Advanced Gasification of Biomass/Waste for …953814/...i Abstract With the current trend of CO 2 mitigation in process industries, the primary goal of this thesis is to promote biomass

Advanced Gasification of Biomass/Waste for Substitution of Fossil Fuels in Steel Industry

Heat Treatment Furnaces

Duleeka Sandamali Gunarathne

Doctoral Dissertation

2016

KTH Royal Institute of Technology School of Industrial Engineering and Management Department of Materials Science and Engineering

Unit of Processes SE-100 44 Stockholm, Sweden

___________________________________________________________________________

Akademisk avhandling som med tillstånd av Kungliga Tekniska Högskolan i Stockholm framlägges

för offentlig granskning för avläggande av teknologie doktorsexamen, Tisdag den 27 September

2016, kl. 10:00 i Sal F3, Lindstedtvägen 26, Kungliga Tekniska Högskolan, Stockholm.

ISBN 978-91-7729-053-7

Duleeka Sandamali Gunarathne. Advanced Gasification of Biomass/Waste for

Substitution of Fossil Fuels in Steel Industry Heat Treatment Furnaces

KTH Royal Institute of Technology School of Industrial Engineering and Management Department of Materials Science and Engineering Unit of Processes SE-100 44 Stockholm, Sweden ISBN 978-91-7729-053-7

Copyright Duleeka Sandamali Gunarathne

i

Abstract

With the current trend of CO2 mitigation in process industries, the primary goal of this thesis is to

promote biomass as an energy and reduction agent source to substitute fossil sources in the steel

industry. The criteria for this substitution are that the steel process retains the same function and

the integrated energy efficiency is as high as possible.

This work focuses on advanced gasification of biomass and waste for substitution of fossil fuels in

steel industry heat treatment furnaces. To achieve this, two approaches are included in this work.

The first investigates the gasification performance of pretreated biomass and waste experimentally

using thermogravimetric analysis (TGA) and a pilot plant gasifier. The second assesses the

integration of the advanced gasification system with a steel heat treatment furnace.

First, the pyrolysis and char gasification characteristics of several pretreated biomass and waste

types (unpretreated biomass, steam-exploded biomass, and hydrothermal carbonized biomass)

were analyzed with TGA. The important aspects of pyrolysis and char gasification of pretreated

biomass were identified.

Then, with the objective of studying the gasification performance of pretreated biomass,

unpretreated biomass pellets (gray pellets), steam-exploded biomass pellets (black pellets), and two

types of hydrothermal carbonized biomass pellets (spent grain biocoal and horse manure biocoal)

were gasified in a fixed bed updraft gasifier with high-temperature air/steam as the gasifying agent.

The gasification performance was analyzed in terms of syngas composition, lower heating value

(LHV), gas yield, cold gas efficiency (CGE), tar content and composition, and particle content and

size distribution. Moreover, the effects on the reactions occurring in the gasifier were identified

with the aid of temperature profiles and gas ratios.

Further, the interaction between fuel residence time in the bed (bed height), conversion, conversion

rate/specific gasification rate, and superficial velocity (hearth load) was revealed. Due to the effect

of bed height on the gasification performance, the bed pressure drop is an important parameter

related to the operation of a fixed bed gasifier. Considering the limited studies on this relationship,

an available pressure drop prediction correlation for turbulent flow in a bed with cylindrical pellets

was extended to a gasifier bed with shrinking cylindrical pellets under any flow condition. Moreover,

simplified graphical representations based on the developed correlation, which could be used as an

effective guide for selecting a suitable pellet size and designing a grate, were introduced.

Then, with the identified positive effects of pretreated biomass on the gasification performance,

the possibility of fuel switching in a steel industry heat treatment furnace was evaluated by effective

ii

integration with a multi-stage gasification system. The performance was evaluated in terms of

gasifier system efficiency, furnace efficiency, and overall system efficiency with various heat

integration options. The heat integration performance was identified based on pinch analysis.

Finally, the efficiency of the co-production of bio-coke and bio-H2 was analyzed to increase the

added value of the whole process.

It was found that 1) the steam gasification of pretreated biomass is more beneficial in terms of the

energy value of the syngas, 2) diluting the gasifying agent and/or lowering the agent temperature

compensates for the ash slagging problem in biocoal gasification, 3) the furnace efficiency can be

improved by switching the fuel from natural gas (NG) to syngas, 4) the gasifier system efficiency

can be improved by recovering the furnace flue gas heat for the pretreatment, and 5) the co-

production of bio-coke and bio-H2 significantly improves the system efficiency.

Keywords: Biomass; Pretreatment; Gasification; Pressure drop; Steel industry; Fuel switch; Energy

efficiency

iii

Acknowledgements

I would like to thank my supervisors, Docent Weihong Yang and Prof. Wlodzimierz Blasiak, for

giving me the opportunity to join the Energy and Furnace Technology research group and offering

me continuous guidance throughout my PhD studies. In addition, I would like to thank Dr. Jan

Chmielewski and all my colleagues for help with experimental work and for sharing enjoyable times.

I would also like to thank all the collaborators, including Prof. Thomas Kolb, Dr. Sabine Fleck and

Mr. Andreas Mueller from Karlsruhe Institute of Technology, Dr. Magnus Pettersson from

Höganäs AB, and Mr. Rolf Ljunggren and Mr. Marko Amovic from Cortus Energy AB. The

Swedish Steel Producers’ Association (Jernkontoret) is also acknowledged. All the biomass

providers: Boson Energy SA, Zilkha Biomass Energy, and AVA-CO2, are highly appreciated.

I am grateful to the European Commission for funding a scholarship for my PhD research under

the Erasmus Mundus AREAS Program. I am also thankful to Jernkontoret for an additional

scholarship through Prytziska fonden nr 2. The financial providers KIC InnoEnergy (under the

project xGaTe) and Swedish Energy Agency –Energimyndigheten- (under the project Probiostål)

are gratefully acknowledged.

I am grateful to all my Sri Lankan friends living in Sweden, who made me feel at home. Above all,

I am grateful to my family for their endless love and care, making me comfortable with my studies.

Duleeka Sandamali Gunarathne

2016-06-12 Stockholm, Sweden

iv

List of Papers in the Thesis

I. Gunarathne, D. S., Mueller, A., Fleck, S., Kolb, T., Chmielewski, J. K., Yang, W., and

Blasiak, W. (2014). Gasification characteristics of steam exploded biomass in an updraft

pilot scale gasifier. Energy, 71, 496-506.

II. Gunarathne, D. S., Mueller, A., Fleck, S., Kolb, T., Chmielewski, J. K., Yang, W., and

Blasiak, W. (2014). Gasification characteristics of hydrothermal carbonized biomass in an

updraft pilot-scale gasifier. Energy and Fuels, 28(3), 1992-2002.

III. Gunarathne, D. S., Chmielewski, J. K., Yang, W., and Blasiak, W. Performance of high

temperature air/steam gasification of hydrothermal carbonized biomass. 22nd European

Biomass Conference and Exhibition (EUBC&E 2014). Hamburg, Germany. June 23-26,

2014.

IV. Gunarathne, D. S., Chmielewski, J. K., and Yang, W. (2014). Pressure drop prediction of

a gasifier bed with cylindrical biomass pellets. Applied Energy, 113, 258-266.

V. Gunarathne, D. S., Mellin, P., Yang, W., Pettersson, M., and Ljunggren, R. (2016).

Performance of an effectively integrated biomass multi-stage gasification system and a steel

industry heat treatment furnace. Applied Energy, 170, 353-361.

Contribution Statement:

In papers I–III, I participated in the gasification experiments, analyzed all of the results, and wrote

the manuscripts;

In paper IV, I performed the theoretical studies and wrote the manuscript;

In paper V, I developed the models, performed the simulations, and wrote the manuscript.

v

List of Papers not in the Thesis

I. Gunarathne, D. S., Mellin, P., Yang, W., Pettersson, M., and Ljunggren, R. System

integration of the heat treatment furnace in steel plant with biomass gasification process.

Nordic Flame Days 2015. Copenhagen, Denmark. October 6-7, 2015.

II. Gunarathne, D. S., Cuvilas, C. A., Li, J., Yang, W., and Blasiak, W. Biomass pretreatment

for large percentage biomass co-firing. 12th International Conference on Boiler Technology

(ICBT 2014). Szczyrk, Poland. October 21-24, 2014.

III. Zhou, C., Stuermer, T., Gunarathne, R., Yang, W., and Blasiak, W. (2014). Effect of

calcium oxide on high-temperature steam gasification of municipal solid waste. Fuel, 122,

36-46.

IV. Gunarathne, D. S., Chmielewski, J. K., and Yang, W. High temperature air/steam

gasification of steam exploded biomass. Finnish-Swedish Flame Days 2013. Jyväskylä,

Finland. April 17-18, 2013.

vi

Contents 1 INTRODUCTION ............................................................................................................... 1

1.1 Introduction to the dissertation .................................................................................... 1

1.2 Objectives ......................................................................................................................... 3

1.3 Structure of the dissertation ........................................................................................... 4

2 BACKGROUND .................................................................................................................. 6

2.1 Biomass pretreatment ..................................................................................................... 8

2.2 Biomass gasification and gasifier types ....................................................................... 10

2.3 Advanced gasification technologies ............................................................................ 12

2.4 Gasification of pretreated biomass ............................................................................. 14

2.5 Integration of a steel industry furnace with biomass gasification ........................... 15

3 METHODOLOGY ............................................................................................................ 16

3.1 Experimental basis ........................................................................................................ 16

3.1.1 Feedstock materials............................................................................................... 16

3.1.2 TGA experiments ................................................................................................. 18

3.1.3 Gasification experiments ..................................................................................... 20

3.2 Application-oriented process simulations .................................................................. 24

3.2.1 System description ................................................................................................ 24

3.2.2 Scenario description.............................................................................................. 26

3.2.3 Process models ...................................................................................................... 26

3.2.4 Performance analysis ............................................................................................ 31

4 BIOMASS CHARACTERIZATION .............................................................................. 33

4.1 Mass loss behavior ........................................................................................................ 33

4.2 Pyrolysis behavior ......................................................................................................... 34

4.3 Char gasification behavior ............................................................................................ 36

vii

5 GASIFICATION PERFORMANCE IN UPDRAFT HTAG .................................... 38

5.1 Gasification of unpretreated and steam-exploded biomass .................................... 38

5.1.1 Temperature distribution ..................................................................................... 38

5.1.2 Syngas composition .............................................................................................. 39

5.1.3 LHV, gas yield, and efficiency ............................................................................. 40

5.1.4 Syngas purity .......................................................................................................... 42

5.2 Gasification of hydrothermal carbonized biomass ................................................... 43




5.2.4 Syngas purity .......................................................................................................... 46

5.3 Gasification of hydrothermal carbonized biomass with controlling measures for

ash slagging ............................................................................................................................... 47




5.3.4 Syngas purity .......................................................................................................... 51

6 PARAMETERS OF FIXED BED GASIFICATION .................................................. 54

6.1 Bed height, conversion, and conversion rate ............................................................ 54

6.2 Gasifier design parameters: Specific gasification rate and superficial velocity ...... 55

6.3 Bed pressure drop ......................................................................................................... 57

7 INTEGRATION OF BIOMASS GASIFICATION PROCESS WITH THE STEEL

INDUSTRY ................................................................................................................................. 60

7.1 Application of syngas in the furnace .......................................................................... 60

7.1.1 Comparison of experimental and modeling results .......................................... 60

7.1.2 Gasifier system performance ............................................................................... 61

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7.1.3 Heat integration performance ............................................................................. 63

7.1.4 Furnace performance ........................................................................................... 64

7.1.5 Energy balance and the overall system performance ....................................... 68

7.2 Co-production of bio-coke and bio-H2 ...................................................................... 69

8 CONCLUSIONS ................................................................................................................. 70

9 FUTURE WORK ................................................................................................................ 74

10 REFERENCES ................................................................................................................ 75

ix

Nomenclature

Abbreviations

ASTM American Society for Testing and Materials

BTEX Benzene, toluene, ethylbenzene, and xylenes

CGE Cold gas efficiency

DTG Derivative thermogravimetry

ER Equivalence ratio

GC Gas chromatography

GCC Grand composite curve

GHG Greenhouse gas

HGI Hardgrove grindability index

HHV Higher heating value

HiTAC High-temperature air combustion

HTAG High-temperature agent gasification

HTC Hydrothermal carbonization

LHV Lower heating value

LPG Liquefied petroleum gas

LPI Low-pressure impactor

NG Natural gas

PAH Polycyclic aromatic hydrocarbon

PSA Pressure swing adsorption

SEM Scanning electron microscope

SNG Synthetic natural gas

SPA Solid phase adsorption

TGA Thermogravimetric analysis

x

Symbols

𝐴 Cross sectional area m2

𝐴′ Modified Ergun constant for the viscous term -

𝐵′ Modified Ergun constant for the inertial term -

𝐶 Carbon content of biomass %

𝐶𝑡𝑎𝑟 Concentration of tar species g/Nm3

𝐶𝑝,𝑖 Constant pressure heat capacity of the ith component J/kg °C

𝑐 Length of a cylindrical particle at any time m

𝑐0 Initial length of a cylindrical particle m

𝐷𝑒 Equivalent diameter of a tortuous passage m

𝑑 Diameter of a cylindrical particle at any time m

𝑑0 Initial diameter of a cylindrical particle m

𝑑𝑝 Equivalent particle diameter m

𝐸𝐶ℎ Chemical energy MJ

𝐸𝑆 Sensible energy MJ

�̇� Fuel feed rate kg/min

�̇�𝑐 Fuel consumption rate kg/min

�̇� Gas flow rate m3/s

𝐺𝑚𝑖𝑥 Gibbs energy for a mixture of ideal gases J

�̅�𝑖,𝑇 Molar Gibbs energy of the ith component at T J/mol

�̅�𝑖,𝑇0 Molar Gibbs energy of the ith component at T and

standard pressure

J/mol

ℎ Shrinking length m

ℎ𝑓0 Standard enthalpy of formation J/mol

𝛥ℎ̅ Enthalpy change J/mol

𝐾1 Constant for the inertial term 1/m

𝐾2 Constant for the viscous term 1/m

𝐾3 Constant for Rep 1/m

𝐾𝑡 Constant depending on the roughness of the particle

and packing tortuosity

-

𝐿 Bed height m

xi

𝐿𝐻 Hearth load m3/cm2 h

𝐿𝐻2𝑂 Latent heat of vaporization J/kg

𝐿𝐻𝑉𝑓𝑢𝑒𝑙 LHV of fuel MJ/kg

𝐿𝐻𝑉𝐺 LHV of pyrolysis gas (excluding tar) MJ/kg

𝐿𝐻𝑉𝑖 LHV of the ith component in the gas mixture MJ/Nm3

𝑙 Equivalent length of a tortuous passage m

𝑀𝑖 Molecular weight of the ith component g/mol

𝑚 Mass g

𝑚0 Initial mass g

𝑚𝑓 Final mass g

�̇�𝑖 Mass flow rate of the ith component kg/s

𝑁𝑖 Number of moles of the ith component mol

�̇�𝑖 Inlet molar flow rate mol/s

�̇�𝑜 Outlet molar flow rate mol/s

𝑂2/𝐹 Oxygen to feed ratio mol/mol

𝑃0 Standard pressure Pa

𝑃𝑖 Pressure Pa

𝑃𝑠𝑣(𝑇) Saturation vapor pressure at T atm

∆𝑃 Pressure drop Pa

�̇�𝐿 Rate of heat loss J/s

𝑅 Universal gas constant J/mol °C

𝑅𝑒𝑝 Particle Reynolds number -

𝑟 Conversion rate 1/min

𝑟𝑏ℎ Rate of change of bed height m/min

𝑟𝑠𝑔 Specific gasification rate kg/m2 h

𝑆/𝐶 Steam to carbon ratio mol/mol

𝑆𝑝 Particle surface area m2

𝑇 Temperature °C

𝑡𝑟 Residence time min

𝑢 Superficial velocity m/s

xii

�̇� Volumetric gas flow rate Nm3/min

𝑉𝑝 Particle volume m3

𝑣 Velocity of flow through a tortuous passage m/s

𝑋 Fuel conversion -

𝑥 Char conversion -

𝑥𝑖 Volume fraction of the ith species in the gas mixture -

𝑌 Volumetric gas yield Nm3/kg

Greek letters

𝛼 Residual mass ratio -

𝛼𝑖 Stoichiometric coefficients i = 1–8 -

𝛽𝑖 Stoichiometric coefficients i = 1–5 -

𝛾𝑖 Stoichiometric coefficients i = 1–7 -

𝜀 Porosity -

𝜂𝐹 Efficiency of the furnace %

𝜂𝐺 Efficiency of the gasifier system %

𝜂𝐼 Efficiency of the integrated system %

𝜇 Viscosity Ns/m2

𝜌 Density kg/m3

𝜌𝑏 Bulk density kg/m3

𝛷 Sphericity -

1

CHAPTER I

1 INTRODUCTION

1.1 Introduction to the dissertation

The iron and steel industry is the largest industrial source of CO2 emissions in Sweden due

to its reliance on fossil fuels and reductants and the large volume of steel production. The

European Union has ambitious plans to reduce the CO2 emissions of today's best steel

processing routes by at least 50% by 2050 [1].

