autothermal hybrid gasification CCHP system with an indirectly radiative reactor

Accepted Manuscript Thermodynamic assessment of a solar/autothermal hybrid gasification CCHP system with an indirectly radiative reactor Xian Li, Ye Shen, Xiang Kan, Timothy Kurnia Hardiman, Yanjun Dai, Chi-Hwa Wang PII:

S0360-5442(17)31649-3

DOI:

10.1016/j.energy.2017.09.149

Reference:

EGY 11664

To appear in:

Energy

Received Date: 30 June 2017 Revised Date:

26 September 2017

Accepted Date: 26 September 2017

Please cite this article as: Li X, Shen Y, Kan X, Hardiman TK, Dai Y, Wang C-H, Thermodynamic assessment of a solar/autothermal hybrid gasification CCHP system with an indirectly radiative reactor, Energy (2017), doi: 10.1016/j.energy.2017.09.149. This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.

ACCEPTED MANUSCRIPT

Thermodynamic assessment of a solar/autothermal hybrid

2

gasification CCHP system with an indirectly radiative reactor

3

Xian Li a, Ye Shen b, Xiang Kan b, Timothy Kurnia Hardiman b,

4

Yanjun Dai c, Chi-Hwa Wang b,*

5

a

NUS Environmental Research Institute, National University of Singapore, Singapore

6

138602, Singapore b

SC

7

RI PT

1

Department of Chemical and Biomolecular Engineering, National University of

8 9

c

M AN U

Singapore, Singapore 117585, Singapore

School of Mechanical Engineering, Shanghai Jiao Tong University,

10

Shanghai 200240, China Abstract

12

The solar/autothermal hybrid gasifier (SAHG) is an attractive approach to provide

13

continuous production of the syngas via coupling autothermal and solar gasification

14

together, where the SAHG mainly includes fully solar, hybrid, and fully autothermal

15

modes. An ICE CCHP system driven by the SAHG with an indirectly irradiative two-cavity

16

reactor introduced conceptually and investigated thermodynamically. Considering the

17

effects of solar flux inputs and various reactant ratios, a zero-dimensional steady-state

18

model of the SAHG was established by using Gibbs free energy minimization, and was

19

validated

20

oxygen-to-feedstock ratios has been achieved based on the restrictions of temperature

21

over 1000 K and minimization of steam input. The results of two consecutive days

22

indicate mole flow rates of H2 and CO were increased by over 38.8% and 11.8%,

23

respectively, leading to an increment in LHVs by 51.7%. An increment in primary energy

24

ratio by 11.5% can be achieved by using the SAHG-CCHP system. The yearly assessment

AC C

EP

TE D

11

with

the

reported

data.

The

1

optimal

steam-to-feedstock

and

ACCEPTED MANUSCRIPT of the SAHG-CCHP system shows that the yearly average increments in heat, power, and

26

cooling for the SAHG system were reached by 19.5%, 23.8%, and 4.5%, respectively. A

27

yearly average increment of 14.2% in primary energy ratio can be obtained under the

28

solar radiation condition of Singapore.

29

Keywords: Concentrated solar energy; gasification; Steam/air; CCHP; Thermodynamics.

30

* Corresponding Author at: Department of Chemical and Biomolecular Engineering,

31

National University of Singapore, 4 Engineering Drive 4, Singapore 117576

32

E-mail Address: [email protected] (Chi-Hwa Wang)

33

Nomenclature

34

A

surface area, m2

35

C

mean concentration ratio

36

cp

specific heat capacity, J/(mol K)

37

F

mass fraction of the element in the feedstock, %

38

h

heat transfer coefficient, W/(m2 K)

39

H

enthalpy, J/(g K)

40

HHV

higher heating value, MJ/kg

41

IDNI

direct normal insolation, W/m2

42

LHV

43

M

44

m

oxygen-to-feedstock ratio

45

mሶ f

reaction rate, mol/s

46

n

moles of the species in the product syngas releasing from the gasifier,

47

mol

AC C

EP

TE D

M AN U

SC

RI PT

25

lower heating value, MJ/kg

steam-to-feedstock ratio

2

ACCEPTED MANUSCRIPT ܰ୬ୣ

nominal electrical power of a natural gas fired ICE, kW

49

ܰ௘

power output of internal combustion engine, kW

50

PER

primary energy ratio

51

q

energy flux, kW

52

Q

energy, MW h

53

T

temperature, K

54

U

solar upgraded ratio

55

ܸሶ

volume flow rate, m3/s

56

Subscripts

57

air

air

58

b

spectral band

59

c

cold fluid

60

cond

conductive heat transfer

61

col

coolant water

62

conv

convective heat transfer

63

chw

chilled water

64

CG

cold gas

65

e

electricity

66

eq

67

ext

68

f

feedstock

69

h

hot fluid

70

hw

hot water

71

HE

heat exchanger

AC C

EP

TE D

M AN U

SC

RI PT

48

equilibrium

exhaust gas/heat

3

ACCEPTED MANUSCRIPT in

inlet

73

loss

heat loss

74

N

natural gas

75

out

outlet

76

pow

electric power

77

rad

radiative heat transfer

78

s

syngas

79

solar

solar power

80

tran

radiative energy leaving from a surface to the environment

81

Greek symbols

82

∆‫ܪ‬ோ

specific enthalpy change, MJ/kg

83

η

efficiency

84

ηHE

effectiveness

85

Abbreviations

86

CCHP

combined cooling, heat, and power

87

CPC

compound parabolic concentrator

88

DEAC

double-effect absorption chiller

89

ICE

internal combustion engine

90

IGCC

91

SAHG

solar/autothermal hybrid gasifier

92

SAHG-CCHP

combined cooling, heat, and power (CCHP) system employed the SAHG

AC C

EP

TE D

M AN U

SC

RI PT

72

integrated gasification combined cycle

93

4

ACCEPTED MANUSCRIPT 94

1. Introduction Thermochemical gasification [1‒3] has been proved to be an attractive approach to

96

convert carbonaceous biomass/solid waste [4] to energy-rich and clean synthesis gas

97

(i.e. syngas, primarily composed of H2 and CO) for more effective utilizations, e.g.,

98

integrated gasification combined cycle (IGCC) power generation [5], conversion to

99

transportable liquid fuels by using catalytic process of Fischer-Tropsch synthesis [6].

100

Currently, researchers focus on autothermal [7,8] and allothermal [9] gasification, such

101

as co-gasification of biowaste and fossil fuels [10], and the hybrid waste-to-energy

102

system combined anaerobic digestion and gasification [11]. For conventional

103

autothermal gasification, a certain proportion of feedstock is required to be combusted

104

with pure oxygen serving as the driving process heat of the endothermic reaction. The

105

product syngas of autothermal gasification is inherently contaminated by the

106

by-products (e.g. SO2 and NOx) of internal combustion.

TE D

M AN U

SC

RI PT

95

Concentrated high-flux solar thermal energy [12], e.g., solar tower [13] and solar

108

beam-down [14,15] systems, has been proposed as a high quality thermal source to

109

supply process heat for allothermal gasification. The most frequent types of solar

110

gasification reactors [16] employ the cavity configuration to harvest the incident

111

high-flux solar radiation hitting onto the aperture, which are classified as: (1) directly

112

irradiative reactor [17–22] associated with reactants directly exposed to the high-flux

113

radiation, and (2) indirectly irradiative reactor [23–26] employed with an opaque wall

114

or emitter capturing the effective solar radiation and then transferring it to reactants by

115

convection and radiation. Directly irradiative configuration provides a high efficient

116

heat and mass transfer as a result of direct transmission from solar radiation to

117

feedstock, whereas it is suffering from the troublesome of glass transparent glazing at a

AC C

EP

107

5

ACCEPTED MANUSCRIPT high-pressure running and pollution by gasification by-products. In contrast, indirectly

119

irradiative reactor can overcome the above issues via scarifying the heat transfer

120

performance. The merits of solar steam/CO2 gasification are: (1) free of O2 requirement

121

and internal combustion of a certain proportion of feedstock leading to significant

122

reductions in contaminates and CO2, and (2) tar reduction due to higher reaction

123

temperatures contributed by concentrated solar fluxes. The most important aspect is

124

that the effective solar energy is chemically stored into the product syngas associated

125

with an upgraded calorific value.

SC

RI PT

118

When it comes to the growing interest in continuous production of syngas, solar

127

gasification is facing the inherent issue of intermittent solar radiation affected by cloud

128

and rain. Consequently, the concept of solar/autothermal hybrid gasifier (SAHG) [27,28]

129

was proposed to couple autothermal and solar gasification together fulfilling the

130

requirement of continuous running and applications. The operation modes of the SAHG

131

mainly include three modes: (1) fully solar mode, where the process heat is completely

132

derived from concentrated solar radiation, (2) hybrid mode, where low irradiance solar

133

energy combined with internal combustion of feedstock provide the essential heat for

134

endothermic reaction, and (3) fully autothermal mode without solar power input. The

135

relevant research has been reported in terms of many aspects. Muroyama et al. [29]

136

developed a simplified dynamic model to automatically control a fluidized steam SAHG

137

process. The typical and seasonal assessment was also conducted to reveal the

138

difference between autothermal and hybrid modes. Van Eyk et al. [30] proposed a

139

dynamic mathematical model of a hybrid entrained-flow reactor for coal gasification to

140

study the mechanism of high-flux solar radiation affecting the gasification process of

141

coal particles. They found that for most cases of incident solar irradiances, the overall

AC C

EP

TE D

M AN U

126

6

ACCEPTED MANUSCRIPT cold gas efficiency of the hybrid gasifier is less than that of autothermal due to

143

additional re-radiative heat loss. So far, such previous efforts have not addressed the

144

optimized oxygen-to-feedstock and steam-to-feedstock ratios for various feedstock

145

materials to maximize cold gas efficiency in the complex gasification process of SAHG. In

146

addition, gasification-based combined cooling, heat and power (CCHP) system with

147

carbonaceous feedstock is a promising tri-generation approach of converting product

148

syngas to cooling, heat, and power, which has been widely investigated [31–41].

