiffiffiffi

3Θsi

qπφρsαsi

g0;sisi

δusi ;w

δnð8Þ

Θsi¼ −

ksiΘsi

γw

δΘsi ;w

δnþ

ffiffiffi3

pπφρsαsi

u2si ;slip

g0;sisi Θsi

32

6αs;maxγwð9aÞ

γw ¼ffiffiffi3

pπ 1−e2si ;w� �

ρsαsig0;sisi Θsi

32

4αs;maxð9bÞ

where φ is the specularity coefficient and esi ;w is the particle–wallrestitution coefficient. In order to reasonably match the particlesize distribution used in the experiments, the simulations werecarried out assuming the solid mixture to consist of three differentparticle sizes. The fraction of each particle size group was estimatedfrom the experimental size distribution given in Fig. 4. The simula-tion particle sizes and percentages are given in Table 3. For the gasphase, the velocity at the wall was assumed zero (no slip condition).Table 4 summarizes the various operating conditions considered inthe simulations. Some of these conditions were carefully selectedto allow for the comparison of the model predictions with the corre-sponding experimental data.

4. Results and discussion

4.1. Mechanism of gas–solid separation

It is postulated that the drag and gravity forces, the last two termsin the left hand side of Eq. (3b), are the main forces dominating thehydrodynamic behavior of the gas–solid phases within the separatorzone. Fig. 6 gives an overall description of the flow structure withclose zoom-in at the cone deflector region. The gas velocity in thegap between the deflector and the wall is very high due to the con-siderable pressure drop, similar to gas expansion through a throt-tling device. In the region under the cone and below the gas exitpipe there is an upward gas drag force due to the high reverse gasphase velocity. However, the extremely dilute solid concentrationin this zonemeans that very little solids are being entrained. It is there-fore desirable to minimize the upward gas drag force in this region inorder to achieve high separator efficiency. On top of the cone deflector,the solid phase is diverted radially towards the wall and then acceler-ates through the gap between the deflector and the downer wall,pushed by a strong gas drag force. The influence of the gravity forcein this region is also significant due to the high solid concentration.Accordingly, it is believed that any particles entrained throughthe exit pipe are falling under the influence of two different dragmechanisms:

i. Reverse gas flow (upward) under the cone deflector due to theabrupt gas pressure drop at the tip of the gas exit pipe. Thismakes the tip of the exit pipe act as a vacuum to the surroundingsolids.

ii. Radial gas flow from the walls towards the core in the region justbelow the cone deflector. This results in the solids being firstdragged towards the core, and then further dragged/sucked bythe gas leaving through the exit pipe.

Fig. 10. The effect of operation conditions on the gas disengagement height. The default simulation conditions (Table 4) were used in all simulations, unless otherwisespecified.

373X. Yu et al. / Fuel Processing Technology 126 (2014) 366–382

This is to some extent similar to the solid–gas separationmechanismin a cyclone, where in both cases the reverse gas flow in the core is re-sponsible of solid entrainment. However, in the cyclone, the particlesmove radially towards the walls under the influence of centrifugalforces, while in the cone deflector, the particles are deflected radiallyby the cone wall to fall under the strong downward gas drag force inthe “throttling” gap, as described earlier. It is worth noting thatthe modeling results reported by the co-authors from the ICFAR sug-gested that the solid rebound upon hitting the walls, investigatedthrough changing the wall–particle restitution coefficient, may have adominant role in themechanism of gas–solid separation in the cone de-flector. This hypothesiswill be discussed in somedetails in the followingsections.

Fig. 7 shows the predicted solid concentration and the gas velocityprofiles at the level of 3.9 cm below the tip of the gas exit pipe (seeFig. 6b for the sampling line level). These profiles reveal a very inter-esting hydrodynamic behavior where the solid concentration profileis shown to take the shape of a dense-wall and dilute-core, while thegas velocity takes the shape of an upward parabolic flow profile atthe core and a downward flow at the walls. It appears that, due tothe existence of the cone deflector, the overall flow pattern belowthis device has been completely changed from the classic gas–soliddown flow pattern, commonly observed in downer reactors, to amore complex flow similar to that existing in a turbulent solid–gasflow riser.

