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THE AMERICAN SOCIETY OF MECHANICAL ENGINEERS
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BLACK LIQUOR-GASIFIER/GAS TURBINE COGENERATION
III 1 10111 11 111111111111
Stefano Consonni
Departimento di Energetica
Politecnico di Milano
20133 Milano, Italy
Eric D. Larson'
Center for Energy and Environmental Studies
School of Engineering and Applied Science
Princeton University, Princeton, NJ 08544
BREAK
Niklas Berglin
Department of Heat and Power Technology
Chalmers University of Technology
412 96 Goteborg, Sweden
ABSTRACT
The kraft process dominates pulp and paper production worldwide.
Black liquor, a mixture of lignin and inorganic chemicals, is generated in
this process as fiber is extracted from wood. At most knit mills today,
black liquor is burned in Tomlinson boilers to produce steam for on-site heat
and power and to recover the inorganic chemicals for reuse in the process.
Globally, the black liquor generation rate is about 85.000 MW (or 0.5
million tonnes of dry solids per day). with nearly 50% of this in North
America. The majority of presently-installed Tomlinson boilers will reach
the end of their useful lives during the next 15 to 20 years. As a
replacement for Tomlinson-based cogeneration. black liquor-gasifier/gas
turbine cogeneration promises higher electrical efficiency, with prospective
environmental, safety, and capital cost benefits for kraft mills. Several
companies arc pursuing commercialization of black liquor gasification for
gas turbine applications. This paper presents results of detailed performance
modeling of gasifier/gas turbine cogeneration systems using different black
liquor gasifiers modeled on proposed commercial designs.
INTRODUCTION
Black liquor is the lignin-rich byproduct of fiber extraction from wood
in kraft pulp production. In 1994. the U.S. pulp and paper industry
consumed 1.2 El ( I 0" 3), or 38.000 MW, of black liquor. This exceeded
the 1.0 El of total fossil fuel used by the industry (AFPA, 1996). The
industry bums black liquor today in Tomlinson recovery boilers that feed
back-pressure steam turbine cogeneration systems supplying process steam
and electricity to mills. A critical second objective of Tomlinson boilers is
the recovery of pulping chemicals (sodium and sulfur compounds) from the
black liquor for reuse (Adams et al., 1997). Technologies for gasifying
black liquor are under development to replace Tomlinson boilers. This
paper assesses the prospective energy performance of black liquor
gasifier/gas turbine technologies that are targeted for commercial
application by early in the next decade.
This work is motivated by the interest of the pulp and paper industry in
black liquor gasification technology arising from prospective improvements
in energy, environmental, safety. and capital-investment characteristics
compared to Tomlinson technology (Larson et al.. 1996: Industra. 1996).
The industry sees some urgency in realizing the promise of gasification
technology because its fleet of Tomlinson boilers is aging. Some 80% of
all currently-operating recovery boilers in the U.S. were built or rebuilt
before 1980 (Fig. 1). Typical recovery boiler lifetimes are around 30 years
(Weyerhaeuser et al, 1995), so most of the recovery boilers in the U.S. (and
Canada, as well) will need major attention or replacement over the next 15
to 20 years. As the most capital-intensive of U.S. manufacturing industries,
the papa industry is keenly interested in minimizing capital expenditures for
replacements, but would like to do so while simultaneously meeting
increasingly stringent environmental regulations, improving energy
performance. and reducing the risk of major recovery boiler explosions that
occur periodically (Grant. 19%). The need to replace Tomlinson boilers
over the next two decades provides a once-in-thirty-years window of
opportunity for the industry to make a major technology change (to
gasification) to address these concerns.
Black liquor gasifier development received considerable attention in the
U.S. in the early 1980s (Kelleher. 1985: Empie. 1991). but the prospects for
commercializing technology appear considerably improved at present. In
fact, the first fully-commercial black liquor gasifier (Smith et al.. 1996) has
begun startup testing at Weyerhaeuser's New Bern mill in North Carolina.
This unit will provide the mill with incremental capacity for chemicals
recovery from black liquor, with the gasifier product gas being burned in a
boiler. For the longer-term, there is industry interest in full replacement of
Tornlinson-based cogeneration with gas turbine-based systems.
Black liquor gasification technologies under development can be
classified according to operating temperature, or equivalently, according to
the physical state in which the majority of the inorganic content of the feed
black liquor leaves the reactor. High-temperature gasifiers operate at 950°C
or higher and produce a molten smelt of inorganic chemicals. Lowtemperature gasifiers operate at 700°C or lower in order t6 insure that the
inorganics leave as dry solids. Kvaerner and Noell are two companies
developing high-temperature gasifiers for gas turbine applications. ABB
Author to whom correspondence should be addressed.
Presented at the International Gas Turbine & Aeroengine Congress & Exhibition
Orlando, Florida — June 2—June 5, 1997
This paper has been accepted for publication in the Transactions of the ASME
Discussion of it will be accepted at ASME Headquarters until September 30,1997
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35000
Fs'
In
Developing Regions
P. 30000 3
'0
- 5000
0Japan
0 Europe, Aus/NZ
25000
-
X
• North America
- 4000 a.
0.
