chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805
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Chemical Engineering Research and Design
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ydrodynamic properties of a cold model of dualuidized bed gasifier: A modeling andxperimental investigation
iddhartha Shresthaa, Brahim Si Alia,∗, Badrul Mohamed Jana,ookTzeng Limb, Khalid El Sheikha
Department of Chemical Engineering, Faculty of Engineering, University of Malaya, 50603 Kuala Lumpur, MalaysiaFuels & Combustion Group, Generation Unit, TNB Research Sdn. Bhd., 43000 Kajang, Selangor, Malaysia
r t i c l e i n f o
rticle history:
eceived 18 June 2015
eceived in revised form 30 March
016
ccepted 1 April 2016
vailable online 8 April 2016
eywords:
ual fluidized bed gasifiers (DFBGs)
ydrodynamic model
ressure loop
olid circulation rate
emi-empirical model
a b s t r a c t
Gasification in dual fluidized bed gasifiers (DFBGs) has been acknowledged to be one of the
promising technologies. The hydrodynamic characteristics play a vital role in the design
and development of such reactors, therefore, their modeling is worthwhile. In this study a
hydrodynamic model considering each section of the DFBG system – riser, cyclone, stand-
pipe, loop-seal, bubbling fluidized bed (BFB), connection with L-valve (lower connection) –
was developed to predict the pressure drops and solid circulation rates. The model param-
eters were fitted with the experimental data obtained from the cold model of DFBG in order
to improve the predicted results. A power law model was developed for the friction factor to
predict the pressure drop in an inclined chute pressure drop. The predicted results from the
model were found to be in reasonable agreement with the experimental results except for
the pressure drop in the lower connection. The lower connection pressure drop was found
to be affected by the riser velocity and L-valve aeration velocity. A correlation was devel-
oped to account for these effects, and was incorporated into the model, which improved the
asification in dual fluidized bed gasifiers (DFBGs) has beenegarded as a simple, reliable and cost-effective clean energy
ethod for the production of high quality syngas (Nguyent al., 2012; Wang et al., 2014). DFBGs, using steam as theasifying agent, yield a syngas composed of high H2 and COontent with calorific values ranging from 12 to 20 MJ/m3 thatould also be nearly free from N2. Syngas, also known asroducer gas, is a valuable product and bears numerous appli-ations, such as-electric power generation in a gas engine or
as turbine, hydrogen for fuel cells, liquid fuels or dimethyl
Abbreviations: BFB, bubbling fluidized bed; CaL, calcium looping; Cual fluidized bed gasifier; DFBG, dual fluidized bed gasifiers; DME, dim∗ Corresponding author. Tel.: +60 3 79676896; fax: +60 3 79675319.
ether (DME) synthesis by Fischer–Tropsch synthesis and theproduction of gaseous products, such as synthetic natural gas(SNG) (Purohit, 2009; Saxena et al., 2009). In addition, the qual-ity of syngas in DFBGs can be improved by optimizing thedesign and operation of the gasifier, by using catalytic bedmaterials to reform tars and methane in situ, and, finally, bya downstream cleaning process (Göransson et al., 2011).
Fig. 1 shows process block diagrams for different DFB (dualfluidized bed) reactors serving numerous purposes such asgasification (DFBGs), gasification with in situ CO2 capture
792 chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805
Fig. 1 – Concepts of dual fluidized bed, extended from Göransson et al. (2011).
[calcium looping (CaL)]. The basic concept of the DFB is tohave two fluidized beds as shown in Fig. 1 with the provi-sion of a solid transfer between them. In DFBGs, one reactoracts as a gasifier whereas other reactor acts as a combustor(Hofbauer et al., 1997). Several reactor arrangements capa-ble of being operated as DFBGs have been identified in theliterature. However, the reactor configuration of the riser, asa combustor, and bubbling fluidized bed (BFB), as gasifier, isconsidered to be the superior technical choice in terms ofparticle circulation, fuel conversion and tar elimination (Xuet al., 2006; Zhang et al., 2013). In this type of configuration,solid fuel is fed into the BFB gasifier where it is gasified with agasifying agent to produce the syngas that flows out from thecyclone and particulate filters before it can be consumed. Inaddition, the solids (char, bed materials, such as-sand, and un-reacted biomass and by-products of the gasification process,such as-tar) are transferred from the BFB to the riser wherethey are combusted along with the additional fuels to produceheat and flue gas, primarily N2, CO2, excess O2 and H2O. Theheated bed materials from the riser flows into the gas–solidseparators like cyclone and are circulated back to the BFB inorder to supply the essential heat required for the endother-mic gasification process. The BFB and the riser are generallycombined with the aid of the auxiliary systems, such as stand-pipe and non-mechanical valves (loop-seal and L-valve) sothat the bed material can be circulated amongst them. Thesenon-mechanical valves also prevent the mixing of gases fromthe gasifier to the riser and vice versa (Yazdanpanah et al.,2012a,b).
In AER also known as sorbent enhanced reforming (SER)CO2 is captured in situ in the BFB-gasifier during gasification(Bidwe et al., 2014; Cormos et al., 2014). In this process (Fig. 1),CO2 adsorbent bed material (such as limestone, dolomite, etc.)captures CO2 (Eq. (1)) simultaneously with gasification andheat transfer. This continuous removal of CO2 during gasifi-
cation enhances production of H2 (Eq. (2)) due to a shift in CO2
concentration on the R.H.S. of reaction (Eq. (2)):
CaO + CO2 → CaCO3 (1)
CO + H2O ↔ CO2 + H2 (2)
Chemical looping combustion (CLC), which offers efficient,low cost CO2 capture (Fan et al., 2012) and the Calcium Looping(CaL) process, which is for post combustion CO2 capture (Deanet al., 2011; Shimizu et al., 1999) (refer Fig. 1) are other emergingDFB technologies.