4.2. Gas disengagement height (GDH)

Fig. 8 shows that there are four distinct flow zones each withcharacteristic flow behavior. These are mainly arising from thechanges induced by the cone deflector and these can be describedas follows:

Zone I: This is where fully developed flow and uniform distribution ofthe solid and gas phases take place, typical to that observed ina conventional downer reactor.

Zone II: This is where both of the solid and gas phases are first hit-ting the inclined plane to create a dense moving solidlayer at the cone walls before being pushed by a strong

gas drag force through the gap between the deflector andthe downer walls. Under the cone, the lowest solid concen-tration in the whole system exists and the gas is removedthrough the exit pipe driven by the rapid pressure drop atthe exit pipe tip.

Zone III: This is where the disengagement of gas from the gas–solidflow mixture takes place. The overall flow hydrodynamicsin this region is very complex due to the effect of sharpchanges in pressure, which consequently leads to reversegas flow towards the top and radial solid movement fromthe dense walls towards the dilute core.

Zone IV: This is where the solid phase is mainly concentrated at thewalls. The radial flow diminishes and the particles fallunder the strong influence of the gravity force before en-tering the solid collection tank at the bottom of the downersystem.

As described earlier, the main objective of the cone separator is toallow for fast and efficient separation of the gas from the downwardgas–solid flow stream. In a biomass pyrolysis downer reactor, thisseparation should ideally take place immediately at the level of thegas outlet pipe tip and with zero solid entrainment. The first reasonis to prevent undesirable secondary gas reactions by removing thegas from the reactor hot zone, and second, to prevent catalytic charcracking by limiting the contacts between the solid and gas. Howev-er, in reality, the gas separation from the solid–gas stream takesplace a little further down beyond the level of the tip of the gas re-moval pipe. It is therefore particularly interesting to quantify theheight of Zone III, in which the gas separation takes place. This is de-fined here as the gas disengagement height (GDH), analogous to thedefinition of the transport disengaging height (TDH) in gas–solidfluidized beds. The method used in this study to estimate the GDHis demonstrated in Fig. 9. The GDH is defined at the intercept of thelines tangential to the low pressure gradient curve and the steepchanging pressure gradient curve, or alternatively, the GDH can beestimated from plotting the axial gas velocity against height asshown in Fig. 9-b. The pressure gradient method is similar to themethod used by Geldart et al. [24] in determining the transportdisengaging height (TDH).

Fig. 11. Overall sensitivity analysis of the effect of the operating conditions on the separa-tor efficiency. The default simulation conditions (Table 4) were used in all simulations,unless otherwise specified.

Fig. 12. Effect of the separation length on the separator efficiency. The simulation wascarried out using the default operating conditions.

Fig. 13. Gas velocity vectors at various normalized separation lengths. To allow visualizationof the low ranges, themagnitude of the velocity vectorwas restricted to amaximumof 5m/s.The color code bar indicates the range of gas velocity.

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Fig. 10 shows the result of a sensitivity analysis of theGDH to a rangeof operating conditions. Please note that the y-axis in Fig. 10 representsthe summation of the GDH and the separation distance Ls, where theseparation distance is defined as the distance from the cone rim to thetip of the gas exit pipe. The GDH range under the various operation con-ditions considered in the simulation was found to fall between 2.5 cmand 6 cm. It is clear that the GDH is most sensitive to the solidloading and the solid flow rate. The increase in the separation distancefrom 0 cm to 7 cm and the solid flow rate from 0.004 kg/s to 0.08 kg/scaused a corresponding increase in GDH of around 30% for both cases.Clearly, the cone angle and gas flow rate appear to cause negligibleeffect in this regard. It should be noted that, while it is desirable todecrease the GDH as discussed earlier, this does not necessarily meanimproving the separation efficiency, as will be demonstrated in thenext section.