CO
15 20000
"5
Ct)
u
m
After 1988,
3" O
only units in
.
the U.S.A. are
included here. - 2000 o
63
10000
.o
_
1000 2
I
0
..
fl
0
1935 1940 1945 1950 1955 1960 1965 1970 1975 1980 1985 1990 1995 2000
Startup Year
Fig. 1. Age and capacity of 540 operating recovery boilers. (Record is incomplete after 1988.) Cumulative total black liquor processing capacity
shown here is 390,000 tonnes of dry solids per day (tds/d), or 64,000 MW
This total represents about 75% of total global capacity in 1995.
gasifier, the gas turbine, the steam cycle, the heat exchanger network), but
treats simplistically other components which, despite their technological
relevance, have minor impacts on the plant mass and energy flows (e.g., a
hot gas filter). The modeling work reported here ignores altogether the
important impacts that gasification of black liquor might have on chemical
recovery steps outside of a pulp mill's cogeneration plant, e.g. on the lime
kiln (NUTEK. 1992; Industra, 1996; Larson et al., 1996). These impacts
have little or no bearing on the heat and mass balances of the gasifierfgas
turbine system, but would need to be considered at any actual mill.
and MTCI are developing low-temperature gasifiers for such applications.
Pressurized operation (at about 25 bar) is being pursued for the hightemperature gasifiers. ABB has proposed a milder pressurization (perhaps
up to 5 bar), and MTCI is an atmospheric-pressure design. Additional
details regarding these gasifier technologies are available elsewhere (Larson
et al., 1996; Grace and Timmer, 1995; Stiggson and Hessebom. 1995;
Larson et al., 1996; Dahlquist and Jacobs, 1992; Aghamohammadi et at.
1995). Gasifier development by companies other than these four is at a
much-reduced level of intensity (Industra, 1996).
Results of calculations of the MI-load performance of gasifier/gas
turbine systems incorporating gasifier designs based on three of the four
designs under commercial development are reported here. Calculations
based on the MTC1 gasifier design have yet to be completed.
Reference Black Liquor
Mass and energy balance calculations with the model require only a
correct estimate of the heating value, the dry solids content and ultimate
composition of the input fuel (Consonni and Larson. 1996: Larson et al..
1996). The dry solids content and ultimate composition determine the flux
of atomic species into the power plant. The heating value allows closing the
heat balances. Black liquor characteristics adopted for all calculations are
shown in Table I. Dry solids can include up to I% chlorine (not shown in
Table I). Minor chlorine content has practical relevance, but a negligible
impact on heat and mass balances, and so is neglected here.
MODELING BLACK - LIQUOR COGENERATION
General Aporoach
The starting point for the calculations is a computation model originally
developed to predict the full-load, design-point performance of complex gassteam power cycles (Consonni, 19921. In the model, the system of interest
is defined as an ensemble of components, which can be of twelve basic
types: pump, compressor. combustor, gas turbine expander, heat exchanger.
mixer. splitter, steam cycle (which includes HRSG. steam turbine, pumps
and all auxiliaries), air separation plant, shaft (which accounts for
turbomachinery interconnections and electric losses), saturator, and
chemical converter. Operating characteristics and mass and energy
balances of each component are calculated sequentially and iteratively until
the conditions (pressure, temperature, mass flow. etc.) at all
interconnections converge toward stable values. The basic model is
extensively described elsewhere (Consonni et at. 1991: Lozza, 1990; Loan
et al., 1993; Chiesa et al., 1993; Chiesa et al.. 1994].
The model accurately simulates full-load performance of plant
components that are crucial to the energy and mass balances (e.g., the
Table 1. Reference black liquor, with 75% dry solids (ds) content
Heat Value (Makg cis)
Ultimate Composition (dry weigh %)
H
0
5
Na
K
Hater
Net'
12.409
14 363
2.5
34.4
3.7
18.6
37.2
3.6
a) The on hewing value (NHV) is commonly used amide of North Amnia. It
accounts cc unrecovaed latent heat at water vapor and umansweed beat ded up in
reduced sulfur compounds leaving • reactor in practice. The M4V is calculated bat
using the ormula of Adams, a al. (1997): NHV (Mlitg) • 1•CHV (hU/kg) •
2.44x0 WAN - 12.9x(78/32)xSs%. where H and S are the mass fractions of hydrogen
and sulfur in the ca cao) black liquor and 11, is the assorted efficiency of sat reduction.
101% sulfur reduction is assumed in the above table.
2
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950
Table 2. Equilibrium and vendor-reported gas compositions.'
Ar
Equilibrium (mol tr) 0.47
Vendor1mb( 'kr
-
CH,
CO
0.05
11.16
2.5
10.5
CO,
H,
NO
H,S
12.94 20.27 14.58 0.74
13.5
19.4
14.6
Mill process steam and power demands,
excluding cogen plant consumption.
N,
r,
39.79
1.4
.0 900
38.1
0
(a) Values assume 72% dry solids bla k liquor with composition (weight % di). C.
a.
to
• 1350
3604: H 3.3. O. 34.23: 5.4,9 .. Na. 17.72: K. 1 . Liquor feed at 115Y. air at SCOT.
stactor iemperarun and press= of 71:0C and I tor. 99%C conversion, heat loss of
1.065% of feed HHV,
(hl Penonal communication from D. Task ABB. Windsor. CT. ha 1996. .
Swedish 'best practice' mill, 1988
(AF-IPN. 1993)
o U.S. mill (anonymous)
800
Older U.S. mill
(Subbiah at al., 1995)
Gasification
#3•
.0 750
The high sodium and potassium content of black li quor help catalyze
g asification reactions, so that even at relativel y low temperatures, the
composition of the product gas is close to equilibrium, except for methane
and hydrogen sulfide (1-1 2S). e.g.. see Table 2. This simplifies modelin g the
gasifiers for heat and mass balance purposes. Since discrepancies in H :S are
much smaller than those with CH,, and the concentrations of H :S are small
regardless. i gnoring the discrepancy in FI 2S yields a negligible error in the
overall heat and mass balance of the plant. Corrections for non-equilibrium
methane content are made based on compositions reported b y Bcrglin
(1996) and Backman and Solmenoja (1994). While there is some
uncertaint y as to what the methane concentration would be in practice.
sensitivity calculations confirm that variations around the concentrations
assumed here have little impact on overall plant heat and mass balances.