A good understanding of gas and solids hydrodynamics iscrucial for the design and development of the fluidized bedprocess since the performance of these fluidized bed reac-tors largely depends upon their hydrodynamic characteristics.Moreover, the hydrodynamic parameters have been found tosignificantly influence the rates of heat transfer, mass trans-fer, chemical reaction and erosion (Berruti et al., 1995; Issangyaet al., 1999; Lim et al., 1995; Zhu, 2005).
Modeling can be one of the valuable tools that can assistin understanding and analyzing these hydrodynamic parame-ters. Inherently, the flow structure of gas–solid mixture is verycomplex. Models designed for fluidized bed systems have beenclassified into three groups by Harris and Davidson (1994):
1. Type I models predict the axial solid fraction.2. Type II models predict the axial and radial solid fraction
distribution.3. Type III models employ the fundamental equations of fluid
dynamics to predict the two phase gas–solid flow.
Type III models are highly complex and require high com-putational time since they incorporate the system’ behaviorfrom fundamental conservation equations held at boundaryconditions. Nevertheless, these models have been developedand studied for DFBGs (Manchasing et al., 2013; Nguyen et al.,2012; Peng et al., 2015; Seo et al., 2011; Wang et al., 2014).
On the other hand, type I and type II models are simpler innature, easier to apply and can also be coupled with reaction
chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805 793
Table 1 – Dimension of the experimental setup.
Parameter Value Unit
Height of the riser (Hr) 3 mDiameter of the riser (Dr) 0.0508 mHeight of the BFB in narrow section (HBFB1) 1.04 mDiameter of the BFB in narrow section (DBFB1) 0.0762 mHeight of the BFB in wide section (HBFB2) 0.1143 mDiameter of the BFB in wide section (DBFB2) 0.1143 mHeight of standpipe (Hsp) 0.955 mDiameter of standpipe (Dsp) 0.0127 mLength of loop-seal (Lls) 0.0318 mWidth of loop-seal (Wls) 0.0159 mDiameter of upper connection (Dcon1) 0.0127 m
mca1f2ie
heLd(iomiiDmsrisbt
dihcdict
2
Efiocspwosr
Table 2 – Particle properties.
Bed material Silica sand
Size range (�m) 50–400Particle diameter (�m) 222Particle density (kg/m3) 2600Bulk density (kg/m3) 1440Sphericity (ϕ) 0.92Archimedes Number 927Minimum fluidization velocity (Umf) (m/s) 0.04a
Terminal velocity (Ut) (m/s) 1.59b
Geldart group of classification B
a Remf =√
C21 + C2Ar − C1, C1 = 33.7, C2 = 0.0408.
b Haider and Levenspiel (1989).
Table 3 – Operating conditions.
Parameters Value Unit
Gas density 1.16 kg/m3
Gas viscosity 1.8680E−5 Pa sGas velocity in riser (Ur) 1.83–2.17 m/sGas velocity in the BFB (UBFB) 0.07 m/sGas velocity in loop-seal (Uls) 0.07 m/sGas velocity in L-valve (Ulv) 0.07–0.145 m/s
Diameter of lower connection (Dcon2) 0.0127 m
odels (Pugsley and Berruti, 1996). In addition, type I modelsan be employed as design tools to study the effect of oper-ting parameters (Kaiser et al., 2003; Smolders and Baeyens,998). The employment of semi-empirical approaches can beound in recent literature (Guedea et al., 2013; Noubli et al.,015). However, they are considered to be highly empiricaln nature, lacking in generality, and often require additionalxperimental data as input (Lim, 2012).
Type I models have been developed previously to study theydrodynamics in the cold model of DFB technology (Diegot al., 2012; Kaiser et al., 2003; Karmakar and Datta, 2010;isbona et al., 2013). However, only a few of them involve aual fluidized bed gasifier (Shrestha et al., 2015a). Kaiser et al.
2003) implemented a hydrodynamic model to study the stabil-ty of operation, the pressure profile and the solid circulationf the loop. The model was validated in a scaled cold flowodel DFBG system incorporating a loop-seal and chute. The
nventory, geometry and solid circulation rate were found tonfluence the stability and pressure balance. Karmakar andatta (2010) also employed a similar kind of hydrodynamicodel and studied the axial voidage profile in the riser, pres-
ure drop across various components and the solid circulationate. The results were validated in a cold model of DFBGncorporating L-valves. The aeration flow in the L-valve, theecondary air-flow in the riser and particle diameter of theed material were shown to have the strongest influence onhe system pressure drop and solid circulation rate.
The aim of this work is to develop a semi-empirical hydro-ynamic model for a cold model of a dual fluidized bed gasifier
ncorporating a loop-seal and L-valve to predict the basicydrodynamic properties such as pressure drop and solid cir-ulation rate. In the process, various model parameters, whichepend upon the particle properties, the geometry and the flu-
dization conditions were adjusted. The predicted results wereompared with the experimental results in order to validatehe hydrodynamic model.
. Experimental
xperiments were conducted in a dual fluidized bed gasi-er cold model as shown in Fig. 2. The test rig consistedf two interconnected fluidized beds of configuration riser-ombustor and the BFB-gasifier with an extended upperection along with other components, such as cyclones, stand-ipe, loop-seal, L-valve and connections. Except the cyclone,hich was made of steel, all the other components were madef acrylic plastic. The dimensions of the DFBG components are
hown in Table 1. Solids from the BFB could flow toward theiser from any of the three ball valves (V3, V2 and V1) placed
Inventory (Iv) 1.5 kg
at different heights: 1.5*DBFB1, 2*DBFB1 and 2.5*DBFB1, from theBFB distributor via the lower connection. In this study onlyvalve V3 was opened while valves V1 and V2 were always shutoff. Dry silica sand was used as the bed material. The particlesize distribution (PSD) of the bed material is shown in Fig. 3,and their properties are presented in Table 2. The particleswere supported on a sintered metal distributor. The fluidiz-ing medium used was compressed air, which entered the airbox through the air filter and regulator. From the air box, airflowed into the riser, the BFB, loop-seal and L-valve via flowmeters. The fluidization velocity in the riser was varied from1.67 to 2.17 m/s (1.05–1.36 Ut) whereas the fluidizing veloci-ties in the BFB as well as the aeration velocity in the loop-sealwere kept constant at 1.75 Umf throughout the experiment. TheL-valve aeration velocity was adjusted from 0.07 to 0.145 m/s(1.75–3.62 Umf) in order to regulate the solid circulation rate.The operating conditions are listed in Table 3.