149

However, the performance and operation features of the CCHP system driven by such

150

type of SAHG has not been clearly investigated and assessed.

M AN U

SC

RI PT

142

To address these needs, in this paper, an existing two-cavity solar steam reactor

152

with indirectly irradiative configuration was conceptually extended to be a SAHG via

153

including the autothermal process with the air reactant. We developed a simplified

154

zero-dimensional steady-state model that was validated with the reported data to

155

analyze the effects of the oxygen-to-feedstock and steam-to-feedstock ratios on

156

gasification performance, e.g., cold gas efficiency, lower heating value, and reaction

157

temperature, and to further achieve the optimal parameters at various solar fluxes. In

158

addition, the operation behaviors of the CCHP system driven by the proposed SAHG was

159

studied by considering the direct normal irradiances of two consecutive days of

160

Singapore. Finally, the potential assessment of the SAHG-CCHP system operating in

161

Singapore was conducted with respect to the performance index of primary energy

162

ratio.

163

2. System configuration

164

An internal combustion engine (ICE) CCHP system, driven by a solar/autothermal

AC C

EP

TE D

151

7

ACCEPTED MANUSCRIPT hybrid gasifier, is depicted in Fig. 1 by illustrating flow streams. The proposed system is

166

composed of two main subsystems, namely, syngas produced subsystem and CCHP

167

subsystem, which can simultaneously supply cooling, heat, and power. The redwood

168

feedstock was fed into the SAHG to produce syngas. Ultimate and proximate analyses of

169

feedstock are listed in Table 1. In view of energy saving and high-graded heat recovery, a

170

gas-water heat exchanger (HE-1) was equipped to harvest the sensible heat from the

171

product syngas exiting the SAHG. The cooled syngas is purified by a cyclone removing

172

particles, and is further cooled through a filter accompanied by the tar separating

173

process.

M AN U

SC

RI PT

165

The product syngas drives an ICE to generate power accompanied by waste heat of

175

coolant water (i.e. jacket water) and exhaust gas. The exhaust gas with an outlet

176

temperature of approximately 734 K is released from the ICE, and then drives a

177

double-effect absorption chiller (DEAC) to supply chilled water for cooling. Sequentially,

178

the exhaust gas passing through the DEAC with a temperature around 421 K is further

179

used to produce hot water with a temperature of 334 K by a gas-water heat exchanger

180

(HE-3). In order to conduct thermodynamic assessment of the CCHP system, the

181

nominal design parameters of all flow streams are listed in Table 2 based on a peak solar

182

power input of 300 kW. The standard state in the thermodynamic analysis is defined as

183

1 bar and 298 K.

184

3. System modelling

185

3.1. Solar/autothermal hybrid gasifier

AC C

EP

TE D

174

8

ACCEPTED MANUSCRIPT 186

Equilibrium composition of redwood pellets was computed via Gibbs free energy

187

minimization using the Aspen Plus code. Some reasonable assumptions, adopted in this

188

paper, are listed as follows:

190 191

chemical equilibrium. •

192

194 195 196

Only C, H, O contents of feedstock are considered neglecting other mineral contents and low species mole fractions (e.g., H2S, HCN).

•

The produce syngas from the gasifier is comprised of H2, CO, CO2, H2O, CH4, and N2 neglecting other higher hydrocarbon.

M AN U

193

The residence time of feedstock in the gasifier is long enough to achieve

RI PT

•

SC

189

Based on the above assumptions, the overall chemical reaction of the complex steam-air hybrid gasification can be represented by

CH୶ O୷ + ݉ሺOଶ + 3.76Nଶ ሻ + ‫ܯ‬Hଶ O + ‫ݓ‬Hଶ O

TE D

→ ݊ୌమ Hଶ + ݊େ୓ CO + ݊େ୓మ COଶ + ݊ୌమ ୓ Hଶ O + ݊େୌర CHସ + 3.76݉Nଶ

ሺ1ሻ

where x and y represent elemental mole ratios of H/C and O/C in feedstock materials,

198

respectively. ‫ ݓ‬and M are defined as the mole numbers of moisture and steam reactant

199

(i.e. steam-to-feedstock ratio) per mol of feedstock, respectively. ݊ୌమ , ݊େ୓ , ݊େ୓మ , ݊ୌమ ୓ ,

200

and ݊େୌర denote the moles of the species in the product syngas releasing from the

201

SAHG. m is defined as the mole number of oxygen reactant per mol of feedstock (i.e.

202

oxygen-to-feedstock ratio).

AC C

EP

197

203

In the actual steam or/and air gasification process, a series of competing

204

intermediate reactions need to be considered, which has been summarized and

205

elaborated in Refs. [7,9]. All these reactions are forcefully dependent on the pressure，

206

temperature, and C/O ratio.

9

ACCEPTED MANUSCRIPT 207

In view of conversion efficiency assessment of the feedstock-to-syngas process,

208

higher heating value ‫ܸܪܪ‬୤ of carbonaceous feedstock can be calculated using the

209

following correlation [42]:

RI PT

‫ܸܪܪ‬୤ = 0.3491‫ܨ‬େ + 1.1783‫ܨ‬ୌ − 0.1034‫ ୓ܨ‬− 0.0151‫ ୒ܨ‬+ 0.1005‫ܨ‬ୗ − 0.0211‫ܨ‬୅ ሺ2ሻ 210

where ‫ܨ‬େ , ‫ܨ‬ୌ , FO, FN, and FA are defined as the mass fractions of carbon (C), hydrogen

211

(H), oxygen (O), nitrogen (N), sulfur (S), and ash (A), respectively.

SC

213

The lower heating value ‫ܸܪܮ‬୤ of feedstock [43] with respect to higher heating value is calculated by

M AN U

212

‫ܸܪܮ‬୤ = ‫ܸܪܪ‬୤ − 21.978‫ܨ‬ୌ

ሺ3ሻ

For the feedstock material used in this paper, the corresponding higher heating

215

value and lower heating value are mentioned in Table 1. Fig. 2 indicates the equilibrium

216

compositions of the proposed feedstock material as a function of temperature at an

217

absolute pressure of 1 bar. CH4, CO, H2, and CO2 are stable components

218

thermodynamically below 600 K. Within the temperature range of 700‒1100 K, the

219

species in gasification were competed based on the intermediate reactions

220

aforementioned. All gasification processes except for the case of redwood with M = 0, m

221

= 0.329 (see Fig. 2b) trends to completion when the temperature beyond 1200 K. The

222

ratio of H2/CO indicates the quality of product syngas. In autothermal gasification (see

223

Fig. 2b), the air was fed into the gasifier which led to highly exothermic combustion that

224

releases a large amount of CO2 and N2. Consequently, the quality of syngas is notably

225

lower than the one obtained from fully steam gasification.

AC C

EP

TE D

214

10

ACCEPTED MANUSCRIPT Fig. 3 shows the enthalpy change ∆‫ܪ‬ோ of reaction equation (1) as a function of

227

temperature for the redwood at various steam-to-feedstock and oxygen-to-feedstock

228

ratios. The overall reaction of the redwood proceeds endothermically at the temperature

229

above 730 K. Under the condition of M = 0 and m = 0.329 with O2 participation, the

230

reaction was exothermic below around 973 K. This is mainly due to the low energy

231

species (CO2 and H2O) favored in the equilibrium composition. Compared steam

232

gasification to autothermal gasification, the combustion reaction results in a kind of

233

phenomenon that the enthalpy change shifts to a lower value during the whole

234

temperature range.

M AN U

SC

RI PT

226

A conceptual SAHG is schematically depicted in Fig. 4. The SAHG is adopted the

236

concept of indirectly irradiated two-cavity solar reactor [25] that includes the upper and

237

lower cavities, and is extended to couple conventional autothermal gasification in the

238

lower cavity. The detailed design and corresponding experiment of the solar reactor

239

have been reported [23–26]. The two-cavity reactor was designed for high-flux

240

concentrated solar irradiation collected by beam-down solar tower, avoiding the

241

negative effect of deposited gaseous and small-size particles on the quartz window

242

located at the entrancing aperture of the upper cavity. The major issue of the

243

degradation is the structural material of the absorber, in terms of maximum operating

244

temperature, thermal conductivity, radiative absorptivity, inertness to the chemical

245

reaction, and resistance to thermal shocks. The stable performance of the ceramic

246

material i.e. SiC-coated graphite, used in the solar gasifier with the temperature span of

247

1000–1500 K, have been demonstrated in the lab-scale test [24] for thermal cycling and

248

thermal shocks. Considering the magnification of the tower reflector such as

249

hyperboloid or ellipsoid, a three-dimensional compound parabolic concentrator (CPC) is

AC C

EP

TE D

235

11

ACCEPTED MANUSCRIPT placed at the aperture of the quartz window to decrease the size of aperture and

251

enhance the flux of incident rays. The solar beams, concentrated by CPC, penetrate the

252

quartz glass accompanied by absorbed and reflected losses, then hit on the absorber

253

made of SiC-coated graphite (i.e. emitter plate) that radiant emitter to lower cavity

254

supplying the necessary reaction process heat.

RI PT

250

Since the 1-D dynamic model of the two-cavity solar reactor has been proposed

256

[25,26], a simplified steady zero-dimension model was considered herein in the

257

thermodynamic analysis in view of the complex reaction processes of solar, autothermal,

258

and hybrid modes in the SAHG. The significant difference of the proposed SAHG is

259

highlighted by the assumptions:

260

•

261

•

265 266 267

Feedstock in the reaction bed is kept a constant height via controlling the feeding

EP

and removing rates.

•

TE D

removed automatically.