700
Swedish mills
(Berglin, 1996)
•
El 650
P 2 is
Newer U.S. mill
(Larson, 1992)
#1 .
Swedish 'best practice' mill, 1988
(AF-IPK, 1993)
cs 600
•
Swedish model mill. 2000 (Wamquist. 1989)
CO
CD
° 550
2
Gas Turbine
o Integrated unbleached !craft linerboard mills
• Bleached icraft market pulp mills
500
An important feature of the computation model used here is its
capability to predict the hill-load desi gn-point performance of actual
commercial gas turbines without the, need for proprietar y manufacturers
data. For specificit y. the Siemens KWU 64.3a is selected here for all
calculations. The crucial parameters that determine turbomachiner y
efficiencies and cooling flows in the model are calibrated such that the
model reproduces published performance of the actual commercial en gine
operating on natural gas fuel (Table 3). The good match with quoted
performance provides a basis for confidence in the results for the
gasification-based calculations.
The same gas turbine is selected here .for all calculations to help
illustrate intrinsic differences among alternative gasifier desi gns. The black
liquor fuel requirements for a 64.3a turbine correspond (approximatel y) to
the black li quor flow at typical modern kraft mills having production
capacities ran ging from 1100-1800 air-dry metric tonnes of final product
per day. A perfect matching between the quantity of gasified black li quor
available at a mill and the fuel requirements of a specific gas turbine will be
rare, because the fuel
availability is largely
determined b y
Table 3. Calculated and vendor-quoted
considerations related
performance of the Siemens KWU 64.3a
to pulp and paper
natural gas simple c ycle gas turbine.
production (and,
Predicted
Quoted
possibly, the costInput Assumptions
effectiveness
• of
1280
Turbine inlet T.
not avail.
generating
excess
Pressure ratio
16.6
16.6
power for external
Inlet air now. kg/s
191.9
191.8
sale),
while the size of
Calculated input*
the gas turbine is
Exhaust flew.
196.3
194
determined by the few
563
Exhaust T. *C
565
models available on
36.4
Efficiency.%
36.8
the market. A practial
Net Power, MW
70.1
70
operating strategy at a
00
5.0
10.0
15.0
20.0
25.0
Process Steam Demand (CJ/air-dry metric tonne of product)
Fig. 2. Kratt-mill energy demands. "Swedish model mill" Is
an estimate of economically achievable level In a greenfield
mill built in year 2000.
mill (not considered here) mi ght involve supplementin g the available
gasified black liquor with natural gas or gasified biomass to provide the full
fuel requirement of a gas turbine.
While results are shown for onl y one gas turbine, the si gnificance of the
results is not restricted to this turbine alone: ver y similar results would be
obtained for turbines of other manufacturers that are of the same basic t ype
(heavy-duty industrial), same power output class, and same generation of
technology. For example, the General Electric 6001FA is ver y similar to
the Siemens KW1J64.3a
The use of gasified black li quor would require some adjustment in the
operating parameters of an actual turbine. In particular, the fuel flow
required to reach the desi gn turbine inlet temperature will be hi gher than
with natural gas, due to the lower volumetric ener gy density. All other
factors constant, this larger fuel flow increases the flow throu gh the turbineexpander, and thereb y the required expander inlet pressure. Increasin g the
engine's compression ratio to achieve this ma y cause compressor stall. To
avoid simulating unrealistic compressor operation, the engine's compression
ratio is allowed to increase modestl y (from 16.6 to 17), but further increase
is limited by simulated closin g of the compressor inlet guide vanes, which
reduces air flow to the en gine to avoid stall.
Some hardware modifications ma y also be re q uired, including enlarging
the fuel nozzle flow area. The calculations assume that such hardware
changes can be made without si gnificantly affecting thermodynamic
performance.
The accurate inte grated modeling of actual commercial gas turbines is
3
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an important distinguishing
feature of the present work
relative to other studies that
have examined black liquor
gasifier/gas turbine
applications in the kraft
pulp and paper industry.
e.g. Berglin (1996). Ihren
(1994), and McKeough et
al. (1995).
Table 4. Tomlinson performance.
This Adams.
work
1997
Ambient air temp.. *C
15.0
27.0
Feedwater temp., °C
89
110
Air preheat temp., °C
150
150
INPUT ASSUMPTIONS
Smelt exit temp.. °C
850
850
Carbon cony, to gas, %
99.00
98.48
Heat loss, % liquor HHV
2.0
2.5
Stack temp.. 'C
177
210
Flue 0, (% dry vol.)
3.0
2.8
0.18
Table 6. Calculated gasifier heat and mass balances.
Reactor type
-
Low P.1
air-blown
High P. T
air-blown
High PT
0,-blown
High P,T
0,-blown
700
950
IMO
1400
2
25
25
25
75
75
75
75
115
Assumptions
Reactor temp., °C
Reactor pressure, bar
Liquor feed. Sc cls
Product Gas Cleaning
Liquor feed temp.. °C
115
115
115
Air feed. Vtds
1.67
1.91
n.a.
n.a.
0, feed, tads
n.a.
n.a.
0.471
0.580
Oxidant temp., IC
268
390
135
135
2.543
2.784
1.344
1.477
Calculated Results
For all plant
Soot blow steam. Utds
0.18
configurations. cleaning of
Slowdown, Inds
0.11
0.11
the syngas is assumed to be
Steam pressure, bar
60
carried out by an isothermal
62
Steam temp.. •C
450
scrubbing at 10°C.
