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Coal Energy Conversion with Aquifer-Based Carbon Sequestration:
An Approach to Electric Power Generation with
Zero Matter Release to the Atmosphere
Investigators
Reginald E. Mitchell (PI), Associate Professor, Mechanical Engineering; Christopher
Edwards, Associate Professor, Mechanical Engineering; Scott Fendorf, Professor,
Department of Environmental Earth System Sciences; Ben Kocar, Post-doctoral Student;
Adam Berger, Eli Goldstein, John (J.R.) Heberle, BumJick Kim, Andrew Lee, and Paul
Mobley, Graduate Researchers, Stanford University
Abstract
The overall objective of this project was to provide the information needed to develop
a coal-based electric power generation scheme that takes water from a deep saline
aquifer, desalinates and uses it to process coal at supercritical water conditions, and
returns the water, which will contain dissolved CO2, back to the aquifer along with the
salts that were removed during desalination. The sensible energy of the hot conversion
products is transferred to the working fluid of a heat engine for power generation. The
only process effluents are supercritical water containing dissolved coal conversion
products and solids, composed of ash and salts that precipitate from the supercritical
water.
In efforts to date, we have constructed a system-level model of the proposed plant and
used it to evaluate the viability of the overall concept in terms of efficiency. Calculations
indicate that efficiencies as high as 42%, on a lower heating value basis, can be obtained
when account is made the energy penalties associated with oxygen separation, non-ideal
pumps, compressors and turbines, and carbon dioxide sequestration.
We have also completed the development of a software library for the thermodynamic
properties of CO2 and H2O, gases that behave non-ideally at the temperatures and
pressures employed in the process units. In order to determine the thermodynamic
properties of mixtures, a mixing model was also developed that models the departure
from an ideal Helmholtz solution using a modified corresponding-states approach. The
model has been used to determine thermodynamic properties when predicting the
composition of the synthesis gas produced in the coal reformer and the composition of
the combustion products.
Wyodak coal, a low-ash, low-sulfur coal, was selected as the base-case coal for study.
This is a sub-bituminous coal from Wyoming and is typical of the coals from the Powder
River Basin that are currently being used for electric power generation. We have already
determined the reactivity of the coal char to oxygen and carbon dioxide, and tests to
determine the char reactivity to water have begun. The experimental facility being built
to determine coal conversion rates under supercritical water conditions is complete as is
the experimental facility being built to demonstrate stable combustion in supercritical
water.
In addition, geochemical models have been used to assess the fates of potentially
toxic trace elements in the coal that may be dissolved in the water being returned to the
aquifer. Results suggest that relatively high concentrations of As, Cr, Cu, Hg, Pb, Zn,
and Cd in the brine being returned to the aquifer are likely to be mitigated by secondary
precipitation/adsorption reactions that occur within aquifers. The extent of these
reactions will depend on the mineralogical makeup of the aquifer.
Introduction
This project had the objective of providing the information needed to develop a coal-
based electric power generation process that involves coal conversion in supercritical
water with CO2 capture and storage in an inherently stable form. The only process
effluents are supercritical water, containing dissolved coal conversion products, and
solids, composed of fly ash and salts that precipitated from the water. Having no gaseous
emissions, such a system eliminates the need for costly gas cleanup equipment. In
addition, it provides a carbon dioxide sequestration option that may be more acceptable to
the public.
Coal-fired power plants have the potential to emit undesirable substances into the
environment such as nitrogen and sulfur oxides, particulate matter, mercury, arsenic,
lead, and uranium. Clean coal technologies that have been developed to remove these
substances from flue gases before they are emitted into the atmosphere have significantly
increased the cost of coal-derived energy, reducing the economic attractiveness of an
otherwise inexpensive fuel. The economic benefit is further reduced when CO2 must be
removed from the flue gases and sequestered because of its potential impact on climate
change. Constructing power plants that have no gaseous emissions at all but instead, that
sequester the entire effluent stream would end the expensive cycle of continually
identifying and managing the next-most-harmful coal conversion product.
