Hydrogen production of solar-driven steam gasification of sewage sludge in an indirectly irradiated fluidized-bed reactor

Applied Energy 261 (2020) 114229

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Hydrogen production of solar-driven steam gasification of sewage sludge in an indirectly irradiated fluidized-bed reactor

T

Xian Lia, Ye Shena, Liping Weia, Chao Heb, Alexei A. Lapkinc,d, Wojciech Lipińskie, Yanjun Daif, ⁎ Chi-Hwa Wangg, a

NUS Environmental Research Institute, National University of Singapore, Singapore 138602, Singapore School of Chemical and Biomedical Engineering, Nanyang Technological University, 62 Nanyang Drive, Singapore 637459, Singapore c Cambridge Centre for Advanced Research and Education in Singapore Ltd., 1 Create Way, CREATE Tower #05-05, Singapore 138602, Singapore d Department of Chemical Engineering and Biotechnology, University of Cambridge, Philippa Fawcett Drive, Cambridge CB3 0AS, United Kingdom e Research School of Electrical, Energy and Materials Engineering, The Australian National University, Canberra, ACT 2601, Australia f School of Mechanical Engineering, Shanghai Jiao Tong University, Shanghai 200240, China g Department of Chemical and Biomolecular Engineering, National University of Singapore, Singapore 117585, Singapore b

HIGHLIGHTS

indirectly irradiative fluidized-bed solar steam gasifier was proposed. • An two-phase flow model coupled with the solar cavity model was developed. • AA transient pyrolysis kinetic model with the particle temperature and size was developed. • Solar • upgraded ratio and solar-to-fuel efficiency of sewage sludge were analyzed. ARTICLE INFO

ABSTRACT

Keywords: Steam gasification Fluidized bed Sewage sludge Concentrated solar radiation Solar-to-fuel efficiency Solar upgraded ratio

A fluidized-bed based solar steam gasification of sewage sludge for production of high-quality syngas with a high content of hydrogen is numerically modeled and validated by experimental data generated from a lab-scale fluidized bed. The solar gasifier is mainly composed of a fluidized bed and a concentrically tubular cavity. A transient model coupling a two-phase fluidization model (in terms of reaction kinetics and hydrodynamics) and a solar cavity receiver model is established to conduct the parametric investigation of the proposed solar gasifier, including the effects of the direct normal irradiance, gasifying agent composition, and spatial flux distribution at the freeboard wall on the performance criteria of solar gasification i.e. solar upgraded ratio and solar-to-fuel efficiency. The transient simulation of the solar gasifier with ~2.2 MW solar power input is performed. A H2 yield range of 61.2–67.6 g/kg(sludge) can be achieved through solar steam gasification of sewage sludge, which can be adjusted by modifying the steam content of the gasifying agent and the direct normal irradiance. Under the condition of the direct normal irradiation of 1000 W/m2, the mean concentration ratio of 1000 suns at the dense bed, and 100 vol% N2 content, a maximum solar upgraded ratio of 1.0 and solar-to-fuel efficiency of 0.26 can be achieved.

1. Introduction Hydrogen is potentially a key energy vector in the future clean energy systems [1]. It is conventionally produced from natural resources such as coal, oil, natural gas and water via gasification and steam reforming. Today, around 96% of hydrogen is produced from fossil fuels, and this is accompanied by a significant emission of CO2

into the atmosphere. Considering the urgent need to reduce the anthropogenic impact on climate, alternative routes for generating hydrogen from renewable sources are attracting significant attention. Steam gasification of biomass is an alternative thermochemical method of hydrogen generation [2]. Compared to fast pyrolysis, a higher yield of hydrogen can be produced via steam gasification due to the steam reforming reaction with char and methane. The fundamental,

⁎ Corresponding author at: Department of Chemical and Biomolecular Engineering, National University of Singapore, 4 Engineering Drive 4, Singapore 117576, Singapore. E-mail address: [email protected] (C.-H. Wang).

https://doi.org/10.1016/j.apenergy.2019.114229 Received 20 August 2019; Received in revised form 7 November 2019; Accepted 22 November 2019 0306-2619/ © 2019 Elsevier Ltd. All rights reserved.

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Nomenclature A Ar C Cb Ce cp db ds DNI Ea F g h Hs Hmf HHV k Kbe LHV

Mi m Nu Pr Q r R Ra Re

T U V

surface area, m2 Archimedes number concentration ratio, suns gas concentration in the bubble phase, kmol/m3 gas concentration in the emulsion phase, kmol/m3 specific heat capacity, J/(kg K) bubble diameter, m solid particle diameter, m direct normal irradiance, W/m2 activation energy, kJ/mol mass fraction of the element gravity acceleration, m/s2 heat transfer coefficient, W/(m2 K) static bed height, m bed height at the minimum fluidization velocity, m higher heating value, MJ/kg for feedstock or MJ/m3 for producer gas conductivity, W/(m K) the mass transfer coefficient between the bubble and emulsion phase in the dense zone, 1/s lower heating value, MJ/kg for feedstock or MJ/m3 for producer gas molar mass of ith species feeding rate, kg/(m3 s) Nusselt number Prandtl number energy flux, W reaction rate, 1/s universal gas constant, 8.314 kJ/(kmol K) Rayleigh number Reynolds number

i, jg g

x s, l

temperature, K solar upgraded ratio volume flow rate, m3/s stoichiometric coefficient of ith species in reaction j mass fraction of species l

Subscripts b be conv cond e g i in mf out s rad solar steam w

bubble phase bubble to emulsion convection conduction emulsion phase gas species i feeding minimum fluidization velocity releasing from the bed solid phase radiation solar radiation steam agent into the gasifier wall

Greek symbols

HR specific enthalpy change, MJ/kg ηsolar-tofuel solar-to-fuel efficiency µ dynamic gas viscosity, Pa s voidage density, kg/m3

experimental and simulated studies of hydrogen generation from biomass (e.g. organic solid waste) by using steam gasification have been widely reported, covering the externally-heated lab-scale and smallscale fluidized beds [3,4]. Table 1 summarizes the conventional steam gasification of biomass. An alternative source of hydrogen, which is being also actively developed is ‘solar hydrogen’, which mainly consists of solar electric water splitting, solar photon water splitting, and solar thermochemical splitting water. However, these methods are still characterized by low efficiency. For a short-term path to efficient hydrogen generation, solar thermochemical steam gasification of biomass, forming a nexus of solar and biomass is a potential alternative technology which enables to overcome the major drawback of steam gasification – low reactivity and high energy intensity [2].