A
482
CALCULATED
variety of commercially
Steam produced. [Ads
proven processes are
136
3.11
Heat to steam, Sc'
available for H,S removal
65.0
62.5
(a) Hem in steam delivered to turbine less hem in
from syngas. but the
soot-blow smart divided by input liquor HHV.
practical feasibility of
scrubbing gasified black
liquor to meet turbine
specifications and also recovering sulfur for recycle to the pulp mill remains
to be demonstrated. Within the model adopted here for the gasifier. only
H„S must be removed from the raw syngas--all sodium and potassium leave
in the condensed phase. In practice. some CO 2 would also be removed
when scrubbing H2S. Given the much higher concentration of CO than
H,S in the raw gas, we have assumed that 2 moles of CO, are removed for
Raw gas. tads
At
(mol Sc)
0.498
0.555
0.870
1.062
2.324
10.532
2.151
2.192
CO
0.908
11.619
CO,
12.405
11.452
23.059
18.571
21.894
20.740
CH,
COS
0.019
0.028
H1
19.083
10.307
0.051
22.371
12.503
14,0
13.005
17.585
31.522
40.927
HA
0.637
0.594
1.081
0.264
NH,
0.007
0.009
0.002
0.000
41.817
46.615
0.321
0.393
N,
Condensed phase. eels
Smelt or dry solid
0.021
0.463
0.461
0.460
0.436
dry solid
smelt
smelt
smelt
Na,..50, (weight Sc)
0.000
0.004
0.002
0.118
81.873
79.946
79.525
62.206
Ha/6
7.668
8.005
7.901
17.398
NaOH
0.0134
1 606
2.119
9.222
K,CO3
9.578
9.630
9,645
10.063
K.,S0,
0.000
0.005
0.002
0.138
0.801
0.805
0.806
0.860
Na,CO3
Table 5. Detailed assumptions used in models.
Gasifier heat losses
0.5% of black lig. HilV .
carbon conversion to gas
99%
gas a condensed phase compositions equilibrium, except CH.
Synge' coolers minimum AT
gas-side AP
Heat exchangers minimum gas-liquid AT
minimum gas-gas AT
AP
heat kisses
HRSG minimum AT at pinch points
10°C
each mole of H.S. which may overestimate the CC! removal that would
result in practice (McKeough eta].. 1995). The CO, removal rate has an
important impact on cattsticizing requirements during recovery of the
pulping chemicals (NUTEK, 1992), but has a negligible impact on
cogeneration performance, the focus of this work.
2%
10•C
25•C
2%, unless otherwise shown
Zero
10•C
Heat intearation
minimum AT at approach points
25°C
minimum subcoding AT at drum inlet
10"C
For each plant configuration modeled here, an effort has been made to
optimize the heat integration among components so as to maximize
efficiency within practical cost (and material) constraints. With this in mind.
the network of heat exchangers in each system has been designed following
two guidelines. First, high-temperature gas streams transfer heat only to
water or steam-water mixtures (evaporators): due to the high heat transfer
coefficients achievable with water and two-phase mixtures, this arrangement
guarantees acceptable heat exchanger metal temperatures. Second, to the
extent possible in practice, heat is transferred across relatively small
temperature differences and between flows having similar thermal
capacities. This reduces heat transfer irreversibilities and increases overall
system efficiency.
superheater AP
10%
economizer AP
25%
maximum steam superheat
gas-side
ars
520fC
300 rnm H,0
heat losses
0.7% of heat released by gas
condensate return from mill
80% at 110•C
2 to 4 bar
Steam cycle dearator pressure
condenser pressure
heat rejection parasitic power
0.056 bar (35 °C)
2% of rejected heat
pump efficiencies
f(volume flow)
steam turbine efficiencies
f (volume flow, AA, rpm, etc.)
Electric generator efficiency
I (power output)
Thermodynamic Data
Biomass boiler flue oxygen
4% (dry volume)
For all computations requiring thermodynamic property data. the
JANAF tables are used (Stull and Prophet, 1971; Gardiner, 1984) in
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three different gasifier designs. including
pressurized, air-blown and pressurized, oxygenblown high-temperature reactors (based on
designs of Kvaemer and Noe) and one lowpressure. low-temperature air-blown reactor
(based on the design of ABB). Figures 3-6 show
sample flow schematics.
High-Temperature Air-Blown Gasifier
The basic plant configuration with the
pressurized, air-blown, high-temperature gasifier
is shown in Figure 4. Black liquor fed at list
is gasified at 25 bar in air bled from the gas
turbine compressor. The product gas passes
through an integral quench bath and is further
cooled
by raising low pressure steam and
Fig. 3. Heat and mass balance for Toml nson recovery boiler plus integrated biomass boiler
preheating makeup and condensate return water.
meeting process steam demand of 16.3 aliadmt For additional details. see Table 7.
Water condenses from the product gas in this
process and is recirculated to the quench bath.
The quench bath water preheats the recirculated condensate. Weak wash
combination with SI steam tables (Schmidt. 1982).
(containing 30 grants/liter of NaOH) is used in the low-temperature
scrubber to capture H 2S. The clean syngas is preheated before being fired
Tomlinson Boiler
in the gas turbine. Preheating does not appreciably improve cycle efficiency.