Deep saline aquifers have been recognized as suitable locations for the storage of
CO2. In the United States, the sites identified have the potential capacity to store about
86 years of CO2 generated in coal-fired power plants at the rate of coal consumption for
electric power generation in 2005. Preliminary estimates indicate that a 30-m thick
aquifer having 1-Darcy permeability and 20% porosity can store the effluent of a 500
MWe power plant over its nominal 40-year lifetime. The proposed scheme for electric
power generation under investigation produces CO2 that is in equilibrium with the aquifer
environment, eliminating the possibility of it migrating back to the surface through
fissures or well bores once injected into the aquifer. The stream that will be returned to
the aquifer will be in thermal and mechanical and close to chemical equilibrium with the
water already in the aquifer. This injected stream will be less buoyant than liquid CO2 at
reservoir conditions, allaying any concerns about selective CO2 release.
The advantages of this aquifer-based coal-fired power plant relative to current and
other proposed power generation systems include (i) maximally efficient power
production while storing CO2 products in indefinitely stable forms, (ii) zero traditional air
pollutant emission and stack elimination, and (iii) size reduction of reactor vessel
(compared to pulverized-coal systems). Although the thrust of the project is directed to
clean coal utilization, the process being developed applies in general to all stationary
thermal processors (gasifiers and combustors) using all types of fuel (coal, natural gas,
oil, biomass, waste, etc.). Our investigation aims to lay the foundation for an efficient
coal energy option with no matter release to the atmosphere and in which all fluid
combustion products, particularly carbon dioxide, are pre-equilibrated in aquifer water
before injection into the subsurface.
Overview of the Proposed Process
In the proposed energy conversion scheme, a coal slurry, oxygen and water obtained
from a deep saline aquifer are fed to a gasification reactor (a reformer) that is maintained
at conditions above the critical temperature and pressure of water (Tc,H2O = 647 K, Pc,H2O
= 221 bar). Due to the solvation properties of supercritical water (SCW), the small polar
and nonpolar organic compounds released during coal extraction and devolatilization are
dissolved in the water and the larger ones are hydrolyzed, yielding dissolved H2, CO,
CO2, and low molecular weight hydrocarbons, without tar, soot or polyaromatic
hydrocarbon formation. In essence, a synthesis gas dissolved in SCW is produced in the
reformer. The syngas is sent through a solids separator in which the solids are returned to
the reactor and the syngas is directed to a combustor where it reacts with oxygen to
produce combustion products dissolved in supercritical water. The process stream that
exits the combustor is a single-phase, supercritical solution of hot combustion products.
The stream is passed through a heat exchanger, transferring its sensible energy to the
working fluid of a heat engine that produces electrical power in a combined cycle
scheme. The cooled water stream, saturated with combustion products, is returned to the
same or nearby aquifer.
Since inorganic salts are insoluble in supercritical water, the aquifer water must be
desalinated before use. If not, the ability of the water to absorb coal conversion products
would be reduced and the salts would precipitate in the gasification reactor, mixing with
the ash. The salts would have to be separated from the ash if the ash were to be used, for
instance, as aggregate material.
Sulfur, nitrogen and many of the trace elements in coals are converted to insoluble
salts in SCW. The salts formed will precipitate from the fluid mixture in the reformer,
and will be removed from the system with the coal ash. In this coal conversion scheme,
all trace species introduced with the coal (such as mercury, arsenic, etc.), as well as all
coal conversion products, are sequestered in the aquifer along with the CO2. The only
matter not directed to the aquifer is the solid material consisting of the ash and
precipitated inorganic salts. Although the proposed system still has an energy cost of air
separation, the trade-off of carbon separation for air separation is advantageous since in
the process all fluid coal combustion products are sequestered, including sulfur and trace
metals, not just carbon dioxide.
Research Objectives and Tasks
The overall objective of this research project was to provide the information needed
to design and develop the key process units in the proposed aquifer-based coal-to-
electricity power plant with CO2 capture and sequestration. The project was divided into
four research areas - Area 1: Systems Analysis, Area 2: Supercritical Coal Reforming,
Area 3: Synthesis Fluid Oxidation and Heat Extraction, and Area 4: Aquifer Interactions.