263 264

Feedstock is fed consecutively and solid residue of the by-product of gasification is

Feedstock inside the reactor has a uniform temperature field via omitting the heat

AC C

262

M AN U

SC

255

resistance from the top of bed to the bottom of bed.

•

In autothermal or hybrid mode, the exothermic reaction of combustion can be well proceeded in the SAHG.

268

The upper cavity requires purge via injecting a 2 lN/min Ar flow from the

269

water-cooled surface [26], which leads to a series of convective heat fluxes including the

270

forced convective heat flux (‫ݍ‬ଶିଷ,conv ) from outer surface of the emitter plate to the

12

ACCEPTED MANUSCRIPT purged gas (Ar), the forced convective heat flux (‫ݍ‬ଷିସ,conv ) from Ar gas to insulation cone,

272

and the forced convective heat flux (‫ݍ‬ଷି଺,conv ) from Ar gas to the surface-6 of the quartz

273

glass. Based on the above convective heat fluxes and other radiative heat fluxes, the

274

energy balance equation of the emitter plate that combined the upper and lower cavities

275

is given by

RI PT

271

‫ݍ‬ଵିଵ଴,conv + ‫ݍ‬ଶିଷ,conv = ‫ݍ‬ଵ,rad + ‫ݍ‬ଶ,rad

ሺ4ሻ

where ‫ݍ‬ଵିଵ଴,conv indicates the convective heat flux from surface-1 of the emitter plate to

277

the gas reactants. ‫ݍ‬ଵ,rad and ‫ݍ‬ଶ,rad are defined as net radiative fluxes on the surface-1

278

and surface-2 of the emitter plate, respectively. The emissivity of 0.88 for the gray

279

diffuse emitter plate was applied in the thermodynamic analysis.

M AN U

SC

276

The quartz glass temperature was determined by the energy balance in terms of

281

absorbed flux of concentrated solar irradiation as well as various radiative and

282

convective heat fluxes between enclosure surfaces. A temperature of 298 K is on the

283

water-cooled surface (surface 11) [26], which causes a significant heat sink,

TE D

280

EP

‫଺ݍ‬,rad + ‫଻ݍ‬,rad = ‫ି଺ݍ‬ଷ,conv + ‫଼ି଻ݍ‬,conv

ሺ5ሻ

where ‫ି଺ݍ‬ଷ,conv represents the convective heat flux from the surface-6 of the quartz

285

glass to the Ar gas, and ‫଼ି଻ݍ‬,conv indicates the convective heat loss from the surface-7 of

286

the quartz glass to the environment. ‫଺ݍ‬,rad and ‫଻ݍ‬,rad are defined as net radiative fluxes

287

on the surface-6 and surface-7 of the quartz glass, respectively. The hemispherical

288

band-approximated property of the quartz glass, reported in Ref. [44], was used in the

289

radiative transmitting calculation.

AC C

284

13

ACCEPTED MANUSCRIPT 290 291

The energy balance of the inner surface of the insulation cone in the upper cavity is expressed as ‫ݍ‬ସିହ,cond = ‫ݍ‬ଷିସ,conv + ‫ݍ‬ସ,rad

ሺ6ሻ

where ‫ݍ‬ସିହ,cond is the conductive heat flux from surface-4 to surface-5 for the

293

insulation cone. ‫ݍ‬ସ,rad is defined as the net radiative heat flux on the surface-4 of the

294

insulation cone.

SC

296

Corresponding energy equation of the outer surface of the insulation cone is calculated by

M AN U

295

RI PT

292

‫ݍ‬ସିହ,cond + ‫ݍ‬ହ,rad = ‫ݍ‬ହି଼,conv

ሺ7ሻ

where ‫ݍ‬ହି଼,conv represents the convective heat flux from the surface-5 of the insulation

298

cone in the upper cavity to the environment, and ‫ݍ‬ହ,rad denotes the net radiative heat

299

flux on the surface-5 of the insulation cone.

TE D

297

The lower cavity (reactor) includes the reaction bed, SiC wall, and insulation. As

301

mentioned in the previous assumptions, taking into account of the complexity of hybrid

302

gasification mode, we assumed the feedstock bed has a uniform temperature field via

303

omitting the heat resistance from the top of bed to the bottom of bed. Effective

304

convective heat transfer ‫ݍ‬ଵିଵ଴,conv was absorbed by the gas reactants (steam-air) and

305

then released into the feedstock bed to become the gasification process heat. The steady

306

energy balance equation of the reaction bed is:

AC C

EP

300

‫ݍ‬ଵିଵ଴,conv + ‫ݍ‬ଽ,rad + ‫ݍ‬ଵଷିଵ଴,conv +‫ݍ‬ଵସିଽ,cond = mሶ f ∆ℎR

14

ሺ8ሻ

ACCEPTED MANUSCRIPT where ‫ݍ‬ଽ,rad represents the net radiative heat flux on the surface-9 (top of the feedstock

308

bed). mሶ f is the reaction rate that is dependent on the kinetic of the feedstock material.

309

‫ݍ‬ଵଷିଵ଴,conv denotes the convective heat flux from the wall-13 to the steam-air reactant.

310

‫ݍ‬ଵସିଽ,cond denotes the conductive heat flux from the wall-14 to feedstock.

311

RI PT

307

The steady energy equation of the upper bed wall is formulated by

ሺ9ሻ

SC

‫ݍ‬ଵଷ,rad = ‫ݍ‬ଵଷିଵ଴,conv +‫ݍ‬ଵଷିଵହ,cond +‫ݍ‬ଵଷିଵସ,cond

where ‫ݍ‬ଵଷିଵହ,cond represents the conductive heat loss from wall-13 to surface-15 of the

313

insulation. ‫ݍ‬ଵଷ,rad is defined as the net radiative heat flux on the wall-13. ‫ݍ‬ଵଷିଵସ,cond

314

represents the conductive heat flux from wall-13 to wall-14.

315

M AN U

312

The steady energy equation of the lower bed wall is formulated by

TE D

‫ݍ‬ଵଷିଵସ,cond = ‫ݍ‬ଵସିଽ,cond +‫ݍ‬ଵସିଵଶ,cond

ሺ10ሻ

where ‫ݍ‬ଵସିଵଶ,cond represents the conductive heat loss from wall-14 to surface-12 of the

317

insulation.

319

320

Finally, the energy balance equations of the upper and lower bed insulation are

AC C

318

EP

316

descried as

‫ݍ‬ଵଷିଵହ,cond +‫ݍ‬ଵହ,rad = ‫ݍ‬ଵହି଼,conv

ሺ11ሻ

‫ݍ‬ଵସିଵଶ,cond +‫ݍ‬ଵଶ,rad = ‫ݍ‬ଵଶି଼,conv

ሺ12ሻ

where ‫ݍ‬ଵଶି଼,conv is the convective heat loss from the surface-12 of insulation to the

15

ACCEPTED MANUSCRIPT 321

environment. ‫ݍ‬ଵଶ,rad is the net radiative heat flux on the insulation surface-12.

322

‫ݍ‬ଵହି଼,conv is the convective heat loss from the upper insulation (surface-15) to the

323

environment.

325

The aforementioned convective heat flux from surface i to surface j can be

RI PT

324

uniformly expressed by

ሺ13ሻ

SC

‫ݍ‬௜ି௝,conv = ℎ௜ି௝,conv ‫ܣ‬௜ ൫ܶ௜ − ܶ௝ ൯

Where ‫ܣ‬௜ means the area of surface i. ܶ௜ and ܶ௝ represent the temperatures of

327

surface i and surface j, respectively. ℎ௜ି௝,conv represents convective heat transfer

328

coefficient.

M AN U

326

The corresponding convective heat transfer coefficients (see Table 3) were

330

determined [26] and used in this paper. The convective heat loss from insulation to the

331

environment was computed by the Nusselt correlation [26], and its radiative heat losses

332

(i.e. ‫ݍ‬ଵଶ,rad, ‫ݍ‬ହ,rad , and ‫ݍ‬ଵହ,rad) were calculated by Stefan–Boltzmann law [2].

TE D

329

In this work, the net radiative flux of each surface i of the upper cavity or the lower

334

cavity, in terms of detailed expression and calculation, was computed via the method

335

proposed by Piatkowski and Steinfeld [26]. The surfaces participating in radiative heat

336

transfer, given in Table 4, were considered in the calculation based on Eqs. (14) and (15).

337

The view factors in the radiative heat transfer were calculated by using Monte Carlo ray

338

tracing method. Here, only the primary equations are given:

AC C

EP

333

ே

‫ݍ‬௜,rad = ෍ ‫ݍ‬௜,௕ ௕ୀ଴

16

ሺ14ሻ

ACCEPTED MANUSCRIPT 339

with ‫ݍ‬i,b = ‫ݍ‬solar,i,b + ‫ݍ‬in,i,b − ‫ݍ‬out,i,b − ‫ݍ‬trans,i,b

ሺ15ሻ

where ‫ݍ‬i,b is the net radiative flux on a surface i at the spectral band b. ‫ݍ‬out,i,b is

341

defined as the spectral radiosity emitted from the surface i to the enclosure inside.

342

‫ݍ‬trans,i,b is the spectral radiative flux leaving from the surface i to the environment.

343

‫ݍ‬solar,i,b denotes the external flux contributed by solar radiation to the surface i. ‫ݍ‬in,i,b

344

denotes the reflection flux of incident solar rays hitting on the surface i.

345

3.2. Internal combustion engine

M AN U

SC

RI PT

340

It is well-known that the design of a conventional gas ICE is based on natural gas.