Improving the energy efficiency and reducing the dependency on fossil fuels are the main

strategies for reducing CO2 emissions. Replacing fossil sources with biomass is one area of

growing interest in the steel industry, with direct replacement of coal or coke with biomass

as the most common option [2, 3]. Substitution of the furnace fuel with biomass-derived

syngas is another option. The replacement of fossil fuels with renewable gaseous fuels

should address the following important issues:

Temperature: The fuel or combustion technology should provide a high enough

temperature.

Efficiency: The furnace should have the same or better efficiency.

Emissions: Under high-temperature operation, NOx emission should be minimized.

Impurities: If the heat transfer is direct, it should not affect the steel quality, and if

indirect firing is applied, the heat transfer surface should be tolerable.

Syngas produced by typical biomass gasification has a low heating value and hence requires

further processing to synthetic natural gas (SNG) for use in furnaces. An alternative

approach involves the improvement of biomass by means of pretreatment before

gasification, which enables the production of high calorific gas with a high adiabatic

temperature. As will be addressed in the thesis, this technology is advantageous in many

ways, for example:

Better performance in the gasification process.

2

Possibility of furnace flue gas heat recovery for pretreatment.

Co-production of valuable bio-based reduction agents to maximize the value chain.

In particular, for the iron and steel industry, pretreated biomass can be used for co-

production of coke, which makes the system more diverse. Further, as the production of

coke is an exothermic process, integration with the endothermic gasification process gives

a synergetic effect in terms of the energy balance.

As this thesis will show, the syngas produced from pretreated biomass results in lower flue

gas losses from the furnace, and hence improves the furnace efficiency. Moreover, steam

gasification results in a high content of H2 in the syngas, which makes the separation of

some H2 feasible, while maintaining the required adiabatic temperature. Coke gas with a

high content of H2 could also be combined with the H2 separation process. H2, which has

many end-use applications, could also be an alternative reduction agent for the iron and

steel making process.

By effective integration of biomass pretreatment, gasification, and the steel making process,

it would be possible to reduce CO2 emissions considerably. Figure 1-1 shows such a

concept of integration.

Figure 1-1. Conceptual integration of biomass pretreatment, gasification, and a

steel plant to maximize the biomass value chain

3

1.2 Objectives

The general objective of this work was to promote biomass as the energy and reduction

agents as a substitute for fossil sources in the steel industry, with the aim of reducing CO2

emissions, as well as achieving a potentially higher value chain by integrating the biomass

gasification process with steel plants.

The specific objective of this work was an assessment of advanced gasification of biomass

and waste for substitution of fossil fuels in steel industry heat treatment furnaces, with a

focus on incorporating pretreatment and gasification for biomass conversion. The

following sub-objectives were chosen to reflect the anticipated objective:

Identify the pyrolysis and char gasification behavior of pretreated biomass.

Experimentally determine the gasification performance of pretreated

biomass with air/steam as the gasifying agent.

Understand the effect of bed height on the gasification performance and

theoretically predict the pressure drop of a gasifier bed with cylindrical

pellets.

Evaluate the possibility of fuel switching in steel industry furnaces by

effective integration (with waste heat recovery) with biomass pretreatment

and gasification.

Explore the possibility of co-production of bio-based reduction agents, such

as bio-coke and bio-H2, from such an integrated system.

4

1.3 Structure of the dissertation

This thesis is organized into nine chapters. Chapter 1 provides a general overview, the

objectives, and the scope of the thesis. Chapter 2 provides a background review of the field

of study. The methodology used to achieve the objectives is discussed in Chapter 3.

Chapters 4–7 summarize and discuss the results, which are presented in detail in the

supplements. The overall conclusions are presented in Chapter 8, with recommendations

for future work in Chapter 9.

The content of this thesis can be divided into four main categories: thermogravimetric

analysis (TGA) study, gasification study, theoretical study, and system study, as

demonstrated in Figure 1-2.

Figure 1-2. Scope of the thesis, including the study area of each supplement

In general, supplements I–III examine the gasification process performance of pretreated

biomass. Based on the identified effect of bed height on such performance, supplement IV

develops methods to predict the bed induced pressure drop. Supplement V investigates the

application of the produced syngas in a heat treatment furnace.

5

Table 1-1 provides an overview of the supplements and their specific objectives.

Table 1-1. Overview of supplements

Supplement Title Objectives

I Gasification characteristics of

steam exploded biomass in an

updraft pilot scale gasifier

• Characterization of normal and steam-exploded biomass by

thermogravimetric analysis (TGA)

• Study of air/steam gasification behavior in terms of the

temperature profile, syngas composition, lower heating value

(LHV), gas yield, cold gas efficiency (CGE), and tar content

II Gasification characteristics of

hydrothermal carbonized

biomass in an updraft pilot-

scale gasifier

• Characterization of hydrothermal carbonized biomass by

TGA

• Study of air gasification behavior in terms of temperature

profile, syngas composition, LHV, gas yield, and CGE

• Identification of the effect of bed height on conversion and

conversion rate

• Study of main design parameters: specific gasification rate

and superficial gas velocity

III Performance of high

temperature air/steam

gasification of hydrothermal

carbonized biomass

• Study of air/steam gasification behavior of two types of

hydrothermal carbonized biomass in terms of temperature

profile, syngas composition, LHV, gas yield, CGE, tar, and

particle content

IV Pressure drop prediction of a

gasifier bed with cylindrical

biomass pellets

• Development of a correlation to predict bed pressure drop

for gasification of cylindrical pellets

• Introduction of a simplified graphical method to reduce the

calculation effort for pressure drop predictions

V Performance of an effectively

integrated biomass multi-stage

gasification system and a steel

industry heat treatment

furnace

• Study of an integrated system based on multi-stage biomass

gasification and the steel process

• Determination of the system-level energy and material

balances

• Assessment of different possibilities for waste heat

recovery based on individual and integrated efficiencies

6

CHAPTER II

2 BACKGROUND

Figure 2-1 shows the greenhouse gas (GHG) emissions from industrial processes and

energy use from the iron and steel industry in Sweden [4, 5]. The highest GHG emissions

related to process industries are recorded for metal production. Further, coal processing

has the highest emissions and is the source for more than 90% of CO2 emissions. Electricity

in Sweden mainly comes from renewable sources, and hence the next highest emissions are

from gaseous fuels such as natural gas (NG) and liquefied petroleum gas (LPG), which are

now extensively used to replace oil.

Figure 2-1. (left) GHG emissions from industrial processes; (right) energy use

in the iron and steel industry

Considering the above facts, biomass can be used in the steel industry through either

biomass-derived reduction agents or biomass-derived furnace fuels.

Biomass can be used as a reducing agent mainly in three forms: bio-char, bio-oil, and

syngas/SNG, which can be obtained by slow pyrolysis, fast pyrolysis, and gasification,

respectively [6]. Most studies have focused on bio-char injection, which is claimed to be

more feasible in terms of economy and operability. However, due to the insufficient

strength of bio-char compared with coke, it is limited to use as an injection fuel rather than

a coke substitute. It is reported that there is no technical restriction on the use of charcoal

up to 200 kg/ton of hot metal, resulting in very low coke consumption of around 260

kg/ton of hot metal in the blast furnace [7]. Alternatively, a limited content of biomass or

7

bio-char (up to 5%) can also be used as a raw material for coke production along with

coking coal without affecting the coke quality [8].

When bio-oil is considered as a blast furnace injectant, the high water and oxygen content

make it less suitable as a reducing agent. The reported coke replacement ratio is low (around

0.25), whereas that of char is high (around 1) [6]. However, the utilization of charcoal

byproducts from fast pyrolysis as a reducing agent would make this process interesting.

The CO- and H2-rich reducing gas (above 90% of the total volume) obtained from the

biomass oxygen/steam gasification process (after conditioning) can be used as a reducing

agent. A pellet reduction rate of 99% is reported at a reduction temperature of 1323 K after

30 min of reaction [9]. The limited capacity of commercial biomass gasifiers hinders the

effective use of such technology if not co-gasified with coal [6], although high-pressure

fluidized bed gasification with O2 followed by hot desulfurization has also been proposed

in the literature [10].

The use of biomass-derived fuels in steel heating furnaces is not yet common, although

coal-derived syngas is occasionally used for this purpose. The production of bio-syngas by

biomass gasification and substitution for LPG as the fuel in reheating furnaces has been

found to be technically feasible if used with an improved combustion technology, such as

high temperature air combustion (HiTAC) or oxy-fuel combustion [11]. If upgraded to

SNG, although not economically attractive, fuel substitution would reduce global CO2

emissions if the marginal biomass were used to produce transportation fuel, but not if it

were used to co-fire a coal power plant [12]. Replacing NG in iron ore pelletizing processes

with syngas derived from circulating fluidized bed gasification of biomass (air, oxygen, or

steam as the gasifying medium) reportedly increases the specific energy consumption, and

hence results in low efficiency. However, upgrading to SNG makes the process more

expensive [13].

In summary, bio-char and bio-coke will continue to be the main bio-based reduction agents

in the near future. To replace furnace fuels with syngas, more economical and efficient

methods that still comply with the process requirements are required.

8

2.1 Biomass pretreatment

Four main objectives are identified for biomass pretreatment:

Extend the use of waste/wet biomass sources

Make transportation, storage, and preprocessing economical

Extend the applications

Increase the gasification performance

Pretreatment plays a vital role when balanced against the associated costs and the gains.

The commonly used pretreatment technologies are mechanical pretreatment, steam

explosion, hydrothermal carbonization, torrefaction, and slow pyrolysis.

Mechanical pretreatment is the simplest and hence commonly applied method, which

incorporates size reduction and subsequent densification. This process improves both the

homogeneity and energy density, without any change of the structure of biomass.

A somewhat more extensive pretreatment, called hydrothermal carbonization (HTC), is

carried out in the presence of water at temperatures of 160–240 °C and pressures of 1–3.5

MPa, with residence times of up to several hours [14]. The reaction path of this process is

hydrolysis followed by dehydration, decarboxylation, condensation polymerization, and

aromatization. HTC has been applied to various feedstocks; however, it is most commonly

used for wet and waste biomass due to the water-based reaction atmosphere.

When the pretreatment process is carried out with high-pressure saturated steam at

temperatures of 160–240 °C and pressures of 0.7–4.8 MPa, followed by a sudden release

of the pressure, it is called steam explosion [14]. The main structural changes associated

with rupture of the cell walls are release of hemicellulose and altered lignin structures [15].

In addition to the improved homogeneity and energy density, hydrophobicity and

grindability are improved. Further, due to the improved hydrophobicity, mechanical

dewatering is possible, which reduces the energy requirement for thermal drying. For

example, a fivefold reduction in drying energy demand is reported after steam explosion

pretreatment [16]. The grinding energy demand is also reportedly reduced by 5–17% due

to steam explosion pretreatment [17]. Further, the high durability minimizes dusting,

9

reduces material losses during, handling and transportation, and reduce safety issues related

to dust explosions.

Torrefaction, slow pyrolysis, and carbonization can be categorized as thermal

pretreatment methods that are similar in principle, differing only in the operating

temperature ranges. During torrefaction, raw biomass is heated in an inert atmosphere at

temperatures of 200–300 °C for several minutes to several hours to produce improved solid

fuel. The slow pyrolysis process produces solid char at temperatures ranging from 300–700

°C [18], which can be termed as carbonization at higher temperatures due to the prioritized

production of solid carbon.

The temperature ranges of each pretreatment process are shown in Figure 2-2.

Figure 2-2. The temperature ranges of various pretreatment processes

As pretreatment improves the energy density of biomass, which is important for

transportation, pretreatment at the source rather than at the point of use (ex situ) can be

economical in some cases, especially for pretreatment at lower temperatures (HTC, steam

explosion, and mild torrefaction combined with mechanical pretreatment). However, for

moderate to high pretreatment temperature ranges, a considerable amount of energy is

10

available in the exhaust gas, which can be recovered if in situ pretreatment is applied (severe

torrefaction, pyrolysis, and carbonization).

For example, autothermal operation is reported at torrefaction temperatures of 270–280

°C (5–20 min residence time) for biomass with 50% moisture [19]. Thus, if this temperature

is exceeded, some energy will be wasted through extra torrefaction gases; hence, integrated

torrefaction becomes more feasible. A comparative study based on integrated and external

torrefaction at 300 °C reported that biomass to syngas efficiency can be increased from

63% to 86% by utilizing the volatiles in an integrated torrefaction process [20].

2.2 Biomass gasification and gasifier types

Biomass gasification is the conversion of carbonaceous material into a gaseous product,

which mainly consists of H2 and CO with lower amounts of CO2, H2O, N2, CH4, and higher

hydrocarbons. The process is carried out with the aid of a gasifying agent (air, O2, steam,

or a mixture of these) at an elevated temperature between 500 and 1400 °C at atmospheric

or elevated pressure. The produced gas can be either combusted for heat and power

generation or further processed to produce synthetic gas, SNG, or value-added chemicals.

Various gasification technologies are shown in Figure 2-3.