In Area 1, thermodynamic analysis of the scheme provided fully qualified cycle
efficiency and process analyses. These analyses were used to determine the design
choices made in investigating component requirements in the proposed scheme.
Research efforts in Area 2 (Supercritical Coal Reforming) were aimed at determining
the supercritical water conditions that maximize the amount of chemical energy from the
coal in the synthesis fluid. Defining the optimum amount of oxygen required to drive the
gasification reactions and at the same time yield a high energy-content synthesis fluid as
a function of coal composition in the SCW environment was one of the goals of this task.
In concert with this was the goal of developing models that can predict accurately coal
conversion rates to synthesis fluid under SCW conditions. This required obtaining the
data needed to characterize coal extraction and pyrolysis rates and char gasification and
oxidation rates in supercritical water environments as functions of temperature, pressure
and properties of the coal and its char.
The research efforts associated with Area 3 (Synthesis Fluid Oxidation and Heat
Extraction) were focused on the design of the oxidation reactor and transfer of the energy
released to a heat engine in order to extract work. The stream exiting the oxidation
reactor, entering the heat exchanger of the heat engine needs to operate as close as
possible to material thermal limits to maximize heat engine efficiency. Thus, a primary
goal of the oxidation reactor design effort was the distancing of oxidation zones from
reactor walls. An additional requirement was control of reducing and oxidizing streams
to avoid liner corrosion. Under consideration was the design of a combustor in which
hydrodynamics and water injection were used to control reaction, mixing, and wall
interactions.
Research activities associated with Area 4 (Aquifer Interactions) were concerned with
characterizing the impact of dissolved constituents in the water being returned to the
aquifer on aquifer ecology. Of interests were the fates of contaminants prevalent within
coal, such as arsenic, mercury, and lead. Geochemical conditions in the deep subsurface
are likely to lead to the partitioning of elements such as As and Hg to the solid phase.
These elements are subject to migration should physical isolation be disturbed. Another
concern was the possible oxygenation of the aquifer, potentially destabilizing the sulfidic
minerals. A third concern was the potential to develop dramatic fluctuations in pH
resulting from variations in CO2 content, possibly destabilizing aquifer solids and
inducing dissolution or colloidal transport. Geochemical constraints are expected to
diminish the risk imposed by heavy metal discharge into the physically isolated deep
brines but in the research efforts, a combination of equilibrium based predictions and
spectroscopic/microscopic characterization of the energy system products were
performed to verify reaction end-points.
Materials degradation represents one of the most critical issues in the development of
the proposed process. The simultaneous presence of oxygen and ions in the supercritical
fluid forms an aggressively corrosive environment. Consequently, reactor designs must
minimize the contact between walls and SCW that contains oxygen. In our approach, a
perforated liner was considered for use inside the combustor that permits cooling flows of
pure water to protect the combustor surfaces from the hot oxygen-containing SCW.
The sections below focus on the significant research activities undertaken during the
final year that were directed towards meeting the overall project objectives.
Project Status
Area 1: Systems Analysis (Edwards)
Development of Plant Concept.
The construct of a system-level thermodynamic model of the proposed plant has been
completed. The details of the model were presented in last year’s annual status report,
and will be included in the final project report. Model results indicate that for a
combustor exit temperature of 1650 K, overall system efficiencies as high as 42% (on a
lower heating value basis) can be expected, even when energy penalties associated with
non-ideal components, operating an ASU to deliver oxygen to the system, and carbon
sequestration are taken into account. As expected, the overall efficiency increases as the
temperature of the fluid leaving the combustor increases since the product stream is the
heat source for the combined cycle heat engine, and the net output of the combined cycle
dominates the overall efficiency.
The model also indicates that the amount of aquifer water that must be circulated for
total dissolution of the CO2 is not unreasonably high. For a 500 MW coal plant and an
aquifer salinity of 20,000 ppm (as sodium chloride), required water flow rates range from
2,000 to 10,000 kg/s, depending upon the aquifer temperature and pressure. For
reference, the amount of cooling water required by a traditional 500 MW coal plant is
around 12,000 to 13,000 kg/s.