The differential pressure transducers were connected tothe pressure taps at different heights in the riser, the BFB,as well in the other components of the experimental setupto record the differential pressure drop are shown in Fig. 2.The pressure drop in the riser (�Pr), riser exit (�Pre), cyclone(�Pcyc), standpipe (�Psp), loop-seal (�Pls), connection 1 (theupper connection), BFB (�PBFB) and connection 2 (the lowerconnection (�Plc)) was measured by connecting the respectivepressure transducers as shown in Fig. 2. The pressure drop inconnection 1 (P6 and P7) was observed to be negligible at alltimes.
A steady state was achieved when the riser and the BFBpressure drop were observed to be stable for a minimum of10 min. The aeration velocity in the loop-seal was stopped andthe rate of increase of the bed height in the standpipe wasmeasured. The bulk density and geometry were then used tocalculate the mass flow rate in standpipe according to Eq. (3).The obtained mass flow rate was converted to solid circulationrate at any locations according to Eq. (4):
m = �z
�t∗ �b ∗ Asp (3)
794 chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805
Fig. 3 – Cumulative less than particle size distribution of
silica sand.
Gs,sp = m
Asp, Gs,r = m
Ar, Gs,con = m
Acon(4)
The pressure drop and the solid circulation rate were mea-sured:
of experimental setup.
1. By varying the riser velocity while keeping the L-valve aer-ation velocity constant and
2. By varying the L-valve aeration velocity while keeping theriser velocity constant.
More details of the experiments can be found in Shresthaet al. (2015b).
3. Modeling
The model was executed in MATLAB, with the geometry, thefluidization conditions, the particle properties and the riserpressure drop supplied as input parameters. Each section ofthe model – the combustion zone (riser), cyclone separator,standpipe, loop-seal, gasification zone (BFB), connection andL-valve – were calculated iteratively, as shown in Fig. 4. Theiteration was carried out until the pressure balance loop wasfulfilled (ref. Section 3.9).
3.1. Riser
When modeling the riser, the most significant hydrodynamicparameter is the solid fraction or solid hold-up because it
chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805 795
Assumption of height of dense zone in riser and iterate it until pressure drop in the riser is fulfilled
Riser (Gs, Pr)
Exit duct
Cyclone
Loop Sea l
Press ure Balance(Eq. (44))
BFB
Connection & L-valve
Mass Balance(Eq. (43))
Press ure Balance(Eq. (45))
Mass di stributi on, p ress ure drop
Geometry, Fluidization conditions
Standpipe
Ass ume height of bed in BFB
Ass ume vo idage as mf
If Gs,exp - Gs,model>0.1,Gs = Gs,exp
pres
gGctI
3Tsflb2z1raa(tFceb
Fig. 4 – Structure of
reatly influences the pressure drop and solid circulation rate.enerally, on the basis of the solid fraction profile, the riseran be divided into two zones: dense zone and upper zone. Ifhe riser exit is constricted an additional exit zone may occur.n the current study, all three zones were considered.
.1.1. Dense zonehe dense zone is characterized by a time-averaged uniformolid concentration. Although large fluctuations due to theow of bubbles have resulted in different time-averaged denseed voidages in both the riser and BFB (Gungor and Eskin,007; Johansson et al., 2007), it has been shown that the denseone in the riser can be modeled as BFB (Svensson et al.,993). However, the correlations used should be within theiranges of validity (Gómez-Barea and Leckner, 2010; Pallaresnd Johnsson, 2006). In this study, the dense bed was modeledccording to the widely accepted modified two phase theoryJohnsson et al., 1991). The dense bed was further divided intohe: solid-free bubble phase and solid-laden emulsion phase.inally, the overall dense bed voidage and solid fraction werealculated using Eqs. (5) and (6), respectively, assuming the
mulsion phase at the minimum conditions and evaluating ıby Eq. (7). Other required gas solid flow properties and fluid
sure balance loop.
dynamic correlations are listed in Table 4. The minimum flu-idization voidage (εmf) was assumed to be 0.5:
εdz = ıb + (1 − ıb)εe (5)
εs,dz = 1 − εdz (6)
ıb = 11 + (Ubr/Uv)
= 11 + (Ubr/Uo − Ue − Utf )
(7)
3.1.2. Upper zoneThe upper zone of the riser consists of a splash zone and atransport zone. The particles from the dense bed are ejectedtoward the freeboard by erupting bubbles at the surface andby the gas flow. This bubble eruption forms a splash zone,which is characterized by a strong backmixing of particles.Above the splash zone, a transport zone exists where the par-ticles are transported from the core into the wall layers, thusexhibiting a core-annulus flow structure. In the annular walllayers, the gas velocity is significantly lower and the particlesfall downwards (Pallares and Johnsson, 2006). Various semi-empirical models have been developed for predicting the solid
fraction in a riser (Kunii and Levenspiel, 1991; Li and Kwauk,1980; Rhodes and Geldart, 1987; Smolders and Baeyens, 2001).
796 chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805
Table 4 – Fluid dynamic correlation for gas flow pattern in bottom bed of riser.