11

Gasifier

Multi-stage

Gasifier

Entrained

Flow

gasifier

Fluidized

bed gasifier

Fixed bed

gasifier

Other Fixed bed gasifier

Co-current fixed bed

gasifier

Counter-current Fixed bed

gasifier

Stationary fluidised bed

gasifier

External Circulating

fluidised bed gasifier

Internal Circulating

fluidised bed gasifier

Twin-bed fluidized bed

gasifier

Direct

FB

Indirect

FB

Figure 2-3. Overview of various gasification technologies

The main criterion for selecting a gasifier type for a particular purpose is the fuel capacity

to be handled. Depending on the type of gasifier, the fuel capacity can be largely influenced,

as shown in Figure 2-4.

12

100

kW

1000

kW

10 000

kW

100 000

kW

Downdraft-Fixed

bed

Updraft-Fixed bed

Atm. BFB

Atm. Indirect

CFB

Pressuried BFB, CFD

and Indirect CFB

Entrained Flow Gasifier

Fuel

Capacites

Multi-Stage Gasifier

Figure 2-4. Capacity ranges of common gasification technologies

Fixed bed gasifiers are limited to small-scale applications, while fluidized bed and entrained

flow gasifiers can be used for large-scale applications.

2.3 Advanced gasification technologies

Based on the typical gasifier types, several advanced gasification processes have been

developed recently to improve the performance of gasification, thus increasing the

gasification efficiency and reducing the tar content in the syngas.

High-temperature agent gasification (HTAG)

In conventional direct gasification, the required heat for the endothermic gasification

process is mainly supplied by heat release from oxidization of a part of the feedstock. This

process produces non-combustible CO2, which dilutes the produced syngas and lowers its

calorific value. In addition, the presence of a high CO2 content results in reduction of the

partial pressures of the other gas species, which reduces the intensity of important reactions,

such as the water-gas shift reaction to produce H2. In contrast, indirect gasification uses the

heat released from an outside source, for example, hot char recirculation and/or sensible

heat from a preheated gasifying agent.

13

The HTAG process lies between these two extremes, i.e., part of the energy is supplied by

a highly preheated gasifying agent (~1000 °C), while the rest is supplied by oxidation inside

the reactor. Preheating of the agent is achieved using a high-cycle regenerative preheater.

Preheating the gasifying agent to high temperatures has many positive effects, such as

increased gas yield, heating value, and gasification efficiency, and decreased tar content [21].

Multi-stage gasification

In a one-stage gasifier, the overlap of the process steps, such as heating, drying, pyrolysis,

oxidation, and gasification, makes it impossible to independently control and optimize the

different steps. As interaction between volatiles and char can have a negative impact on the

char reactivity, a higher gasification efficiency is expected if char gasification is performed

in the absence of volatiles. Multi-stage gasification, which was developed based on this

concept, results in high gas purity with low levels of tar and high process efficiencies with

high char conversion rates. The complexity of the gasification process is compensated by a

simpler subsequent gas cleaning process. Table 2-1 provides an overview of the

performance of multi-stage gasification processes [22].

Table 2-1. Overview of multi-stage gasification processes

Gasification

technology

Number of stages CGE

(%)

Tar content

(Mg/Nm3)

HHV

(MJ/Nm3)

Viking gasifier 2 93 <15 6.6

FLETGAS process 3 81 10 6.4

LT-CFB process 2 87–93 >4800 5.2–7

Carbo-V process 3 82 Tar free High

In addition to the above-listed processes, the Wood-Roll® multi-stage gasification process

has a unique feature, in which the pyrolysis gas is used only as the heat carrier for the

endothermic steam gasification process, thus enabling indirect gasification. This process

claims to have several advantages, such as raw material flexibility, less tar, high calorific

value gas, and an efficient gasification process [23].

14

2.4 Gasification of pretreated biomass

The main focus of gasification studies of pretreated biomass has been mechanically

pretreated (pelletized) biomass. A few studies can be found on physicochemically or

thermochemically pretreated biomass gasification.

Comparative simulation studies based on entrained flow gasification of hydrothermal

carbonized biomass and fluidized bed gasification of raw wood reported that biocoal

gasification is more efficient. However, the conversion losses in the HTC process were

revealed to outweigh this efficiency gain [24]. Entrained flow gasification of hydrothermal

carbonized biomass was reported to have high carbon and overall conversion at low

residence times (1 s) and high temperatures (1000–1400 °C). Moreover, a slightly lower

conversion of lignite was observed, indicating the high reactivity of biocoal [25].

Recently, HTC pretreatment of rice residues to produce biocoal as a transportable value-

added product has been studied with the purpose of gasifying in remote villages to generate

electricity for rural communities [26].

Steam gasification of torrefied wood at 1400 °C produced 7% more H2 and 20% more CO

compared with the parent wood [27]. However, while the kinetics of gas phase reactions

were comparable at a lower gasification temperature (1200 °C), a lower char gasification

reactivity was reported. Compared with raw biomass, air gasification of torrefied switch

grass gave lower CO and H2 yields, while the CH4 yield was higher. However, the

significantly higher CO and H2 yields reported for densified torrefied material were claimed

to result from the binder and the moisture added during densification [28]. Air gasification

of torrefied pellets was reported to result in a higher calorific value of syngas compared

with that obtained from unpretreated pellets, with a cold gas efficiency (CGE) of around

75%. In addition, less tar was obtained with torrefied pellets. However, due to stability

effects, this type of fuel is much more suitable for co-gasification [29]. Large-scale co-

gasification tests of torrefied biomass and steam-exploded biomass with hard coal have

been carried out by Vattenfall [30].

15

2.5 Integration of a steel industry furnace with biomass gasification

On the system level, the integration of a steel industry furnace with biomass gasification

should also consider

how to use the waste heat from the furnace in the gasification process to achieve a

higher integrated energy efficiency, and

how to increase the value chain of the whole system, for example, extraction of

metallurgical coke/carbon and H2.

There is a large amount of waste heat in steel plants. Waste heat in steel plants is separated

as

High-temperature waste heat: T > 650 °C

Medium-temperature waste heat: 230 °C < T < 650 °C

Low-temperature waste heat: T < 230 °C

To achieve higher integrated energy efficiency by using the waste heat, the high- and

medium-temperature waste heat can be used in the biomass pretreatment; drying process,

and pyrolysis stages. Finally, a waste heat to gas concept can be established, which is one

of the main objectives of this thesis.

16

CHAPTER III

3 METHODOLOGY

3.1 Experimental basis

In this section, the experimental basis of the methodology used in thesis is described, which

includes feedstock materials, TGA experiments, and gasification experiments.

3.1.1 Feedstock materials

Four biomass types, namely gray pellets, black pellets, spent grain biocoal, and horse

manure biocoal, were used in the gasification experiments, as shown in Figure 3-1.

Figure 3-1. Biomass types used in the experiments: (top-left) gray pellets, (top-

right) black pellets, (bottom-left) spent grain biocoal, and (bottom-right) horse

manure biocoal

17

Gray pellets consist of woody-based roadside scrub cuts that were produced and supplied

by Boson Energy SA, Luxembourg [31]. These pellets represent unpretreated biomass in

this study, in the sense of no thermochemical or physicochemical pretreatment was applied.

The only treatment that these pellets have undergone is milling, drying, and pelletizing

(physical pretreatment).

Black pellets, which consist of 75% softwood and 25% hardwood, were produced by steam

explosion pretreatment and were supplied by Zilkha Biomass Energy, Texas, USA [32]. The

main process steps that these pellets have undergone are milling, steam pretreatment,

drying, and pelletizing.

The two types of biocoal pellets types consist of spent grain and horse manure produced

by AVA-CO2 Forschung Gmbh in Karlsruhe, Germany, at their HTC demonstration plant

[33]. The process conditions used for the HTC pretreatment were: temperature around

210–215 °C and residence time around 4 h. The solid product yield was around 67% of the

dry input.

Table 3-1 provides the chemical compositions of the various biomass types, and the

compositions are represented in a van Krevelen diagram in Figure 3-2.

Table 3-1. Chemical compositions of the various biomass types

Gray

pellets

Black

pellets

Spent

grain

biocoal

Horse

manure

biocoal

Proximate analysis Parameter %wt

Moisture 9.8 4.2 10.6 9

Volatile matter, dry basis 81.2 76.8

Fixed carbon, dry basis 16.9 22.3

Ash, dry basis 1.9 0.9 6.7 11.5

LHV (MJ/kgdry) 18.4 20.1 28.9 21.1

Ultimate analysis Element %wtdb

C 49.4 52.6 66.3 57.1

H 5.9 5.8 7 5.5

O 42.6 40.6 16.1 24.4

N 0.2 0.1 3.7 1.3

S - - 0.1 0.2

18

Figure 3-2. Van Krevelen diagram

The significant features observed in the proximate analysis are the high ash content of both

biocoal types and the exceptionally high carbon content (and lower oxygen content) and

LHV of spent grain biocoal. Moreover, the van Krevelen diagram shows that the black

pellets have a significantly reduced H/C ratio, while both biocoal types have significantly

reduced O/C ratios.

3.1.2 TGA experiments

The decomposition characteristics of biomass were analyzed using a laboratory-scale

thermogravimetric analyzer (TG 209 F1 Iris, Netsch, Germany) at Karlsruhe Institute of

Technology, Germany. The system can be operated with both inert and oxidizing

atmospheres with variable CO2 concentrations. Figure 3-3 shows a schematic of the

system.

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7

H/C

O/C

Hard wood

Soft wood

Gray pellets

Peat

Black pellets

Biocoal from horse manure

Biocoal from spent grain

Lignite

Sub-bituminous coal

Bituminous coal

Anthracite

19

Figure 3-3. Schematic of TGA

The sample holder is a ceramic crucible that is connected to a balance system (with 0.1 µg

resolution), and surrounded by a high-temperature oven (temperatures up to 1000 °C). In

order to avoid transfer limitations, a sample mass of 2 mg is typically used.

The experimental procedure was as follows. The ground and homogenized biomass sample

was first pyrolyzed under a N2 atmosphere with a constant heating rate of 30 °C/min from

30 °C to 1000 °C, followed by a 15 min holding time to remove all the volatile matter. The

sample remaining after the pyrolysis step was then cooled to 900 °C at a rate of 30 °C/min,

followed by a 15 min holding time. The sample was then gasified in a N2/CO2 atmosphere

(10 vol% CO2) at 900 °C until a constant final weight was observed. The sample

temperature and mass were measured and recorded with respect to time.

The following definitions were adopted in the data analysis.

For the pyrolysis period, the first derivative of the residual mass (with respect to

temperature) is defined as

𝐷𝑇𝐺 = −𝑑𝛼

𝑑𝑇 (Eq. 3-1)

where the residual mass is given by,

𝛼 =𝑀𝑎𝑠𝑠

𝐼𝑛𝑖𝑡𝑖𝑎𝑙 𝑚𝑎𝑠𝑠 (Eq. 3-2)

For the gasification period, the char conversion is defined as

𝑥 =𝑚0−𝑚

𝑚0−𝑚𝑓 (Eq. 3-3)

20

3.1.3 Gasification experiments

The gasification experiments were performed in a pilot-scale updraft HTAG unit that uses

preheated air/steam as the gasifying medium. The system consists of a fuel feeding system,

a feed gas preheater, an updraft gasifier, and a syngas post-combustion unit, as shown in

Figure 3-4. The gasifier is approximately 3 m high with an internal diameter of 0.4 m.

Figure 3-4. Experimental setup of the HTAG system: (top) schematic view,

(bottom-left) actual view, and (bottom-right) temperature measuring points

21

A screw conveyor transports the biomass pellets stored in the feed tank to the gasifier. The

preheater supplies hot gasifying agent (air/steam) to the gasifier at the side of the bottom

section, below the grate. The 6 mm thick grate is perforated, with 40% open area, and made

from Kanthal steel, which allows high-temperature operation up to 1425 °C. A boiler

supplies steam, if steam is the gasifying agent. The biomass and produced hot gases flow

countercurrently. The syngas leaves the gasifier at the side of the top section. The remaining

unconverted fuel particles (mainly ash) pass through the grate and are collected in the ash

box. The produced syngas is burned in the post-combustion chamber.

The experimental procedure was as follows. First, the feeder was calibrated for each type

of biomass pellet before the experiments. The feeder works at a specific frequency, which

is correlated with the feeding rate. To calculate and examine the relation between the

frequency of the feeder and the feeding rate, a number of trials were conducted with each

biomass type, in which the frequency was varied at a certain time interval and the final

weight of the biomass feed was determined. Then, the relation between the frequency of

the feeder and the biomass feed rate was established, as shown in Figure 3-5.

Figure 3-5. Relation between feed rate and frequency of the feeder

y = 2.1158x

y = 2.4006xy = 2.7558x

y = 1.0064x

0

20

40

60

80

100

120

10 15 20 25 30 35 40 45 50

Fee

d (

kg/

h)

Frequency (Hz)

Gray pellets Black pelletsSpent grain biocoal Horse manure biocoal

22

For each pellet type, experiments were conducted with air/steam as the gasifying medium.

The gasifying agent was preheated by burning NG in a regenerative preheater. Once the

preheated gasifying agent reached the desired temperature, the frequency of the biomass

feeder was adjusted to achieve the required feed rate. The vertical temperature distribution

of the gasifier was measured with eight type-S thermocouples located along the reactor

height (see Figure 3-4) and recorded every minute by a data acquisition system connected

to a computer. The horizontal temperature gradient inside the gasifier was assumed to be

less significant. To calculate the pressure drop over the fuel bed and ensure a negative

pressure at the gasifier top (for safety reasons), four digital manometers were located along

the gasifier height. The values were manually recorded every five minutes.

After cooling and condensing by passing through the water traps, the dry syngas

composition was measured every three minutes by an online gas chromatograph (GC). Tar

sampling was carried out using a solid phase adsorption (SPA) method, in which 100 mL

of gas was collected in a syringe at a constant flow rate for off-line analysis. The particle

size distribution was analyzed using a low-pressure impactor (LPI).

The syngas flow rate was calculated by applying overall N balance. The slightly negative

pressure maintained at the top of the gasifier can cause air infiltration into the gasifier

through the biomass feeding line, and when applying N balance, a correction method was

applied to compensate this. First, the volumetric flow rate of the N-free gas stream was

calculated by applying C balance (negligible C content was assumed in the outgoing ash).

To obtain the total syngas flow rate, the input N2 flow rate was then directly added to the

volumetric flow rate of N-free gas.

The following method and definitions were adopted in the data analysis. A time period of

steady-state operation was selected for each run (20–60 min) based on the temperature and

gas composition recordings, and the measurements were averaged within the selected time

period.

The equivalence ratio (ER) is defined as

𝐸𝑅 =(𝑂2/𝐹)𝑎𝑐𝑡𝑢𝑎𝑙

(𝑂2/𝐹)𝑠𝑡𝑜𝑖𝑐ℎ𝑖𝑜𝑚𝑒𝑡𝑟𝑖𝑐 (Eq. 3-4)

23

Table 3-2 summarizes the stoichiometric O2 requirements calculated for combustion of

the various biomass types.