Contacts (Systems Analysis)
Christopher F. Edwards: cfe@stanford.edu
J.R. Heberle: jrh@stanford.edu
Area 2: Supercritical coal reforming (Mitchell)
Autothermal Reformer Operation.
The thermodynamic model indicates that the residual sensible energy in the stream
leaving the heat exchanger where energy is transferred to the heat engine can be used to
preheat the water fed to the reformer to a temperature as high as 700 K. By preheating
the reactants for gasification, the amount of oxygen required for autothermal operation of
the reformer is reduced and the heating value of the synthesis gas is increased. Shown in
Fig. 1 is the required amount of oxygen per 100 kg of coal to operate an autothermal
reformer at 647.3 K as a function of total pressure when the water is preheated to 700 K.
The required amount of oxygen is relatively insensitive to the total pressure for a given
solids loading (the coal weight fraction in the slurry).
220 240 260 280 300 3201.6
1.8
2
2.2
2.4
2.6
2.8
3
Ptotal
(atm)
O2 a
dd
ed
(km
ol/100kg
of
co
al)
Solids Loading = 0.150
Solids Loading = 0.200
Solids Loading = 0.300
Solids Loading = 0.400
Figure 1: Required oxygen for autothermal operation of the reformer at
647.3K without and with preheating of the reformer feed water.
At any selected reformer operating pressure, as the solids loading is decreased, more
slurry water needs to be heated. With preheated water, for a specified reformer operating
pressure, the amount of oxygen required to keep the gasification temperature at 647.3 K
by partial oxidation of the coal decreases as the solids loading decreases.
By preheating the water supplied to the reformer, less energy needs to be produced
from combustion of coal and volatilized species; a higher concentration of combustible
species exists in the synthesis gas compared to the no-preheat case. Also, the lower the
coal fraction in the slurry, the less oxygen required for autothermal gasification.
Consequently with preheated water, the lower the solids loading the higher the heating
value of the synthesis gas. This is demonstrated in Figs. 2, in which are shown calculated
values of the water partial pressure in the reformer as a function of the total reformer
pressure and solids loading. The ratio of the heating value of the synthesis gas to the
heating value of the coal (the HV ratio) is indicated for each solids loading curve.
When evaluating the energy requirements of the reformer, real gas properties were
used for all species for which data are available (H2O, O2, N2, CO2, CH4) when
determining the equilibrium composition of the synthesis gas. For species for which real
gas data are not available (H2S, HCl, Cl2, CO, COS, C2H2, C2H4, C2H6, NH3, NO, NO2,
SO2) the Peng-Robinson equation of state was used to describe their P-v-T behavior.
Experimental Setup.
An experimental facility was designed, built and assembled that permits the
acquisition of data needed to develop a coal/biomass gasification mechanism suitable for
supercritical water conditions (P > 218 atm and T > 647K). The focus point of the
facility is an entrained flow reactor that can be pressurized up to 340 atm at temperatures
up to 1000 K. The reactor is 15 meters long, sufficient for residence times up to 15
minutes when the coal-water slurry feed contains 20% solids and the total mass flow rate
(slurry plus supercritical water (SCW) flow rates) is 20 g/min. The residence time in the
220 240 260 280 300 32050
100
150
200
250
300
Ptotal
(atm)
PH
2O (
atm
)
Solids Loading = 0.150
Solids Loading = 0.200
Solids Loading = 0.300
Solids Loading = 0.400
(Syngas HV)/(Coal HV) = HV Ratio = 0.97
HV Ratio = 0.92
HV Ratio = 0.87
HV Ratio = 0.84
Figure 2: Heating values of synthesis fuel at 647.3K without and
with water preheating.
reactor is varied by controlling the solids content of the slurry and the flow rate of the
water supplied. The feed water is pressurized and preheated to the test conditions before
the solids slurry and any oxygen are admitted. Oxygen is added so that the heat release
from partial oxidation of the solids can supply energy for autothermal gasification.
The stream leaving the flow reactor is quenched and dumped into a water bath where
the solids are collected and analyzed to determine the extent of conversion. A side-
stream is directed to a gas analyzer, purchased from Rosemount Analytic, which
measures in real time the concentrations of four species of interest: CO, CO2, CH4, and
H2. The side-stream can also be directed to the sampling loop of our gas chromatograph
to permit the measurement of the concentrations of such other species as N2, O2, H2S, and
NH3.