In this work, Eq. (18) (Zenz and Weil, 1958) shown to be validfor DFBGs (Karmakar and Datta, 2010; Löffler et al., 2003) hasbeen used:
εs − ε∗sεs,dz − ε∗s
= exp[−ad(h − hdz)] (18)
‘ad’ in Eq. (18) represents the decay factor. The value ofthis factor depends on the particle properties, operating condi-tions, design and dimensions of the unit. However, the valuediffers extensively in the literature and several correlationshave been put forward. The correlation developed by Adanezet al. (1994), as shown in Eq. (19), has been broadly used(Kaiser et al., 2003; Karmakar and Datta, 2010; Miao et al., 2012;Shrestha et al., 2014) and was followed in this work as well:
ad(Uo − Ut)2D0.6 = 3.5 − 1670dp (19)
In order to evaluate the distribution of the solid fractionfrom Eq. (18), in addition to the decay constant, ad, the solidfraction at the dense bed, εdz and at exit, ε∗s are required. Theformer can be found from the dense bed model whereas thelatter is known as the solid fraction above the transport dis-engaging height (TDH) or saturation carrying capacity of gas(ε∗s ) is ambiguous (Gómez-Barea and Leckner, 2010; Kunii andLevenspiel, 1991). The common method to find ε∗s is describedby Eq. (20):
ε∗s = Gs�p(Uo − Ut)
(20)
In the literature, various correlations are available (Bai and
Kato, 1999) to evaluate the value of the solid fraction at the exit(ε∗s ). However, the equation requires information of the solid
circulation rate (Gs), but in cases where only the pressure dropis known, ε∗s may be determined from the particle elutriationrate constant (K∞) according to Eq. (21) (Löffler et al., 2003):
ε∗s = K∞�p(Uo − Ut)
(21)
Different correlations for K∞ have been proposed whichcan be found summarized in Löffler et al. (2003). In this study,a correlation represented by Eqs. (22)–(25) provided by Wenand Chen (1982) valid for dp = 37–3400 �m, �p = 860–7850 kg/m3,Uo = 0.1–10 m/s, D = 0.034–2.06 m was adopted:
K∞ = �p˛i(Uo − Ut) (22)
with
˛i = 1 −(
1 + fs(Uo − Ut)2
2gD
)−1/4.7
(23)
where fs is the coefficient of friction and is given by,
fs�p
d2p
(�
�g
)2.5
= 5.17
[�g(Uo − Ut)dp
�
]−1.5
D2, for
�g(Uo − Ut)dp�g
≤ 2.38D
(24)
fs�p
d2p
(�
�g
)2.5
= 12.3
[�g(Uo − Ut)dp
�
]−2.5
D, for
�g(Uo − Ut)dp�g
≤ 2.38D
(25)
chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805 797
3t(ic
m
C
t
U
3d
G
ttvta
3d(c
�
us
�
ws�
affH�
ih
�
.1.2.1. Solid velocity. The solid velocity is determined fromhe force balance on the solid particle, as shown by Eqs.26)–(28) (Löffler et al., 2003). Where FB is buoyancy force, FG
s gravitational force and FD is the drag force. CD is the dragoefficient determined from the relation shown in Eq. (27):
sdUsdt
= msUsdUsdz
= FB − FG + FD = (�g − �p)g�d3
p
6
+ CDεn�g
(Ug − Us)|(Ug − Us)|2
�d2p
4(26)
D =
⎧⎪⎪⎪⎨⎪⎪⎪⎩
24Rep
for Rep ≤ 5.8,
10
Re0.5p
for 5.8 < Rep ≤ 540
0.43 for Rep > 540
(27)
As it is assumed that the gas flows upward in the core only,he average gas superficial velocity in the core, Ug, is simply:
g = Uoε
(28)
.1.2.2. Solid circulation rate. The solid circulation rate (Gs) isetermined by the following correlation,
S = �pεsUs (29)
The solid circulation rate from the model is compared withhe experimental values. If the accuracy of the model predic-ion is poor, the experimental values are used instead as inputalues. This is undertaken so that the accuracy of the model inhe sections where the solid circulation rate is required, suchs standpipe and L-valve can be improved.
.1.2.3. Pressure drop. The pressure drop of the bottom zone isetermined by static heads of the bed particles as given by Eq.
30), assuming that the solids acceleration and decelerationompensate each other and neglecting friction forces:
Pdz =∫ Hdz
0
�pεs,dzg dh (30)
Whereas, the pressure drop in the transport zone is madep of four main components (Pugsley and Berruti, 1996) ashown by Eq. (31):
Ptz = �Phs + �Pacc + �Psf + �Pgf (31)
here �Phs is the pressure drop due to the hydrostatic head ofolids, �Pacc is the pressure drop due to solids acceleration andPsf and �Pgf are the pressure drops due to the solids frictionnd gas friction, respectively. Solids friction is defined as therictional force between the solids and the wall, while the gasriction is the frictional force between the gas and the solids.owever, in this study, the pressure drop components �Pacc,Psf and �Pgf were neglected since the maximum gas velocity
n the riser was only 2.17 m/s (Lim, 2012). Only the hydrostaticead of solids was considered and calculated as:
∫ H
Phs =Hdz
�pεsg dh (32)
Finally, the total riser pressure drop is:
�Pr = �Pdz + �Ptz (33)
The height of the dense zone is iterated until the riser pres-sure drop converges according to Eq. (33).
Similarly, the amount of the bed material at any height isby,
mdz =∫ Hdz
0
Adz�pεs,dz dh (34)
mtz =∫ H
Hdz
Atz�pεs dh (35)
mr = mdz + mtz (36)
3.2. Riser exit
The riser exit pressure drop is measured from point P2 to P3in Fig. 2. This includes the top exit zone of the riser as well asthe exit duct connecting the riser to the cyclone.
3.2.1. Exit zoneRisers with an abrupt exit have considerable recirculation ofsolids or internal reflux near the exit zone, causing solid frac-tion to increase in this zone (Charitos et al., 2010; Johanssonet al., 2007). The solid fraction is the sole major contributortoward the pressure drop, and the contribution from solidsfriction or acceleration is negligible in the exit region (Mabrouket al., 2008). A semi-empirical model developed to estimate theratio of solids that exit the riser to solids that recirculate backinto it (Lim et al., 2012), as shown in Eqs. (37)–(38), is adoptedin order to calculate the pressure drop in this zone.
mso = ˚d
(w
�D
)mcore (37)
˚d =(�D
w− 1
)exp(−aεs,core) + 1 (38)
where ˚d is aerodynamic factor, εs,core is solid fraction closerto riser exit, w is the opening width to the cyclone from riserand coefficient, a = 3668.