Table 3-2. Stoichiometric O2 requirement of each biomass type

Biomass type Composition (O2/F)stoichiometric

(mol/mol)

Gray pellets CH1.43O0.65 1.03

Black pellets CH1.32O0.58 1.04

Spent grain biocoal CH1.27O0.18 1.23

Horse manure biocoal CH1.16O0.32 1.13

The performance parameters can be defined as

𝐿𝐻𝑉𝑠𝑦𝑛𝑔𝑎𝑠 = ∑(𝐿𝐻𝑉𝑖 𝑥𝑖) (Eq. 3-5)

𝑌𝑠𝑦𝑛𝑔𝑎𝑠 =�̇�𝑠𝑦𝑛𝑔𝑎𝑠

�̇�𝑐 (Eq. 3-6)

𝐶𝐺𝐸 =𝐿𝐻𝑉𝑠𝑦𝑛𝑔𝑎𝑠

𝐿𝐻𝑉𝑓𝑢𝑒𝑙 𝑌𝑠𝑦𝑛𝑔𝑎𝑠 × 100% (Eq. 3-7)

For a single tar compound, the tar dew point is given by [34]

22400𝐶𝑡𝑎𝑟

𝑀𝑡𝑎𝑟

(273+𝑇)

273

1

𝑃𝑠𝑣(𝑇)= 1 (Eq. 3-8)

For a mixture of tar compounds, the dew point can be calculated in a similar manner from

the sum of each contribution. Based on this, the complete dew point model developed by

Energy Research Centre of the Netherlands [35] was used to calculate the tar dew points in

this study.

24

3.2 Application-oriented process simulations

3.2.1 System description

A multi-stage biomass gasification plant and a steel industry heat treatment furnace were

studied for integration. The example gasifier system is the Cortus WoodRoll® process [36],

whereas the example furnace is a heat treatment belt furnace in the Höganäs steel powder

plant [37]. The details of each process step are described in the next section.

NG firing followed by five different syngas firing cases with different heat recovery options

(as given below) was assessed.

Case 1: No heat recovery

Case 2: Low-temperature heat recovery for steam superheating

Case 3: Low-temperature heat recovery for biomass predrying

Case 4: Low-temperature heat recovery for steam superheating and biomass predrying

Case 5: High-temperature heat recovery for steam superheating and biomass drying

As high-temperature heat recovery from furnace flue gases could be costly and complex,

more focus was given to low-temperature heat recovery. The temperature limits for the

low-temperature and high-temperature heat recovery options (400 °C and 700 °C,

respectively) were chosen by considering the conventional heat exchanger operating range

[38] and the flue gas temperature, respectively.

Figure 3-6 shows the basic integration system without heat recovery (Case 1).

25

Fig

ure

3-6

. B

asi

c i

nte

gra

tio

n o

f th

e g

asi

fier

syst

em

an

d t

he f

urn

ace

26

3.2.2 Scenario description

In addition to the use of syngas in the furnace, the co-production of limited quantities of

coke and H2 (⅓ of the feedstock for coke production and ¼ of the feedstock for H2

separation) was considered as the initial step.

The following four production scenarios were analyzed for energy efficiency.

Scenario 1: Only syngas

Scenario 2: Syngas and coke

Scenario 3: Syngas and H2

Scenario 4: Syngas, coke, and H2

3.2.3 Process models

A steady-state model was developed for the integrated system using Aspen Plus. Details of

such solids modeling with Aspen Plus can be found elsewhere [39]. Only the main sub-

processes and boundary conditions are discussed here.

Predrying

Predrying is incorporated for cases 3 and 4. Raw biomass with approximately 40% moisture

is dried to a lower moisture content of around 20–25% in a predryer prior to the drying

process. This process allows furnace flue gas heat recovery, while enabling smooth

functioning of the gasifier system. To avoid any acid condensation, the flue gas exhaust

temperature is kept above 170 °C.

Drying

The moisture content of the biomass is reduced to approximately 5% in a dryer by using

the hot flue gas from the pyrolyzer exit. In case 5, the furnace flue gas is used instead. The

temperature of both the predryer and dryer is maintained at 105 °C. To remove the evolved

moisture, hot vent air at 100 °C is used with a flow rate of 2 Nm3/kg moisture removed.

27

The mass balance of the drying process is expressed as

�̇�𝑏𝑖𝑜𝑚𝑎𝑠𝑠_𝑤𝑒𝑡 + �̇�𝑣𝑒𝑛𝑡 𝑎𝑖𝑟 = �̇�𝑏𝑖𝑜𝑚𝑎𝑠𝑠_𝑑𝑟𝑦 + �̇�𝑚𝑜𝑖𝑠𝑡 𝑣𝑒𝑛𝑡 𝑎𝑖𝑟 (Eq. 3-9)

The energy balance of the drying process is given by

∑ �̇�𝑖𝑖 ∫ 𝐶𝑝,𝑖𝑑𝑇𝑇𝑓𝑙𝑢𝑒𝑔𝑎𝑠_𝑖𝑛

𝑇𝑓𝑙𝑢𝑒𝑔𝑎𝑠_𝑜𝑢𝑡= �̇�𝑏𝑖𝑜𝑚𝑎𝑠𝑠_𝑑𝑟𝑦 ∫ 𝐶𝑝,𝑏𝑖𝑜𝑚𝑎𝑠𝑠_𝑑𝑟𝑦𝑑𝑇

𝑇𝑏𝑖𝑜𝑚𝑎𝑠𝑠_𝑜𝑢𝑡

𝑇𝑏𝑖𝑜𝑚𝑎𝑠𝑠_𝑖𝑛+

�̇�𝐻2𝑂 [∫ 𝐶𝑝,𝐻2𝑂𝑑𝑇𝑇𝑏𝑖𝑜𝑚𝑎𝑠𝑠_𝑜𝑢𝑡

𝑇𝑏𝑖𝑜𝑚𝑎𝑠𝑠_𝑖𝑛+ 𝐿𝐻2𝑂] + �̇�𝑣𝑒𝑛𝑡 𝑎𝑖𝑟 ∫ 𝐶𝑝,𝑣𝑒𝑛𝑡 𝑎𝑖𝑟𝑑𝑇 + �̇�𝐿

𝑇𝑣𝑒𝑛𝑡 𝑎𝑖𝑟_𝑜𝑢𝑡

𝑇𝑣𝑒𝑛𝑡 𝑎𝑖𝑟_𝑖𝑛

(Eq. 3-10)

in which 10% heat loss is assumed.

Pyrolysis

Pyrolysis of dry biomass is carried out at a temperature that generates just enough pyrolysis

gas and char for the process by using hot flue gas from the gasifier exit. This temperature

depends on the heat recovery option considered. When calculating the yields of the

pyrolysis products, it was assumed that C2H4 is the only significant higher hydrocarbon

present in the pyrolysis gas, and tar is considered as a single component.

𝐵𝑖𝑜𝑚𝑎𝑠𝑠(𝑑𝑎𝑓) → 𝛼1𝐶ℎ𝑎𝑟 + 𝛼2𝐶𝑂 + 𝛼3𝐶𝑂2 + 𝛼4𝐻2 + 𝛼5𝐻2𝑂 + 𝛼6𝐶𝐻4 + 𝛼7𝐶2𝐻4 +

𝛼8𝑇𝑎𝑟 (R. 3-1)

To calculate the stoichiometric coefficients, some available empirical correlations were

coupled with elemental balances and energy balance. The ash content (0.4%wtdb) and

moisture content of dried biomass are then directly added to the output.

The empirical relations are as follows [40]:

𝑀𝐻2 𝛼4/𝑀𝐵𝑖𝑜𝑚𝑎𝑠𝑠

𝑀𝐶𝑂 𝛼2/𝑀𝐵𝑖𝑜𝑚𝑎𝑠𝑠= 3 × 10−4 +

0.0429

1+(𝑇

632)

−7.23 (Eq. 3-11)

𝑀𝐶𝐻4 𝛼6

𝑀𝐵𝑖𝑜𝑚𝑎𝑠𝑠= −2.18 × 10−4 + 0.146

𝑀𝐶𝑂 𝛼2

𝑀𝐵𝑖𝑜𝑚𝑎𝑠𝑠 (Eq. 3-12)

𝑀𝐻2 𝛼4

𝑀𝐵𝑖𝑜𝑚𝑎𝑠𝑠= 1.145 [1 − 𝑒𝑥𝑝 (−0.11 × 10−2 𝑇)]9.384 (Eq. 3-13)

𝐿𝐻𝑉𝐺 = −6.23 + 2.47 × 10−2 𝑇 (Eq. 3-14)

where

𝐿𝐻𝑉𝐺(𝑀𝐵𝑖𝑜𝑚𝑎𝑠𝑠 − 𝑀𝐶ℎ𝑎𝑟𝛼1 − 𝑀𝐻2𝑂𝛼5 − 𝑀𝑇𝑎𝑟𝛼8) = 𝑀𝐶2𝐻4𝛼7 𝐿𝐻𝑉𝐶2𝐻4

+

𝑀𝐶𝐻4𝛼6 𝐿𝐻𝑉𝐶𝐻4

+ 𝑀𝐻2𝛼4 𝐿𝐻𝑉𝐻2

+ 𝑀𝐶𝑂𝛼2 𝐿𝐻𝑉𝐶𝑂 (Eq. 3-15)

28

Char yield [41]

𝑀𝐶ℎ𝑎𝑟𝛼1

𝑀𝐵𝑖𝑜𝑚𝑎𝑠𝑠= 0.0017𝑇2 − 1.426𝑇 + 322.49 (Eq. 3-16)

Composition of biomass CxHyOz [36]

𝑥 𝑀𝐶

𝑀𝐵𝑖𝑜𝑚𝑎𝑠𝑠= 0.499 (Eq. 3-17)

𝑦 𝑀𝐻

𝑀𝐵𝑖𝑜𝑚𝑎𝑠𝑠= 0.062 (Eq. 3-18)

𝑧 𝑀𝑂

𝑀𝐵𝑖𝑜𝑚𝑎𝑠𝑠= 0.439 (Eq. 3-19)

Composition of char CaHbOc [40], [41]

𝑎 𝑀𝐶

𝑀𝐶ℎ𝑎𝑟= −9 × 10−6𝑇2 + 0.0074𝑇 − 0.7377 (Eq. 3-20)

𝑏 𝑀𝐻

𝑀𝐶ℎ𝑎𝑟= −0.41 × 10−2 + 0.1 𝑒𝑥𝑝 (−0.24 × 10−2𝑇) (Eq. 3-21)

𝑐 𝑀𝑂

𝑀𝐶ℎ𝑎𝑟= 1 − 𝑎 − 𝑏 (Eq. 3-22)

Composition of tar CdHeOf [40]

𝑑 𝑀𝐶/𝑀𝑇𝑎𝑟

𝑥 𝑀𝐶/𝑀𝐵𝑖𝑜𝑚𝑎𝑠𝑠= 1.14 (Eq. 3-23)

𝑒 𝑀𝐻/𝑀𝑇𝑎𝑟

𝑦 𝑀𝐻/𝑀𝐵𝑖𝑜𝑚𝑎𝑠𝑠= 1.13 (Eq. 3-24)

𝑓 𝑀𝑂 /𝑀𝑇𝑎𝑟

𝑧 𝑀𝑂 /𝑀𝐵𝑖𝑜𝑚𝑎𝑠𝑠= 0.8 (Eq. 3-25)

The elemental balances are

Carbon 𝑥 = 𝛼1𝑎 + 𝛼2 + 𝛼3 + 𝛼6 + 2 𝛼7 + 𝛼8𝑑 (Eq. 3-26)

Hydrogen 𝑦 = 𝛼1𝑏 + 2 𝛼4 + 2 𝛼5 + 4 𝛼6 + 4 𝛼7 + 𝛼8𝑒 (Eq. 3-27)

Oxygen 𝑧 = 𝛼1𝑐 + 𝛼2 + 2 𝛼3 + 𝛼5 + 𝛼8𝑓 (Eq. 3-28)

The energy balance is


𝑇𝑓𝑙𝑢𝑒𝑔𝑎𝑠_𝑜𝑢𝑡= ∑ �̇�𝑜𝑃 (ℎ𝑓

0 + 𝛥ℎ̅)𝑜 − ∑ �̇�𝑖𝑅 (ℎ𝑓0 + 𝛥ℎ̅)𝑖 + �̇�𝐿 (Eq. 3-29)

with the reactor heat loss set at 10%.

29

Gasification

Prior to gasification, the char is cooled to ambient temperature and crushed to

approximately 100 µm.

The Hardgrove grindability index (HGI) is given as [42]

𝐻𝐺𝐼 = 2.65 𝑒𝑥𝑝 (𝐶

20) (Eq. 3-30)

Gasification is carried out at 1100 °C by indirect heat supplied by burning the pyrolysis gas.

The gasifying agent is saturated or superheated steam at a pressure of 6 bar. Steam is

generated by recovering waste heat from the hot syngas and superheated by recovering

waste heat from the furnace flue gases. An S/C molar ratio of 1.3 is maintained by

considering the transfer limitations.

𝐶ℎ𝑎𝑟 + 1.3𝑎 𝐻2𝑂 → 𝛽1𝐶𝑂 + 𝛽2𝐶𝑂2 + 𝛽3𝐻2 + 𝛽4𝐻2𝑂 + 𝛽5𝐶𝐻4 (R. 3-2)

For a fixed temperature and pressure, the equilibrium gives

𝑑𝐺𝑚𝑖𝑥 = 0 (Eq. 3-31)

where the Gibbs energy for a mixture of ideal gases is

𝐺𝑚𝑖𝑥 = ∑ 𝑁𝑖�̅�𝑖,𝑇 = ∑ 𝑁𝑖 [�̅�𝑖,𝑇0 + 𝑅𝑇 𝑙𝑛 (

𝑃𝑖

𝑃0)] (Eq. 3-32)

To account for possible deviations from equilibrium, a restricted equilibrium method was

applied by defining an approach temperature for the following reactions.

Water gas shift reaction: 𝐶𝑂 + 𝐻2𝑂 ⇌ 𝐶𝑂2 + 𝐻2 (R. 3-3)

Methane steam reforming: 𝐶𝐻4 + 𝐻2𝑂 ⇌ 𝐶𝑂 + 3 𝐻2 (R. 3-4)

With the restricted equilibrium, the equilibrium is evaluated at a temperature lower than the

operating temperature, such that the model results are comparable with the experimental

case. This well-established method has been used for both fluidized bed gasifiers [43, 44]

and entrained flow gasifiers [45].

The energy balance is


𝑇𝑓𝑙𝑢𝑒𝑔𝑎𝑠_𝑜𝑢𝑡= ∑ �̇�𝑜𝑃 (ℎ𝑓

0 + 𝛥ℎ̅)𝑜

− ∑ �̇�𝑖𝑅 (ℎ𝑓0 + 𝛥ℎ̅)

𝑖+ �̇�𝐿 (Eq. 3-33)

The gasifier heat losses are obtained from the experimental conditions.

30

Gas cleaning and cooling

To separate the coarse ash particles from the syngas, a cyclone is used, which is modeled

as a high-efficiency-type cyclone in the design mode.

Gas cooling is carried out by means of a waste heat recovery boiler that generates steam.

The boiler is modeled by two heat exchangers in series for sensible and latent heat,

respectively. The energy balance is as follows.