however, a transparent window is compulsory to fulfill the transmission of solar radiation, which results in a significant materials challenge, especially for high-pressure conditions. High pressures require special window configurations, e.g., hemispherical configuration. Contamination caused by the tar and carbonaceous particles decomposed from biomass is another issue in the directly irradiated reactors. The development of solar steam or/and CO2 gasification during the time period of 1980–2013 has been reported by Puig-Arnavat et al. [11]. In this work, we only reviewed the most recent progress in the representative reactors for solar steam gasification with carbonaceous feedstock as shown in Table 2. Compared to the packed-bed reactors in the directly irradiated scheme, the fluidized-bed reactors have the advantage of superior heat and mass transfer. This leads to a reduction in the formation of hot spots caused by the non-uniform concentrated solar radiation. The directly irradiated fluidized-bed gasifiers have been widely studied since 1983. Most recently, a beam-down secondary-reflection based CO2-gasification with coal coke was experimentally studied by Kodama et al. [16], via a high-flux visible spectrum solar simulator with a 6 kW Xe-arc lamp. In this work, a peak conversion efficiency of 14% was obtained at the particle size of 500–710 μm. The significant issue discussed in that work was the large surface-and-bottom temperature gradient occurring in the bubbling fluidized bed, which led to a low carbon conversion. To solve this issue, Gokon et al. proposed a conceptual design – internally circulating fluidized bed for CO2 gasification of coal coke [17,18]. This research was conducted in a lab-scale prototype via a high-flux solar simulator with a 3 kW solar power output. Similar to other directly irradiated reactors, a quartz glazing window was covered at the top of the reactor to reduce heat losses. The reacting particles of coal coke were exposed in the radiation flux. Due to an improvement of heat and

1.1. Literature review Published studies related to solar steam gasification are mainly focusing on the lab-scale [9] and pilot-scale [10] studies. Solar cavities are most favorable for serving as solar gasification reactors for hightemperature reactions – the concentrated solar flux is internally reflected and absorbed inside a cavity to reach the high-temperature target. Solar gasifiers can be classified according to two categories: (1) directly irradiated reactors – biomass particles are directly exposed to high-flux solar radiation, and (2) indirectly irradiated reactors – which are separated into two cavities i.e. absorber cavity and reaction cavity. High-temperature thermal energy from the absorber cavity is transferred to the biomass particles in the reaction cavity by heat transfer (e.g. conduction, radiation, and convention). Directly irradiated reactors have the advantages of enhancing heat and mass transfer, 2

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Experiment

Experiment Simulation Experiment

1200

1173 1200–1500 1114–1123

[6] [7] [8]

Research type

[5]

mass transfer inside the bed, a peak conversion efficiency of ~12% was observed at 0.9 kW power input, associated with a 73% carbon conversion. More recently, Gokon et al. conducted the experiment of steam gasification of coal coke with the similar devices of the reactor and solar simulator, where a carbon conversion of 60% via steam gasification was two times higher than the conventional fluidized bed reactor [19]. For the indirectly irradiated fluidized-bed reactors, limited reports were published. Muroyama et al. presented an original prototype of the solar/autothermal hybrid gasifier, which was related to using O2/steam to gasifying activated charcoal and lignite coal based on a 1.5 kWth fluidized bed reactor [15]. Their findings indicated that a maximum carbon conversion of 79% and a conversion efficiency of 22.1% were achieved. Adoption of such a hybrid concept may allow to avoid the disadvantages of intermitted solar energy, and to upgrade the overall efficiency by controlling the feeding rates and compositions of the reaction agents.

5.62 3–5 0.6–0.91 N.A. 58–60 30–36 Wastewater sludge Char Wood pellets Lab-scale gasifier Updraft fixed bed Fluidized bed

2–6 35–55

60.4 (exclusive of energy consumption of generating steam) 35 (inclusive of energy consumption of generating steam) N.A. N.A. N.A. Wood chips Updraft fixed bed

H2 fraction [vol.%] Feedstock

Cold gas efficiency [%]

1.2. Motivations and objectives

Gasifier type

Table 1 Summary of the conventional steam gasification.

Steam-to-carbon ratio [–]

Reaction temperature [K]

Source

X. Li, et al.

Analyzing the available literature, we found that there exists a research gap in the field of the indirectly irradiated solar gasifiers with the concentrically tubular cavity structure, which we classified into two aspects: (1) The particle-flow/entrained-flow patterns integrated with the concentrically tubular or multi-tubular solar cavities were mainly focused by previous studies. However, limited research on the fluidized-bed pattern in terms of steam gasification of biomass has been touched, especially for the effect of spatial distribution of solar flux at the bed wall on the gasification performance. (2) Three-dimensional computational fluid dynamics (3D-CFD) numerical simulation models of the fluidized-bed [20] or particle-flow [21,22] gasifiers were applied to accurately design and optimize a solar gasifier but they consumed a lot of computation time thus being not suitable for monthly or yearly performance assessment. So far, one-dimensional two-phase flow models [23,24] of fluidized-bed gasifiers and solar cavity-receiver models [9,25] have been individually proposed and they are a favorable solution to monthly/yearly performance assessment of solar steam gasification of biomass. However, a specific model that combines the solar cavity-receiver model and one-dimensional two-phase fluidization model has not been well developed for the solar gasifier with the concentrically tubular cavity-receiver and the fluidized bed. Motivated by the above research gap, we would develop a program that combines the experimental and simulated studies to parametrically analyze various factors, such as the H2O content of the gasifying agent, direct normal irradiance (DNI), and spatial distribution of solar flux, on gasification efficiencies (i.e. solar upgraded ratio and solar-to-fuel efficiency). The research findings and outcomes are able to give a reference or guideline for the design and optimization of such fluidizedbed based solar gasifiers. Firstly, a concentrically tubular fluidized-bed gasifier would be introduced with detailed discussion of the configuration as well as the optical and thermo-physical properties. Secondly, a hybrid model of combining a transient two-phase fluidization model and a pseudo-one-dimensional model of the solar cavity-receiver would be developed and validated by the experimental data of the pyrolysis process under a lab-scale fluidized-bed reactor. The kinetics model of the pyrolysis process would be derived based on Arrhenius expression coupled with the reaction temperature and the particle size. Finally, the effects of the direct normal irradiance (solar power input), gasifying agent composition and spatial distribution of solar flux at the absorber surface on the gasification performance would be conducted.

3

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Table 2 State-of-the-art development on the major reactors for solar steam gasification with carbonaceous feedstock. Absorber type

Reactor configuration

Gasifying agent

Feedstock

Solar-to-fuel efficiency (%)

Source

Indirectly Directly Spouted bed Indirectly Indirectly

Drop-tube/trickle-bed Packed bed

Steam Argon/steam Woody biomass Steam O2/steam

Brazilian sugarcane bagasse Micro-algae

21 N.A.

[12] [13]

Crop residues Mixture of activated charcoal and lignite coal

17.9 22.1

[14] [15]

Packed bed Fluidized bed

2. Description of a fluidized-bed solar gasifier

Table 3 Optical, mechanical, and thermo-physical parameters of the solar gasifier proposed in this work.