4.3. Separator efficiency

The theoretical separator efficiency was obtained by dividing thepredicted solid flow rate at the gas exit pipe (entrained solids) by theinlet solid mass flow rate, such that,

η ¼ms;exit

ms;in� 100%: ð10Þ

Fig. 14. Effect of the separation length on the (a) radial profiles of axial gas velocity and(b) solid volume fraction (for the selected particle size ds = 124 μm). The data wastaken at the sample level of 3.9 cm below the gas exit pipe tip.

Fig. 15. Effect of the cone deflector angle on the separator efficiency. The experiment datawas obtained from Hurad et al. [10] The simulation was carried out using the default op-erating conditions.

Fig. 16. Gas velocity vectors at various cone deflector angles. To allow visualization ofthe low ranges, the magnitude of the velocity vector was restricted to a maximum of5 m/s. The color code bar indicates the range of gas velocity.

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The separator efficiency was analyzed with respect to various op-erating conditions. This also included a sensitivity analysis of theseparation efficiency towards varying the wall–particle interactionmechanism, through changing the particle–wall restitution coeffi-cient and specularity coefficient. Both parameters appear in thesolid boundary condition of Eqs. (8), (9a), and (9b). The first coeffi-cient is a measure of the degree of energy loss when the particleshit the walls, hence determining the rebound velocity, and thesecond coefficient defines the angle of rebound. It is therefore possi-ble to determine the effect of the wall surface material and particleproperties on the separation efficiency through changing these twoparameters in the model.

In this study, 100% separator efficiency was obtained when operat-ing with: large particle size of dp = 328 μm, separation length Ls =0 cm, cone angle θ = 60°, gas mass flow rate mg = 0.0039 kg/s andhigh solid flow ratems = 0.08 kg/s. This was found to dramatically de-crease when decreasing the particle size. This result is in good

agreement with the experimental study by Huard et al. [10] whereit was shown that the separator efficiency, when using FCC catalyst ofdp = 43 μm and glass beads of dp = 63 μm, is much lower than thatachieved with sand of dp=200 μm. Fig. 11 shows the values of the pre-dicted separator efficiency obtained within the range of operating

Fig. 17. Solid volume fraction for the particle size of ds=324 μmat two different deflectorangles. To allow visualization of the low ranges, the solid volume fractionwas restricted to5 × 10−4. The color code bar indicates the range of solid volume fraction.

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conditions considered in this study. The detailed results and discussionon the effect of each of these parameters on the overall hydrodynamicsand separator efficiency are given in the next sections. According to thisdata, it is concluded that the sensitivity of the separator efficiency to-wards the operating conditions can be ranked in the order of decreasing

Fig. 18. Effect of the gasmassflowrate on the separator efficiency. The simulationwas car-ried out at the default operating conditions.

impact on the separation efficiency as follows; (1) separation length (2)cone angle (3) gas flow rate (4) solid flow rate and (5) particle physicalproperties (expressed in terms of the restitution and specularity coeffi-cients). Note that, the impact of particle size on the separation efficiencycomes on top of all the above parameters.

4.3.1. Effect of the separation length (Ls)The effect of separation distance on the separator efficiency was

studied using three different separation lengths of 0 cm, 3.5 cm and7 cm, which corresponds to the normalized separation length (Ls/D) of0, 0.5 and 1 respectively. All the other parameterswere set at the defaultvalues.