∑ �̇�𝑖𝑖 ∫ 𝐶𝑝,𝑖𝑑𝑇𝑇𝑠𝑦𝑛𝑔𝑎𝑠_𝑖𝑛

𝑇𝑠𝑦𝑛𝑔𝑎𝑠_𝑜𝑢𝑡= �̇�𝐻2𝑂 [∫ 𝐶𝑝,𝐻2𝑂𝑑𝑇

𝑇𝐻2𝑂_𝑜𝑢𝑡

𝑇𝐻2𝑂_𝑖𝑛+ 𝐿𝐻2𝑂] + �̇�𝐿 (Eq. 3-34)

with the heat losses from these heat exchangers set at 10%.

To separates the fine particles from the syngas, a fabric filter is used, which is modeled in

design mode. Then, a water-cooled condenser operating at 35 °C condenses the extra steam

available in the syngas. Clean dry syngas is stored in a gas storage tank at a pressure slightly

above atmospheric pressure.

Combustion and heat transfer inside the furnace

When the syngas is burned in the heat treatment furnace, complete combustion is assumed.

The stoichiometric reactions occurring in the furnace are

𝐶𝑂 + 0.5 𝑂2 → 𝐶𝑂2 (R. 3-5)

𝐻2 + 0.5 𝑂2 → 𝐻2𝑂 (R. 3-6)

𝐶𝐻4 + 2 𝑂2 → 𝐶𝑂2 + 2 𝐻2𝑂 (R. 3-7)

The combustion is carried out with 10% excess air.

The adiabatic combustion process is expressed as

∑ �̇�𝑖𝑅 (ℎ𝑓0 + 𝛥ℎ̅)𝑖 = ∑ �̇�𝑜𝑃 (ℎ𝑓

0 + 𝛥ℎ̅)𝑜 (Eq. 3-35)

where

𝛥ℎ̅𝑖 = ∫ 𝑐𝑝,𝑖𝑇𝑖𝑛

𝑇𝑟𝑒𝑓𝑑𝑇 (Eq. 3-36)

𝛥ℎ̅𝑜 = ∫ 𝑐𝑝,𝑖𝑇𝑎𝑑

𝑇𝑟𝑒𝑓𝑑𝑇 (Eq. 3-37)

The heat generated by the combustion of syngas is used to supply heat for the iron powder

and process gas.

31

The energy balance of the furnace is given by

∑ �̇�𝑖𝑖 ∫ 𝐶𝑝,𝑖𝑑𝑇𝑇𝑎𝑑

𝑇𝑓𝑙𝑢𝑒𝑔𝑎𝑠_𝑜𝑢𝑡= �̇�𝑖𝑟𝑜𝑛 ∫ 𝐶𝑝,𝑖𝑟𝑜𝑛𝑑𝑇

𝑇𝑖𝑟𝑜𝑛_𝑜𝑢𝑡

𝑇𝑖𝑟𝑜𝑛_𝑖𝑛+

�̇�𝑝𝑟𝑜𝑐𝑒𝑠𝑠𝑔𝑎𝑠 ∫ 𝐶𝑝,𝑝𝑟𝑜𝑐𝑒𝑠𝑠𝑔𝑎𝑠𝑑𝑇𝑇𝑝𝑟𝑜𝑐𝑒𝑠𝑠𝑔𝑎𝑠_𝑜𝑢𝑡

𝑇𝑝𝑟𝑜𝑐𝑒𝑠𝑠𝑔𝑎𝑠_𝑖𝑛+ �̇�𝐿 (Eq. 3-38)

The heat loss from the furnace is assumed to be similar to the NG firing case in the real

furnace.

Coke making process

A separate stream of char (equivalent to ⅓ of the feed) is used for the coke making process.

The process is carried out at 1100 °C.

𝐶ℎ𝑎𝑟 → 𝛾1𝐶 + 𝛾2𝐶𝑂 + 𝛾3𝐶𝑂2 + 𝛾4𝐻2 + 𝛾5𝐻2𝑂 + 𝛾6𝐶𝐻4 + 𝛾7𝐶2𝐻4 (R. 3-8)

For preliminary calculations, the yields of the coking process were obtained from the

experimental data. At this stage, the coke gas is recycled back to the gasifier burners to

recover heat.

H2 separation process

Pressure swing adsorption (PSA) is used to separate H2 from part of the syngas (equivalent

to ¼ of the feed). A H2 recovery rate of around 80% is expected for a typical industrial

PSA process, and this value is used as a preliminary value for the calculation of the H2 yield.

3.2.4 Performance analysis

Efficiency

The gasifier system efficiency is defined by

𝜂𝐺 =𝐸𝐶ℎ,𝑠𝑦𝑛𝑔𝑎𝑠

𝐸𝐶ℎ,𝑏𝑖𝑜𝑚𝑎𝑠𝑠× 100% (Eq. 3-39)

The furnace system efficiency is defined by

𝜂𝐹 =𝐸𝑆,𝐹𝑒

𝐸𝐶ℎ,𝑠𝑦𝑛𝑔𝑎𝑠× 100% (Eq. 3-40)

32

The integrated system efficiency is defined by

𝜂𝐼 =𝐸𝑆,𝐹𝑒+𝐸𝐶ℎ,𝑐𝑜𝑘𝑒+𝐸𝐶ℎ,𝐻2

𝐸𝐶ℎ,𝑏𝑖𝑜𝑚𝑎𝑠𝑠× 100% (Eq. 3-41)

All the calculations are performed based on an ambient temperature of 10 °C. The chemical

energy values are based on the LHV.

Pinch analysis

The performance of the heat integration was analyzed by means of pinch analysis using the

IChemE pinch analysis spreadsheet [46] as a tool to develop the graphical representation

of the heat cascade (grand composite curve (GCC)). The minimum temperature difference

was taken as 20 °C.

33

CHAPTER IV

4 BIOMASS CHARACTERIZATION

4.1 Mass loss behavior

The decomposition behavior of both pellet types was studied based on the

thermogravimetric residual mass curve presented in Figure 4-1.

Figure 4-1. TGA mass loss curves

The proximate composition of each biomass obtained from above mass loss curves is

summarized in Table 4-1.

Table 4-1. Proximate analysis of each biomass type obtained with TGA

Biomass type Volatiles (% wt) Fixed carbon (% wt) Ash (% wt)

Gray pellets 87 10 3

Black pellets 87 12 1

Biocoal 81 15 4

0

100

200

300

400

500

600

700

800

900

1000

1100

0

10

20

30

40

50

60

70

80

90

100

0 10 20 30 40 50 60 70 80 90 100 110

Tem

pe

ratu

re [

°C]

Mas

s [%

]

Time [min]

Gray pellets Black pellets Biocoal Temperature

N2 N2/CO2

34

It can be noted that the proximate composition values obtained from this TGA method

are different from those obtained from the ASTM method in Table 3-1. This difference

could be due to two factors: the different heating rates applied and the different oxidizing

atmospheres. The typical oxidizer used in the ASTM method is O2, while CO2 was used in

the TGA method.

4.2 Pyrolysis behavior

The constant heating rate period in the pyrolysis stage was further studied using the

derivative thermogravimetry (DTG) plots given in Figure 4-2, which is the first derivative

of the residual mass curve. The peak temperatures and intensities are summarized in Table

4-2.

Figure 4-2. Derivative thermogravimetric plots for each biomass type

35

Table 4-2. DTG peak temperatures and intensities for each biomass type

Biomass type Peak 1 Peak 2 Peak 3

Temperature

(°C)

Intensity

(Wt%/°C)

Temperature

(°C)

Intensity

(Wt%/°C)

Temperature

(°C)

Intensity

(Wt%/°C)

Gray pellets - 372 0.79 -

Black pellets 241 0.13 367 0.83 -

Biocoal 207 0.26 374 0.35 419 0.19

One of the major differences seen in the DTG plots of pretreated biomass (black pellets

and biocoal) is the early stage peak attached to the main peak compared with the

asymmetrical peak of unpretreated biomass (Gray pellets). Pretreatment has been reported

to release some hemicellulose into the solution, while depolymerizing the remaining

hemicellulose. Due to the depolymerized hemicellulose components, the peak shifts to a

lower temperature, which is otherwise merged with the cellulose peak. This phenomenon

is more prominent (low peak temperature and high peak intensity) with hydrothermal

carbonized biomass, which has undergone severe pretreatment.

Another difference is the peak temperature and peak intensity of cellulose decomposition.

Compared with gray pellets, black pellets show a slightly lower peak temperature and

slightly higher peak intensity, whereas biocoal shows a slightly higher peak temperature and

considerably lower peak intensity. The low thermal stability and high decomposition rate

of cellulose in black pellets could be due to the separation of the hemicellulose peak from

the cellulose peak, which otherwise suppresses cellulose decomposition due to melted

hemicellulose on the cellulose surface. Another reason could be the depolymerization of

cellulose during pretreatment [47]. Such low thermal stability of cellulose due to the steam

explosion pretreatment has been previously reported [47-50]. Black pellets are

predominantly softwood, whereas gray pellets are produced from hardwood. Hardwood

decomposes at a lower temperature than softwood due to a high content of hemicellulose.

Therefore, a more prominent peak shift would be observed if the black pellets and gray

pellets are from the same origin.

The DTG plot for biocoal decomposition shows a unique feature, with a shoulder at 419

°C attached to the cellulose peak. HTC reportedly results in the release of lignin into the

solution, followed by redeposition on the surface, as validated in several studies using

36

scanning electron microscope (SEM) images [51-54]. It was hypothesized in another study

that lignin may undergo a phase transition during aqueous high-temperature pretreatment

[55]. The shoulder observed beyond the cellulose peak in the DTG plot of biocoal may be

due to the presence of redeposited lignin globules on the surface, as observed by other

researchers for biocoal based on wheat straw and dried digestate. The thermal

decomposition of the separated globules has been reported to result in a peak near 400 °C

[52]. The presence of lignin globules on the surface may suppress cellulose decomposition

(high thermal stability and low intensity) by covering the cellulose surface with melted

lignin. A decrease in the peak cellulose decomposition rate and increase in the cellulose

thermal stability due to HTC pretreatment has been reported [52, 53, 56]. The high intensity

of lignin decomposition for biocoal may be due to an enhancement of the lignin content as

a result of the loss of hemicellulose. A slightly higher intensity above 400 °C is also seen

for black pellets compared with gray pellets. The breakdown of β-O-4 linkages and the C

α-Cβ bonds of lignin have also been reported due to HTC pretreatment, which can result

in the early thermal degradation of lignin [53]. Such low lignin decomposition temperatures

have also been reported for steam-exploded biomass [48, 50].

4.3 Char gasification behavior

The char conversion behavior of each biomass obtained from the above mass loss curves

is shown in Figure 4-3.

37

Figure 4-3. Char conversion progress with time

If the period of the char gasification process is considered, black pellets show a slightly

higher conversion rate than gray pellets. This could be due to the high fixed C content that

effectively moves the Boudouard reaction towards CO production. Biocoal shows a

considerably lower conversion rate compared with both black pellets and gray pellets. This

is possibly due to a lower char surface area because of the lower volatile content of biocoal.

0

0.2

0.4

0.6

0.8

1

0 5 10 15 20 25 30

Ch

ar c

on

vers

ion

Time (min)

Gray pellets Black pellets Biocoal

38

CHAPTER V

5 GASIFICATION PERFORMANCE IN UPDRAFT HTAG

5.1 Gasification of unpretreated and steam-exploded biomass

This section summarizes the results from air gasification and steam gasification of

unpretreated gray pellets and steam-exploded black pellets. The ER of air gasification was

maintained at around 0.2 and the steam/biomass ratio for steam gasification was fixed at

1.2. The preheating temperature of the gasifying agent was around 1000 °C.

5.1.1 Temperature distribution

Figure 5-1. Temperature profiles along the gasifier height

According to Figure 5-1, the temperature profiles feature two regions: the bed zone with

a significant temperature drop and the gas phase with an almost flat temperature profile.

The typical peak temperature expected in the combustion zone is not observed, possibly

because the combustion zone is thin, making it difficult to identify with the limited

temperature measuring points.

-20

0

20

40

60

80

100

120

500 700 900 1100 1300

Dis

tan

ce f

rom

th

e g

rate

(cm

)

Temperature (°C)

Black pellets with air

Black pellets with steam

Gray pellets with air

Gray pellets with steam

39

The high C content of black pellets results in a high temperature drop in the bed area,

mainly due to the endothermic Boudouard reaction. In contrast, better temperature

uniformity within the gasifier is seen for gray pellet gasification due to the lower C and

higher volatile content of gray pellets.

5.1.2 Syngas composition

Figure 5-2. Gas compositions and ratios for (left) air gasification and

(right) steam gasification

According to Figure 5-2, both air and steam gasification of black pellets result in a high

CO content in the syngas, whereas a high content of H2 and CO2 is obtained with gray

pellet gasification. The high CO/CO2 ratio and low H2/CO ratio reveal that the Boudouard

reaction is more effective, while the water gas shift reaction is suppressed during gasification

0

10

20

30

40

50

CO H2 CO2 CH4 CxHy*10

Co

mp

osi

tio

n (

%) Black pellets

Gray pellets


0

10

20

30

40

50


Co

mp

osi

tio

n (

%) Black pellets

Gray pellets


0

1

2

3

4

CO/CO2 H2/CO CH4/H2

Ch

arac

teri

stic

rat

io (

mo

l/m

ol)

Black pellets

Gray pellets

CO/CO2 H2/CO CH4/H2

0

1

2

3

4

CO/CO2 H2/CO CH4/H2

Ch

arac

teri

stic

rat

io (

mo

l/m

ol)

Black pellets

Gray pellets

CO/CO2 H2/CO CH4/H2

40

of black pellets. It was observed that black pellets have a significantly lower H/C ratio than

gray pellets (Figure 3-2). The Boudouard reaction at a high bed temperature results in a

comparatively low gas phase temperature, which is not favorable for the water gas shift

reaction. The water gas shift reaction is reportedly more effective in the temperature range

of 730–830 °C [57]. Further, black pellets resulted in a high hydrocarbon content in the

syngas due to limited cracking at the low gas phase temperature.

5.1.3 LHV, gas yield, and efficiency

Table 5-1 shows the LHV of syngas, the syngas yield, and the CGE of the tested cases.

Table 5-1. LHV of syngas, syngas yield, and CGE

Run LHV (MJ/Nm3) Ysyngas (Nm3/kg) CGE (%)

Air gasification of gray pellets 6 2.1 75.9

Steam gasification of gray pellets 8.2 1.5 74.1

Air gasification of black pellets 7.3 2 75.6

Steam gasification of black pellets 10.6 1.4 76.9

Black pellets had a high syngas LHV for both air and steam gasification. The main

contribution to this syngas heating value is the CO and hydrocarbon content. The increased

content of combustibles in the syngas and the reduced dilution of the syngas with N2

resulted in a higher LHV for the steam gasification case.

Compared with the corresponding gray pellet gasification cases, a slightly lower syngas yield

was observed for both black pellet gasification cases. This could be a result of the lower

volatile matter content of black pellets and fewer thermal cracking reactions during

gasification of black pellets. Both steam gasification cases showed lower syngas yields due

to less dilution with N2.