A schematic of the experimental facility is shown below in Fig. 3. A picture is shown
in Fig, 4.
Figure 3: Experimental setup for coal gasification under supercritical conditions.
The facility consists of seven main systems:
1. Two water-pumping lines
a. Pure water supply line: Directed to the preheater for high-pressure water
b. Slurry supply line: Connected to a piston accumulator for slurry injection
2. Oxygen supply line
3. Coil-shaped, Inconel 625 reactor in a sand bath
4. Cool-down chamber
5. Liquid/solid-gas separation chamber
6. Gas analyzer
7. System control and data acquisition box
Details of each of these systems are discussed in the sections that follow.
Figure 4: A picture of the experimental facility.
1. High-pressure water supply lines
Two water supply lines were developed. one for pure water supply and the other for
the coal-water slurry supply. The pure water line is preheated before it is injected into
the reactor. A water pump, made by SC Hydraulic, is used to pressurize and supply
water to the system. This pump is driven by air at 120 psi (8.17 atm) supplied from high-
pressure air tank. The air to drive this pump is regulated by a FESTO air controller to
control the water pressure. The water pump and air controller are shown in Fig. 5.
Since a flowmeter to measure the flow rate of high-pressure water at pressures as high
as 340 atm could not be found, we measure the pressure difference between two pressure
transducers (from DJ Instruments) across a flow restrictor (a 30.0 inch (762 mm) long
stainless steel tubing with 1/16 inch (1.5875 mm) outer diameter and 0.01 inch (0.254
mm) inner diameter). Calibration indicates that with this restrictor, a 500 psi (34 atm)
pressure drop corresponds to 20 grams/min of water flow and a 1500 psi (102 atm)
pressure drop corresponds to 40 grams/min of water flow. The flow restrictor is shown in
Fig. 6.
Figure. 5: The SC Hydraulic water pump (left). The FESTO air controller (right).
Figure 6: The flow restrictor between two pressure transducers.
The water in this line is preheated using three Chromalox MaxiZone cartridge heaters
(0.685 inch (17.40 mm) OD, 20.5 inch (521 mm) length, 240 VAC and 1600 watts)
surrounded by aluminum powder and calcium silicate insulation. The heater is shown in
Fig. 7.
The other high-pressure line is used for the coal-water slurry. The water pump for
this line is made by Haskel and the pump-driving air is controlled by another FESTO air
controller. A picture of the high-pressure water pump is shown Fig. 8.
Figure 7: The high-pressure water preheater.
Figure 8: A Haskel high-pressure water pump.
Due to its high viscosity, slurry cannot flow through the high-pressure water pump
and flow restrictor for high-pressure water flowrate measurements. Therefore, the slurry
is loaded on one side of a piston accumulator by a peristaltic pump before experiments,
and pure high-pressure water from the Haskel pump supplies the loaded slurry to the
system by moving the piston from the pure water side toward the slurry side of the
accumulator. The high-pressure piston accumulator was purchased from High Pressure
Equipment Company. For slurry loadings at low pressures, a Vector peristaltic slurry
pump is used instead of the Haskel pump. The piston accumulator and the Vector
peristaltic slurry pump are shown in Fig. 9.
Figure 9: The high-pressure piston accumulator (left). The Vector slurry pump (right).
The water pumps only have a small volume per stroke compared to the overall
amount of water required for experiments consequently, the pumps need to stroke
frequently to supply the needed flow rates of pressurized water. The fluctuations due to
these pump strokes need to be dampened. For each water supply line, a pulse dampening
accumulator was set up between the water pump and pressure transducers. The
accumulator contains a bladder inside the steel housing, and the bladder is filled with
nitrogen. By expanding and contracting the bladder, the water pressure fluctuations can
be dampened. Pressure gauges were attached so that the nitrogen pressure profiles in the
bladder could be observed. The bladder accumulators were purchased from HyDac; one
is shown in Fig. 10.
Figure 10: A HyDac bladder accumulator.