A fluidized-bed solar gasifier is schematically shown in Fig. 1. It consists of a cylindrical cavity receiver integrated with a fluidized bed that serves as a solar absorber to harvest concentrated solar radiation and to convert it to high-temperature process heat for solar steam gasification of sewage sludge. The bed wall was made of the silicon carbide material and has sufficient durability at high temperature and high chemical inertness, while the Al2O3 ceramic (serving as the cavity wall) has high durability and high reflectivity that is able to maximize the absorption by multi-reflection. The cavity was lined with Al2O3 insulation to minimize the heat loss between the cavity wall and the environment, and was integrated with a quartz window to minimize the convective and radiative heat losses from the absorber wall to the environment. Table 3 lists the major specifications of the proposed solar gasifier including optical, mechanical, and thermo-physical properties. A thermochemical pathway of the steam gasification is illustrated in Fig. 2. It is composed of two mechanisms: (1) the complex optical and thermo-physical behavior within the solar cavity receiver, and (2) the gasification mechanism in the fluidized bed. Since it is an indirectly irradiated reactor, the linkage between the two mechanisms is the bed wall (i.e. absorber wall). Specifically, partially concentrated solar radiation penetrating the quartz window (i.e. aperture) is absorbed by the absorber wall (the bed wall), which is associated with re-radiation loss due to the optical property of the SiC material. The high solar flux produces high-temperature heat that is transferred to the fluidized bed via convection and radiation. The radiation heat exchange is present among the absorber wall, cavity wall and quartz window. The natural convection heat transfer phenomena is observed among the absorber wall, cavity gas (air) and quartz window. The process heat drives the devolatilization of sewage sludge particles to generate the producer gas (mainly composed of CO and H2) that undergoes a multitude of

Specification

Unit

Value

Width of the quartz window Height of the quartz window Thickness of the quartz window Inner diameter of the fluidized bed Thickness of the fluidized bed wall Height of the dense bed Height of the freeboard Inner diameter of the cavity Cavity wall thickness Insulation thickness Emissivity of the quartz window Transmittance of the quartz window Emissivity of the SiC wall Absorptivity of the SiC wall Conductivity of the SiC wall (at 300 K) Conductivity of the quartz window Conductivity of the Al2O3 insulation (at 1200 K) SiC density Specific heat capacity of the SiC wall

m m m m m m m m m m – – – – W/(m K) W/(m K) W/(m K) kg/m3 J/(kg K)

0.5 2 0.005 0.5 0.01 1 1 0.8 0.01 0.25 0.85 0.85 0.93 0.9 150 2 0.24 3100 800

thermochemical reactions e.g. homogeneous reactions in the gas phase (see Table A3) and heterogeneous reactions (see Table A4) between the gas and the particles. The majority of particle mass is released to the producer gas, while the by-product – biochar – is consumed by steam gasification. Since benzene (C6H6) accounted for 60–80% in the tar [26], it was assumed as the sole tar compound in the following models.

Fig. 1. Schematics of a fluidized-bed solar gasifier with a concentrically tubular cavity configuration. 4

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Fig. 2. A thermochemical pathway of the solar-driven steam gasification process in a fluidized bed.

3. Materials and modelling

Qi,b = Qsolar,i,b + Qin,i,b

3.1. Solar cavity-receiver model

where Qi,b denotes the net radiative flux at ith surface under the spectral band b, Qout,i,b is the spectral radiosity emitted from ith surface to other surfaces participating radiation heat transfer, Qtrans,i,b is the spectral radiative flux emitted from ith surface to the environment. Qsolar,i,b represents the solar spectral flux to ith surface. Qin,i,b is the multi-reflection flux of incident solar radiation to ith surface.

3.1.1. Radiation model Gaussian distribution is conventionally adopted as the spatial distribution of solar flux at the aperture surface of the cavity. However, it would be completely different when considering the central tower system with the individual control for each heliostat, since the flux distribution can be adjusted and optimized to achieve a nearly uniform distribution. In this work, different non-uniform distribution patterns would be considered in order to explore their impacts on the gasification performance whereas the uniform solar flux distribution at an individual element/cell of the absorber surface (i.e. the outer surface of the bed wall) was assumed due to the proposed one-dimensional model of the fluidized bed. The radiation heat transfer inside the cavity was calculated as a face-to-face radiation model. The participating surfaces inside the cavity were treated as grey and diffuse due to the surface properties of SiC and Al2O3. The view factors were calculated using the Monte Carlo ray-tracing method [15]. The primary formulas are given as

Qi, b b=0

Qtrans,i,b

(2)

3.1.2. Natural convection model Empirical correlations were used to quantify natural convection inside the cavity. The natural convection heat flux Qconv was calculated by

Qconv = hconv Aw [Tw (z )

Tg]

(3)

with

hconv =

Nuconv kg L

(4)

where hconv represents the heat transfer coefficient between the surfaces inside the cavity and the cavity gas, Aw is the surface area, Tw (z ) is the spatial temperature of the surface at the axial direction, Tg is the cavity gas temperature, kg denotes the conductivity of the cavity gas, and L is the characteristic length. The natural convective Nusselt number [27] between the surfaces inside the cavity and cavity gas was calculated by

N

Qi,rad =

Qout,i,b

(1)

with 5

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without reactions.

2

Nuconv = 0.825 +

• The shrinking unreacted particle model (SUPM) [24] was adopted in the gas-solid conversion of a single char particle. • The cavity gas was assumed as a non-participant in radiation.

1/6 0.387Raconv

1+

( ) 0.492 Prg

9/16 8/27

(5)

3.2.1. Mass and energy conservation equations in the dense bed Mass conservation of species i in the bubble phase of the dense bed is expressed as

where Raconv is the Rayleigh number of natural convection, and Prg represents the Prandtl number of the cavity gas. 3.2. Fluidized-bed gasifier model

(

An unsteady-state two-phase kinetic model was developed to simulate the steam gasification in the fluidized-bed reactor that consists of complicated hydrodynamics, heat and mass transfer, and thermochemical reactions. As shown in Fig. 3, in the two-phase flow pattern, the fluidized bed is conventionally separated into two phases i.e. the bubble phase and the emulsion phase. The gasifier is divided into two main regions – the dense bed and the freeboard. The major assumptions used in the fluidized-bed gasifier model are summarized as the following items:

c b, i ) = t

(

b

ub c b, i ) z

b

Kbe (c b, i

ce, i ) +

b jg g

i, jg g

r b, jg

g

(6) where b denotes the volume fraction of bubbles in the dense bed, c b, i and ce, i respectively refer to the concentration of species i in the bubbles, ub represents the rising velocity of bubbles, Kbe is the mass transfer coefficient between the bubble and emulsion phase in the dense zone, i, jg g is the stoichiometric coefficient of ith species in reaction j, and r b, jg g represents the reaction rate of jth reaction in the bubble phase. Mass conservation of species i in the emulsion phase of the dense bed is given as

• The bubble model is assumed as the spherical shape and free of solid particles. • The freeboard is assumed as a gas–gas reactor. The elutriation char • • •

b

(

and ash are the non-participating media reacting with the producer gas. The uniform temperature of each element of the reactor is assumed due to the one-dimensional model. Non-penetrative and isothermal solid particles are assumed. Char contains pure carbon. Ash is assumed as an inert material

e

=

ce, i ) t (

jg s

e

ue ce, i ) + z

e

Kbe (c b, i

ce, i ) +

e jg g

i, jg g re, jg g

s

i, jg s re, jg s

where ue is the gas velocity in the emulsion phase,

Fig. 3. Schematics of a two-phase flow model of the bubbling fluidized bed driven by concentrated solar radiation. 6