Since the value of the riser solids outflow (mso) is known(from experimental Gs values), mcore (which is the solids flowrate up the core of the riser) is calculated from Eq. (37) byusing Eq. (38). Then the internal reflux is obtained by subtrac-ting the riser outflow mso by mcore. The internal reflux obtainedwas used to calculate the solid fraction at this section (assum-ing particle falling at terminal velocity) in order to obtain thepressure drop in this exit zone (�Pez).
3.2.2. Riser exit ductThe pressure drop in the horizontal section between riser andcyclone has been modeled by Eq. (39) given by Patience et al.(1990).
�Pred = Gred(2.84 + 0.0108U2red) (39)
where Gred and Ured are the solid flux and gas velocity in thesection.
Gred = GsArAred
(40)
798 chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805
0
0.5
1
1.5
2
2.5
0.80.70.60.50.40.30.20.10
Ris
er V
eloc
ity
Pcyc ( kPa)
Fig. 5 – Cyclone pressure drop for different riser velocity.
Now, the total pressure drop in the exit region consists ofthe pressure drop at the riser exit zone and pressure drop atriser duct.
�Pre = �Pez + �Pred (41)
3.3. Cyclone
The pressure drop in the cyclone is directly proportional to thesquare of the cyclone inlet velocity (Ucyc) and is calculated byEq. (42) (Perry et al., 1997):
�Pcyc = kcyc�gU2cyc (42)
where kcyc is a function of cyclone dimension. Based on thecold model pressure drop measurements (Fig. 5), in this studythe kcyc was assumed to be 3.3.
3.4. Standpipe
The dense bed above the loop seal in the standpipe wasobserved to be a non-fluidized transitional packed bed flowwith negative slip velocities (Zhang and Rudolph, 1991). Inthis condition, the pressure drop across the standpipe can beexpressed as a function of the slip velocity as described by Eqs.(43) and (44) (Ergun, 1952; Kim et al., 1999; Yang et al., 2009) byimposing a positive pressure drop (Yazdanpanah et al., 2012a).Basu and Cheng (2000) showed that the superficial velocity ofgas up the standpipe (Usp) is small and is equal to fraction ı ofthe air flow supplied to the supply chamber of the loop-seal.The ı is taken as 0.095 (Basu and Cheng, 2000):
|�P|spL
= 150�g(1 − ε)2
(ϕdp)sε2Usl +
1.75�g(1 − ε)(ϕdp)ε
U2sl (43)
The slip velocity can be expressed as,
Usl = Usp
ε− Us,sp
1 − ε= Usp
ε− Gs, sp
�p(1 − ε)(44)
The voidage across the standpipe, loop-seal, L-valve andorifice increases with Usl, as described by Eq. (45) (Kim et al.,1999; Knowlton et al., 1978; Yang et al., 2009):
ε = εb + UslUmf /εmf
(εmf − εb) (45)
3.5. Loop-seal
In this study the loop-seal velocity was kept constant at 1.75Umf throughout the experiments. The same model used to
describe the dense zone of riser (Section 3.1.1) is used to esti-mate voidage and pressure drop in the loop-seal. The voidageand pressure drop in the loop-seal is estimated conservativelyfrom Eqs. (5) and (30), respectively (Cheng and Basu, 1999;Kaiser et al., 2003).
3.6. Modeling of BFB
The correlation for predicting the pressure drop and voidagein the BFB are similar to those for the dense zone of the riser.The voidage in the BFB can be obtained from the dense bedmodel, as described in Section 3.1.1 by using Eqs. (5) and (30),respectively.
3.7. Modeling of lower connection
The lower connection consists of a connection/chute placed atan angle together with a L-valve. The total connection pressuredrop is given as follows:
�Plc = �PLv + �Pcon (46)
3.7.1. Modeling of L-valveSolids in the vertical section of the L-valve were observedas a packed flow. The pressure drop across this sectionwas calculated as a function of the slip velocity using Eqs.(43)–(45), similar to the standpipe. However, unlike the stand-pipe, the slip velocities obtained were positive for the L-valve,and imposing a positive pressure drop was not required.Yazdanpanah et al. (2012a) suggested that the moving bedvoidage in the packed bed flow is constant and correspondsapproximately to the tapped bed voidage. However, for theoperating conditions of L-valve in this study, voidages by usingEq. (45) would result in voidage values higher than the mini-mum fluidization voidage. This discrepancy might be becausethe range of slip velocities in their study was 0.01–0.05 m/s,whereas, in this study, it was 0.12–0.25 m/s. Chan et al. (2009)stated that the average flow voidage in L-valve is constant(∼0.5) and independent of the rate of air injection. Thus, thevoidage in L-valve was assumed to be εmf (which is 0.5).
3.7.2. ConnectionThe pressure drop across the inclined connection can be esti-mated from Eqs. (47)–(50), as suggested by Kaiser et al. (2003).Considering that the fluidization conditions in the connectionare similar to the L-valve (Yazdanpanah et al., 2012a), the solidvelocity was calculated from solid circulation rates evaluatedfor connection assuming voidage εcon = εmf. In Eq. (47), is thefriction factor, and depends upon the materials used, theirshape and size. Kaiser et al. (2003) estimated the friction factorfrom Eq. (48) with coefficient kCON = 3.5, which was determinedaccording to the pressure drop measurements:
�Pcon = �Pcon,hs − �Pcon.fr (47)
�Pcon,hs = g�p(1 − εcon)lcon sin (48)
�Pcon,fr = U2s,con
2Awall
Across sec�p(1 − εcon) (49)
= kconUs,con
(50)
chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805 799
(a)
(b)
-30.0
-25.0
-20.0
-15.0
-10.0
-5.0
0.0
5.0
10.0
0 20 40 60 80
Pres
sure
dro
p (k
Pa)
Sold circulatio n rate (K g/m2s)
kcon = 3.5
kcon = 0. 80
Exp
y = 48. 807x-1.014
R² = 0. 996 4
0
0.5
1
1.5
2
2.5
3
80706050403020100
kcon
Sold circ ula tion rate - Gs, con (Kg/m2s)
prdnbhlgsdl
3
Tsthm
M
3
Tf
P
Fig. 7 – Model and experimental results for solid circulationrate for different riser velocity.