Fig. 12 shows the separator efficiency as a function of the normalizedseparation length. The overall trend indicates a negative impact on theseparator efficiency. The maximum mean efficiency was 99.986% andthis dropped to 99.633% at Ls/D = 1. The greater separation efficiencyachieved with the particle size of 206 μm compared to the size groupof 324 μmcan be explained by the fact that the concentration of this par-ticle group (60 wt.%) was greater than the latter one (20 wt.%). There-fore, the more frequent particle–particle interaction within the samegroup can neutralized part of radial velocity which may cause entrain-ment of particle. The same phenomenon was observed in Fig. 15. Interms of sensitivity, the effect of the separation length on the separationefficiency is the highest compared to the other parameters investigated,as shown earlier in Fig. 11. Itwas also demonstrated earlier that the sep-aration distance has also a relatively high effect on theGDH. The velocityvectors shown in Fig. 13 indicate that the increase in Ls resulted in thecreation of two vortices in the space between the cone deflector andthe tip of the gas exit pipe. This can be attributed to the strong radialgas flow in this region, resulting from the considerable pressure dropat the exit pipe.

Fig. 14 shows the changes in the solid concentration and velocityprofiles with changing the separation length at the sample level of3.9 cm below the bottom of the cone deflector. It is clear that the solidconcentration increases with increasing Ls, while the vertical upwardgas velocity at the core decreases. This suggests that the upward gasdrag force may have limited influence on the separation efficiency. Itis the increased radial gas velocity (radial drag), the subsequent forma-tion of vortices and the increased solid concentration at the core thatcollectively play the dominant role in decreasing the separator efficien-cy as the separation length increases.

4.3.2. Effect of the cone deflector angle (θ)The effect of the cone deflector angle on the separator efficiency

was studied using various angles θ= 60°, 90° and 120°. All other op-erating conditions were set at the default values. Fig. 15 shows thatthe separator efficiency decreases with increasing the cone angle.The maximum mean separator efficiency (taking into account thethree particle sizes) was 99.986%, this dropped to 99.869% efficiencywhen the angle is increased to 120°. This trend is in satisfactoryagreement with the experimental data of Huard et al. [10]; however,the experiments showed less pronounced changes compared to thepredictions, and this may be attributed to the differences betweenthe particle size distribution in the experiment and the assumedsize mixture in the model.

In Fig. 16 the magnitude and direction of the gas velocity vectorssuggest that as the cone angle increases there is a greater chance theparticles reboundmore in the reverse direction from the cone inner sur-face and normal to the gas exit. This would slow down the particles andmake them easier to be entrained, thus having a negative impact on theseparator efficiency. In Fig. 17 there is a clear increase in the solid con-centration on top of the cone's upper surface due to flattering of thecone external surface as shown in Fig. 17; however this is not expectedto have contributed to the change in the separation mechanism orefficiency.

Fig. 19. Gas velocity vectors at various gas flow rates. The color code bars indicate the range of gas velocity.

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4.3.3. Effect of the gas mass flow rateThe effect of inlet gas mass flow rate on the gas separator efficiency

was investigated at three differentflow rates of 0.0039 kg/s, 0.0239 kg/s

Fig. 20. Solid volume fraction at various gas mass flow rates for the particle size ds =206 μm. To allow visualization of the low ranges, the magnitude of the solid volume frac-tion is restricted here to amaximumof 2 × 10−4. The color code bar indicates the range ofthe solid volume fraction.

and 0.0439 kg/s and a fixed solid flow rate of 0.004 kg/s. This corre-sponds to inlet gas velocities of 0.73 m/s, 4.5 m/s and 8.2 m/s, respec-tively. All other operating conditions were set to the default values.Fig. 18 shows that the effect of the gas mass flow rate on the separatorefficiency is negligible. This is in good agreement with the experimentalobservation reported by Huard et al. [10].