347

When it comes to product syngas, the output performance of ICE will significantly

348

decrease. In order to reasonably assess output performance of electrical power, coolant

349

heat, and exhaust heat of a gas ICE fired with syngas, various modified models of ICE has

350

been proposed and reported [34].

be calculated by

EP

352

The first law of thermodynamic for a control volume ICE under the steady flow can

AC C

351

TE D

346

‫ݍ‬s + ݉ሶୟ୧୰ ‫ܪ‬ୟ୧୰ = ܰ௘ + ‫ݍ‬col + ‫ݍ‬ext + ‫ݍ‬loss

ሺ16ሻ

353

where ‫ݍ‬s denotes the fired syngas energy. ܰ௘ is the power output of ICE. ݉ሶୟ୧୰ is the

354

mass flow rate of the air fed into the ICE. ‫ܪ‬ୟ୧୰ represents the enthalpy of the air. ‫ݍ‬col ,

355

‫ݍ‬ext , and ‫ݍ‬loss are coolant heat, exhaust heat, and heat loss of an ICE, respectively.

356

The syngas energy fed into an ICE can be calculated by ‫ݍ‬s = ܸୱሶ ‫ܸܪܮ‬s 17

ሺ17ሻ

ACCEPTED MANUSCRIPT 357

where ܸୱሶ denotes the volume flow rate of the syngas fed into an ICE, which is

358

determined by the performance of the SAHG. ‫ܸܪܮ‬s is lower heating value of product

359

syngas.

361

The electrical efficiency ߟe of the syngas-fired ICE in terms of the nominal electrical power can be calculated by the following formula [31]: ߟe = 28.08ܰ୬ୣ ଴.଴ହ଺ଷ ൬0.102

RI PT

360

‫ܸܪܮ‬s + 0.897൰ ‫ܸܪܮ‬N

ሺ18ሻ

where ܰ୬ୣ denotes the nominal electrical power of a natural gas fired ICE, which

363

means the maximum power. ‫ܸܪܮ‬N represents the lower heating value of natural gas.

M AN U

SC

362

The exhaust gas composition mainly depends on the air-fuel ratio and the complex

365

combustion features in the ICE. The fired syngas was assumed to be completely

366

combusted associated with equilibrium state. Thus, the corresponding temperature of

367

exhaust gas for the syngas-fired ICE can be calculated by the following formula [32]: ܶext = ൬0.025

‫ܸܪܮ‬s + 0.974൰ ൫2 × 10ିହ ܰ୬ୣ ଶ − 0.0707ܰ୬ୣ + 758.33൯ ‫ܸܪܮ‬N

ሺ19ሻ

EP

The coolant heat from the ICE jacket was assumed to be equal to 17% syngas

369

energy [34], and is calculated by

370

AC C

368

TE D

364

where ݉ሶୡ୭୪ denotes the mass flow rate of coolant water. ܿ௣,col is the specific heat of

371

coolant water. ܶcol,in and ܶcol,out are inlet and outlet temperatures of coolant water for

372

the ICE, respectively.

373

3.3. Double-effect absorption chiller

‫ݍ‬col = ݉ሶୡ୭୪ ܿ௣,col ൫ܶcol,out − ܶcol,in ൯ = 0.17‫ݍ‬s

18

ሺ20ሻ

ACCEPTED MANUSCRIPT The simulation model of a pressure-water double-effect absorption chiller, based on

375

theoretical analysis as well as the energy and mass balance of solution, referred to the

376

one proposed by Kaushik and Arora [33]. The inputs of temperatures and flow rates of

377

hot water, chilled water, and cooling water are necessary parameters. In this work, the

378

inlet temperature of chilled water was kept a constant of 12 °C. The temperature of

379

cooling water entrancing the chiller was affected by the behavior of cooling tower and

380

the local weather condition.

381

3.4. Heat exchanger

SC

RI PT

374

Logarithmic mean temperature difference (LMTD) was adopted in the

383

thermodynamic calculation of heat exchangers. All heat exchangers, proposed in this

384

paper, have the counter-flow structure. The energy balance of heat exchangers can be

385

expressed as

M AN U

382

TE D

݉ሶୡ ܿ௣,c ൫ܶc,out − ܶc,in ൯ = ߟHE ݉ሶ୦ ܿ௣,h ൫ܶh,in − ܶh,out ൯

ሺ21ሻ

where ݉ሶୡ and ݉ሶ୦ are the mass flow rates of cold fluid and hot fluid, respectively. ܿ௣,c

387

and ܿ௣,h are the specific heat of cold fluid and hot fluid, respectively. ܶc,in and ܶc,out

388

are inlet and outlet temperatures of cold fluid, respectively. ܶh,in and ܶh,out are inlet

389

and outlet temperatures of hot fluid, respectively. ߟHE represents effectiveness of heat

390

exchangers.

391

3.5. Performance assessment indexes

392 393

AC C

EP

386

Cold gas efficiency of gasification process (i.e. solar-to-fuel efficiency in solar or hybrid mode) is defined as

19

ACCEPTED MANUSCRIPT ߟCG = 394

݉ሶୱ ‫ܸܪܮ‬s ݉ሶ୤ ‫ܸܪܮ‬f + ‫ݍ‬ୱ୭୪ୟ୰

ሺ22ሻ

With ‫ݍ‬ୱ୭୪ୟ୰ = ‫ܫ‬DNI ‫ܥ‬

ሺ23ሻ

where ߟCG is cold gas efficiency of gasification. ‫ݍ‬ୱ୭୪ୟ୰ is total solar power input at the

396

entrancing aperture of the quartz glass. ‫ܫ‬DNI is direct normal irradiance. C represents

397

the average concentration ratio of solar flux over the aperture.

SC

399

The solar upgrade factor, defined as the ratio of the lower heating value of product syngas to that of the feedstock fed, is expressed as the following formula:

M AN U

398

RI PT

395

ܷ=

401

ሺ24ሻ

In order to assess the performance of the SAHG-CCHP system, the primary energy ratio (PER) is defined as

‫ݍ‬୮୭୵ + ‫ݍ‬୦୵ + ‫ݍ‬ୡ୦୵ ݉ሶ୤ ‫ܸܪܮ‬f

TE D

400

݉ሶୱ ‫ܸܪܮ‬s ݉ሶ୤ ‫ܸܪܮ‬f

ܲ‫= ܴܧ‬

ሺ25ሻ

where ‫ݍ‬୮୭୵ is total electricity output of the ICE, ‫ݍ‬୦୵ is total heat recovered from

403

coolant water by HE-2 and from exhaust gas of ICE by HE-3. ‫ݍ‬ୡ୦୵ is cooling output of

404

the double-effect chiller.

405

4. Results and discussion

406

4.1. Model validation

AC C

EP

402

407

The zero-dimension steady model developed in this work was validated via

408

comparing the deviation between the simulated results and reported data [23] of a

409

scale-up 200 kW solar reactor. To conduct an accurate validation, all parameters of the

20

ACCEPTED MANUSCRIPT reactor in terms of geometry, optical features, and initial inputs were identical with that

411

reported in Ref. [26]. The industrial sludge [26] served as the feedstock material in the

412

model validation. The primary input parameters of the proposed model in the fully solar

413

steam mode are listed in Table 5, where three cases with different bed diameters were

414

considered in this validation. The dynamic reaction rates [26] of industrial sludge in 12h

415

operation time were adopted and served as input values of ݉ሶ୤ considering a time

416

interval of 5 min. In this case, the bed sharking rate is dependent on the reaction rate.

417

The finalized comparison results regarding the mean cold gas efficiency (i.e. mean

418

solar-to-fuel efficiency) are given in Table 6. Comparison results show that the relative

419

deviation between the developed model and reported data are 14.3%, 12.1%, and 10.1%

420

for the cases with bed diameters of 1.5m, 2.0 m, and 2.5 m, respectively. It means the

421

developed zero-dimension model overestimated the overall syngas energy, which is

422

mainly due to the modeling assumption of omitting the heat resistance of feedstock bed.

423

Because the uniform temperature field of the feedstock underestimated the emitter

424

temperature and the top surface temperature of the feedstock bed, it leads to an

425

overestimated net radiative heat flux from the emitter to the feedstock accompanied by

426

a higher cold gas efficiency. In general, the developed model is reasonable and feasible

427

to conduct thermodynamic evaluation of the SAHG.

428

4.2. Input parameters

AC C

EP

TE D

M AN U

SC

RI PT

410

429

In order to carry out the thermodynamic analysis of the SAHG-CCHP system and

430

compare the difference of solar, autothermal, and hybrid modes, the inherent input

431

parameters of the SAHG, ICE, and DEAC are given in Table 7. In the following work, the

432

two-cavity SAHG with the bed diameter of 1.5 m and a reasonable reaction rate of 0.5

433

mol/s were considered. As we mentioned in assumptions, a 0.5 m constant height of the

21

ACCEPTED MANUSCRIPT feedstock bed was initialed herein via controlling the feeding rate of feedstock. In the

435

potential evaluation of the solar upgraded ratio in Singapore, the temperatures of

436

cooling water and chilled water entrancing the chiller are respectively assumed as a

437

constant of 30 °C and 12 °C.

438

4.3. Optimization of steam-to-feedstock and oxygen-to-feedstock ratios

RI PT

434

The previous research, regarding the effects of steam-to-feedstock ratio and

440

equivalence ratio (ER) on the performance of autothermal gasification with and without

441

heat loss, has been well conducted by experiment and simulation [45]. The objective of

442

this section is to analyze the effect of two critical ratios, namely steam-to-feedstock ratio

443

M ( 0 ≤ ‫ ≤ ܯ‬3 ) and oxygen-to-feedstock ratio m ( 0 ≤ ‫ ≤ ܯ‬0.6 ), on the reaction

444

temperature, lower heating value of product syngas, and cold gas efficiency

445

(solar-to-fuel efficiency in the fully solar mode), at different solar flux conditions. The

446

contour maps of reaction temperature as a function of M and m in autothermal and solar

447

modes are depicted in Fig. 5. Here, 300 kW solar power was considered as an input over

448

the aperture of the SAHG in the solar/hybrid mode (see Figs. 5b). Both in autothermal

449

and solar/hybrid modes, the equilibrium reaction temperature Teq increases distinctly

450

with the oxygen-to-feedstock ratio as a result of increasing exothermic reaction of

451

combustion. However, under the same oxygen-to-feedstock ratio, the reaction

452

temperature in the solar/hybrid mode represents significantly higher than that in the

453

autothermal gasification due to the solar power input. The reaction temperature

454

decreases with the increasing steam-to-feedstock ratio, which is mainly due to the

455

steam-reforming endothermic reaction.