Comparable CGE values of between 74% and 77% were observed, irrespective of the type

of biomass (black pellets or gray pellets) or the type of feed gas (air or steam).

41

Figure 5-3. Tar composition based on identified tar species: (top) by category,

(middle) BTEX, phenols, and cresols, and (bottom) PAH

0

2

4

6

8

10

12

Air gasification ofBlack pellets

Air gasification ofGray pellets

Steam gasificationof Black pellets

Co

mp

osi

tio

n (

g/N

m3 )

BTEX PAH Phenol and Cresols

0

0.5

1

1.5

2

2.5

3

Air gasification ofBlack pellets

Air gasification ofGray pellets

Steam gasificationof Black pellets

Co

mp

osi

tio

n (

g/N

m3 )

Benzene Toluene Xylene Phenol Cresol

0

0.5

1

1.5

2

2.5

Co

mp

osi

tio

n (

g/N

m3) Air gasification of Black pellets

Air gasification of Gray pellets

Steam gasification of Blackpellets

42

5.1.4 Syngas purity

The syngas purity was examined using the tar content, composition, and dew point. The

identified tar species of syngas are grouped into three categories:

benzene/toluene/ethylbenzene/xylenes (BTEX), phenols and cresols, and polycyclic

aromatic hydrocarbons (PAH). Figure 5-3 shows the distribution of these three categories

in the obtained tar and the composition of each category by compound.

The tar obtained with air gasification was dominated by PAHs (mainly lighter compounds),

while phenols and cresols were the least abundant. Compared with gray pellets, black pellets

had a lower PAH and higher phenol and cresol content in the obtained tar.

The high phenol and cresol content of black pellet-derived tar could be due to the low gas

phase temperature. The content was further increased with steam gasification. At high

temperatures, secondary phenolic tars are converted to tertiary tars, including benzene and

naphthalene. As a result, the benzene and naphthalene content was nearly doubled during

gray pellet gasification. This type of aromatization at high temperatures and favorable

production of phenolic compounds with high steam content and low temperature has been

reported in the literature [58, 59]. Table 5-2 shows the total tar content1 and tar dew point

for each experimental case.

Table 5-2. Total tar content and tar dew point

Run Tar content (g/Nm3) Tar dew point (°C)

Air gasification of black pellets 15.4 108

Air gasification of gray pellets 13.4 105.4

Steam gasification of black pellets 11.4 93.3

The low syngas temperature observed for black pellet gasification had a limited effect on

the total tar content. This could be due to the low volatile (or hemicellulose) content and

high fixed carbon (or lignin) content of black pellets. Crosslinking reactions between

cellulose and lignin cause the formation of less tar during pyrolysis [60]. Further, steam

gasification produces less tar due to steam reforming reactions. The dew points were

around 100 °C in all cases, giving more freedom to downstream processing. Generally,

1 Total tar content including unidentified tar species

43

PAHs contribute to a high dew point of tar. However, irrespective of a high PAH content,

gray pellet gasification resulted in a slightly lower tar dew point due to a high content of

low boiling point benzene and naphthalene.

5.2 Gasification of hydrothermal carbonized biomass

This section summarizes the results for air gasification of HTC biocoal pellets. Three runs

with ER values varying from 0.2 to 0.22 were considered. The preheating temperature of

the gasifying agent was around 900 °C.


The temperature profiles for three gasification cases with different ER values are depicted

in Figure 5-4. Because the average temperature of the steady-state operating period was

used for the analysis, the standard deviation is plotted as the error bars.

Figure 5-4. Temperature profiles along the gasifier height

500

600

700

800

900

1000

1100

1200

1300

1400

1500

0 0.2 0.4 0.6 0.8 1 1.2

Tem

pe

ratu

re (

0 C)

Distance from the grate (m)

ER=0.2 ER=0.21 ER=0.22

44

Similar to the previous experiments, the temperature profiles show a bed zone with a

significant temperature drop and a gas phase with an almost flat temperature profile. A high

ER value resulted in a higher temperature, which facilitated the endothermic Boudouard

reaction and ultimately reduced the temperature in the bed zone. This effect was more

pronounced in the highest ER case, which could be due to the comparatively high bed

observed (0.5 m compared with 0.4–0.42 m). The slight temperature increase immediately

after the bed zone in the lower ER cases could be due to the slightly exothermic water gas

shift reaction. The highest ER case seems less effective for the water gas shift reaction

because the gas phase temperature is too low. The gas exit temperature showed a decreasing

trend with ER.


Figure 5-5 shows the composition of the syngas and various important gas ratios. Because

the average composition of the steady-state operating period was used, the standard

deviation is plotted as the error bars. The sum of the higher hydrocarbons is represented

by CxHy, where x ranges from 2 to 6.

Figure 5-5. Gas composition and ratios

Increasing the ER value had a positive effect on the CO content and negative effects on

the H2 and hydrocarbon contents of the syngas. As discussed for the temperature profile,

the extraordinary performance of the Boudouard reaction (indicated by the CO/CO2 ratio)

for the highest ER case could be due to the effect of a higher bed. The high H2/CO ratio

observed for low ER cases is an indication of the water gas shift reaction, which was also

0

5

10

15

20

25

30

35

H2 CO CO2 CH4 CxHy

Co

mp

osi

tio

n (

%)

ER=0.2 ER=0.21 ER=0.22

H2 CO CO2 CH4 CxHy

0

1

2

3

4

5

6

CO/CO2 H2/CO CH4/H2

Ch

arac

teri

stic

rat

io

(mo

l/m

ol)

ER=0.2 ER=0.21 ER=0.22

CO/CO2 H2/CO CH4/H2

45

evidenced by the temperature profiles. The high temperature in the gas phase enhances tar

cracking and hence increasing the hydrocarbon content.


Figure 5-6 shows the LHV and CGE of the tested cases.

Figure 5-6. LHV and CGE

Irrespective of the high content of CO, the LHV and CGE were negatively correlated with

the ER as a result of the low hydrocarbon and H2 contents. The gas yield was around 2.9-

3 Nm3/kg. The most promising case has an LHV of 7.9 MJ/Nm3 and CGE of 80%, which

are considerably higher than the values obtained for previously studied gray pellets and

black pellets (with similar gasification conditions). The main reason for this could be a

significantly higher hydrocarbon content. Considering the reported efficiency of the biocoal

production process (greater than 90% [33]), the total system efficiency is not considerably

affected compared with that of unpretreated biomass, still giving a high LHV gas.

70

75

80

85

6

6.5

7

7.5

8

0.19 0.2 0.21 0.22 0.23

CG

E (%

)

LHV

(M

J/N

m3 )

ER

LHV Efficiency

46

5.2.4 Syngas purity

At high temperatures, C2H6 decomposes into C2H4 and C2H2. The severity of this C-2

hydrocarbon cracking has been qualitatively correlated with tar thermal cracking [61-63].

Similarly, the syngas purity in this study was analyzed by comparing the C2H6/(C2H4 +

C2H2) ratio using online GC measurements (Figure 5-7).

Figure 5-7. Syngas purity based on severity of hydrocarbon cracking

The C2H6/(C2H4 + C2H2) ratio increased from run 1 to run 3 with decreasing gas phase

temperature, and hence the tar content is expected to increase. Most of the thermal cracking

reactions (hydrocarbons and tar) are expected to occur in the gas phase, and hence a

temperature drop at the top of the reactor significantly affects the tar content of gas. The

measured tar content of case 3 is 6.36 g/Nm3, which is lower than the values obtained for

previously studied gray pellets and black pellets. This could be due to the low volatile (or

hemicellulose) content and high fixed carbon (or lignin) content of biocoal.

However, although experiments were carried out at low ER values, the combustion zone

temperature increased with ER to reach 1400 °C, as observed in Figure 5-4. This high

temperature caused ash slagging, which resulted in discontinuous operation. Controlling

the temperature inside the gasifier to completely avoid ash slagging was difficult due to the

high carbon content in biocoal.

0

150

300

450

600

750

900

0

0.05

0.1

0.15

0.2

1 2 3

Tem

per

atu

re (

0 C)

C2H

6/(

C2H

4+C

2H

2)

Run

C2H6/(C2H4+C2H2) TexitTexitC2H6/(C2H4+C2H2)

47

5.3 Gasification of hydrothermal carbonized biomass with controlling measures for ash slagging

Two methods of biocoal gasification were tested to achieve temperature control, and hence

smooth operation in non-slagging mode.

• Gasification with a lower O2 content in the gasifying agent (use of a

considerable amount of CO2 and/or H2O in the gasifying agent)

• Gasification with a lower preheating temperature of the gasifying agent

To evaluate these methods, gasification was carried out using two types of biocoal pellets

in a fixed bed updraft HTAG system using the experimental conditions shown in Table 5-3.

Table 5-3. Experimental conditions

Run Biomass

type

Control

method for

ash slagging

Preheating

temperature

of the agent

(°C)

O2 content

in the

agent

(%)

H2O

content in

the agent

(%)

CO2

content in

the agent

(%)

ER Steam/

biomass

1 Spent grain

biocoal

Lower O2

content in

the agent

1000 10 10 5 0.28 0.38

2 12 8 4 0.29 0.27

3 1.5

1.5

75 2 0.03 2.2

4 80 1.5 0.04 2.6

5 Horse

manure

biocoal

Lower

preheating

temperature

of the agent

850 21

21

-

-

-

-

-

-

0.29 -

- 6 0.31

7 -

-

100

100

-

-

2.6

82


Figure 5-8 depicts the axial temperature distribution along the reactor height for each

gasification case.

2 Due to bed accumulation, the biocoal feed rate had to be reduced in the 8th run, which made the calculation of actual steam/biomass ratio infeasible.

48

Figure 5-8. Axial temperature profiles for (top) gasification of spent grain

biocoal and (bottom) gasification of horse manure biocoal

Similar to the previous experiments, the temperature profiles show a bed zone with a

significant temperature drop and a gas phase with an almost flat temperature profile. The

temperature profiles of horse manure biocoal gasification are relatively low compared with

those of spent grain biocoal gasification, which could be due to the lower preheating

temperature and low carbon content of the fuel. The temperature inside the gasifier was

kept below 1200 °C with both control measures. However, it should be noted that for

gasification of spent grain biocoal with the agent temperature around 1000 °C, an O2

0

200

400

600

800

1000

1200

-0.2 0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6

Tem

per

atu

re (

0C

)

Distance from the grate along the vertical axis (m)

1

2

3

4

0

200

400

600

800

1000

1200

-0.2 0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6

Tem

per

atu

re (

0 C)

Distance from the grate along the vertical axis (m)

5

6

7

8

49

content above 10% resulted in operational instability with high temperatures throughout

the reactor.

Increasing the steam content in the gasifying agent resulted in a low gas phase temperature

for spent grain biocoal gasification. In contrast, very similar gas phase temperatures were

seen for each run of horse manure biocoal gasification, irrespective of the steam content.


Figure 5-9 shows the syngas composition and important gas ratios obtained for gasification

of the two biocoal types.

Figure 5-9. Gas composition and gas ratios of (left) spent grain biocoal and

(right) horse manure biocoal

A comparatively low amount of combustibles was observed in the syngas obtained by spent

grain biocoal gasification due to the high agent flow rate applied. The amounts of all the

0

10

20

30

40

50

1 2 3 4

Syn

gas

com

po

siti

on

(%

)

H2 CO CO2 CH4 CxHyH2 CO2 CH4 CxHy

0

10

20

30

40

50

5 6 7 8

Syn

gas

com

po

siti

on

(%

)

H2 CO CO2 CH4 CxHyH2 CO2 CH4 CxHy

0

1

2

3

4

5

6

1 2 3 4

Ch

arac

teri

stic

rat

io (

mo

l/m

ol)

CO/CO2 H2/CO CH4/H2CO/CO2 H2/CO CH4/H2

0

1

2

3

4

5

6

5 6 7 8

Ch

arac

teri

stic

rat

io (

mo

l/m

ol)

CO/CO2 H2/CO CH4/H2CO/CO2 H2/CO CH4/H2

50

combustibles (H2, CO, and hydrocarbons) were higher with steam gasification of both

biocoal types due to less dilution with N2. The comparatively low amounts of hydrocarbons

obtained for high ER and high steam cases could be due to thermal cracking and steam

reforming reactions, respectively. This explanation is further validated by the CH4/H2

ratios.

The extraordinary performance of the Boudouard reaction (high CO/CO2 ratio) in run 4

and run 6 was due to the high temperature observed in the bed. The increasing H2/CO

ratio with steam gasification of horse manure biocoal was due to the dominance of the

water gas shift reaction. However, the H2/CO ratio did not improve with steam gasification

of spent grain biocoal because, even if the H2 content increased, the CO content was also

high.


Table 5-4 shows the LHV of syngas, the syngas yield, and the CGE of the tested cases.

Table 5-4. LHV of syngas, syngas yield, and CGE

Run Biomass type LHV (MJ/Nm3) Ysyngas (Nm3/kg) CGE (%)

1 Spent grain

biocoal

4.6 5.36 85.5

2 3.8 5.1 67.2

3 10.2 2.37 83.3

4 10.1 2.33 81.7

5 Horse manure

biocoal

7.1 1.1 38

6 8.6 1.4 56

7 11.1 0.65 34

83 10.4

In general, steam gasification gave a high LHV and low gas yield for both biocoal types.

Gasification of spent grain biocoal resulted in a lower LHV and higher syngas yield due to

the high agent flow rate applied. Overall, gasification of spent grain biocoal with a CGE

greater than 80% (with the exception of run 2) seems to be more feasible. The poor fuel

3 Due to bed accumulation, the biocoal feed rate had to be reduced in the 8th run, which made the calculation of syngas yield and CGE infeasible.

51

quality of horse manure biocoal (high ash, lower C, and high O content) limited the CGE

to a maximum of 56%.

5.3.4 Syngas purity

The tar composition of the tested cases in terms of BTEX and PAH, and the detailed

composition by compound are shown in Figure 5-10. There were no phenolic compounds

observed in the syngas tar from biocoal gasification, which could be due to the high

temperature. PAHs were dominant in the tar obtained from gasification of spent grain

biocoal, whereas BTEX was the dominant tar category for horse manure biocoal

gasification. This result is due to the considerable difference in the gas phase temperatures:

650–1000 °C and 450–600 °C, respectively. In general, benzene, toluene, and naphthalene

were the major tar components observed. Table 5-5 shows the total tar content and tar

dew point of each case.

Table 5-5. Total tar content and tar dew point

Run Biomass type Tar content (g/Nm3) Tar dew point (°C)

1 Spent grain

biocoal

10.98 113.1

2 1.02 46.8

3 9.38 104.2

4 4.46 97.2

5 Horse manure

biocoal

3.07 93.2

6 1.54 68.8

7 1.57 73.7

8 0.94 48.4

Gasification of spent grain biocoal resulted in more tar compared with gasification of horse

manure biocoal, with the exception of run 2, which had high temperatures throughout the

gasifier. Both the high ER and steam gasification cases showed lower tar contents due to

thermal cracking and steam reforming with both biocoal types. The dominance of PAHs

in the tar obtained by gasification of spent grain biocoal resulted in a high tar dew point of

around 100 °C (with the exception of run 2) compared with that of horse manure biocoal,

which ranged from 50–90 °C.