To gain further understanding on the effect of gas flow rate on theoverall hydrodynamics, Fig. 19 shows the gas velocity vectors asfunction of the gas mass flow rate. It is clear that there is a significantchange in the magnitude of the gas velocity but little change in theflow pattern. There is also evidence of a significant change in thesolid concentration around the cone deflector as shown in Fig. 20.Despite this, such a dramatic change caused no effect on the separa-tor efficiency due to counterbalance of forces, which are described asfollows:

i. At a high gas velocity, there is considerable increase in the pressuredrop between the gas exit pipe and its surroundings, hence high up-ward gas velocity (drag force), as shown in Fig. 21a. However, this iscounterbalanced by the considerable reduction in the solid concen-tration in the wall and the core region below the exit pipe, asshown in Fig. 21b.

ii. At a low gas velocity, there is high solid concentration at the wall(i.e. high gravity force), as shown in Fig. 21b. This is associatedwithlow pressure drop between the wall and the tip of the exit pipe.Hence, there is reduction in the solid migration from the wall tothe core (i.e. low radial gas drag force) or solid carry over by the re-versing gas (i.e. low upward gas drag force).

According to the above analysis, it is concluded that the gas velocityhas little effect on the separator efficiency, at least within the operatingconditions considered here. In biomass pyrolysis, however, the gas ve-locity has a critical effect on the product quality due to its effect on thegas and solid residence time. The residence time can be quantifiedthrough the average gas velocity, particularly within the GDH region,as discussed in Section 4.2. The interrelation between the gas velocity,

Fig. 21. Effect of the gasflowrate on the (a) radial profiles of axial gas velocity and (b) solidvolume fraction (for the selected particle size ds = 206 μm). The data was taken at thesample level of 3.9 cm below the gas exit pipe tip.

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gas/solid residence times and the GDH in a downer pyrolysis reactor is acomplex one and requires careful optimization in order to achieve thebest product quality.

4.3.4. Effect of solid mass flow rateThe effect of solid mass flow rate on the separator efficiency was in-

vestigated using four different flow rates of ms = 0.004 kg/s, 0.02 kg/s,0.04 kg/s and 0.08 kg/s at a fixed gas mass flow rate of mg =0.0039 kg/s. This corresponded to solid to gas flow ratios (solid loading)of ms/mg = 1, 5, 10 and 20 respectively. All the other operating condi-tions were set to the default values. The experimental and predicted re-sults, shown in Fig. 22, suggest that the separator efficiency improves asthe solid loading increases within the range of ms=mg b 10, beyondwhich the efficiency appears to be independent of solid loading. Thistrend is less pronounced in the predicted data, which show very limitedchanges. Quantitatively, there is an over-prediction of separator

efficiency when compared with the experiment data; particularly atlow solid loading.

Fig. 23 shows the velocity vectors as a function of the solid loading. Inthe wall region below the cone deflector, there is a clear change in themagnitude and direction of the gas velocity vector, particularly in theright hand side below the gas exit pipe. This implies an increased down-ward gas drag force, which positively adds to the solid gravity force. It istherefore concluded that as the solid loading increases the amount ofsolid entrained by the reversing gas at the central region below thecone deflector reduces. The solid concentration in the wall region mas-sively increaseswhile the core region remains relatively constant whichcan be seen in Fig. 24. The solid concentration and velocity profiles at thesample level, shown in Fig. 25, indicate considerable hydrodynamicchanges below the cone deflector as the solid loading increased. Thegas velocity, however, shows exactly the opposite behavior with theaxial velocity in the core region more than doubled when increasingthe solid flow rare from 0.004 to 0.08 kg/s, while the velocity near thewalls is slightly increased. Because the increase in the axial upwardgas velocity in the center takes place in a region that is at extremelylow in solid concentration, the separator efficiency remains almost in-dependent of the increase in solid loading. Accordingly, it is recom-mended to operate this downer reactor at a high solid flow rate forthe following three main advantages:

i. Increased upward gas velocity towards the gas outlet pipe withinthe GDH region, therefore reducing the gas residence time in the re-actor.

ii. Improved separator efficiency, as evident from the experimental andpredicted results.

iii. Increasing the reactor processing capacity for biomass pyrolysis.