AC C

EP

TE D

M AN U

SC

439

456

The contour map of the cold gas efficiency as a function of M and m is shown in Fig.

457

6. It is observed that the cold gas efficiency (i.e. solar-to-fuel efficiency) in solar/hybrid 22

ACCEPTED MANUSCRIPT mode is significantly lower than the cold gas efficiency in the autothermal gasification.

459

The primary reason is a large amount of incident solar power was remitted or lost to the

460

environment and to the cooling water. However, the syngas quality was definitely

461

upgraded by solar, which refers to the solar upgraded ratio that will be discussed in the

462

following sections. As shown in Fig. 6a, in the autothermal mode, the high efficiency

463

region can be observed at the upper-left corner due to the high content of CH4

464

contributed by a high steam input. In the solar/hybrid mode (see Figs. 6b), the

465

optimized region of cold gas efficiency shifted to the left corner without oxygen input

466

(i.e. fully solar steam gasification). Another significant finding is that a minimum value

467

of steam-to-feedstock ratio can be obtained at the high efficiency region, since the

468

increasing steam input has no distinct effect on the cold gas efficiency.

M AN U

SC

RI PT

458

Fig. 7 illustrates the effects of M and m on the lower heating value of dry product

470

syngas. The lower heating value decreased with the increasing oxygen-to-feedstock ratio

471

due to the increasing by-product yield of combustion such as CO2 and N2. In the

472

autothermal gasification mode, the high quality of syngas located at the upper-left

473

boundary (high M and low m) as a result of a high content of CH4 at low equilibrium

474

reaction temperatures. Whereas, it can be observed that the lower-left corner with the

475

low steam-to-feedstock and oxygen-to-feedstock ratios produced high quality of syngas

476

in the solar/hybrid mode.

EP

AC C

477

TE D

469

It

is

feasible

to

determine

the

optimized

steam-to-feedstock

and

478

oxygen-to-feedstock ratios via establishing a certain restriction condition [45] of

479

reaction temperature under various solar power inputs. In this work, in order to boost

480

char conversion and minimize tar production, a reaction temperature of over 1000 K

481

was expected to be achieved. It means the optimized parameters are required to obtain

23

ACCEPTED MANUSCRIPT the maximum cold gas efficiency under the precondition of satisfying the temperature

483

over 1000 K and minimizing steam input. Consequently, the corresponding optimal

484

ratios of M and m as a function of the normalized incident solar power with the peak

485

value of 300 kW is shown in Fig. 8a. The oxygen-to-feedstock ratio decreased from 0.47

486

to 0 with the increasing normalized solar power accompanied by the reducing portion

487

of feedstock combustion in the reactor. In contrast, the steam-to-feedstock ratio

488

gradually increased from 0 to 0.6, and then decreased from 0.6 to 0.28, presenting a

489

peak value at qsolar/qsolar,peak = 0.7, which is mainly due to the temperature restriction of

490

over 1000 K. In fact, different restriction of temperature definitely leads to different

491

optimal values of M and m. Noted that the optimized steam-to-feedstock ratio at the

492

peak solar power input (qsolar/qsolar,peak = 1) is equal to the stoichiometry reported in the

493

overall reaction equation of solar steam gasification [9].

M AN U

SC

RI PT

482

The cold gas efficiency and solar upgraded ratio of redwood as a function of the

495

normalized solar power input is shown in Fig. 8b. It can be seen that an increment in the

496

portion of solar power input from autothermal mode (i.e. qsolar /qsolar,peak = 0) to fully

497

solar mode (i.e. qsolar /qsolar,peak = 1) contributes to the increasing solar upgraded ratio

498

from 0.67 to 1.32. However, the cold gas efficiency of solar/hybrid mode was always

499

lower than that of the autothermal mode as a result of a large amount of solar power

500

remitted and convective heat losses to the environment and cooling water.

AC C

EP

TE D

494

501

The cold syngas composition as a function of the normalized solar power is

502

depicted in Fig. 9a. It can be seen that the contents of N2 and CO2 both dramatically

503

decreased with the increasing solar power input accompanied by the increasing

504

contents of H2 and CO. The content of CH4 presented a peak yield around qsolar /qsolar,peak

505

= 0.7 which is corresponding to the peak of steam-to-feedstock ratio shown in Fig. 8a.

24

ACCEPTED MANUSCRIPT Fig. 9b shows the variation of mole flow rates of hot syngas and cold syngas as well as

507

reaction temperature vs the normalized solar power. The mole flow rates both

508

decreased as the normalized solar power increased, where the hot gas includes steam

509

component. The equilibrium reaction temperature of the SAHG shows a constant value

510

of 1000 K when the normalized solar power is within the range of 0 to 0.7. It means sole

511

solar power input was insufficient to reach the minimum reaction temperature of 1000

512

K during this normalized solar power span and the gasifier operated at the hybrid mode

513

for supplements via adjusting the steam-to-feedstock and oxygen-to-feedstock ratios.

514

However, it increased hugely when the normalized solar power was beyond 0.7

515

contributed by more solar power input, which means sole solar power input exceeded

516

the critical process heat demand for the reaction temperature of 1000 K.

517

4.4. Potential evaluation of the SAHG-CCHP system in Singapore

M AN U

SC

RI PT

506

To analyze the operation behavior of the SAHG-CCHP system, two consecutive days

519

in the typical meteorological year (TMY) of Singapore were determined. In this work, we

520

only considered DNI neglecting the effect of the ambient temperature. Fig. 10 shows the

521

direct normal irradiances of the two consecutive days and whole typical year in

522

Singapore, derived from TRNSYS weather database [46]. As shown in Fig. 10a, the first

523

day was a representative sunny day with few fluctuation, while the second day was a

524

real cloudy day. DNI in Singapore was limited below 800 W/m2 most of the time due to

525

its cloudy and rainy climate.

AC C

EP

TE D

518

526

The mole flow rates of steam, oxygen, and redwood fed into the SAHG in the

527

consecutive days are depicted in Fig. 11a. The feeding rate of redwood maintained a

528

constant of 0.5 mol/s during the autothermal or hybrid mode. The feeding mole flow

529

rates of the gas reactants were controlled by the optimized ratios proposed in Fig. 8a 25

ACCEPTED MANUSCRIPT according to the solar power input. The system switched automatically during the

531

autothermal, fully solar, and hybrid modes to reach the reaction temperature presented

532

in Fig. 9b. The transient variations of solar upgraded ratio and lower heating value of

533

syngas are shown in Fig. 11b. In general, the variation trend of solar upgraded ratio and

534

lower heating value of syngas abided by the change of DNI. In autothermal gasification

535

mode, an average solar upgraded ratio (i.e. cold gas efficiency) of 0.76 was reached.

RI PT

530

The corresponding mole flow rates of syngas species for the autothermal and SAHG

537

are shown in Fig. 12. It is observed that the CO2 and N2 emissions of the SAHG were

538

significantly reduced with the increasing DNI compared to autothermal gasification. The

539

shaded regions denote alleviated CO2 and N2 emissions contributed by incident solar

540

power. The results of numerical integration of the two consecutive days indicate that

541

reductions in mole flow rates of CO2 and N2 by 17.8% and 28.6% were achieved,

542

respectively, compared to the autothermal gasification (the baseline data). In contrast,

543

H2 and CO mole flow rates were increased by over 38.8% and 11.8%, respectively,

544

leading to an increment in LHVs by 51.7% as shown in Fig. 11b.

TE D

M AN U

SC

536

Fig. 13 reveals the variations of heat, cooling and power of the SAHG-CCHP system

546

in the two days. It can be seen that power and heat produced in the system distinctly

547

followed the variation trend of DNI due to the increments in LHVs and total syngas

548

energy fed into the ICE. Noted that the cooling capacity of the DEAC is primarily

549

dependent on the inlet temperature of hot water (state 8 in Fig. 1) that is directly

550

affected by the temperature and flow rate of hot gas (state 3 in Fig. 1) for HE-1 as well as

551

the temperature and flow rate of exhaust gas (state 5 in Fig. 1) for HE-4. As discussed

552

previously, the flow rate of hot gas decreased with the increasing normalized solar

553

power, but its temperature increased when the normalized solar power was over 0.7. In

AC C

EP

545

26

ACCEPTED MANUSCRIPT addition, the flow rate of exhaust gas definitely decreased with DNI due to the reduced

555

flow rate of syngas, however, its temperature increased as a result of increased LHVs.

556

Therefore, the variation of cooling capacity is determined by the trade-off between

557

reduction in flow rates and increment in temperatures for hot syngas and exhaust gas,

558

which justified the phenomenon that during some periods of high DNI, the cooling

559

capacity exhibited a significant reduction. However, the primary energy ratio well

560

followed the variation of DNI. PER = 1.13 at the peak DNI of 802 W/m2 was obtained.

561

Compared to the CCHP system driven by autothermal gasification, an increment in

562

primary energy ratio by 11.5% can be achieved by using the SAHG-CCHP system. Heat,

563

cooling, and power were respectively increased by 24%, 1.3%, and 27.3% due to the

564

hybrid operation mode.

M AN U

SC

RI PT

554

The annual performance of the SAHG-CCHP system with respect to the monthly

566

distribution of cooling, heat, power, feedstock energy, and incident solar power is shown

567

in Fig. 14, taking into account of a 15 min time interval. The simulated operation period

568

of the system is from 7 am to 7 pm each day. As expected, the monthly incident solar

569

power exhibited slight difference due to the climate near the equator, leading to the

570

results of small monthly differences in heat, power and cooling distributions. Specifically,

571

Qhw, Qchw, and Qpow slightly varied within the range of 9.2–10.1, 37.1–40.8, and 20.6–22.6

572

MW h, respectively. Compared to the autothermal system, the yearly average increments

573

in heat, power, and cooling were reached by 19.5%, 23.8%, and 4.5%, respectively. It

574

indicates the most significant improvement in SAHG-CCHP system is electric power

575

output.