Fig. 6 – Fitting procedure to estimate the kcon.
Fig. 6a shows the model results for the lower connectionressure drop with kCON = 3.5 and 0.80. The value of kCON = 3.5esulted in a higher value of , and a higher frictional pressurerop that leads to a negative total pressure drop for the con-ection. Positive pressure drops were only obtained for kCON
elow 0.80. However, with kCON < 0.80, the pressure drop wasigher for the solid circulation rates at the lower level. At the
ower rate of solid circulation rate, kCON is higher. This sug-ested that the value of kCON is dependent on the operatedolid circulation rate. Thus, kCON was fitted to the pressurerop results and a relation between kCON and the solid circu-
ation rate was obtained, as shown in Fig. 6b.
.8. Mass balance
he mass of the solids inside the riser, BFB, standpipe, loopeal, connection and L-valveare calculated with the informa-ion of the geometry (volume) and the solid fraction. Theeight of the bed material in the BFB was iterated until theass balance of solids according to the Eq. (51) was satisfied:
T = mBFB + mr + msp + mls + mcon + mLv (51)
The total inventory (MT) was 1.4 kg.
.9. Pressure loop
he height of the solids in the standpipe was determined toulfill the pressure balance loop of the loop-seal and standpipe:
OUT,riser − POUT,gasifier = �Psp − �Pls (52)
The pressure balance of the dual fluidized bed is given byEq. (51). The riser pressure drop was calculated iteratively untilthis balance is satisfied:
In this study the riser pressure drop was supplied to themodel as input and was used for model convergence.
4. Results and discussion
4.1. Solid circulation rate
The riser velocity has been found to influence the solid cir-culation rate significantly and is often used as a controllingparameter. With the increase in riser velocity, the solid circu-lation rate increases (Karmakar and Datta, 2010; Kronbergeret al., 2005). However, the enhancement in solid circulationrate by the riser velocity is limited i.e. the solid circulation ratefalls after reaching a certain maximum value while increas-ing riser velocity. This maximum value of solid circulationrate is termed as the saturation carrying capacity of the gasin the riser for the bed materials (Bai et al., 1992). Moreover,limitation in the solid circulation rate also occurs when thefeed to the riser is controlled. The riser outflow will increaseup to a certain level with the riser velocity increased, butwhen the riser velocity reaches where the feed is less thanthe riser outflow, the solid circulation will start to decreasesince the feed is not capable to replenish the material that istransported out. In the current study, the later behavior wasobserved.
The obtained solid circulation rate from the model andexperiments for different riser velocity at L-valve aerationvelocity of 0.09 m/s, 0.12 m/s and 0.145 m/s is shown in Fig. 7.The solid circulation rate escalated with increasing riser veloc-ity and L-valve aeration velocity until riser velocity 2 m/s. Thesolid circulation rate fell increasing the riser velocity further to2.17 m/s. The observed limitation in the solid circulation ratewas found to be caused by a slower material recirculation rateback into the riser for riser velocity 2.17 m/s for L-valve aera-tion velocity of 0.145 m/s. This limitation in solid circulation
rate at riser velocity 2.17 m/s can be overcome by increasingthe L-valve aeration velocity to 0.17 m/s which has been shown
800 chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805
Fig. 8 – Model and experimental results for riser pressuredrop for different riser velocity at Ulv = 0.12 m/s, and
Fig. 9 – Ratio of model to experimental results on height ofdense zone in riser for different riser velocity atUlv = 0.12 m/s and Ulv = 0.145 m/s.
show this trend. However, the model prediction of the riserexit was always found to be under predicted for all conditions
Ulv = 0.145 m/s.
by Shrestha et al. (2015b), however up to a certain L-valve aer-ation velocity known as maximum L-valve aeration velocity(Chan et al., 2009; Shrestha et al., 2015b).
Compared to the experimental results the model resultswere not capable to predict the solid circulation rate accu-rately. This was mainly because the developed model for solidcirculation rate described by Eq. (29) only took an account ofriser velocity but did not include the effect by a variation inthe L-valve aeration velocity. Clearly, this suggests that theequation considering only riser velocity is only not suitablefor estimating solid circulation rate in riser feed controlledDFBGs. Aeration in the L-valve or any other non-mechanicalvalve employed should be taken into consideration while pre-dicting the solid circulation rate. The aeration velocity in solidflow control valve has been classified as a necessary param-eter while analyzing the gas–solid flow in CFB risers (Xu andGao, 2003) while in DFBGs they are essential because they cancontrol the global circulation rates1 (Bidwe et al., 2014). In thisstudy, aeration velocity in L-valve is main controlling param-eter for solid circulation rate. A model has been developed toincorporate the effect L-valve aeration velocity together withriser velocity and published elsewhere (Shrestha et al., 2015b).However, for the purposes of the better prediction and furtheranalysis, the value of solid circulation rate obtained experi-mentally was used as inputs to the model in the current study.
4.2. Riser pressure drop
The total riser pressure drop was observed to decrease withthe increase in riser velocity, as shown in Fig. 8, at a constant L-valve aeration velocity. With the increase in the riser velocity,the solid entrainment from the riser was increased, resultingin a loss in the height of the bed that was formed above theriser distributor, thus reducing the riser pressure drop.