To provide a consistent comparison between gasification-based systems
but because of the low heating value of the fuel gas it is important in
and Tomlinson boiler cogeneration systems, the Tomlinson technology is
increasing combustion stability. Steam is raised at 90 bar in the HRSG from
modeled at a comparable level of detail. Table 4 compares model results
the relatively clean turbine exhaust.
with an industry-standard performance estimate. The difference in overall
performance between the two is due primarily to the difference in assumed
stack gas temperature and black liquor heating value (14.363 Gl/tds in this
High-Temperature, Oxygen-Blown Gasifier
work and 13.950 Glltds in Adams et al). In all calculations here. the
This plant configuration (Fig. 5) is conceptually similar to the previous
Tomlinson steam pressure is set at 60 bar, a common level in practice to
one. Oxygen is used in place of air, permitting higher reactor temperatures
minimize corrosion in the furnace and superheater. Pressures
of 80 bar or higher are found at a few mills worldwide. With
pressures above 60 bar. power production efficiencies would
be higher than calculated here, but the comparisons between
calculated performance of Tomlinson-based systems and
gasifier-based systems would not be qualitatively different.
Process Steam Demands
The demand for process steam at a market-pulp or
integrated pulp and paper mill sets the requirement for steam
to be supplied by the cogeneration plant. Depending on many
factors (end product mill location and age, installed process
equipment. etc.), mill steam demands can vary significantly
(Fig. 2). Calculations are carried out here for a range of
process steam demands. For steam demand levels that cannot
be met using black liquor alone, supplemental consumption of
biomass in a boiler is included.
A combined-cycle configuration is considered with each
gasifier. Steam at 90 bar feeds the bottoming back-pressure
steam cycle. The requisite amount of 10 bar process steam is
extracted, with the balance of steam exhausting at 4 bar. (One
case without a steam bottoming cycle is also considered. In
this case, process steam is generated directly at 10 bar and 4
bar.) In all cases, substantial quantities of low grade heat are
rejected to the environment. No effort is made to find useful
applications of this heat. e.g. as process hot water for the mill.
PROCESS FLOW DESCRIPTIONS
Performance calculations are reported here for
cogeneration systems based around Tomlinson boilers and
Fig. 4. Heat and mass balance for high-temperature, air-blown gasifier/combined
cycle with Integrated biomass boiler. Process steam demand of 16.3 G.liadmt Is
provided by the system. See Table 7 for additional details.
5
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A final observation is that adiabatic flame temperatures (AFTs) for the
clean syngas streams fed to the gas turbine combustor in all cases (Table 7)
are well above current turbine inlet temperature limits, so that future
increases in firing temperatures can probably be achieved with gasified black
liquor. An expected correlation between AFT and syngas heating value
(molar basis) is observed in Table 7. Interestingly, the 1400't 0 2-blown
gasifier gives the lowest AFT, while the low-temperature, air-blown gasifier
gives the highest. With the 0 2-blown gasifier the turbine fuel is rich in H 20
due to the saturator (Fig. 5). bringing down the AFT, while nitrogendilution with the air-blown gasifier is minimized by the low reaction
temperature. thereby raising the AFT.
The final two rows in Table 7 are of interest in a situation where a mill
is considering replacing an existing Tomlinson-based cogeneration system.
a likely common situation over the next two decades (Fig. I). A baseline
alternative in this situation would be the installation of a new Tomlinson
recovery boiler plus a supplemental-biomass boiler to augment Stearn
delivery to a back-pressure steam turbine. If the mill has an opportunity to
export power, then the incremental fuel chargeable to power (IFCP) shown
in Table 7 provides a measure of the marginal fuel costs associated with the
',induction of exported power using each of the gasification options shown
in Table 7 in lieu of the Tomlinson baseline. The numerator of the IFCP is
the biomass consumption required in excess of that in the Tomlinson case
to meet the same process steam demand. The denominator of the IFCP is
the amount of power generated in excess of the Tomlinson power
production. The high IFCP values indicate that export power from any of
the five systems (especially those based on air-blown gasification) could be
generated at low (or negative) marginal fuel costs. Obviously capital
investment and operating and maintenance costs would also be considered
in any full evaluation of alternative cogeneration technology options.
Table 7. Summary results for cogeneration with black liquor and
biomass that meets steam demands per air-dry metric tonne of
pulp (admt) corresponding to those for "Mill #2" In Fig. 2. Figs.
3-6 show detailed flows for 4 of the 6 eases summarized here.
Gasification • Biomass Boiler
High
Tomlin. low P.7
.
a ir.bi-
bio-
boiler
Corresponding Figure *
3
CC
•
Pres.. Temp.
Air-blown
CC
6
4
SC
0,-blown
1000
1400
5
Mill production. aCmticlay
1241
1098
1251
1312
1307
1751
Black liquor tdsladmt
1.74
1.74
1.74
1.74
1.74
1.74
90
79.6
90.1
95.1
94.8
127
359
318
362
380
378
506
tdsthr
MW (HHV)
Biomass fuel My
0.318
al/admt (HMV)
Fuel to turbine HMV. MU/kg
0.700 0.428 0.228 0.738 0.763
4.56
15.7
6.35
13.99
8.56
n.a.
4.64
3.698 3.698 6.678 4.966
15.3
LHV. Iclimol
n.a.
101
85.5
85.5
106
78.7
Flame T.' •C
n.a.