AC C

EP

TE D

565

576

Based on the compiled results shown in Fig. 14, the finalized comparison results of

577

PER for the autothermal-based and SAHG-based CCHP systems are listed in Table 8.

27

ACCEPTED MANUSCRIPT Since the effect of ambient temperature was neglected in this work, the PER of the

579

autothermal system remained as a constant of 0.782. In contrast, due to the small

580

difference in accumulated monthly solar power input, the PER of the SAHG system

581

slightly varied from 0.871 to 0.921 accompanied by increments of 11.4% and 17.8%,

582

respectively. It can be observed that the SAHG-CCHP system has distinct benefits of

583

energy saving through operating the hybrid mode at all times of the whole year. A yearly

584

average increment of 14.2% in primary energy ratio of the proposed SAHG-CCHP system

585

can be obtained under the solar radiation condition of Singapore.

586

5. Conclusions and future work

M AN U

SC

RI PT

578

A specific ICE CCHP system driven by a solar/autothermal hybrid gasifier has been

588

introduced and studied thermodynamically, where the ICE CCHP system contains

589

gas-water waste heat exchangers and a double-effect absorption chiller. The SAHG

590

combined with the concept of indirectly irradiative two-cavity reactor has been

591

introduced and investigated theoretically, which can automatically switched during the

592

fully solar, hybrid, and autothermal modes according to the incident solar power input.

593

To conduct the performance analysis of the SAHG, a simplified zero-dimensional

594

steady-state model was developed and validated with the reported data. The potential of

595

solar upgraded ratio for the SAHG-CCHP system has been assessed under the solar

596

radiation data of Singapore.

AC C

EP

TE D

587

597

Both in autothermal and solar/hybrid modes, the equilibrium reaction temperature

598

increases distinctly with the oxygen-to-feedstock ratio as a result of increasing

599

exothermic reaction of combustion. The reaction temperature decreases with the

600

increasing steam-to-feedstock ratio, which is mainly due to the steam-reforming

601

endothermic reaction. The cold gas efficiency (i.e. solar-to-fuel efficiency) in 28

ACCEPTED MANUSCRIPT solar/hybrid mode is significantly lower than the cold gas efficiency in the autothermal

603

gasification. Because solar power was remitted or lost to the environment and to the

604

cooling water. However, the syngas quality is definitely upgraded by solar energy

605

evidenced by an upgraded lower heating value of the product syngas.

RI PT

602

The optimized steam-to-feedstock and oxygen-to-feedstock ratios of the SAHG for

607

redwood pellets were obtained under various incident solar fluxes and a reaction

608

temperature condition of over 1000 K. Considering a 300 kW peak solar power input,

609

the oxygen-to-feedstock ratio decreased from 0.47 to 0 with the increasing normalized

610

solar power accompanied by the reducing portion of feedstock combustion in the

611

reactor. In contrast, the steam-to-feedstock ratio gradually increased from 0 to 0.6, and

612

then decreased from 0.6 to 0.28, presenting a peak value at qsolar/qsolar,peak = 0.7. An

613

increment in the portion of solar power input from autothermal mode (i.e. qsolar

614

/qsolar,peak = 0) to fully solar mode (i.e. qsolar /qsolar,peak = 1) contributes to the increasing

615

solar upgraded ratio from 0.67 to 1.32. The contents of N2 and CO2 both dramatically

616

decreased with the increasing solar power input accompanied by the increasing

617

contents of H2 and CO.

EP

TE D

M AN U

SC

606

The operation results of two typical days indicate the variation trend of solar

619

upgraded ratio and lower heating value of syngas abided by the change of DNI. In

620

autothermal gasification mode, an average solar upgraded ratio (i.e. cold gas efficiency)

621

of 0.76 was reached. Reductions in mole flow rates of CO2 and N2 by 17.8% and 28.6%

622

were achieved, respectively, compared to the autothermal gasification. In contrast, H2

623

and CO mole flow rates were increased by over 38.8% and 11.8%, respectively, leading

624

to an increment in LHVs by 51.7%. Compared to the CCHP system driven by autothermal

625

gasification, an increment in primary energy ratio by 11.5% can be achieved by using

AC C

618

29

ACCEPTED MANUSCRIPT 626

the SAHG-CCHP system. Heat, cooling, and power were respectively increased by 24%,

627

1.3%, and 27.3% due to the hybrid operation mode. The yearly assessment of the SAHG-CCHP system shows that the monthly incident

629

solar power exhibited slight difference due to the Singapore’s climate near the equator,

630

leading to the results of small monthly differences in heat, power and cooling

631

distributions. Compared to the autothermal system, the yearly average increments in

632

heat, power, and cooling for the SAHG system were reached by 19.5%, 23.8%, and 4.5%,

633

respectively. A yearly average increment of 14.2% in primary energy ratio of the

634

proposed SAHG-CCHP system can be obtained under the solar radiation condition of

635

Singapore.

M AN U

SC

RI PT

628

It has been well proven that the SAHG system has distinct benefits of energy saving

637

through operating the hybrid mode at all times of the whole year. The solar/autothermal

638

hybrid gasification process promises to be competitive with conventional autothermal

639

and solar gasification, mainly benefited from lower feedstock consumption and the

640

elimination of additional inefficiencies and costs in start-up and shutdown processes

641

caused by the intermittency of solar radiation, respectively. Future work will focus on

642

the prototypical design of the hybrid solar/autothermal gasifier, and investigation of the

643

complex dynamic behaviors of the SAHG-CCHP system.

644

Acknowledgement

AC C

EP

TE D

636

645

This research is supported by the National Research Foundation (NRF), Prime

646

Minister’s Office, Singapore under its Campus for Research Excellence and Technological

647

Enterprise (CREATE) programme (Grant Number R-706-001-101-281).

648

References 30

ACCEPTED MANUSCRIPT 649

[1] Ruiz J, Juárez M, Morales M, Muñoz P, Mendívil M. Biomass gasification for electricity

650

generation: review of current technology barriers. Renewable and Sustainable

651

Energy Reviews 2013;18:174–83.

653

[2] Steinfeld A. Solar thermochemical production of hydrogen—a review. Solar energy

RI PT

652

2005;78(5):603–15.

[3] Service, R. Biomass fuel starts to see the light. Science 2009;326(5959):1474.

655

[4] Lee U, Chung JN, Ingley HA. High-Temperature Steam Gasification of Municipal Solid

657

Waste, Rubber, Plastic and Wood. Energy & Fuels 2014;28(7):4573–87.

M AN U

656

SC

654

[5] Bonalumi D, Giuffrida A. Investigations of an air-blown integrated gasification

658

combined cycle fired with high-sulphur coal with post-combustion carbon capture

659

by aqueous ammonia. Energy 2016;117(Part 2):439–49.

662

Netherlands, 2004.

TE D

661

[6] Steynberg A. Dry, M. Fischer−Tropsch Technology. Elsevier BV: Amsterdam, The

[7] Ong Z, Cheng Y, Maneerung T, Yao Z, Tong YW, Wang CH. Co-gasification of woody

EP

660

biomass and sewage sludge in a fixed-bed downdraft gasifier. AIChE Journal

664

2015;61(8):2508–21.

665 666 667 668 669

AC C

663

[8] Kumabe K, Hanaoka T, Fujimoto S, Minowa T, Sakanishi K. Co-gasification of woody biomass and coal with air and steam. Fuel 2007;86(5):684–9. [9] Piatkowski N, Wieckert C, Weimer AW, Steinfeld A. Solar-driven gasification of carbonaceous feedstock—a review. Energy & Environmental Science 2011;4:73–82. [10]

Howaniec N, Smoliński A. Biowaste utilization in the process of co-gasification

31

ACCEPTED MANUSCRIPT 670 671

with bituminous coal and lignite. Energy 2017;118(Supplement C):18–23. [11]

Yao Z, Li W, Kan X, Dai Y, Tong YW, Wang C-H. Anaerobic digestion and

gasification hybrid system for potential energy recovery from yard waste and woody

673

biomass. Energy 2017;124(Supplement C):133–45.

675 676

[12]

Kalogirou SA. Solar thermal collectors and applications. Progress in Energy and

Combustion Science 2004;30(3):231–95. [13]

SC

674

RI PT

672

Noone CJ, Torrilhon M, Mitsos A. Heliostat field optimization: A new

computationally efficient model and biomimetic layout. Solar Energy

678

2012;86(2):792–803.

679

[14]

M AN U

677

Li X, Dai YJ, Wang RZ. Performance investigation on solar thermal conversion of

a conical cavity receiver employing a beam-down solar tower concentrator. Solar

681

Energy 2015;114:134–51.

682

[15]

TE D

680

Dai Y, Li X, Zhou L, Ma X, Wang R. Comparison-based optical study on a

point-line-coupling-focus system with linear Fresnel heliostats. Optics express

684

2016;24(10):A966–A73.

686 687

[16]

AC C

685

EP

683

Puig-Arnavat M, Tora EA, Bruno JC, Coronas A. State of the art on reactor designs

for solar gasification of carbonaceous feedstock. Solar Energy 2013;97:67–84. [17]

Müller F, Poživil P, van Eyk PJ, Villarrazo A, Haueter P, Wieckert C. A pressurized

688

high-flux solar reactor for the efficient thermochemical gasification of carbonaceous

689

feedstock. Fuel 2017;193:432–43.

690

[18]

Gokon N, Izawa T, Kodama T. Steam gasification of coal cokes by internally

32

ACCEPTED MANUSCRIPT 691

circulating fluidized-bed reactor by concentrated Xe-light radiation for solar syngas

692

production. Energy 2015;79:264–72.