The model results for riser pressure drop, as described inSection 3.1, also resulted in a similar trend. A decrease in theriser pressure drop in the model is signified by decreasing theheight of the dense bottom zone (Kaiser et al., 2003). Fig. 9
shows the ratio of the model to the experimental results on the
1 Global circulation rates: in DFB systems the solids exchangebetween the two fluidized bed are often referred to as globalcirculation rates (Pröll et al., 2009).
height of the dense zone in the riser for increase in riser veloc-ity for L-valve aeration velocities of 0.12 m/s and 0.145 m/s,respectively. It can be seen that the model over-predicts theheight of the dense zone in the riser. This might be because themodel is not capable of significantly justifying the gas–solidmixing mechanism above the dense zone and in the lowerregions of the riser (Lim, 2012). This also suggests that themodel under-predicted the riser upper section pressure drop.
4.3. Riser exit and cyclone
Fig. 10a and b shows the model and experimental results forthe pressure drop in the riser exit and cyclone for increasein riser velocity with L-valve aeration velocity of 0.12 m/s and0.145 m/s, respectively. The pressure drop at the riser exit wasfound to increase with increase in the riser velocity, up toa riser velocity of 2.17 m/s, signifying the reduction in thesolid circulation rate. Both the model and experimental results
Fig. 10 – Model and experimental results on pressure dropin riser exit and cyclone for different riser velocity atdifferent riser velocity (a) Ulv = 0.12 m/s and (b)Ulv = 0.145 m/s.
chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805 801
Fig. 11 – Model and experimental results on the BFBpressure drop for different riser velocity at (a) Ulv = 0.12 m/sa
wu
wkac
4
TiisitwIdoTttoti(i
4
Ftacifods2T
E
Fig. 12 – Ratio of model to experimental results on height ofbed in the BFB for different riser velocity at Ulv = 0.09 m/s,Ulv = 0.12 m/s, and Ulv = 0.145 m/s.
Fig. 13 – Model and experimental results on lowerconnection pressure drop for different riser velocity at (a)
nd (b) Ulv = 0.145 m/s.
ith a relative deviation2 of 32%. This might be due to thender-prediction in the reflux ratio.
The cyclone pressure drop was always found to increaseith the increase in riser velocity. The assumed parameter
cyc for cyclone pressure drop yielded a good prediction (rel-tive deviation of 7–22%) for the cyclone pressure drop whenompared to the experimental data.
.4. Standpipe and loop-seal
he pressure drop in the standpipe was increasing with thencrease in riser velocity until the solid circulation rate wasncreasing. Same was obtained from the model as well. Also,ince the pressure drop in the standpipe incorporated thenfluence of the solid circulation rate, the effect of a reduc-ion in the solid circulation was noticed. The predictions wereithin the acceptable range, with a relative deviation of 7–22%.
n contrast, the gas solid flow behavior in the loop-seal wasescribed as a dense bed using the modified two-phase the-ry and did not include the effect of the solid circulation rate.he pressure drop solely depended upon the loop-seal aera-
ion velocity, which was constant at all conditions. Therefore,he loop-seal pressure drop obtained from the model wasbserved to remain constant. However, the experiments showhat the loop-seal pressure drop increases with an increasen the riser velocity. This means that the models (Eqs. (5) and30)) should have incorporated the additional weight from thencrease in the solids circulation.
.5. BFB pressure drop
ig. 11a and b shows the model and experimental results forhe BFB pressure drop for increasing riser velocity at L-valveeration velocity of 0.12 m/s and 0.145 m/s, respectively. Whenompared to the experimental results after altering the totalnventory, the model results for the BFB pressure drop wasound to be within the tolerable range, with relative deviationsf 10–20%. With the increase in riser velocity, the BFB pressurerop increases because the riser entrainment increases andhifts the solids mass from the riser to the BFB (Kaiser et al.,003). This shift in the mass increases the BFB pressure drop.
he model also predicts a similar trend with the experiments.
2 Relative deviation % = (abs(Experimental Value − model value)/xperimental value) * 100.
Ulv = 0.09 m/s, (b) Ulv = 0.12 m/s and (c) Ulv = 0.145 m/s.
The mass balance loop in the model is the criterion that influ-ences the BFB pressure drop by varying the height of the bedin the BFB. The decrease in the riser pressure drop with theincrease of riser velocity is caused by the decrease in the riserdense zone height and also total riser mass. Thus, in order tofulfill the total mass balance loop, this decrease in riser massis over-come with the increase in the solids mass in the BFBby increasing the BFB bed height. This, in turn, increases thepressure drop in the BFB. However, the ratio of the BFB heightpredicted by the model to the experimental is under-predictedfor all other riser velocities except 1.67 m/s, as can be seen inFig. 13. The over-prediction of the BFB pressure drop, as seenin Fig. 12, and under-prediction in height of bed in the BFB,as seen in Fig. 13 suggests that the solid fraction was overpredicted.
4.6. Lower connection pressure drop
Fig. 13a–c shows the model and experimental results on thelower connection pressure drop for increasing the riser veloc-ity at the L-valve aeration velocity of 0.09 m/s, 0.12 m/s and0.145 m/s, respectively. The model predictions were accuratewhen compared to the experimental results, thus validating,the correlation for kcon in Fig. 6b for the conditions investi-
gated (Fig. 14a). The model results obtained increased withthe solid circulation rate and L-valve aeration velocity. This
802 chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805
Fig. 14 – (a) Pressure drop over lower connection. (b) Model prediction of hydrostatic pressure drop and frictional pressure m/s
drop in connection and pressure drop in L-valve at Ulv = 0.12
can be further illustrated by Fig. 14b, which shows the modelprediction of hydrostatic pressure drop and frictional pressuredrop in the connection, and pressure drop in the L-valve atUlv = 0.12 m/s and Ulv = 0.145 m/s, for different solid circulationrates. The hydrostatic pressure drop was obtained constantfor all operating conditions since the voidage was assumed tobe constant, as minimum fluidization voidage. Although, thefrictional pressure drop was varied slightly with the solid cir-culation rate, the variation was within the same range. Thepressure drop in the L-valve was found to increase with theincrease in the L-valve aeration velocity and varied little withthe solid circulation rate. The increase in the pressure dropwas due to the rise in the slip velocity as the L-valve aera-tion velocity increased. This increment in the L-valve pressuredrop led to an increase in the lower connection pressure dropwith the increase in the L-valve aeration velocity, as shown inFig. 13. However, with the increment of solid circulation ratealthough the solid velocity increased, its influence was neg-ligible in terms of slip velocity, therefore the effect was notinfluential.