4.3.5. Effect of the particle restitution and specularity coefficientsIt is understood that the particle size plays a major role on the sepa-

rator efficiency such that the larger the particle size the higher theseparator efficiency. Another important parameter of interest here isthe degree of particle momentum loss or rebound upon hitting thesolid surfaces, which is defined in the model through the restitution co-efficient and specularity coefficient. The effective particle–wall restitu-tion coefficient (es,w) was determined experimentally by the co-authors from the ICFAR by measuring the rebound velocities of silicasand when hitting various types of solid surfaces; giving restitution co-efficients ranging from 0.73 (Plexiglas surface) to 0.48 (paper surface).In this study, the same range of particle restitution coefficient was im-plemented in the model to investigate the effect of this parameter onthe separator efficiency. The effect of the specularity coefficient (φ)was investigated by using values of φ = 0, 0.1 and 1.0, thus coveringthe two extreme ends of particle–wall interaction; free slip conditionat φ = 0 and no slip condition at φ = 1. Reported studies (e.g. [25])have shown that the restitution coefficients have an effect on the solidvelocity, gas velocity and solid concentration. It is also understood thatas the restitution coefficient increases there is a corresponding increasein the wall shear stress. The specularity coefficient, on the other hand,has been reported to have a pronounced effect on the solid concentra-tion, as increasing this parameter results in reducing the solid concen-tration at the wall.

Fig. 26 shows a comparison between the predicted and measuredseparator efficiency as a function of the particle restitution coefficient.It is clear that the predicted separator efficiency is a very weak functionof this parameter. This is in good agreement with some of the reportedliterature (e.g. [25]) which suggest that the particle restitution coeffi-cient (in the range 0.6–0.99) has limited effects on the solid velocity,gas velocity and the solid concentration in circulating fluidized bed re-actors. The experimental data shows a slight decrease in the efficiencyas the restitution coefficient decreases; however, this is still within avery limited range.

Fig. 22. Effect of the solid loading on the separator efficiency. The simulation was carried out at the default operating conditions.

379X. Yu et al. / Fuel Processing Technology 126 (2014) 366–382

Fig. 27 shows the changes in the predicted separator efficiencywith changing the value of the secularity coefficient. While there isclear drop in the efficiency as the specularity coefficient increases,this is still within a very limited range. This change is believed tobe a result of the increase in the wall shear stress (no-slip condition),which in turn results in hindering the downward flow of the densewall layer and hence giving rise to particle migration from the wallto the core followed by entrainment by the reversing gas towardsthe exit pipe. This phenomenon is demonstrated by the changes inthe solid velocity and concentration profiles shown in Fig. 28. The

Fig. 23.Gas velocity vectors at various solid flow rates. To allow visualization of the low ranges,bar indicates the range of gas velocity.

specularity coefficient appears to have a significant effect on thegas velocity and solid concentration at the wall regions, which is ingood agreement with the observation reported by Jin et al. [25],and in spite of this there is a negligible effect on the separatorefficiency.

5. Conclusions

The hydrodynamics in a downer pyrolysis reactor equipped witha novel gas–solid separator has been investigated theoretically using

themagnitude of the velocity vector was restricted to amaximumof 5m/s. The color code

Fig. 24. Solid volume fraction at various solid flow rates for the selected particle size ds = 206 μm. The color code bars indicate the range of solid volume concentration.

380 X. Yu et al. / Fuel Processing Technology 126 (2014) 366–382

an Eulerian–Eulerian (two-fluid) CFD model. The novel separator,which consists of a cone deflector and a gas outlet pipe, was designedby the co-authors from the ICFAR (Canada). The model predictionswere compared with experimental measurements of separator effi-ciency. This study revealed interesting hydrodynamic featuresaround the cone deflector, where due to the restriction of the flowpassage and solid deflection towards the walls, the region belowthe deflector in the downer reactor was completely transformed tobehave like a riser, characterized by distinct upward gas flow at thecore and dense falling solid layer at the walls. These distinct hydro-dynamic features allowed for high efficiency of gas–solid separationup to 100%. A new method for estimating the gas disengagementheight (GDH) was developed to help in estimating the gas residencetime in this novel reactor. This study also included detailed sensitiv-ity analysis of the separator efficiency towards the various operatingconditions, including the effect of particle restitution and secularitycoefficients. In the second part of this study, the present hydrody-namic model will be extended to include reaction kinetics and heattransfer to simulate the reactor thermochemical performance duringthe pyrolysis of biomass.