1770
1634
1587
1690
1468
16.3
16.3
16.3
16.3
16.3
16.3
alladm410-bar
5.5
5.5
5.5
5.5
5.5
5.5
GJ/admt, 4-bar
10.8
10.8
10.8
10.8
10.8
10.8
Process steam,' total
Gross Power Output. MW,
48.4
136.7
104.7 72.8
Auxiliary consumption, MW,
1.64
22.04
4.19
4.03
19.77 32.29
Net Power Output, MW,
46.8
114.7
100.5 68.7
132.0 147.7
2425 2024
151.8 180.0
kWh/sem(
904
2507
1928
1258
Net Excess Power, MW,
12.8
84.7
66.3
32.9
96.3
99.6
kWinadmI
248
1851
1272
602
1769
1368
Fuel HHV to electricity
steam
Fuel MN to electricity
10.4
23.2
20.7
15.3
21.5
18.1
52.0
41.8
48.6
55.1
40.2
40.6
12.0
26.8
23.9
17.7
24.9
21.0
60.2
48.4
56.2
63.8
46.5
46.9
Total efficiency, 111-1V
62.4
65.0
69.2
70.4
61.7
58.7
NHV
72.2
75.2
80.1
81.5
71.4
67.9
0.20
0.55 _ 0.43
Incremental Fuel ChargeiLib
le to Power'
0.28
0.53
0.45
. steam
Electricity/steam ratio
SA.1 fuel (HHV) per kWh
-
4.8
2.2
-5.1 IEl
%
--
76
167
-71
CONCLUSIONS
Black liquor gasification systems offer the possibility for Strait-based
market pulp or integrated pulp and paper mills to generate far more
electricity than at present while still meeting process steam demands.
Depending an the gasification technology and cycle design, power-to-steam
ratios can vary widely, providing flexibility in meeting mill requirements.
The present work has focussed on better understanding the prospective
energy benefits of black-liquor gasifier/gas turbine systems. Prospective
benefits of gasification with regard to capital-investment, environmental
impact, flexibility in preparation of pulping chemicals, safety, and reliabilty
have not been addressed.
Black liquor gasification for gas turbine applications is not yet proven
commercially. The level of development of several gasification technologies
appears advanced sufficiently that large-scale demonstrations could be
launched toward demonstrating commercial viability of gasifier/gas turbine
systems. In particular, key features that must be successfully demonstrated
include gas cleanup to meet gas turbine specifications, stable gas turbine
combustion of syngas, heat recovery from syngas streams and (in the case
of the low-temperature gasification process) from solids, cost-effective
recovery of pulping chemicals, and overall-plant thermal integration.
8.0
59
45
(a) CC is combined cycle with beck pressure stem =bane. SC is simple cycle gas
turbine. Results tor /007C and 1407C muiticacioo ormnerarre are shown for the rain
with 05 -blown gasifiers.
(0) Maximum achievable temperature with fuel and oxidant temperatures and pressures
shown in Figs. 3-6
(c) Assuming 80% condensate return at 110C and fresh make-up at 15•C. the heat pm
to steam in the cogeneration plans is 2.315 GI/tonne for 10-bar SICIUTI and 2.276 Glhoore
for 4-tor =urn.
(d)rse numerator is the biomass fuel consumed in excess of that required to greet steam
demand with the Tomlinson syvem. The denominator is the power produced in excess
of that produced in the Tomlison case.
primarily because with a lower gasifier outlet temperature, a larger fraction
of the gasifier output enters the inherently more-efficient brayton cycle
rather than the rankine bottoming cycle. (With higher gasifier outlet
temperatures, a larger fraction of the gasifier output is recovered as steam,
which can only be used in the rankine cycle. Stated another way, more of
the liquor fed to a high-temperature gasifier must be fully oxidized to reach
reaction temperature, with the result that less liquor is converted into
chemical energy in the product gas.) Also, the irreversibilities involved in
cooling the gasifier product gas are smaller with the low-temperature design.
Pressurization of the high-temperature gasifiers partly offsets the electricefficiency advantage enjoyed by the low-temperature system.
ACKNOWLEDGEMENTS
For cost-share contributions to this work, the authors thank the
Weyerhaeuser Company. We also thank Tom Kreutz for his contributions.
For helpful comments at the draft stage, we thank Terry Adams. Russ
Andrews, Thorn Bemtsson, Craig Brown. Billy Davis, Patricia Hoffman. Ed
}Cachet, Bob Kinstrey, George McDonald. Bill Nicholson. Del Raymond.
Rolf Ryham. Al Sutb, Gunnar Svedberg, and Bjdm Wamqvist. For
assistance with artwork, we thank Roberto Biscuola and Ryan Hayward.
For financial support, the authors thank the Office of Industrial
8
Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 12/29/2014 Terms of Use: http://asme.org/terms
••
mill convening logs into pulpable chips might amount to 0.25 dry tonnes of
biomass per tonne of final pulp, or about 5 al/admt. Many mills may have
access to much more biomass. One detailed study around a Weyerhaueser
market-pulp mill in North Carolina identified a sustainable supply of up to
3 dry tonnes of biomass per tonne of pulp (or 60 al/admt) at reasonable
cost in the form of harvest residues and self- and externally-generated mill
residuals (Weyerhaeuser et al., 1995).
The curves for the gasification-based technologies in Fig. 7 are steeper
than for the Tomlison technology because of a higher biomass boiler
pressure--the pressures are asssumed to match those of the gas turbine
HRSG (90 bar) and the Tomlinson boiler (60 bar). Also, as the biomass
share of total energy input increases, differences in power output from one
technology to the next diminsh in most cases. (At the highest level of
biomass input shown, the biomass energy input approaches the black liquor
energy input.)
The gasifier/gas turbine systems produce considerably greater
kWhiadmt than the Tomlinson-based system. However, more supplemental
biomass must be consumed with these systems to meet the same process
steam demand. At lower levels of process steam demand (e.g., as at Mill
#1). the gasification-based systems produce two (with simple cycle gas
turbine) to three (with combined cycle) times the kWh/admt as the
Tomlinson system. At higher process steam production (e.g., as needed at
Mill #3), the simple-cycle gas turbine provides comparable kWhketmt as the
Tomfmson, while the combined-cycle systems produce about twice as much
power. In all cases, including with the Tondison-based system. power
production is in excess of a typical pulp mill's process power needs. In
contrast, most North American pulp mills today use Tomlinson plus
supplemental-fuel boilers with back-pressure steam turbines, but operate
relatively inefficiently and so generate little or no excess power.