693

[19]

Kruesi M, Jovanovic ZR, dos Santos EC, Yoon HC, Steinfeld A. Solar-driven

steam-based gasification of sugarcane bagasse in a combined drop-tube and

695

fixed-bed reactor–Thermodynamic, kinetic, and experimental analyses. biomass and

696

bioenergy 2013;52:173–83. [20]

SC

697

RI PT

694

Ermanoski I, Siegel NP, Stechel EB. A new reactor concept for efficient

solar-thermochemical fuel production. Journal of Solar Energy Engineering

699

2013;135(3):031002.

700

[21]

M AN U

698

Gokon N, Ono R, Hatamachi T, Liuyun L, Kim H-J, Kodama T. CO2 gasification of

coal cokes using internally circulating fluidized bed reactor by concentrated Xe-light

702

irradiation for solar gasification. International Journal of Hydrogen Energy

703

2012;37(17):12128–37. [22]

Kodama T, Gokon N, Enomoto S-i, Itoh S, Hatamachi T. Coal coke gasification in a

EP

704

TE D

701

windowed solar chemical reactor for beam-down optics. Journal of Solar Energy

706

Engineering 2010;132(4):041004.

707

[23]

AC C

705

Wieckert C, Obrist A, Zedtwitz Pv, Maag G, Steinfeld A. Syngas production by

708

thermochemical gasification of carbonaceous waste materials in a 150 kWth

709

packed-bed solar reactor. Energy & Fuels 2013;27(8):4770–6.

710 711

[24]

Piatkowski N, Wieckert C, Steinfeld A. Experimental investigation of a

packed-bed solar reactor for the steam-gasification of carbonaceous feedstocks. Fuel

33

ACCEPTED MANUSCRIPT

713 714 715

Processing Technology 2009;90(3):360–6. [25]

Piatkowski N, Steinfeld A. Solar-driven coal gasification in a thermally irradiated

packed-bed reactor. Energy & Fuels 2008;22(3):2043–52. [26]

Piatkowski N, Steinfeld A. Solar gasification of carbonaceous waste feedstocks in

RI PT

712

a packed-bed reactor—Dynamic modeling and experimental validation. AIChE

717

Journal 2011;57(12):3522–33.

718

[27]

SC

716

Kaniyal AA, van Eyk PJ, Nathan GJ. Dynamic Modeling of the Coproduction of

Liquid Fuels and Electricity from a Hybrid Solar Gasifier with Various Fuel Blends.

720

Energy & Fuels 2013;27(6):3556–69.

721

[28]

M AN U

719

Kaniyal AA, van Eyk PJ, Nathan GJ, Ashman PJ, Pincus JJ. Polygeneration of

Liquid Fuels and Electricity by the Atmospheric Pressure Hybrid Solar Gasification of

723

Coal. Energy & Fuels 2013;27(6):3538–55.

724

[29]

TE D

722

Muroyama A, Shinn T, Fales R, Loutzenhiser PG. Modeling of a

Dynamically-Controlled Hybrid Solar/Autothermal Steam Gasification Reactor.

726

Energy & Fuels 2014;28(10):6520–30. [30]

AC C

727

EP

725

Van Eyk PJ, Ashman PJ, Nathan GJ. The effect of high-flux solar irradiation on the

728

gasification of coal in a hybrid entrained-flow reactor. Energy & Fuels

729

2016;30(6):5138‒47.

730

[31]

Kalina J. Integrated biomass gasification combined cycle distributed generation

731

plant with reciprocating gas engine and ORC. Applied Thermal Engineering

732

2011;31:2829‒40.

34

ACCEPTED MANUSCRIPT 733

[32]

Skorek-Osikowska A, Bartela Ł, Kotowicz J, Sobolewski A, Iluk T, Remiorz L. The

influence of the size of the CHP (combined heat and power) system integrated with a

735

biomass fueled gas generator and piston engine on the thermodynamic and

736

economic effectiveness of electricity and heat generation. Energy 2014;67:328–40.

737

[33]

RI PT

734

Kaushik SC, Arora A. Energy and exergy analysis of single effect and series flow

double effect water–lithium bromide absorption refrigeration systems. Int. J. Refrig.

739

2009;32:1247–58. [34]

Wang J, Mao T, Sui J, Jin H. Modeling and performance analysis of CCHP

M AN U

740

SC

738

741

(combined cooling, heating and power) system based on co-firing of natural gas and

742

biomass gasification gas. Energy 2015;93:801–15.

745 746 747

Bai Z, Liu Q, Lei J, Hong H, Jin H. New solar-biomass power generation system

integrated a two-stage gasifier. Applied Energy 2017;194:310‒9. [36]

TE D

744

[35]

Yun KT, Cho H, Luck R, Mago PJ. Modeling of reciprocating internal combustion

engines for power generation and heat recovery. Applied Energy 2013;102:327–35. [37]

EP

743

Gao P, Li W, Cheng Y, Tong Y, Dai Y, Wang R. Thermodynamic performance

assessment of CCHP system driven by different composition gas. Applied Energy

749

2014;136:599–610.

750

[38]

AC C

748

Gao P, Dai Y, Tong Y, Dong P. Energy matching and optimization analysis of waste

751

to energy CCHP (combined cooling, heating and power) system with exergy and

752

energy level. Energy 2015;79:522–35.

753

[39]

Patuzzi F, Prando D, Vakalis S, Rizzo AM, Chiaramonti D, Tirler W, et al.

35

ACCEPTED MANUSCRIPT 754

Small-scale biomass gasification CHP systems: Comparative performance

755

assessment and monitoring experiences in South Tyrol (Italy). Energy

756

2016;112:285–93. [40]

Zheng CY, Wu JY, Zhai XQ, Yang G, Wang RZ. Experimental and modeling

RI PT

757

investigation of an ICE (internal combustion engine) based micro-cogeneration

759

device considering overheat protection controls. Energy 2016;101:447–61.

760

[41]

SC

758

Marbe Å, Harvey S, Berntsson T. Biofuel gasification combined heat and

power—new implementation opportunities resulting from combined supply of

762

process steam and district heating. Energy 2004;29(8):1117–37.

765 766 767

Channiwala SA, Parikh PP. A unified correlation for estimating HHV of solid,

liquid and gaseous fuels, Fuel 2002;81:1051–63. [43]

Bilgen S, Kaygusuz K, Sari A. Second law analysis of various types of coal and

TE D

764

[42]

woody biomass in Turkey. Energy Sources 2004;26:1083–94. [44]

Müller R, Steinfeld A. Band-approximated radiative heat transfer analysis of a

EP

763

M AN U

761

solar chemical reactor for the thermal dissociation of zinc oxide. Solar Energy

769

2007;81(10):1285–94.

770 771 772 773

[45]

AC C

768

Yoon HC, Cooper T, Steinfeld A. Non-catalytic autothermal gasification of woody

biomass. International Journal of Hydrogen Energy 2011;36(13):7852–60. [46]

TRNSYS – Official Website – University of Wisconsin-Madison 2016.

http://sel.me.wisc.edu/trnsys/[Accessed 1 May 2016].

36

ACCEPTED MANUSCRIPT Table 1. Ultimate and proximate analyses of the redwood feedstock. Feedstock

Redwood pellets

47.11

H (wt %)

5.47

O (wt %)

45.0

N (wt %)

0.50

SC

C (wt %)

RI PT

Ultimate analysis

0.10

M AN U

S (wt %) Proximate analysis Moisture (wt %)

9.30

Volatile (wt %)

87.9 1.40

TE D

Ash (wt %) Fixed carbon (wt %)

LHVf (MJ/kg)

EP

HHVf (MJ/kg)

AC C

Chemical composition

1.40 18.20 17.00 CH1.39O0.72

ACCEPTED MANUSCRIPT Table 2. Nominal design parameters of the SAHG-CCHP system. Unit Value

Air inlet temperature (ܶଵ,ୟ୧୰ )

K

400

Steam inlet temperature (ܶଵ,ୱ୲ୣୟ୫ )

K

400

Feedstock temperature (ܶଶ )

RI PT

Parameter

K

298

K

1180

K

498

K

734

Outlet temperature of exhaust gas from HE-4 (ܶ଺ )

K

421

Outlet temperature of exhaust gas from HE-3 (ܶ଻ )

K

395

Outlet temperature of product syngas from SAHG (ܶଷ )

M AN U

Outlet temperature of exhaust gas from ICE (ܶହ )

SC

Outlet temperature of product syngas from HE-1 (ܶସ )

426/417

Inlet/outlet temperature (ܶଵ଴ /ܶଵଵ) of cooling water for DEAC

K

303/308

K

285/280

TE D

Inlet/outlet temperature of pressure hot water for DEAC (଼ܶ /ܶଽ ) K

Inlet/outlet temperature (ܶଵଷ /ܶଵଶ) of chilled water for DEAC COP of the DEAC

1.21 K

298

Water outlet temperature from HE-2 (ܶଵହ )

K

332

Water outlet temperature from HE-3 (ܶଵ଺ )

K

334

Outlet temperature of jacket water from ICE (ܶଵ଻ )

K

347

Inlet temperature of jacket water from ICE (ܶଵ଼ )

K

333

Outlet temperature of pressure hot water from HE-4 (ܶଵଽ )

K

420

AC C

EP

Water temperature (ܶଵସ )

Effectiveness of heat exchangers (ߟHE )

0.8

ACCEPTED MANUSCRIPT Table 3. The determined convective heat transfer coefficients of solar gasifier [26].