2. Oxygen supply
Oxygen is pressurized using a SC Hydraulic gas booster (shown in Fig. 11). This
booster is driven by air from a high-pressure air tank at 150 psi (10.2 atm). The
pressurized oxygen is stored in a stainless steel cylinder and flows through a high-
pressure oxygen flow controller made by Bronkhorst High-Tech. The controller controls
the oxygen mass flow rate up to 10 slpm (14.3 grams/min).
Figure 11: The SC Hydraulic oxygen booster.
The preheated pure water and the slurry from the piston accumulator are mixed with
high-pressure oxygen at a mixing junction, located at the starting point of the coiled
reactor.
3. Reactor
The coil-shaped reactor, shown in Fig. 12, is made of Inconel 625 to endure high-
pressure, high-temperature environments. The reactor is 15 meter long, with 0.3125 inch
(7.9375 mm) ID and 0.5450 inch (13.843 mm) OD. It is submerged in a bath filled with
sand (aluminum oxide, 120 grit (=116 microns)) and sixteen Watlow Firerod cartridge
heaters (0.375 inch (9.525 mm) OD, 7 inch (118 mm) length, 240VAC and 300 watts).
The heaters are used to heat the air used to circulate the sand particles in order to
maintain a uniform sand bath temperature. These heaters are capable of maintaining the
sand bath at temperatures up to 1000 K.
A two-channel heating system control station was developed to control temperatures
of the preheater and the reactor sand bath. This station contains two temperature
controllers, two over-temperature controllers, 40 amp solid state relays and high speed
fuses.
Figure 12: The 15-m coiled tubular reactor.
4. Cool-down chamber
The flow leaving the flow reactor needs to be quenched as rapidly as possible to
prevent further reactions. To decrease the cool-down time, the tube diameter at the
reactor exit is decreased to 0.06 inch (1.524 mm) ID to increase the velocity of the flow
and improve the overall heat transfer by convection. Cold water flows around the cool-
down tubing in the chamber. Our calculations indicate that a 1 meter long cool-down
section is sufficiently long to freeze all the reactions in less than 20 seconds. The cool-
down tube is pictured in Fig. 13 along with the cold water chamber.
Figure 13: The cool-down tube and cold water chamber.
5. Liquid/Solid-Gas separator
The flow contains liquid, gas products and solid particles, which must be separated.
The gas is separated from the liquids and solids in the separation chamber, a high-
pressure cylinder made by High Pressure Equipment. At this stage, the flow is still at a
high-pressure but it is at a low temperature after being cooled in the cooling chamber.
Due to the limited volume of the separation chamber, during warm-up the flow bypasses
the separator. The flow is directed through the separator only during actual tests. High-
pressure HIPCO valves are used to direct the flow. The separation chamber is shown in
Fig. 14.
Figure 14: Separation chamber (left) and HIPCO control valves (right).
The entire system pressure is controlled by a Tescom back-pressure-regulator. A
Marsh Bellofram air controller supplies air at pressures up to150 psi (0~10.2 atm) to the
Tescom back-pressure-regulator, sufficient to maintain system pressures up to 5500 psi
(374 atm). The Tescom back-pressure regulator and Marsh Bellofram air controller are
pictured in Fig. 15.
6. Gas analyzer
The gas leaving the separation chamber is directed to an analyzer where the gas
composition is measured. Gas analysis is made using an Emerson Process Management
X-STREAM gas analyzer that has four channels: CO (0~40%), CO2 (0~40%), CH4
(0~20%), and H2 (0~50%). For measurements of gaseous products other than these four
products, our Varian gas chromatograph is used. The analyzer is pictured in Fig. 16.
Figure 15: The Tescom back-pressure regulator (left). A Marsh
Bellofram air controller (right).
Figure 16: An Emerson Process Management X-STREAM gas analyzer.