+

(7) e

and

s

are

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the cavity, Q w,out,rad is the radiative heat flux emitted from the bed wall, Q we,conv is the heat flux transferred from the bed wall to emulsion gas by convection, Q ws,conv is the heat flux transferred from the bed wall to the solid particles by convection, Q ww,cond,in is the heat flux transferred from the adjacent wall element by conduction, Q ww,cond,out is the heat flux transferred to the adjacent wall element, and Q wcg,conv is the heat flux transferred from the bed wall to the cavity gas.

respectively the gas volume fraction and solid volume fraction in the emulsion phase, re, jg g is the reaction rate of jth reaction in the emulsion phase, i, jg s is the stoichiometric coefficient of ith species in gas-solid reaction j in the emulsion phase, and re, jg s refers to the reaction rate of jth gas-solid reaction in the emulsion phase. Mass conservation of solid species l is calculated by

(

s x s, l )

s

t

= m in, l

mout, l +

s

Mi jg s

i, jg s re, jg s

3.2.2. Mass and energy conservation equations in the freeboard Mass conservation of species i in the freeboard is expressed as

(8)

where m in, l is the feeding rate of species l, mout, l denotes the outlet mass flow rate (attrition rate) of species l, Mi is the molar mass of ith species, x s, l is the mass fraction of species l, and s is the average density of the solid particles. Energy conservation for species i in the bubble phase:

t

i

(

=

b

cg, i t

ub

b

b

c b, i cp, i Tb i

hbe (Tb dz

Te )

b jg g

r b, jg

g

HR, jg

(

=

e

b

t

re, jg

g

HR, jg

Te )

s

s

g

jg s

hse (Te dz

re, jg

s

Ts )

HR, jg

e

+

h we (Tw dz

(

s

e

Mi jg s

re, jg

mout cp,s Ts + s

HR, jg

s

h se (Te

s

Ts) +

s

h ws (Tw

Ts )

s

(11)

t

Q ww,cond,out

Q w,out,rad

Q wcg,conv

Q we,conv

Tg )

dz

i

jg g

rf, jg

g

HR, jg

g

(m fw cp,fw Tfw )

Qfw

Qfw,out,rad

Qfw

g,conv

+ Qfw,cond,in (15)

cg,conv

(mcg cp,cg Tcg ) = Q wcg,conv + Q fw

cg,conv

Qcg

cw,conv

Qcg

quartz,conv

(16)

where mcg is the gas mass, cp,cg is the specific heat capacity of the cavity gas, Tcg is the temperature of the cavity gas, Qcg cw,conv is the heat flux transferred from the cavity gas to the cavity wall by natural convection, Qcg quartz,conv is the heat flux transferred from the cavity gas to the quartz window by natural convection. Energy conservation of the subdivision of the quartz:

[m w cp,w Tw (z )] = Q w,solar,rad + Q w,in,rad

g (Tfw

3.2.3. Energy conservation equations of the cavity receiver Energy conservation of the cavity gas:

where cp,s is the specific heat capacity of the particles, Ts is the solid temperature, m in is the total feeding rate for all solid species, mout is the total elutriation rate for all solid species, h ws is the heat transfer coefficient between the bed wall and solid particles, and Mi is the molar mass of the species i. Energy conservation of the subdivision of the bed wall:

t

h fw

where m fw is the bed wall mass in the freeboard, cp,fw denotes the specific heat capacity of the freeboard bed wall, Qfw,solar,rad is the heat flux contributed by incident solar radiation, Qfw,in,rad is the radiative heat flux from other surfaces inside the cavity, Qfw,out,rad is the radiative heat flux emitted from the bed wall, Qfw g,conv is the heat flux transferred from the bed wall to the freeboard gas by convection, Qfw,cond,in is the heat flux transferred from the adjacent wall element by conduction, Qfw,cond,out is the heat flux transferred to the adjacent wall element, and Qfw cg,conv is the heat flux transferred from the bed wall to the cavity gas.

(10)

s

cg, i cp, i Tg +

Qfw,cond,out

s cp,s Ts )

= m in cp,s Ts

ug

= Qfw,solar,rad + Qfw,in,rad

Te )

where hse is the heat transfer coefficient between the particles and emulsion gas, h we is the heat transfer coefficient between the bed wall and emulsion gas, Tw is the bed wall temperature, and HR, jg s is the reaction enthalpy of the gas-solid heterogeneous reactions. Energy conservation in the solid:

t

z

(14)

ce, i cp, i

hbe (Tb dz

(13)

where Tg is the gas temperature in the freeboard, h fw g is the heat transfer coefficient between the bed wall and the gas in the freeboard which was given as wall-gas Nusselt number in Table A2, and Tfw is the bed wall temperature in the freeboard. Energy conservation of the subdivision of freeboard wall:

i

Te +

jg g

ue

e

i, jg g rf, jg g

jg g

(cg, i cp, i Tg )

=

ce, i cp, i Te)

z

t

i

where hbe denotes the heat transfer coefficient between the bubble gas and emulsion gas, cp, i is the specific heat capacity of the species i, Tb and Te are the temperatures of the bubbles and emulsion gas, respectively, and HR, jg g is the reaction enthalpy of the gas-gas homogenous reactions. Energy conservation for species i of gas in the emulsion phase:

t

+

g

(9)

i

z

where cg, i is the concentration of the species i in the freeboard, ug is the velocity of the gas, and rf, jg g is the reaction rate of jth reaction in the freeboard. Energy conservation of the gas phase in the freeboard is expressed as

c b, i cp, i Tb )

z

(ug cg, i )

=

Q ws,conv + Q ww,cond,in

t

(12)

(mquartz cp,quartz Tquartz) = Qcg

where m w is the bed wall mass, cp,w denotes the specific heat capacity of the bed wall, Q w,solar,rad is the heat flux contributed by incident solar radiation, Q w,in,rad is the radiative heat flux from other surfaces inside

quartz,conv

Qquartz

+ Qquartz,solar,rad + Qquartz,rad,in

sky,rad

Qquartz

amb,conv

Qquartz,rad,out (17)

where mquartz is the mass of the quartz window, cp,quartz is the specific 7

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heat capacity of the quartz window, Tquartz is the temperature of the quartz window, Qcg quartz,conv is the heat flux transferred from the cavity gas to the quartz window by natural convection, Qquartz,solar,rad is the heat flux gain contributed by multi-reflection solar radiation inside the cavity and absorption of the incident solar radiation penetrating the window, Qquartz,rad,in is the heat flux transferred from other surfaces inside the cavity by radiation, Qquartz,rad,out is the radiative heat flux emitted from the quartz window, Qquartz amb,conv is the heat loss from the quartz window to the environment by convection, and Qquartz sky,rad is the heat loss from the quartz window to the sky by radiation. Energy conservation of the cavity wall:

t

capacity of the cavity wall, Tcw is the temperature of the cavity wall, Qcg cw,conv is the heat flux transferred from the cavity gas to the cavity wall by natural convection, Qcw,solar,rad is the heat flux gain contributed by multi-reflection solar radiation inside the cavity, Qcw,rad,in is the heat flux transferred from other surfaces inside the cavity by radiation, Qcw,rad,out is the radiative heat flux emitted from the cavity wall, and Qcw amb,cond is the heat loss from the quartz window to the environment by conduction. To conduct the performance assessment, the higher heating value HHVf of carbonaceous feedstock was calculated via the following correlation [28]:

HHVf = 34.91FC + 117.83FH

(mcw cp,cw Tcw ) = Qcg

cw,conv

+ Qcw,solar,rad + Qcw,rad,in

10.34FO

1.51FN + 10.05FS

2.11FA (19)

Qcw,rad,out

Qcw

where FC , FH , FO, FN, and FA are defined as the mass fractions of carbon (C), hydrogen (H), oxygen (O), nitrogen (N), sulfur (S), and ash (A), respectively. The lower heating value LHVf of feedstock related to HHVf was

amb,cond

(18) where mcw is the mass of the cavity wall, cp,cw is the specific heat

Fig. 4. Solution algorithm of the two-phase and pseudo-one-dimensional model of the high-temperature solar gasifier coupled the fluidized bed and high-flux solar receiver. 8

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3.2.5. Numerical solution As illustrated in Fig. 4, we combined the unsteady-state two-phase flow pattern of a fluidized gasifier with the pseudo-one-dimensional model of the solar cavity receiver. The time-scale analysis method [30] of the biomass fluidized bed gasification was embedded into the solution algorithm to judge the sequence of devolatilization and gasification processes and to assist in the design of the parametric configuration of a fluidized bed. The correlations of the time-scale analysis are given in Table A1. The numerical discretization of the two-phase flow model of fluidized-bed relied on the implicit volume finite method with a stepby-step method. The partial differential equations in each cell were solved by using the upwind method implemented in Matlab. The time step was 0.001 s while the space step is 0.01 m. The convergence criterion with a residual of 10−5 was used in the calculations.

Table 4 Ultimate and proximate analyses of dried sewage sludge. Feedstock

Sewage sludge

Ultimate analysis C (wt.%) H (wt.%) O (wt.%) N (wt.%) S (wt.%)

35.0 4.8 24.7 5.2 1.7

Proximate analysis Moisture (wt.%) Volatile (wt.%) Ash (wt.%) Fixed carbon (wt.%) HHVf (MJ/kg) LHVf (MJ/kg)

5.8 58.6 22.8 12.8 14.9 13.8

3.3. Performance criteria

calculated by

LHVf = HHVf

21.978FH

The solar-to-fuel efficiency is calculated by

(20)

solar-to-fuel

3.2.4. Devolatilization model Devolatilization is the most important step in the biomass gasification process. The kinetics and products of devolatilization significantly affect the producer gas composition and yield of gasifiers. For a specific type of biomass, the temperature and heating rate are two key parameters of pyrolysis. Kinetic measurements are conventionally carried out by thermogravimetric analysis (TGA) at various heating rates and temperatures. The first-order reaction rate, commonly applied in the biomass kinetics, is expressed as

dX = rsludge (1 dt

X)

m0

m s (T , t ) m0 mf

Vs LHVs m f LHVf + Qsolar + Qsteam

Qsolar = DNI·C

(23) (24)

where Qsolar denotes the concentrated solar power hitting at the aperture of the solar gasifier, Vs is the volume flow rate of the cold gas released from the freeboard, LHVs is the lower heating value of the cold gas, m f is the feeding rate, and Qsteam is the energy consumption of steam generation. Another performance criterion is solar upgrade factor defined as the ratio of the lower heating value of the cold gas to that of the feeding biomass, which is calculated by

(21)

U=

where rsludge denotes the reaction rate of the sludge pyrolysis which is related to the conversion X at time t, pre-exponential factor A, and activation energy Ea, particle temperature Ts, and particle size ds. Conversion X at time t of a biomass sample is calculated by

X=

=

Vs LHVs m f LHVf

(25)

4. Results and discussion 4.1. Experimental setup of the pyrolysis process

(22)

The raw sewage sludge particles were ground and sieved into 300–500 μm by using a variable-speed ball grinder and mesh screen, respectively. After that, standard cylindrical pellets with a dimension of 2 mm (diameter) and 2 mm (length) were prepared from the small particles. Ultimate and proximate analyses of feedstock are given in Table 4. The devolatilization kinetics of the sewage sludge pellets was experimentally studied using a lab-scale fluidized-bed gasifier as shown in

where ms (T , t ) is defined as the transient sample mass at time t and temperature T, m 0 and m f represent the initial sample mass and the final sample mass, respectively. Based on the Kissinger Method [29], the pre-exponential factor A and activation energy Ea were experimentally determined. The devolatilization kinetics (of the sewage sludge sample) given in Table A4 would be discussed in the following section.

Fig. 5. Schematics of the experimental setup of the fluidized-bed gasifier used to explore the devolatilization kinetics of sewage sludge pellets. 9

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Fig. 5. The rig comprises a tubular reactor inside an electrical furnace, a manual pellets feeder, a water-scrubbing gas-cleaning device used to purify producer gas, and mass spectrometer (MS) for quantifying realtime gas compositions. The fluidized bed is a stainless-steel tube with dimensions of 320 mm (bed height), 19.8 mm (inner diameter), and 22.2 mm (outer diameter). A K-type thermocouple was fixed at the location of ~5 cm above the bottom of the dense bed to monitor the reaction temperature inside the bed. A mass flow controller (Bronkhorst, EL-FLOW) controlled an Ar flow rate of 20 LN/min. The pyrolysis experiment of sewage sludge is discontinuous. The sewage sludge feedstock was fed at the top of the fluidized bed by the simplified manual pellets feeder that includes a manual valve and a hopper. In a typical experiment, the reactor was preheated and maintained the required temperatures such as 600, 700, and 800 °C, and then 1 g sample of the sewage sludge pellets was fed into the reactor for each experimental condition. Since it is a discontinuous reactor, there’s no surplus particle stored inside the hopper. Fig. 6 shows the transient molar flow rates of sludge pyrolysis, which were used to quantify each gas content. The reaction rate and the molar flow rate both increased with the reaction temperature increased. Arrhenius expression described as Eq. (21) represents the temperature dependency of the rate constant. Based on the method of the constants and minimization of root-of-meansquare-deviation, TGA data at three heating rates (10, 20 and 30 °C/ min) were fitted to Arrhenius expression. Thus, the final rate expression (formula R7 given in Table A4) for the sewage sludge particle was formed with the modified factors of the particle temperature and the particle size. Based on the elemental analysis of the solid residue from the sewage sludge gasification [31], the O content in the solid residue at 700 °C was less than 0.5 wt.%. In this work, we assumed that the solid residue was free of O and H. Thus, the content of H2O can be calculated via oxygen conservation. After that, the C6H6 content was obtained according to the element conversation of H. The gas compositions and mass losses are listed in Table 5, corresponding to different pyrolysis temperatures, in which the mass loss shown in Fig. 7 was quantified by TGA under an intermittent heating pattern with an isothermal process of 20 min and a heating rate of 20 °C/min.