Several interesting comparisons among alternative gasifier-based
systems are illuminated by a more detailed examination of the results in Fig.
7 for a fixed mill process steam demand. Table 7 summarizes performance
calculations for alternative cogeneration systems that provide a level of
process heat and power per admt that characterizes Mill #2.
One interesting comparison is between a simple-cycle and a combinedcycle system with the same gasifier (high-temperature, air-blown is
considered hem). With a high back-pressure turbine (4-bar) only about 20%
of the power output of the combined cycle fired only with gasified black
liquor is provided by the steam turbine (Fig. 7). The fraction is larger than
this for Mill #2 (see Fig. 4) due to the use of biomass as a supplemental
boiler fuel. Correspondingly, the difference in power output between the
simple and combined cycles meeting Mill #2's steam demand is larger than
would be the case without supplemental firing. The marginal efficiency of
producing additional power with the combined cycle in place of the simple
cycle is about 60%. On the ether hand, the simple cycle requires only about
half the biomass fuel of the combined cycle to meet the same process steam
demand.
A second comparison is between two combined cycles using hightemperature gasifiers. one air-blown (operating at 950°C) and the other
oxygen-blown (operating at I000 °C). The 02-blown system is a more
efficient electricity producer, but a considerably less efficient steam producer
(Table 7). This can be attributed largely to the use of the saturator with the
03-gasifier in place of a low-pressure evaporator. (The difference in
electricity output would be still greater if oxygen were produced in an
integrated fashion from air bled from the gas turbine compressor.) The net
result is that to meet the same process steam demand. the O z-blown system
must consume considerably more biomass than the air-blown system.
Electricity production per adnit is correspondingly much higher (Fig. 7).
A third comparison is between two 0 2-blown systems operating with
different gasification temperatures (Table 7). The steam production
3.000
LOW-templwatUre
air Wen
combined cycle
2,900
2,800
2.700
2,600
2,500
2.400
2,300
2,200
2,100
v
to
2,000
tri 1,800
zi 1,700
2
• 1,600
2
1,500
SI 1,400
cr)
-
1,300
1
H4SMmpOMWMI
oxygen-010ml
arteined cycle
1,200
1,100
1,000
900
800
700
600
500
400
5
7.5 10 12.5 15 17.5 20 22.5 25
Heat to process, GJ/admt
Fig.?. Performance summary for alternative black Ilguor-based
cogeneration systems. For reference, stars represent process
steam and power demands for corresponding mill Is In Fig. 2,
efficiency of the system operating at I400°C is comparable to that for the
1000°C system, but the electricity production efficiency is considerably
lower. This can be attributed primarily to the greater irreversibilities
generated in the raw syngas quench step from 1400°C to 200°C, instead of
I000° C to 200C. In the range 100X1400 C, higher gasification
temperature may provide benefits relating to chemical recovery due to the
lower sodium carbonate levels in the smelt (Table 6) (Larson et al.. 1996),
but energy efficiency will suffer.
A fourth comparison is between combined cycles with hightemp:mune/high-pressure gasification and one using low-temperaturenowpressure gasification. The latter system provides the highest electrical
efficiency of any systems considered here (Fig. 7 and Table 7). This results
7
Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 12/29/2014 Terms of Use: http://asme.org/terms
A final observation is that adiabatic flame temperatures (AFTs) for the
clean syngas streams fed to the gas turbine combustor in all cases (Table 7)
are well above current turbine inlet temperature limits, so that future
increases in firing temperatures can probably be achieved with gasified black
liquor. An expected correlation between AFT and syngas heating value
(molar basis) is observed in Table 7. Interestingly, the 1400't 0 2-blown
gasifier gives the lowest AFT, while the low-temperature, air-blown gasifier
gives the highest. With the 0 2-blown gasifier the turbine fuel is rich in H 20
due to the saturator (Fig. 5). bringing down the AFT, while nitrogendilution with the air-blown gasifier is minimized by the low reaction
temperature. thereby raising the AFT.
The final two rows in Table 7 are of interest in a situation where a mill
is considering replacing an existing Tomlinson-based cogeneration system.
a likely common situation over the next two decades (Fig. I). A baseline
alternative in this situation would be the installation of a new Tomlinson
recovery boiler plus a supplemental-biomass boiler to augment Stearn
delivery to a back-pressure steam turbine. If the mill has an opportunity to
export power, then the incremental fuel chargeable to power (IFCP) shown
in Table 7 provides a measure of the marginal fuel costs associated with the
',induction of exported power using each of the gasification options shown
in Table 7 in lieu of the Tomlinson baseline. The numerator of the IFCP is
the biomass consumption required in excess of that in the Tomlinson case
to meet the same process steam demand. The denominator of the IFCP is
the amount of power generated in excess of the Tomlinson power
production. The high IFCP values indicate that export power from any of
the five systems (especially those based on air-blown gasification) could be
generated at low (or negative) marginal fuel costs. Obviously capital
investment and operating and maintenance costs would also be considered
in any full evaluation of alternative cogeneration technology options.
Table 7. Summary results for cogeneration with black liquor and
biomass that meets steam demands per air-dry metric tonne of
pulp (admt) corresponding to those for "Mill #2" In Fig. 2. Figs.
3-6 show detailed flows for 4 of the 6 eases summarized here.
Gasification • Biomass Boiler
High
Tomlin. low P.7
.
a ir.bi-
bio-
boiler
Corresponding Figure *
3
CC
•
Pres.. Temp.