7. System control box
Air supplies to drive the SC Hydraulic water pump, the Haskel water pump, the SC
Hydraulic oxygen booster, and the HIPCO valves are controlled by ASCO solenoid
valves, which are operated by OPTO22 modules on our data acquisition box. The data
acquisition/system control box was developed by combining a data acquisition board
(DaqScan 2001) with extension boards (DBK 207/CJC, DBK 208 and DBK 11A) for
analog inputs/outputs and digital inputs/outputs from IOtech. Signals from the
Bronkhorst oxygen flow controller, the FESTO air controllers, and the thermocouples are
also acquired on these boards. All inputs and outputs are managed remotely from a
dedicated computer. The automated system control and data measurement programs are
written in LabVIEW (National Instruments). The ASCO solenoid valve and system
control box are shown in Fig. 17. A picture of the automated system control panel is
shown in Fig. 18.
Figure 17: ASCO solenoid valves (left). The system control box (right).
Figure 18: The automated system control program.
Continuing Research (Supercritical water reforming)
Coal conversion tests in supercritical water environments are currently underway.
Extents of coal conversion are being measured at selected residence times in various
SCW environments. The data are being used to develop a model of char gasification in
supercritical water. In order to help define the rates of the key reaction pathways,
reactivity tests are also being performed in our pressurized thermogravimetric analyzer.
References (Supercritical water reforming) 1. Churchill, S.W., and M. Bernstein, J. Heat Transfer, 99, 300, 1977.
Contacts (Supercritical water reforming)
1. Reginald E. Mitchell: remitche@stanford.edu
2. B. J. Kim: bjkim@stanford.edu
Area 3: Synthesis Fluid Oxidation and Heat Extraction (Edwards) Supercritical Combustor Development
The next task of the project is the development of a supercritical water combustor.
Our goal is to demonstrate stable combustion in supercritical water with near-
stoichiometric amounts of fuel and oxidizer and with high volumetric firing rate. These
conditions are desirable for a combustor integrated into the plant described in the
Introduction.
Design and Layout of Experimental Apparatus
Wellig has identified six tasks that must be performed by a SCWO experimental
apparatus: reactant storage, pressurization, preheating, reaction, pressure let down, and
effluent discharge1. Due to the firing rates we intend to use and the amount of fluid we
will preheat, cooling of effluent is also a concern, since otherwise the ambient
temperature in the laboratory could rise to an unacceptable level.
Figure 19 is a schematic of the major hardware for the supercritical combustor
experiments. At left are components to store and deliver pressurized fuel, oxygen, and
water. In the center there is an electric resistance heater ahead of the reactor vessel.
Combustion occurs above the heater in the lower section of the high-pressure stack. The
upper section of the stack is a dilution cooler, where ambient temperature water is
introduced to begin cooling the combustor effluent. At right is a condenser to complete
cooling of products and reject thermal energy from the lab. Each subsystem is discussed
in detail below.
The fuel system is shown in Fig. 20. Prior to running the combustor, a low-pressure,
air-powered pump transfers fuel into a bladder accumulator in the laboratory. This
accumulator is sized to hold enough fuel for a four hour run at a maximum firing rate of
50 kW. During operation, the gas side of the bladder is pressurized with nitrogen to feed
fuel into the combustor. This strategy avoids the use of a high-pressure fuel pump. The
nitrogen is supplied from a six-pack of gas cylinders, pressurized with a two-stage, air-
powered booster, and regulated to a pressure above the combustor operating pressure.
Most of the (known) pressure drop from the accumulator pressure to the combustor
pressure occurs across a metering valve. This needle valve is operated remotely with a
stepper motor, so that fine changes in the valve’s flow coefficient allow for adjustment of
the fuel flow rate. The fuel line also has a normally-closed solenoid valve for positive
fuel shut-off. The fuel then flows through a spur gear flow meter and a check valve.
Next, the fuel mixes with an adjustable amount of water so that pyrolysis products do not
plug the fuel line through the heaters.
Figure 20: Liquid fuel delivery system.
Figure 19: Diagram of the supercritical combustor experiment. Reactant storage
and pressurization components are shown to the left of the vertically oriented high
pressure stack. At right are components to remove waste heat from the lab and
separate the effluent components for disposal.