reaction temperatures, being related to the time-scale analysis [30] that considers the heat transfer resistance of the feedstock particles and the heat and mass transfer between the solid particle and gasifying agent inside the fluidized bed. Due to the merits of the fluidized bed reactor, the fast heat and mass transfer inside the reactor led to the slight differences among such time. Table 7 shows the comparison of the mass fractions of the major gas components. A reasonable agreement in the mass fraction indicates that the developed model is acceptable to predict the pyrolysis process in the fluidized bed. This model and code were utilized for the parametric analysis on a scale-up fluidized bed integrated with a concentric solar cavity receiver in the following sections. To minimize the difference in operation conditions between the scale-up and lab-scale fluidized beds, the same size and property of the sewage sludge pellets were utilized. 4.2.2. Validation of the fluidized-bed based solar cavity-receiver model Since it is difficult to find an identical configuration of the fluidizedbed based solar cavity-receiver with the indirectly irradiated and quartz-window structure, the experimental data from a windowless fluidized-bed solar cavity gasifier were utilized to validate the accuracy of the fluidized-bed based solar cavity-receiver model. The optical, geometric and thermos-physical properties of the fluidized bed and solar cavity receiver as well as operation parameters in the validation were set to be identical with the data reported by Muroyama et al. [15]. The radiation model was acceptable for both the windowless and window configurations, however, the natural convection model for the heat loss from the solar cavity to the environment was totally different. Thus, the natural convection model proposed by Paitoonsurikarn et al. [32] for the windowless solar cavity was applied in this validation. Due to a lack of the pyrolysis kinetics of the charcoal particles in the fluidized-bed reactor, we only focused on the preheated stages with granular Al2O3 particles. Fig. 9a shows the dynamic data of two key temperatures inside the fluidized bed, in which Tbed,1 was located at the top surface of the inert bed and Tbed,2 monitored the local temperature at ~30 mm above the focal point. The radiative solar flux map [15] at the outer surface of the fluidized bed was set as the flux boundary input for two operation stages. At the first stage (i.e. the preheated stage), the center xenonlamp delivered ~656 W effective radiation power hitting at the aperture with a 20 mm diameter. An overestimation in the temperatures was observed during the time period of 0–4 min whereas an underestimation presented from 4 min to 10 min. It is mainly attributed to a deviation in the spatial flux distribution between the experiment and simulation. At the second stage (10–20 min), the top xenon-lamp was added to increase the effective power input up to ~1184 W, which led to a sharp increase in the temperatures. An accurate prediction in Tbed,2 and an obvious underestimation in Tbed,1 were found, which is mainly due to the deviation in the spatial radiation flux as well. A significant increase in Tbed,1,exp was noted due to a higher local flux contributed by the top xenon-lamp. Compared the experimental data to the simulated results (see Fig. 9b), a relative deviation span of −18 to 12% was determined for the fluidized-bed based solar cavity-receiver model, which

4.2. Model validation 4.2.1. Validation of the fluidized-bed gasification model Due to a lack of reported experimental data using the identical sewage sludge in a fluidized-bed reactor, we utilized the experimental data from the lab-scale fluidized bed to validate the accuracy of the established two-phase fluidized-bed model. A specific code suitable for the lab-scale fluidized bed was developed based on the specifications listed in Table 6. The comparison results of the total molar flow rate were shown in Fig. 8. It indicates that the developed model accurately predicted the pyrolysis kinetics of sewage sludge in the lab-scale fluidized bed. The achieved time of the peak molar flow rates at different reaction temperatures was slightly different through experiment, such as ~30 s, ~35 s, and ~50 s for 600, 700, and 800 °C. It is mainly affected by the different heat and mass heat transfer at different

Fig. 6. Molar flow rate evolutions of the producer gas: (a) 600 °C, (b) 700 °C, and (c) 800 °C. 10

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Table 5 Gas compositions and mass losses as a function of the pyrolysis temperature. Pyrolysis temperature [°C]

CO2 [wt. %]

CO [wt. %]

H2 [wt. %]

CH4 [wt. %]

H2O [wt. %]

C6H6 [wt.%]

Solid residue [wt. %]

Mass loss of the sample [wt.%]

Sample mass [g]

600 700 800

4.0 7.3 12.8

8.9 15.2 24.4

0.6 1.1 1.9

2.5 4.0 6.4

24.2 18.0 7.3

10.6 5.9 1.4

48.8 46.4 43.8

51.2 53.6 56.2

1.0 1.0 1.0

led us to conclude that the developed model can be used for the performance assessment on such a fluidized-bed solar gasifier. 4.3. Transient behaviors of the solar gasifier To parametrically analyze the gasification performance of the solar gasifier, considering the optical, mechanical, and thermos-physical properties given in Table 3, the input parameters of the dense bed and sewage sludge are shown in Table 8, where the geometric parameters of the dense bed were determined by the time-scale method listed in Table A1. Typically transient behavior of the solar gasifier with 100 vol% N2 is shown in Fig. 10. A fast heating rate of ~5 K/s was observed during the heat-up stage (0–200 s). Owing to a high heat transfer coefficient between the bed wall and sewage sludge particles, the particle temperature was higher than gas temperature inside the dense-bed region, which led to a fast pyrolysis process started from ~100 s as the molar flow rate of producer gas increased. Note that a significant decrease in the molar flow rate of CO was evident at ~230 s associated with the corresponding temperature decrease in the bed wall (see Tw,bed in Fig. 10a), which is due to the water-gas shift reaction in the gas phase. Fig. 11 shows the gas composition distribution along the height of the fluidized bed at the reaction time of 400 s. The gas concentrations of producer gas from the gasification of sewage sludge accumulated along the flow direction in the dense bed, accompanied by the variations of the bubble fraction and bubble velocity. Compared to the concentration increment in the dense bed, the gas concentrations of CO2 and H2 in the freeboard (~0.9–2 m) presented a distinct increase whereas CO, CH4, and tar concentrations decreased due to the reactions of methane steam reforming, tar steam reforming and water gas shift. The descent rate of the CO concentration was faster than that of the CH4 and tar concentrations, which is mainly attributed to the reaction rate of water-gas shift reaction (R2) is significantly faster than that of methane steam reforming (R1) and tar steam reforming (R3), under the reaction conditions at t = 400 s.