52

Figure 5-10. Tar composition (top) by category and (bottom) by compound

0

4

8

12

1 2 3 4

Tar

con

ten

t (g

/Nm

3)

PAH BTEX

0

4

8

12

5 6 7 8

Tar

con

ten

t (g

/Nm

3)

PAH BTEX

0

1

2

3

4

5

6

Co

mp

osi

tio

n (

g/N

m3)

1

2

3

4

0.0

0.2

0.4

0.6

0.8

1.0

1.2

Co

mp

osi

tio

n (

g/N

m3 )

5

6

7

8

53

Figure 5-11 demonstrates the particle content in the syngas and the particle size distribution

for each case of biocoal gasification.

Figure 5-11. Particle content and size distribution for gasification of (left) spent

grain biocoal and (right) horse manure biocoal

Similar to the tar content, the gasification of spent grain biocoal resulted in more particles

in general compared with the gasification of horse manure biocoal and was dominated by

the fine fraction (<1 µm). It has been reported that nucleation and condensation of tar

components result in the formation of fine particulates [64]. The comparatively high

content of coarse particles (>1 µm) observed with the gasification of horse manure biocoal

could be due to entrainment of ash, the content of which is comparatively high in horse

manure biocoal.

The high temperature of run 2 resulted in a high gas velocity inside the gasifier and,

ultimately, entrainment of ash and unburnt char in the gas stream, leading to the formation

of some coarse particles. Further, steam gasification of both biocoal types resulted in a

comparatively high content of coarse particles, which may also be due to entrainment of

ash and unburnt char.

0

500

1000

1500

2000

2500

1 2 3 4

Par

ticl

e co

nte

nt

(mg/

Nm

3 )

<0.1 0.1-1 1-10 >10Particle size µm

0

500

1000

1500

2000

2500

5 6 7 8

Par

ticl

e co

nte

nt

(mg/

Nm

3 )

<0.1 0.1-1 1-10 >10Particle size µm

54

CHAPTER VI

6 PARAMETERS OF FIXED BED GASIFICATION

It is important to understand the major parameters of the gasification process and their

interactions to achieve better control over the outcomes. Figure 6-1 shows the possible

interactions of three main parameters related to fixed bed gasification and the detailed

studies are summarized in the following subsections.

Figure 6-1. Interactions between fixed bed gasification parameters

6.1 Bed height, conversion, and conversion rate

To minimize the unreacted residue after gasification, the fuel feed rate should be sufficiently

higher than the gasification rate to maintain a fuel bed with a long enough residence time

for the biomass to react.

In general, the overall fuel conversion is given by

𝑋 =𝐹�̇�

�̇� (Eq. 6-1)

The fuel consumption rate can be calculated as

𝐹�̇� = �̇� − 𝑟𝑏ℎ𝜌𝑏𝐴 (Eq. 6-2)

Conversion

Conversion rate

Bed height

55

Furthermore, the residence time of the fuel in the gasifier bed is given by

𝑡𝑟 =𝐴𝐿

𝐹�̇�/𝜌𝑏 (Eq. 6-3)

The fuel conversion rate can be calculated as

𝑟 =𝑋

𝑡𝑟 (Eq. 6-4)

Figure 6-2 graphically represents the relation between these parameters.

Figure 6-2. Relations between residence time, conversion, and conversion rate

Figure 6-2 (left) shows that the overall fuel conversion is positively correlated with the

residence time of fuel in the bed (bed height), irrespective of the ER (the difference in ER

was too low to have a considerable effect). Fuel conversion of 90–99% was observed with

a residence time varying from 1 to 2 h. A conversion of around 90% has been reported for

biocoal gasification in a lab-scale entrained flow reactor at a temperature of 1000 °C [25].

Figure 6-2 (right) shows that the conversion rate decreases as the conversion progresses.

These behaviors are consistent with the results obtained with TGA.

6.2 Gasifier design parameters: Specific gasification rate and superficial velocity

To normalize the fuel consumption rate over the reactor dimensions, a general parameter

called the specific gasification rate is commonly used. This parameter is defined as a

measure of the fuel consumption rate per unit cross sectional area of the gasifier:

𝑟𝑠𝑔 =60 𝐹�̇�

𝐴 (Eq. 6-5)

0.85

0.9

0.95

1

0 50 100 150

Ove

rall

fue

l co

nve

rsio

n X

Fuel residence time in the bed tr (min)

0

0.004

0.008

0.012

0.016

0.02

0.85 0.9 0.95 1Fue

l co

nve

rsio

n r

ate

r(1

/min

)

Overall fuel conversion X

56

A similar normalized parameter for the gas production rate is the superficial gas velocity,

which is considered to be an important parameter that affects heat transfer, as well as the

char and tar production rates. The superficial gas velocity is defined as the gas generation

rate per unit cross sectional area of the gasifier, resulting in units similar to that of velocity

(m/s).

𝑢 =�̇�

𝐴 (Eq. 6-6)

A more commonly used variation of the superficial gas velocity is the hearth load, which is

defined using the practical units of the cross sectional area (cm2) and gas flow (m3/h). The

unit conversion shows the relationship between these two parameters:

𝐿𝐻 = 0.36 𝑢 (Eq. 6-7)

Table 6-1 shows these three design parameters for each run, along with the fuel conversion

rate.

Table 6-1. Design parameters

Run 𝑟

(min-1)

𝑟𝑠𝑔

(kg/m2 h)

𝑢

(m/s)

𝐿𝐻

(m3/cm2 h)

1 0.015 275.5 0.23 0.08

2 0.017 319.8 0.3 0.11

3 0.009 259.4 0.22 0.08

The specific gasification rate follows the trend of the conversion rate, which is inversely

related to the conversion and hence the residence time (or bed height). Further, the specific

gasification rate is positively correlated with the superficial gas velocity or hearth load.

Wood gasification in an updraft gasifier has been reported to have a specific gasification

rate and superficial velocity in the range of 170–190 kg/m2 h and 0.16–0.2 m/s, respectively

[65]. The observation of higher values in the present study could be due to two reasons:

the high-temperature operation and specific fuel type used. High-temperature operation

results in more gas with enhanced tar cracking. Furthermore, the high bulk density of

biocoal (around 800 kg/m3) accommodates a high fuel load. Combined with the high

energy value of biocoal, a considerable power output is expected, even in a compact reactor.

This means that coupling biomass pretreatment with HTAG can extend the capacity range

of fixed bed gasifiers, which is otherwise a limitation of this technology.

57

6.3 Bed pressure drop

Considering the effect of bed height on the performance of the gasification process, an

attempt was made to predict the bed pressure drop accurately. The total pressure drop

through a gasifier bed is the sum of the pressure drops through the particle bed and the

grate. Considering the high grate opening area of around 40% and low grate thickness of

around 6 mm, the grate resistance was found to be negligible in the present study.

A correlation has been proposed for pressure drop prediction for a bed with cylindrical

particles in turbulent flow conditions based on the equivalent tortuous passage method

[66].

∆𝑃 = 𝐿𝐾𝑡𝜌𝑢2 (1−𝜀)5

4⁄

𝜀3 (1

2𝑐+

1

𝑑)

54⁄

(𝜇

𝜌𝑢)

14⁄

(Eq. 6-8)

where

𝐾𝑡 = 112𝜀3.2 (Eq. 6-9)

The correlation can be extended to laminar and transitional flow conditions by

incorporating a viscous term.

The Hagen-Poiseuille equation for pressure drop in a laminar flow is

∆𝑃 =32𝜇𝑣𝑙

𝐷𝑒2 (Eq. 6-10)

Assuming an equivalent inclined passage with an angle to the direction of the mean flow,

the Hagen-Poiseuille equation can be rearranged as

∆𝑃 = 32𝜇𝑢𝐿(1−𝜀)2

𝜀3 (1

2𝑐+

1

𝑑)

2

(Eq. 6-11)

Combining and rearranging these equations gives the modified correlation as

(∆𝑃

𝐿) (

1

𝜌𝑢2) =

𝐾1

𝑅𝑒𝑝1/4 +

𝐾2

𝑅𝑒𝑝 (Eq. 6-12)

where 𝑅𝑒𝑝 =(

𝜌𝑢

𝜇)

𝐾3 (Eq. 6-13)

𝐾1 = 112 𝜀0.2(1 − 𝜀) (1

2𝑐+

1

𝑑) (Eq. 6-14)

𝐾2 = 32(1−𝜀)

𝜀3 (1

2𝑐+

1

𝑑) (Eq. 6-15)

𝐾3 = (1 − 𝜀) (1

2𝑐+

1

𝑑) (Eq. 6-16)

58

The porosity of a cylindrical particle bed is given by [67]

𝑙𝑛 𝜀 = 𝛷5.58 𝑒𝑥𝑝[5.89(1 − 𝛷)] 𝑙𝑛0.4 (Eq. 6-17)

where the sphericity is defined as

𝛷 =(36𝜋𝑉𝑝

2)1/3

𝑆𝑝= [

2.25×(𝑐

𝑑)

2

0.5+(𝑐

𝑑)3

]

1/3

(Eq. 6-18)

The particle size changes along the bed due to the reaction occurring in the gasifier.

Neglecting the wall effect and the thickness effects on porosity variation due to the scale

of the reactor, the particle size of the reacting bed can be calculated by applying mass

balance for one particle.

𝑚𝑐ℎ𝑎𝑟 = (1 − 𝑋)𝑚𝑏𝑖𝑜𝑚𝑎𝑠𝑠 (Eq. 6-19)

Neglecting fragmentation due to the high density of pellets and taking the surface

conversion into account by assuming the density of biomass particle is constant throughout

the conversion period, the volume of a char particle is given by

𝑐𝑑2 = (1 − 𝑋)𝑐0𝑑02 (Eq. 6-20)

Further, assuming all the dimensions of the particle are reduced by a uniform thickness h

for a certain time period [68], the new dimensions are given by

𝑐 = 𝑐0 − 2ℎ (Eq. 6-21)

𝑑 = 𝑑0 − 2ℎ (Eq. 6-22)

Resulting in

𝑐 − 𝑑 = 𝑐0 − 𝑑0 (Eq. 6-23)

By knowing the initial particle size distribution and the expected conversion, it is possible

to calculate the size distribution of the converted particles, which is then used to calculate

the distribution of the sphericity, and ultimately the bed porosity. These values, along with

the flow properties such as velocity, density, and viscosity, can be used in in the developed

correlation to calculate the pressure drop along the gasifier bed. The calculation procedure

is discussed in detail in supplement IV.

For cylindrical particles, some researchers [69] have obtained a relationship with the

sphericity and Ergun constants as given below.

59

∆𝑃 = 𝐿 [𝐴′(1−𝜀)2𝜇

𝜀3𝛷2𝑑𝑝2 𝑢 + 𝐵′

(1−𝜀)𝜌

𝜀3𝛷𝑑𝑝𝑢2] (Eq. 6-24)

where

𝐴′ = 150

𝛷3/2 (Eq. 6-25)

𝐵′ = 1.75

𝛷4/3 (Eq. 6-26)

By incorporating the shrinking effect, as discussed above, this equation can be improved

for a reacting bed by using modified Ergun constants.

The pressure drop results calculated with the developed correlation and the above empirical

correlation were compared with experimental data, as shown in Figure 6-3.

Figure 6-3. Comparison of pressure drop results

The method formulated in the present study gives a better approximation, with a maximum

error of only 7% with respect to the two performed runs. The available empirical equation

was able to predict the pressure drop within a 20% interval after the shrinking effect was

taken into account. Graphical representations of the correlation constants of the modified

equation for typical biomass pellet sizes are available in supplement IV.

0

300

600

900

1200

1500

Run 1 Run 2

Bed

pre

ssu

re d

rop

(P

a)

Experimental

Modified equation

Empirical correlation

60

CHAPTER VII

7 INTEGRATION OF BIOMASS GASIFICATION PROCESS WITH THE STEEL INDUSTRY

7.1 Application of syngas in the furnace

7.1.1 Comparison of experimental and modeling results

To check the validity of the developed models, the pyrolysis model and the gasification

model were compared with the experimental data for the gasification of Salix and vine

prunes [36]. Figure 7-1 shows a comparison of the C, O, and H distribution in the char

and pyrolysis gas predicted by the pyrolysis model and the reported experimental values at

a pyrolysis temperature of 380 °C.

Figure 7-1. Experimental and model results for the C, O, and H distribution of

char and pyrolysis gas for pyrolysis of (left) Salix and (right) vine prunes

Table 7-1 shows a comparison of the dry syngas composition predicted by the gasification

model and the reported experimental values at a gasification temperature of 1100 °C.

0

20

40

60

80

100

C O H

Co

mp

osi

tio

n (

%)

Component

Char-experimental Char-model

Pyrolysis gas-experimental Pyrolysis gas-model

0

20

40

60

80

100

C O H

Co

mp

osi

tio

n (

%)

Component

Pyrolysis gas-model Pyrolysis gas-experimental

Char-model Char-experimental

61

Table 7-1. Experimental and modeling results of dry syngas composition

Component Biomass type: Salix

S/C molar ratio: 1.35

Biomass type: Vine prunes

S/C molar ratio: 1.5

Experimental Model Experimental Model

CO (%) 30 30.3 27.6 27.9

CO2 (%) 11.4 11.9 12.4 13.3

H2 (%) 56.3 55.4 57 57

CH4 (%) 2.3 2.4 3 1.8

As shown in Figure 7-1 and Table 7-1, the model predicts the composition of the pyrolysis

products and dry syngas fairly well. Therefore, the model is considered suitable for

evaluating the integration cases.

7.1.2 Gasifier system performance

Table 7-2 summarizes the pyrolysis temperatures required for sustaining the process in

each integration case.

Table 7-2. Pretreatment condition for each case

Case

number

Integration option Pyrolysis

temperature

(°C)

1 Integration with no heat recovery 375

2 Integration with low-temperature heat recovery for steam superheating

up to 400 °C

370

3 Integration with low-temperature heat recovery for biomass predrying 360

4 Integration with low-temperature heat recovery for steam superheating

up to 400 °C and biomass predrying

355

5 Integration with high-temperature heat recovery for steam superheating

up to 700 °C and biomass drying

345

According to the results summarized in Table 7-2, any form of flue gas heat recovery (high

temperature or low temperature) reduces the required pyrolysis temperature. By

62

incorporating high-temperature heat recovery (for steam superheating and biomass drying),

a 30 °C reduction of the pyrolysis temperature was achieved. The reduced pyrolysis

temperature resulted in a maximum 20% higher char yield and syngas yield. Due to the

narrow range of pyrolysis temperatures applied and the fixed S/C molar ratio used in the

gasification process, the changes of the syngas compositions were not that significant.

Approximately 55% H2, 31% CO, 11% CO2, and 3% CH4 is observed, which accounts for

almost 90% combustibles, giving an LHV of 10.8 MJ/Nm3. The specific biomass input

requirement was reduced by 16% at maximum heat recovery. Figure 7-2 shows the

efficiency and heat demand of the gasifier system.