Air-blown
CC
6
4
SC
0,-blown
1000
1400
5
Mill production. aCmticlay
1241
1098
1251
1312
1307
1751
Black liquor tdsladmt
1.74
1.74
1.74
1.74
1.74
1.74
90
79.6
90.1
95.1
94.8
127
359
318
362
380
378
506
tdsthr
MW (HHV)
Biomass fuel My
0.318
al/admt (HMV)
Fuel to turbine HMV. MU/kg
0.700 0.428 0.228 0.738 0.763
4.56
15.7
6.35
13.99
8.56
n.a.
4.64
3.698 3.698 6.678 4.966
15.3
LHV. Iclimol
n.a.
101
85.5
85.5
106
78.7
Flame T.' •C
n.a.
1770
1634
1587
1690
1468
16.3
16.3
16.3
16.3
16.3
16.3
alladm410-bar
5.5
5.5
5.5
5.5
5.5
5.5
GJ/admt, 4-bar
10.8
10.8
10.8
10.8
10.8
10.8
Process steam,' total
Gross Power Output. MW,
48.4
136.7
104.7 72.8
Auxiliary consumption, MW,
1.64
22.04
4.19
4.03
19.77 32.29
Net Power Output, MW,
46.8
114.7
100.5 68.7
132.0 147.7
2425 2024
151.8 180.0
kWh/sem(
904
2507
1928
1258
Net Excess Power, MW,
12.8
84.7
66.3
32.9
96.3
99.6
kWinadmI
248
1851
1272
602
1769
1368
Fuel HHV to electricity
steam
Fuel MN to electricity
10.4
23.2
20.7
15.3
21.5
18.1
52.0
41.8
48.6
55.1
40.2
40.6
12.0
26.8
23.9
17.7
24.9
21.0
60.2
48.4
56.2
63.8
46.5
46.9
Total efficiency, 111-1V
62.4
65.0
69.2
70.4
61.7
58.7
NHV
72.2
75.2
80.1
81.5
71.4
67.9
0.20
0.55 _ 0.43
Incremental Fuel ChargeiLib
le to Power'
0.28
0.53
0.45
. steam
Electricity/steam ratio
SA.1 fuel (HHV) per kWh
-
4.8
2.2
-5.1 IEl
%
--
76
167
-71
CONCLUSIONS
Black liquor gasification systems offer the possibility for Strait-based
market pulp or integrated pulp and paper mills to generate far more
electricity than at present while still meeting process steam demands.
Depending an the gasification technology and cycle design, power-to-steam
ratios can vary widely, providing flexibility in meeting mill requirements.
The present work has focussed on better understanding the prospective
energy benefits of black-liquor gasifier/gas turbine systems. Prospective
benefits of gasification with regard to capital-investment, environmental
impact, flexibility in preparation of pulping chemicals, safety, and reliabilty
have not been addressed.
Black liquor gasification for gas turbine applications is not yet proven
commercially. The level of development of several gasification technologies
appears advanced sufficiently that large-scale demonstrations could be
launched toward demonstrating commercial viability of gasifier/gas turbine
systems. In particular, key features that must be successfully demonstrated
include gas cleanup to meet gas turbine specifications, stable gas turbine
combustion of syngas, heat recovery from syngas streams and (in the case
of the low-temperature gasification process) from solids, cost-effective
recovery of pulping chemicals, and overall-plant thermal integration.
8.0
59
45
(a) CC is combined cycle with beck pressure stem =bane. SC is simple cycle gas
turbine. Results tor /007C and 1407C muiticacioo ormnerarre are shown for the rain
with 05 -blown gasifiers.
(0) Maximum achievable temperature with fuel and oxidant temperatures and pressures
shown in Figs. 3-6
(c) Assuming 80% condensate return at 110C and fresh make-up at 15•C. the heat pm
to steam in the cogeneration plans is 2.315 GI/tonne for 10-bar SICIUTI and 2.276 Glhoore
for 4-tor =urn.
(d)rse numerator is the biomass fuel consumed in excess of that required to greet steam
demand with the Tomlinson syvem. The denominator is the power produced in excess
of that produced in the Tomlison case.
primarily because with a lower gasifier outlet temperature, a larger fraction
of the gasifier output enters the inherently more-efficient brayton cycle
rather than the rankine bottoming cycle. (With higher gasifier outlet
temperatures, a larger fraction of the gasifier output is recovered as steam,
which can only be used in the rankine cycle. Stated another way, more of
the liquor fed to a high-temperature gasifier must be fully oxidized to reach
reaction temperature, with the result that less liquor is converted into
chemical energy in the product gas.) Also, the irreversibilities involved in
cooling the gasifier product gas are smaller with the low-temperature design.
Pressurization of the high-temperature gasifiers partly offsets the electricefficiency advantage enjoyed by the low-temperature system.
ACKNOWLEDGEMENTS
For cost-share contributions to this work, the authors thank the
Weyerhaeuser Company. We also thank Tom Kreutz for his contributions.
For helpful comments at the draft stage, we thank Terry Adams. Russ
Andrews, Thorn Bemtsson, Craig Brown. Billy Davis, Patricia Hoffman. Ed
}Cachet, Bob Kinstrey, George McDonald. Bill Nicholson. Del Raymond.
Rolf Ryham. Al Sutb, Gunnar Svedberg, and Bjdm Wamqvist. For
assistance with artwork, we thank Roberto Biscuola and Ryan Hayward.
For financial support, the authors thank the Office of Industrial
8
Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 12/29/2014 Terms of Use: http://asme.org/terms
Chicago. 1996.
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