Fig. 7. TGA data with 20-mins steady state used to quantify the mass loss of the sewage sludge sample. Table 6 Input parameters for model validation. Specification

Value 3

Particle density (kg/m ) Particle diameter (mm) Particle shape Feeding rate of feedstock Particle diameter of the bed material [mm] Particle density of the bed material [kg/m3] Static bed height [m] Inner diameter of the fluidized bed [mm] Outer diameter of the fluidized bed [mm] Total height of the bed [mm] Gasifying agent Gas flow rate [LN/min] Bed material

1800 2.0 Ideal/standard cylinder 1 g for each sample 0.297 2650 0.05 (bed material) 19.8 22.2 320 Ar 20 Granular SiO2

4.4. Effects of the DNI and gasifying agent composition For a solar gasifier, the most critical factor that affects the gasification performance is the intermittent solar irradiance variation, thus the effects of the DNI and gasifying agent composition on performance criteria of solar steam gasification were studied and the results are

Fig. 8. Comparison of the experimental and simulated data at different pyrolysis temperatures: (a) 600 °C, (b) 700 °C, and (c) 800 °C. 11

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Table 7 Comparison of experimental data and simulated results in terms of total mass fractions under different pyrolysis temperatures. Pyrolysis temperature [°C]

600 700 800

Experiment Composition [wt.%]

Simulation

CO2

CO

H2

CH4

H2O

C6H6

CO2

CO

H2

CH4

H2O

C6H6

4.0 7.3 12.8

8.9 15.2 24.4

0.6 1.1 1.9

2.5 4.0 6.4

24.2 18.0 7.3

10.6 5.9 1.4

4.6 7.7 13.3

8.8 14.7 25.5

0.7 1.2 2.0

2.3 3.9 6.7

24.1 18.0 7.2

10.6 5.9 1.5

Fig. 9. Validation of the fluidized-bed based solar cavity-receiver model: (a) dynamic data of two key temperatures (Tbed,1 located at the top surface of the inert bed and Tbed,2 located at ~30 mm above the focal point at the absorber tube).

shown in Fig. 12. As depicted in Fig. 12a, the H2O content has a significant impact on the solar upgraded ratio U at various solar irradiances (500–1000 W/m2). A decrease in the yields of CO and CH4 may be associated with a large amount production of CO2. As expected, the solar upgraded ratio decreased significantly with the decreased DNI. A maximum solar upgraded ratio of ~1.0 can be obtained at DNI = 1000 W/m2 and 100 vol% N2. It is interesting to note that optimum values of the H2 yield were observed for different DNIs. 30, 30, 40, and 50 vol% H2O contents were the most favorable ratios for DNIs of 500, 700, 900, and 1000 W/m2, respectively. A higher solar irradiance boosted the H2 yield due to the higher reaction temperatures. Compared to Fig. 12d, the lower heating value of the cold producer gas decreased with the increased N2 content in Fig. 12c, whereas an opposite trend can be observed from the volumetric yield of the cold producer gas. This finding indicates that introducing steam enables to upgrade the lower heating value of the cold gas with a lower gas yield due to the steam condensation. A trade-off between the lower heating

Table 8 Input parameters of the sewage sludge particle and bed in the solar gasifier. Specification

Value 3

Particle density [kg/m ] Particle diameter [mm] Particle shape Static porosity of the dense bed Static dense bed height [m] Gasifying agent Superficial velocity [m/s] Minimum fluidization velocity [m/s] u0/umf DNI [W/m2] C (average concentration ratio at a cell of the bed wall) [suns]

1800 2.0 Ideal/standard cylinder 0.5 0.6 N2 and/or H2O 3.0 0.78 3.8 500, 700, 900, 1000 1000

Fig. 10. (a) Transient behaviors of the temperatures and (b) molar flow rates under the conditions (DNI = 1000 W/m2, C = 1000 suns for the dense bed, C = 0 for the freeboard, and 100 vol% N2). 12

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Fig. 13. The solar-to-fuel efficiency for various DNIs and H2O contents.

Fig. 11. Gas composition along the height of the fluidized bed.

decreased dramatically with the decrease in DNI which is attributed to the low pyrolysis rate and high amounts of water and tar generated in the low-temperature pyrolysis. Fig. 13 shows the solar-to-fuel efficiency decreased as the H2O content increased due to the same reason with the solar upgraded ratio. A maximum solar-to-fuel efficiency of 25.9% can be achieved at DNI = 1000 W/m2 and 100 vol% N2. As depicted in

value and the volumetric yield of the producer gas led to the variation of solar upgraded ratio with the H2O content (see Fig. 12a). Due to the low temperatures of the particle and gas in the dense bed affected by the low solar radiation, the lower heating value and gas yield both

Fig. 12. (a) solar upgraded ratio U, (b) H2 yield, (c) lower heating value LHVs, and (d) volumetric yield Vs of producer gas as a function of the H2O content and DNI. 13

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reaction heat required by endothermic reactions. To reach a reasonable trade-off between the solar upgraded ratio and the H2 yield, a mean concentration ratio of above 700 suns with DNI = 1000 W/m2 is recommended for solar steam gasification of sewage sludge in such a fluidized-bed reactor. A decrease in the solar-to-fuel efficiency was observed even if C > 700 suns, which is due to the fact that, although total energy of the producer gas was solar upgraded in the freeboard (C > 700 suns), an increase in the total energy consumption of steam generation and solar power input overcame the energy increment of the producer gas. It should be noted that the decrease in the solar-to-fuel efficiency was mainly affected by the low gas-wall heat transfer coefficient (in the freeboard region). 5. Conclusions A fluidized-bed based solar gasifier with the concentrically tubular configuration has been introduced for steam gasification of sewage sludge to produce high quality of syngas or hydrogen production. The kinetic model of the pyrolysis process of the sewage sludge has been developed by using a lab-scale fluidized-bed reactor, which integrates the particle temperature and size for more accurate prediction. Based on the kinetic model of the pyrolysis, a dynamic two-phase and pseudoone-dimensional model of the high-temperature solar gasifier coupled the fluidized bed and concentrically tubular solar cavity receiver has been established and validated with the experimental data from the labscale setup and the reference. The developed solar gasification model was used to investigate the effects of concentrated solar irradiance (direct normal irradiance), gasifying agent composition, and spatial distribution of solar flux on the gasification performance (i.e., solar upgraded ratio, solar-to-fuel efficiency, and H2 yield). The major findings of this work are summarized below:

Fig. 14. The gas composition as a function of the H2O content at DNI = 1000 W/m2.