Figure 7-2. (left) Gasifier system efficiency and (right) heat demand of gasifier

system

The gasifier system efficiency increased with the flue gas heat recovery. Low-temperature

flue gas heat recovery for biomass predrying is more feasible than steam superheating due

to high efficiency. Simultaneous low-temperature heat recovery for predrying and steam

superheating gives higher efficiency. However, high-temperature heat recovery, which is

13% more efficient than the case with no heat recovery, is the most energy efficient option.

The gasifier system energy requirement is reduced by 5% from case 1 to case 5 due to the

lower pyrolysis temperature. The endothermic steam gasification process consumes

approximately 50% of the total heat. Further, the gasifier heat demand is balanced by the

69.9 71.7 75.978.2

83.3

0

20

40

60

80

100

1 2 3 4 5

Gas

ifie

r sy

ste

m e

ffic

ien

cy (

%)

Case number

174 171 161 157 148

57 54 47 44 37

385 365 386 366 338

133 156 132 155 185

0

100

200

300

400

500

600

700

800

1 2 3 4 5He

at d

em

and

pe

r u

nit

th

erm

al e

ne

rgy

ou

tpu

t (k

J/M

J)

Case number

Predrying and drying Pyrolysis

Gasification Steam generation

63

degree of steam superheating applied. The predrying, drying, and pyrolysis heat demands

follow the trend of the specific biomass feed requirement.

7.1.3 Heat integration performance

To determine how well the proposed heat integration options fit with the theoretically

achievable limit, a pinch analysis was performed, which is exemplified by the GCCs for

cases 1 and 5 (per MJ output) in Figure 7-3.

Figure 7-3. GCCs for cases 1 and 5

The pinch analysis showed few features. No external heat requirements for either case,

whereas case 1 possessed a considerable amount of waste heat (cooling load of 292 kJ/MJ),

and two thirds of the waste heat was already recovered in case 5, accounting for the lower

amount of waste heat (cooling load of 95 kJ/MJ). Theoretically, the system is capable of

handling biomass with a higher moisture content.

The critical moisture content of biomass for the two extreme pyrolysis conditions

considered in this study is plotted in Figure 7-4 at a range of flue gas temperatures.

0

400

800

1200

1600

2000

0 1 0 0 2 0 0 3 0 0 4 0 0 5 0 0 6 0 0 7 0 0

Shif

ted

tem

per

atu

re (

oC

)

Net heat flow (kJ/MJ)

Case 1 Case 5

64

Figure 7-4. Critical moisture content

The system can handle biomass with up to 55% moisture content if the pyrolysis

temperature is low (around 345 °C). For fairly dry biomass (e.g., wood pellets), low-

temperature heat recovery is adequate under this pyrolysis condition. The critical moisture

content can be increased to 65% if the pyrolysis temperature is increased to 375 °C. Under

this condition, low-temperature flue gas heat recovery is sufficient for processing typical

woody biomass.

7.1.4 Furnace performance

The adiabatic temperature is a critical parameter that can affect the steel quality. Because

the adiabatic temperature of syngas combustion (~2035 °C) is higher than that of NG

combustion (1914 °C), a small portion of the flue gas was recycled back to the furnace to

maintain a similar adiabatic flame temperature. Further, the heat input to the furnace was

fixed, so that the flue gas temperature was similar to that of the NG firing case.

Because the syngas compositions do not differ significantly, cases 1–5 were considered as

one case for evaluating the furnace performance.

5

15

25

35

45

55

65

75

200 250 300 350 400 450 500 550 600 650 700 750

Cri

tica

l mo

istu

re c

on

ten

t (%

)

Flue gas temperature (oC)

Pyrolysis at 345°C Pyrolysis at 375°C

65

Figure 7-5. (left) Input flow rates to the furnace and (right) energy distribution

of the furnace

Figure 7-5 (left) shows the fuel, air, and recycle stream flow rates for the furnace. The low

energy density of the syngas results in more than three times the gas flow rate compared

with the NG case. However, the air flow rate is reduced by more than 20% as a result of

the stoichiometric difference in the combustion of the two fuels. The increase of the total

input flow rate to the furnace is only 9%. However, the higher composition of H2 in the

syngas may affect the burning velocity, which would ultimately require a dedicated burner

design.

Figure 7-5 (right) shows the energy distribution of the furnace. It should be noted that the

example furnace is fairly old, and the design is not optimized for energy efficiency, resulting

in high furnace heat losses. Irrespective of the similar flue gas temperature, the flue gas loss

for syngas firing is decreased by 10% compared with that for NG firing due to the low flue

gas flow rate, as shown in Table 7-3. As a result, the furnace efficiency can be improved

from 37.3% to 38.7% by the fuel switch. The flue gas compositions are shown in Table

7-3.

33116

379294

0 38

0

100

200

300

400

500

Natural gas SyngasFue

l, ai

r, a

nd

re

cycl

e f

lue

gas

fl

ow

rat

e(N

m3 /

ton

of

ste

el)

Furnace fuel

Fuel flow Air flow Recycle flow

485.1 485.1

354.9 354.9

461.4 414.1

0

300

600

900

1200

1500

Natural gas Syngas

Ene

rgy

dis

trib

uti

on

of

furn

ace

(M

J/to

n o

f st

ee

l)

Furnace fuel

Flue gas losses Furnace losses Sensible heat to iron powder

66

Table 7-3. Flue gas composition and flow rate

Component NG case Syngas case

H2O (%) 16.9 19.4

CO2 (%) 9.1 14.6

O2 (%) 1.7 1.6

N2 (%) 72.3 64.4

Total flow

(Nm3/ton of steel)

415 398

Syngas firing gives more CO2 and H2O (and less O2 and N2) in the flue gas compared with

NG firing. This is due to the principal compositional difference and the combustion

stoichiometry of the two fuels. The total flue gas flow rate is also reduced by 4%, and hence

the fuel switch will not affect the size of the flue gas handling system. Higher CO2 and H2O

contents in the flue gas may be advantageous in heat transfer due to increased radiation

effects with high flame emissivity, which may lead to increased capacity of the furnace after

fuel switching. Considering biomass as CO2 neutral, the on-site CO2 emission reduction by

avoiding the burning of fossil fuel NG is 74 kg/ton of steel produced.

67

Fig

ure

7-6

. E

nerg

y b

ala

nce p

er

ton

of

steel

pro

cess

ed

(case

5)

68

7.1.5 Energy balance and the overall system performance

Figure 7-6 demonstrates the energy balance of integration case 5, which is the most

promising integration case studied in terms of energy. This Sankey diagram was developed

by converting each material stream into its energy value, including recycle streams that

represent the energy recovery measures applied. Exhaust streams, such as flue gases, were

calculated by balancing the input and output streams.

The energy balance revealed that almost 75% of the heat recovered from the furnace flue

gas was used for biomass drying, and only 25% was required for steam superheating due to

the high moisture content of the woody biomass used.

Figure 7-7 shows the integrated system efficiency of the analyzed cases.

Figure 7-7. Integrated system efficiency

The maximum integrated system efficiency was achieved in case 5 with high-temperature

flue gas heat recovery by biomass drying and steam superheating. The efficiency was

improved by 5%.

27.0 27.8 29.4 30.3 32.3

0

10

20

30

40

1 2 3 4 5

Inte

grat

ed

sys

tem

eff

icie

ncy

(%

)

Case number

69

7.2 Co-production of bio-coke and bio-H2

Figure 7-8 shows the integrated system efficiency of the analyzed production scenarios.

Figure 7-8. Efficiency comparison for various scenarios

The maximum integrated system efficiency was achieved in scenario 4, which considered

full integration, including coke and H2 production. The efficiency was improved by 10%.

A considerable efficiency improvement was also achieved if only coke or H2 was produced

along with syngas (scenarios 2 and 3, respectively). H2 production, which is a physical

separation process, was more efficient than coke making, which had some losses due to

burning of the produced coke gas. If the coke gas were instead used in the H2 separation

process, an even higher efficiency could be expected.

3238 40 42

0

10

20

30

40

50

1 2 3 4

Inte

grat

ed

sys

tem

eff

icie

ncy

(%

)

Scenario number

70

CHAPTER VIII

8 CONCLUSIONS

In this thesis, advanced gasification of biomass and waste for substitution of fossil fuels in

steel industry heat treatment furnaces was performed. The conclusions from the

experimental investigation of advanced biomass gasification and the assessment of the

integration of the advanced gasification system with steel heat treatment furnaces are

summarized here.

Conclusions from experimental investigation of advanced biomass gasification

TGA of pretreated biomass:

The decomposition characteristics of unpretreated biomass (gray pellets), steam-exploded

biomass (black pellets), and hydrothermal carbonized biomass (biocoal) were analyzed with

TGA at a constant heating rate in a N2 atmosphere, followed by a constant temperature in

a N2/CO2 atmosphere. The TGA experiments provided a necessary preliminary

understanding for real thermal applications.

Both types of pretreated biomass feature early thermal decomposition of

hemicellulose due to depolymerization during pretreatment.

Steam explosion pretreatment results in a lower thermal stability of the cellulose

fraction, possibly due to early decomposition of the hemicellulose fraction and

depolymerization of the cellulose fraction during pretreatment. The opposite trend

was observed for biocoal.

The shoulder attached to the cellulose peak was attributed to the decomposition of

lignin globules redeposited on the surface after hydrothermal pretreatment. This

overlapping and high intensity lignin decomposition can affect the cellulose

decomposition adversely.

Steam-exploded biomass showed a comparatively high char gasification rate,

whereas hydrothermal carbonized biomass showed a low rate.

71

Gasification of pretreated biomass:

Gasification experiments of typical unpretreated biomass pellets (gray pellets), steam-

exploded biomass pellets (black pellets), and hydrothermal carbonized biomass pellets

(biocoal) were performed with air and steam in an updraft HTAG unit.

Gasification of pretreated biomass resulted in a more divergent temperature profile

in the gasifier due to the high fixed carbon and lower volatile content.

An effective Boudouard reaction and suppressed water-gas shift reaction,

hydrocarbon reforming, and cracking reactions are features of the gasification of

black pellets, which results in more CO and hydrocarbons, but less H2 in the syngas.

Steam gasification of black pellets is more feasible if a syngas with a high energy

value is desired, whereas if a higher H2 yield is preferred, gasification of gray pellets

is more attractive.

As the carbon content of biocoal is too high (e.g., 66% in spent grain biocoal), air

gasification (even with a low ER of around 0.2) resulted in a too high temperature,

which created ash slagging problems (coupled with the high ash content of biocoal)

and hence operational instability. Dilution of the gasifying agent (with CO2 and/or

H2O) or reducing the agent temperature showed better stability at a high ER.

However, dilution of the gasifying agent resulted in a low LHV of the syngas and a

high gas yield.

Overall, air gasification of pretreated biomass produced syngas with an LHV of

around 7–9 MJ/Nm3 (with the exception of the diluted cases, which gave an LHV

of around 4–5 MJ/Nm3), whereas steam gasification of pretreated biomass

produced better syngas with an LHV of around 10–11 MJ/Nm3. Spent grain biocoal

gasification was the most efficient (CGE of 80–85%), whereas horse manure biocoal

gasification was the least efficient.

Depending on the temperature, the tar composition of the syngas obtained with

gasification of steam-exploded biomass and horse manure biocoal was dominated

by BTEX compounds, whereas that obtained from spent grain biocoal was

dominated by PAHs (similar to unpretreated biomass). Phenols were less abundant

in the tar obtained from HTAG, which was readily converted to benzene and

naphthalene. A low tar dew point of around 100 °C was observed in all cases, except

72

for the gasification of horse manure biocoal, which showed an even lower tar dew

point of around 50–90 °C.

The syngas obtained by gasifying horse manure biocoal resulted in significant

amounts of coarse particles, whereas that for spent grain biocoal predominately

contained fine particles.

Important parameters for fixed bed gasification and prediction of bed pressure

drop:

The interaction between bed height, conversion, conversion rate/specific gasification rate

and superficial velocity was identified. A correlation for pressure drop prediction in a

gasifier bed with cylindrical particles was proposed and verified with experimental data.

Graphical representations for commonly available pellet sizes were developed based on the

correlation to simplify the calculations.

The fuel conversion is positively correlated with the fuel residence time in the bed

and hence with the bed height. The conversion rate (and specific gasification rate)

drops as the conversion progresses. The superficial velocity (or hearth load) is

positively correlated with the specific gasification rate.

The superficial velocity (or hearth load) and specific gasification rate of biocoal

gasification are higher than the reported values of updraft gasifiers due to the high-

temperature operation and high bulk density of biocoal. Coupled with the high

energy value of biocoal, it should be possible to extend the capacity range of fixed

bed gasifiers.

The developed correlation predicts the bed pressure drop fairly well. The

correlation/plots can be used as an initial guide for selecting a suitable pellet size

and designing a grate.

Conclusions from integration of gasification with the steel industry

An integrated system based on a multi-stage gasification process and a steel heat treatment

furnace was proposed, and a model for the proposed integrated system was developed using

Aspen Plus and validated with experimental data. Five heat recovery cases were evaluated

73

in terms of energy efficiency, and a pinch analysis was used to evaluate the heat integration

performance. The operational limits of the system were identified in terms of the critical

moisture content of the feedstock. A preliminary study on the co-production of bio-coke

and bio-H2 was also performed.

The process conditions of the furnace can be achieved with syngas firing without

any major changes to the furnace, combustion technology, or flue gas handling

system. The high H2 content of the syngas may require revamping of the burner

system. Further, a flue gas heat recovery network for predrying and/or steam

superheating should be introduced.

Both the furnace and gasifier system efficiencies were increased by integration, with

a maximum 13% increase of the gasifier system energy efficiency and 10% decrease

of flue gas loss from the furnace. Moreover, the flue gas possessed more radiative

gases.

The most energy-efficient integrated system gave an overall efficiency of

approximately 32%, and 74 kg CO2/ton of steel could be eliminated by the fuel

switch.

Although the studied biomass feedstock had around 40% moisture, the system can

sustain up to 65% feedstock moisture content.

Co-production of bio-coke and bio-H2 increase the overall efficiency significantly.

74

CHAPTER IX

9 FUTURE WORK

This study revealed some important information about biomass pretreatment, gasification,

and industrial applications. However, future work is encouraged regarding the following

aspects.

An ash melting problem was observed with air gasification of high-carbon biocoal.

Although dilution of the gasifying agent and lowering the agent temperature are

possible solutions for this problem, the application of a leaching method coupled

with HTC pretreatment would lower the ash content of the produced biocoal and

may facilitate air gasification.

Syngas combustion is expected to be significantly different from NG combustion

due to the high content of H2 in the syngas, high content of radiative gases in the

flue gas, etc. Further analysis of the effects on burning velocity, flame emissivity,

etc. are required to design a suitable burner and understand the heat transfer

properties. The emissions from syngas combustion, particularly NOx emissions,

also need to be studied. Further, experimental trials are required to evaluate any

effects on the quality of the produced iron.

Preliminary studies on coke making and H2 separation revealed that the overall

efficiency of the system could be significantly improved with these value-added

products. A detailed study of these processes is required, including experimental

trials to determine the suitability of the produced coke for metallurgical purposes.

A detailed environmental and economic analysis is necessary before potential

industrial application of this system.

75

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Advanced Gasification of Biomass/Waste for …953814/...i Abstract With the current trend of CO 2 mitigation in process industries, the primary goal of this thesis is to promote biomass