In addition, the model prediction of lower connection wasnot affected by the change in riser velocity. The pressure dropin the lower connection was observed to decrease with theincrease in the riser velocity at the constant L-valve aerationvelocity (Fig. 13). This signifies the tendency of the gas bypassfrom the riser to the BFB, as well as a reduction of the solidflow rate through the lower connection (Lim et al., 2014) with
the increase in riser velocity. Since it is challenging to con-trol the solid flux and maintain the pressure drop over this
and Ulv = 0.145 m/s.
connection the non-mechanical valves are used in the riserfeed section to overcome this pressure drop. In this work, aL-valve is embedded within the lower connection. The exper-iments were conducted at constant L-valve aeration velocitybut with increasing riser velocity. Therefore, a reduction inthe pressure drop can be related to the riser velocity. Exper-iments with increasing L-valve aeration velocity at constantriser velocity showed that the lower connection pressure dropincreased with the increase in L-valve aeration velocity. Aneffort was made correlate the pressure drop with the riser andL-valve aeration velocity. In order to develop a correlation, thedifference between the model result and experimental resultfor the respective riser velocity and L-valve aeration velocitywas investigated based on regression analysis, which is shownin Eq. (54):
This term �Prp(reduced pressure drop) was subtracted onthe right hand side of Eq. (55), as follows and the model wasmodified:
�Plc = �Pcon + �PLv − �Prp (55)
Fig. 15 shows the comparison of the model predictionversus the experimental results after the modification of the
model. It can be seen to be good estimate, as shown by the ratioof the predicted values to the experimental results in Fig. 16.
chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805 803
Fig. 15 – Modified model and experimental results on lowerconnection pressure drop for different riser velocity at (a)Ulv = 0.09 m/s, (b) Ulv = 0.12 m/s and (c) Ulv = 0.145 m/s.
Fig. 16 – Ratio of model to experimental results on lowerconnection pressure drop for different riser velocity afterm
tli
5
GadtstcmmswnptLw
odification.
Eq. (55) was unable to be compared with the experimen-al results from the literature, since the pressure drop in theower connection was not widely reported. Thus, further works required to improve the general applicability of Eq. (55).
. Conclusion
asification in dual fluidized beds has been regarded to ben efficient and emerging technology. In this study, a hydro-ynamic model for a dual fluidized bed gasifier was developedo study the behavior of the hydrodynamic properties in theystem. The dual fluidized bed gasifier consists of a riser (ashe combustor), a BFB (as the gasifier), a loop-seal as the upperonnection between the two reactors (to circulate the bedaterial), and an L-valve as the lower connection. Differentodel parameters were obtained fitting with cold model mea-
urements. The riser pressure drop and solid circulation rateere supplied as the input values. The riser and lower con-ection pressure drop were found to decrease whereas theressure drop in the riser exit, cyclone and BFB were foundo increase with the increase in riser velocity at a constant
-valve aeration velocity. The model results for the standpipeere satisfactory and also varied with solid circulation rate
whereas in case of loop-seal, a model to incorporate the addi-tional weight of solid circulation rate is required. A powerlaw model for kcon (instead of constant value) was developedbased on the solids circulation rate to predict the inclinedchute connection pressure drop. In addition, the lower con-nection pressure drop was found to be influenced by the riservelocity and L-valve aeration velocity, and a correlation wasestablished to calculate this pressure drop. The developed cor-relation was embedded in the pressure balance loop, and themodel results improved in accuracy when compared with theexperimental data.
List of symbols
ad decay constantA area, m2
Ar Archimedes number, –CD drag coefficientdbo initial bubble diameter, mdbm maximum bubble diameter, mD diameter, mDb bubble diameter, mdp particle diameter, �mfs friction factorg acceleration of gravity, m/s2
Gs solid circulation rate, kg/(m2 s)h height, mH height, mkcyc cyclone coefficientkcon friction coefficient in the connection.K∞ elutriation rate constantl length, mm mass flow rate, kg/sms solid mass flow rate, kg/sm mass load, kgMT total inventory, total particle mass load, kgNt no. of orifices in the distributorP pressure, PaRe Reynolds number, –U velocity, m/sUo superficial gas velocity, m/sUs solids velocity, m/sV volume flow rate, m3/hrw opening width to the cyclone from riser, m�Px pressure drop through component x, Pa�t time, s�z height of accumulated material, m
Greek letterε voidage, –εs concentration of solidε∗s solid density above transport disengaging height
(TDH) or saturation carrying capacity of gasıb bubble fraction� density, kg/m3
� viscosity, Pa s sphericity, – dimensionless visible bubble flow, –d aerodynamic factor friction factor
Subscriptsb bubble/bulk (for density)
804 chemical engineering research and design 1 0 9 ( 2 0 1 6 ) 791–805
br bubble riser velocitycon connectioncyc cyclonedz dense zonee emulsionez exit zonefr frictiong gasgf gas frictionhs hydrostatic solidsls loop-seallc lower connectionLv L-valvemf minimum fluidizationp particlet terminal (applied to velocity)tf throughflowtz transport zones solidssf solid frictionsl slipso solid outflowsp standpiper riserre riser exitred riser exit ductrp reduce pressure drop
Acknowledgement
This research is financially supported by University ofMalaya, Ministry of Higher Education High Impact Research(UM.C/HIR/MOHE/ENG/30).
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