(1) Increasing the H2O content of the gasifying agent is able to increase the lower heating value (from 1.54 MJ/m3 to 9.73 MJ/m3) of the cold gas but reduces the gas yield due to the steam condensation. With the H2O content increased from 0 vol% to 100 vol%, the lower heating value of the cold gas increased from 1.54 MJ/m3 to 9.73 MJ/m3 under the direct solar irradiance of 1000 W/m2. However, the cold gas yield decreased from 294.7 m3 to 36.6 m3. (2) Optimum values of 30, 30, 40, and 50 vol% for the H2O content were observed for the direct normal irradiances of 500, 700, 900, and 1000 W/m2, respectively, under the mean concentration of 1000 suns at the dense-bed wall. A H2 yield range of 61.2–67.6 g/kg (sludge) was achieved by solar steam gasification of sewage sludge, which can be adjusted by modifying the H2O content and solar radiation. (3) Increasing direct normal irradiance always improved the solar upgraded ratio and solar-to-fuel efficiency. However, increasing the H2O content reduced both the solar upgraded ratio and solar-to-fuel efficiency. With the H2O content increased from 0 vol% to 100 vol %, the solar upgraded ratios decreased by 5.7%, 5.8%, 6.9%, and 8.4% at the direct normal irradiances of 500, 700, 900, and 1000 W/m2, respectively, while the solar-to-fuel efficiency decreased by 32.9%, 22.6%, 18.6%, and 18.5%. (4) Adding solar power input into the freeboard region further reduced the solar-to-fuel efficiency. Under the direct normal irradiances of 1000 W/m2, the solar-to-fuel efficiency decreased from 0.19 to 0.14 as the mean concentration ratio increased from 100 suns to 1000 suns. Considering both the solar upgraded ratio and the hydrogen yield, a mean concentration ratio of over 700 suns is recommended for the direct normal irradiances of 1000 W/m2. (5) Under the parameters (a direct normal irradiance of 1000 W/m2, a mean concentration ratio of 1000 suns at the dense bed, a mean concentration ratio of 0 sun at the freeboard wall, and 100 vol% N2 content), a maximum solar upgraded ratio of 1.0 and solar-to-fuel efficiency of 0.26 can be harvested.

Fig. 15. The solar upgraded ratio and solar-to-fuel efficiency as a function of the mean concentration ratio C at the freeboard wall (100 vol% H2O, DNI = 1000 W/m2, C = 1000 suns for the dense-bed wall).

Fig. 14, the H2 and CO2 contents significantly increased as the H2O content increased. 4.5. Effect of the spatial distribution of solar flux Another key parameter for the solar gasifier is the solar flux distribution at the absorber surface which has a significant impact on the fluidized bed along the height direction. Due to the fast heat and mass transfer as well as two-phase flow phenomena inside the dense bed, the hot spot of the absorber wall in the dense bed can be mitigated. Thus, in this work, we only focused on the spatial distribution of solar flux at the freeboard wall. The effect of the mean concentration ratio C on the performance indexes was analyzed under constant parameters (DNI = 1000 W/m2, C = 1000 suns for the dense-bed wall, and 100 vol % H2O). Fig. 15 shows that the solar upgraded ratio decreased within a concentration ratio range of 100–700 suns, whereas it increased when the concentration ratio was over 700 suns. It indicates that high-temperature heat at the freeboard wall contributed by concentrated solar irradiation (C > 700 suns) enabled to provide a critical amount of 14

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For future work, we will aim to utilize the developed model to assess the yearly gasification performance based on the different local weather conditions. In addition, this model will be extended to cover the solar/ autothermal hybrid gasification process.

Acknowledgement This research is supported by the National Research Foundation (NRF), Prime Minister’s Office, Singapore, under its Campus for Research Excellence and Technological Enterprise (CREATE) programme (Grant Number R-706-001-102-281).

Declaration of Competing Interest The authors declared that there is no conflict of interest. Appendix A See Tables A1–A4.

Table A1 Semi-empirical correlations used for hydrodynamics of a fluidized-bed gasifier. Parameter

Correlation

Minimum fluidization velocity, umf (m s−1) Bubble velocity, ub (m s−1) Bubble diameter, db (m)

Source

µg

umf =

ub = u 0 db = dbm

The porosity at minimum fluidization, mf Solid porosity, s Axial mixing time of particles, taxial (s) Lateral mixing time of particles, tlateral (s)

b

=

mf

s

dbo ) e

2 dR (u 0 4

u mf )

g

[37] [33]

( ) 4

b )(1

[38] [39]

mf )

H bed 0.6(u0 umf )

tlateral =

[40,41]

2 d bed (2DL )

umf ) Hmf (dbed/ Hmf )0.5Fr

DL = 0.013(u 0

Hs =

[35,36]

0.3z dR 0.4

umf )2

3.69(u0

Fr = Static bed height, Hs (m) Bed height at minimum fluidization, Hmf (m)

[34]

umf ub 1 1/3

= (1

taxial =

g )g

µg2

(dbm

2.59 g 0.2

u0

=

[33]

33.7]

u mf + 0.711 gdb

dbo = b

ds3 g ( s

Ar =

dbm =

Bubble porosity,

[ 1135.7 + 0.0408Ar

g ds

0.15

(u0 umf )2 (gHmf )

[30]

ms s AR (1

m)

[30]

ms

Hmf =

s AR (1

mf )

Table A2 Semi-empirical correlations of heat and mass transfer. Parameter

Correlation

Gas-solid Nusselt number, Nug

Nugp = 2 + 0.6

Wall-solid Nusselt number

Source

Nu ws = 0.525Res

Remf 0.5 mf 0.75

Wall-gas Nusselt number

Nu wg = 0.525Reg 0.75

Solid-solid Nusselt number

Nus =

Reynolds number Prandtl number

Remf = Pr =

hs d s kg

= 0.85Ar 0.19 + 0.006Ar 0.5Pr 1/3

g umf d s µg

µg cs,g kg

15

1

Pr 3

[42] [43] [43] [24] [30] [30]

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Table A3 Kinetics of the major homogeneous reactions in biomass gasification. Reactions

Stoichiometric Chemical Equations

R1: Methane steam reforming

CH4 + H2 O

R2: Water gas shift

H2 O + CO

R3: Tar steam reforming

C6 H 6 + 6H2 O

Rate equation [C in kmol/m3, T in K, R in J/(mol K)]

3H2 + CO

1.25 × 105 RT

r 1 = 3 × 10 8exp

H2 + CO2

r2 = 2.78 × 103exp

9H2 + 6CO

Source [44]

CCH4 CH2 O

1.256 × 10 4 RT

CCO CH2O

[45]

CCO2 CH2 /exp(3958/T ) 0.0265

[46]

0.4 0.2 r3 = 2.0 × 1016exp[ 4.43 × 105/(RT )] CC1.3 6 H6 CH2 CH2 O

Table A4 Kinetics of the major heterogeneous reactions in biomass gasification. Reactions

Stoichiometric Chemical Equations

Rate equation [C in kmol/m3, T in K, R in J/(mol K)]

Source

R4: Boudouard Reaction R5: Char steam Reaction R6: Methane Reaction R7: Pyrolysis

C + CO2

2CO

0.6 r4 = 4.4T exp[ 1.62 × 108/(RT )] CCO 2

[47]

C + H2 O

H2 + CO

r5 = 1.33T exp[ 1.47 × 108/(RT )] CH0.6 2O

[47]

C + 2H2

CH 4

r6 = 4.189 × 10 3exp[ 1.92 × 107/(RT )] PH22

[48]

Sludge

nCH4 CH4 + nCO2 CO2 + nCO CO + nH2 H2 + n H2O H2 O + nTar Tar + n char Char

1

rsludge = 2.1 × 109exp

1.30965 ×

105 (RT )

d s(2.25 × 10

3T 0.759)

1

nCH4 , nCO2 , nCO , nH2 , nH2 O , nTar , and nchar are the stoichiometric coefficients of species which can be calculated based on the mass fractions given in Table 5 at different pyrolysis temperatures.

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