Notation

a Gap between conical deflector and reactor wall (m)A Model parameter (−)B Model parameter (−)C Courant number (−)CD Drag coefficient (−)Cμ, C1ε, C2ε Constants (−)Cfr;sis j Friction coefficient between solid phase i and phase j (−)dsi Particle diameter of solid phase i (m)DC Reactor diameter (m)Dgo Diameter of gas outlet pipe (m)esis j Particle–particle restitution coefficient (−)esi ;w Particle–wall restitution coefficient (−)g Gravity (m s−2)

g0 Radial distribution function (−)Gk,g Production of turbulent kinetic energy (kg m−1 s−2)I Unit vector (−)I2D Second invariant of the deviatoric stress tensor (s−2)L1, L2 Reactor dimension (m)Ls Separation length (m)kg Turbulence kinetic energy (m2 s−2)mcollected,min Mass of collected and fed solid particles respectively (kg)mg ; ms Mass flow rate of gas and solid respectively (kg s−1)P Pressure (pa)S Strain rate (s−1)Resi Reynolds number of solid phase i (−)t Time (s)u*g ; u

*si Gas and solid velocity vector (m s−1)

usi ;w Particle velocity at wall (m s−1)vr;si Terminal velocity correlation (−)X10, X50, X90 Particle size at accumulative volume fraction at 10%, 50%,

and 90%

Greek symbolsα Angle of conical deflector (degree)αg ;αsi Volume fraction of gas and solid phase i respectively (−)β Momentum exchange coefficient (kg m−3 s−1)γΘsi

Collisional energy dissipation (kg m−1 s−3)εg Turbulent dissipation rate (m2 s−3)η Separation efficiency (−)Θsi Granular temperature of solid phase i (m2 s−2)κΘsi

Diffusion coefficient of granular energy (kg m−1 s−1)λsi Particle bulk viscosity (kg m−1 s−1)μl,g, μt,g Viscosity of gas phase due to laminar, turbulent flow

(kg m−1 s−2)μsi ;col Viscosity of solid phase i due to collision (kg m−1 s−1)μsi ;kin Viscosity of solid phase i due to kinetics (kg m−1 s−1)μsi ;fr Viscosity of solid phase i due to friction (kg m−1 s−1)

Fig. 25. Effect of the solid mass flow rate on (a) radial profiles of axial gas velocity and (b)solid volume fraction (for the selected particle size ds = 206 μm). The data was taken atthe sample level of 3.9 cm below the gas exit pipe tip.

Fig. 26. Effect of the particle–wall restitution coefficient (es,w) on the separator ef

Fig. 27. Effect of specularity coefficient on the separator efficiency. The simulation wascarried out at the default operating conditions.

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Πk,g Influence of solid phases on gas phase (m2 s−3)Πε,g Influence of solid phases on gas phase (m2 s−4)ρs, ρg Solid and gas densities respectively (kg m−3)τ Shear stress tensor (kg m−1 s−2)σk, σε Constants (−)ϕ Angle of internal friction (degree)ϕksi Energy exchange between phase k and solid phase i

(kg m−1 s−1)φ Specularity coefficient (−)

Acknowledgment

The authors from the UK thank the Leverhulme Trust for a researchgrant (RPG-410).

ficiency. The simulation was carried out at the default operating conditions.

Fig. 28. Effect of the specularity coefficient on the (a) radial profiles of axial gas velocityand (b) solid volume fraction (for the selected particle size ds = 206 μm). The data wastaken at the sample level of 3.9 cm below the tip of the gas exit pipe.

382 X. Yu et al. / Fuel Processing Technology 126 (2014) 366–382

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