Unit

Value

ℎଵିଵ଴,conv

W m−2 K

17

ℎଵଷିଵ଴,conv

W m−2 K

17

ℎଶିଷ,conv

W m−2 K

ℎଷିସ,conv

W m−2 K

ℎ଺ିଷ,conv

W m−2 K

ℎଵଵିଷ,conv

W m−2 K

ℎ଻ି଼,conv

W m−2 K

RI PT

Convective heat transfer coefficients

1.5

AC C

EP

TE D

M AN U

SC

3.0 6.0 6.0 15

ACCEPTED MANUSCRIPT Table 4. Objects participating in radiative heat transfer within the upper and lower cavities.

Radiative flux items

Objects participating in radiative heat

1, 9, 13

‫ݍ‬ଶ,rad

2, 4, 6, 11

‫ݍ‬ସ,rad

2, 4, 6, 11

‫ݍ‬ହ,rad

5, 8

‫଺ݍ‬,rad

M AN U

2, 4, 6, 11

‫଻ݍ‬,rad

7, 8

‫ݍ‬ଵଶ,rad

8, 12

‫ݍ‬ଵଷ,rad

EP

TE D

1, 9, 13

AC C

‫ݍ‬ଵହ,rad

RI PT

‫ݍ‬ଵ,rad

SC

transfer

8, 15

ACCEPTED MANUSCRIPT

Solar power over the aperture (kW)

200

Steam-to-feedstock ratio M

0.75

Oxygen-to-feedstock ratio m

0

Height of the upper cavity (m)

0.65

Height of the lower cavity (m)

0.85

Aperture diameter (m)

0.55

Bed height (m)

0.5

SC

Value

M AN U

Parameters

RI PT

Table 5. Simulated input parameters of the SAHG in the fully solar steam mode [26].

Insulation thickness (m)

0.18

1.5 (case-1), 2.0 (case-2), 2.5 (case-3)

AC C

EP

TE D

Bed diameter (m)

ACCEPTED MANUSCRIPT

Table 6. Comparison of the model predictions and reported data. Mean cold gas

Model

Data [26]

Relative error (%)

ߟCG (case-1)

0.48

0.42

14.3

ߟCG (case-2)

0.65

0.58

ߟCG (case-3)

0.76

0.69

AC C

EP

TE D

M AN U

SC

RI PT

efficiency

12.1 10.1

ACCEPTED MANUSCRIPT Table 7. Input parameters for simulation. Value

Concentration ratio C over the aperture

1430

Reaction rate (feeding rate) ݉ሶ୤ (mol/s)

0.5

RI PT

Parameter

Height of the upper cavity (m)

0.65

Height of the lower cavity (m)

0.85

0.55

SC

Aperture diameter (m) Bed height (m)

0.5

0.18

Bed diameter (m) Quartz glass thickness (mm) Cold surface height (mm)

Emissivity of emitter Conductivity of SiC (W/m K)

TE D

Emissivity of insulation

M AN U

Insulation thickness (m)

EP

Emissivity of industrial sludge

1.5 3 5 0.6 0.88 25 [25] 0.94 130

Nominal cooling capacity of the DEAC (kW)

120

Effectiveness of heat exchangers ߟHE

0.8

AC C

Nominal electrical power ߟ୬ୣ (kW)

ACCEPTED MANUSCRIPT Table 8. Comparison of the upgraded PER by solar for redwood.

1 2 3

PERSAHG

Increment (%)

0.898

14.8

0.921

17.8

0.902

4

0.897

5

0.890

6

0.894

8 9

12

AC C

Yearly average

EP

TE D

10 11

15.3 14.7

M AN U

0.782

7

RI PT

PERautothermal

SC

Month

13.8 14.3

0.902

15.3

0.891

13.9

0.883

12.9

0.887

13.4

0.879

12.4

0.871

11.4

0.893

14.2

ACCEPTED MANUSCRIPT

Figures

Fig. 1. Schematic of a solar/autothermal hybrid gasification CCHP system.

RI PT

Fig. 2. Equilibrium compositions as a function of temperature for stoichiometric system of redwood pellets at 1 bar: (a) M = 0.28, m = 0, and (b) M = 0, m = 0.329. Fig. 3. Specific enthalpy change ∆‫ܪ‬ோ of redwood pellets with gas reactants (Eq. 1) fed at 1 bar and 400 K and feedstock fed at 1 bar and 298 K based on the equilibrium composition.

SC

Fig. 4. Schematic of thermo-physical model of the SAHG.

M AN U

Fig. 5. Reaction temperature as a function of M and m in the autothermal and solar/hybrid modes: (a) the autothermal mode, (b) the 300 kW solar/hybrid mode. Fig. 6. Cold gas efficiency as a function of M and m in the autothermal and solar/hybrid modes: (a) the autothermal mode, and (b) the 300 kW solar/hybrid mode. Fig. 7. Lower heating value of product syngas as a function of M and m in the autothermal and solar/hybrid modes: (a) the autothermal mode, and (b) the 300 kW solar/hybrid mode.

TE D

Fig. 8. The optimized parameters as a function of the normalized solar power input (qsolar,peak = 300 kW): (a) the optimal steam-to-feedstock and oxygen-to-feedstock ratios, and (b) the corresponding cold gas efficiency and solar upgraded ratio.

EP

Fig. 9. Mole fraction based on the optimized ratios under various solar irradiances: (a) cold syngas composition, and (b) mole flow rates of hot and cold syngas as well as reaction temperature.

AC C

Fig. 10. Direct normal irradiances in Singapore: (a) two consecutive days, and (b) whole year. Fig. 11. Mole flow rates, solar upgraded ratio, and lower heating value of product syngas converted from redwood as functions of local time: (a) mole flow rates, and (b) solar upgraded ratio and lower heating value of syngas. Fig. 12. Mole flow rates of syngas species as a function of local time: (a) N2, (b) CO2, (c) H2, and (d) CO. Fig. 13. Cooling, heat and power, and primary energy ratio as a function of local time. Fig. 14. Monthly energy distribution of the SAHG-CCHP system.

M AN U

SC

RI PT

ACCEPTED MANUSCRIPT

AC C

EP

TE D

Fig. 1. Schematic of a solar/autothermal hybrid gasification CCHP system.

SC

RI PT

ACCEPTED MANUSCRIPT

M AN U

Fig. 2. Equilibrium compositions as a function of temperature for stoichiometric system of

AC C

EP

TE D

redwood pellets at 1 bar: (a) M = 0.28, m = 0, and (b) M = 0, m = 0.329.

SC

RI PT

ACCEPTED MANUSCRIPT

M AN U

Fig. 3. Specific enthalpy change ∆‫ܪ‬ோ of redwood pellets with gas reactants (Eq. 1) fed at 1 bar

AC C

EP

TE D

and 400 K and feedstock fed at 1 bar and 298 K based on the equilibrium composition.

EP

TE D

M AN U

SC

RI PT

ACCEPTED MANUSCRIPT

AC C

Fig. 4. Schematic of thermo-physical model of the SAHG.

SC

RI PT

ACCEPTED MANUSCRIPT

Fig. 5. Reaction temperature as a function of M and m in the autothermal and solar/hybrid

AC C

EP

TE D

M AN U

modes: (a) the autothermal mode, (b) the 300 kW solar/hybrid mode.

SC

RI PT

ACCEPTED MANUSCRIPT

Fig. 6. Cold gas efficiency as a function of M and m in the autothermal and solar/hybrid modes:

AC C

EP

TE D

M AN U

(a) the autothermal mode, and (b) the 300 kW solar/hybrid mode.

SC

RI PT

ACCEPTED MANUSCRIPT

Fig. 7. Lower heating value of product syngas as a function of M and m in the autothermal and

AC C

EP

TE D

M AN U

solar/hybrid modes: (a) the autothermal mode, and (b) the 300 kW solar/hybrid mode.

SC

RI PT

ACCEPTED MANUSCRIPT

M AN U

Fig. 8. The optimized parameters as a function of the normalized solar power input (qsolar,peak = 300 kW): (a) the optimal steam-to-feedstock and oxygen-to-feedstock ratios, and (b) the

AC C

EP

TE D

corresponding cold gas efficiency and solar upgraded ratio.

SC

RI PT

ACCEPTED MANUSCRIPT

Fig. 9. Mole fraction based on the optimized ratios under various solar irradiances: (a) cold

M AN U

syngas composition, and (b) mole flow rates of hot and cold syngas as well as reaction

AC C

EP

TE D

temperature.

ACCEPTED MANUSCRIPT (b) 1000

600 400 200 0 0

RI PT

IDNI (W/m2)

800

2000

4000

6000

8000

SC

Local time (h)

AC C

EP

TE D

M AN U

Fig. 10. Direct normal irradiances in Singapore: (a) two consecutive days, and (b) whole year.

SC

RI PT

ACCEPTED MANUSCRIPT

Fig. 11. Mole flow rates, solar upgraded ratio, and lower heating value of product syngas

M AN U

converted from redwood as functions of local time: (a) mole flow rates, and (b) solar upgraded

AC C

EP

TE D

ratio and lower heating value of syngas.

EP

TE D

M AN U

SC

RI PT

ACCEPTED MANUSCRIPT

AC C

Fig. 12. Mole flow rates of syngas species as a function of local time: (a) N2, (b) CO2, (c) H2, and (d) CO.

M AN U

SC

RI PT

ACCEPTED MANUSCRIPT

AC C

EP

TE D

Fig. 13. Cooling, heat and power, and primary energy ratio as a function of local time.

M AN U

SC

RI PT

ACCEPTED MANUSCRIPT

AC C

EP

TE D

Fig. 14. Monthly energy distribution of the SAHG-CCHP system.

ACCEPTED MANUSCRIPT

Highlights A solar/autothermal hybrid gasifier was introduced and studied thermodynamically.




The optimal steam-to-feedstock and oxygen-to-feedstock ratios were obtained.




A zero-dimensional steady model based on equilibrium composition was developed.




Operation behaviors of the CCHP system with the hybrid gasifier was investigated.




A yearly average primary energy ratio was upgraded by 14.2% in Singapore.

AC C

EP

TE D

M AN U

SC

RI PT