Figure 21 shows the oxygen system. Oxygen is fed from a six-pack of cylinders in
the oxidizer galley. Like the fuel bladder, this quantity of oxygen was chosen since it has
capacity for a four hour run. The oxygen is regulated to 600 psi and flows into the lab,
where it is pressurized by a two-stage booster similar to the nitrogen booster in the fuel
system. To provide a steady flow of gas, the output of the booster is connected to a
plenum and a regulator. By pressurizing this plenum above delivery pressure, the
regulator can deliver oxygen at an approximately constant pressure while the piston in the
booster cycles. The tube after the regulator is fitted with a cooling jacket. The jacket is
used to reduce temperature variations in the oxygen as the plenum blows down,
facilitating repeatable metering and flow measurement downstream. Next are valves for
manually isolating and bleeding different parts of the system. The remaining components
are similar in function to those in the fuel system.
There is a remote solenoid valve for positive oxygen shut-off, followed by a stepper-
driven needle valve to control the flow rate. A thermocouple is located just upstream of
the flow meter so that the latter’s readings may be corrected for temperature variations.
The flow meter consists of a 0.5μm filter in parallel with a differential pressure
transducer. This arrangement is similar in operation to a laminar flow element. Finally,
there is a check valve and a mixing tee. While water will almost always be mixed with
the fuel, initial combustion experiments will involve injection of neat oxygen. The tee
gives the option of mixing water with the oxygen later, which is a way to affect mixing in
the combustor.
Figure 21: Oxygen delivery system.
The low-pressure water system can be seen in Fig. 22. At the top left of the diagram,
water enters from a domestic cold water supply and proceeds through a 20 micron
particle filter. The line goes to a manual three-way valve that accepts either fresh water
from the supply or recycled water from the drain tank. This valve directs the water to a
needle valve and rotameter to ensure that the flow rate through the deionizer is not more
than 2 gpm. This flow rate is the recommended limit for effective operation of the
deionizer. Deionization is used to reduce corrosion in the heaters and high-pressure
stack. The other line that feeds into the manual three-way valve before the deionizer is
for direct recycling while the system is running water through the heaters and stack for
warm-up and cool-down. This flow rate, averaging about 0.5 gpm, is controlled
elsewhere and thus does not need to run through the needle valve and rotameter before
entering the deionizer. After the water is treated in the deionizer, it enters a 400 gallon
supply tank. This tank is filled before each set of combustion experiments and holds
enough water for a four hour experiment.
The supply tank contains a submersible pump that sends water up into the head tank
with a flow rate higher than is necessary for the experiment. The excess water overflows
back into the supply tank, while water for the experiment is gravity fed through a 20
micron particle filter and valve before entering a triplex plunger pump rated for 5000 psi
(345 bar). This positive-displacement pump provides all of the high-pressure water for
the system. From the pump, the water goes to the high-pressure water manifold shown in
Fig. 23. Much like the head tank, the water manifold is supplied by the triplex plunger
pump with more water than is necessary for the experiment. Excess water returns to the
supply tank from the manifold back pressure regulator (BPR) described later. There is a
liquid level switch in the supply tank that controls the submersible pump to keep it from
running dry as the water level drops in the supply tank. The head tank also has a liquid-
Figure 22: Low-pressure water handling system.
level switch that controls the motor driving the triplex plunger pump to keep it from
running dry.
The outlet stream from the high-pressure apparatus re-enters the low pressure system
in the middle of Fig. 22 after depressurization in the stack BPR. The main components of
the fluid here will range from water at ambient temperature during start-up and shut-
down to a mixture of saturated steam (maximum vapor fraction of 0.5) and carbon
dioxide during combustion. The flow will pass through the main condenser to cool the
stream to near ambient temperature. The main and residual condensers will use process
cooling water (PCW) from the building at 15°C and 30 gpm. The CO2 and water vapor
pass into the residual condenser to cool them below room temperature. This prevents
water vapor from condensing in the exhaust line running to the roof. There is also a
condensate drain in the vent line in case there is any condensation due, for example, to
cold outdoor temperatures. Thermocouples in both the gas and liquid lines exiting the
condensers are used to make sure that the condensers are operating correctly. The lines
connecting the two condensers are arranged to form a p-trap and maintain a high water
level in the main condenser for efficient heat transfer. The bottom outlets from the two
condensers send cooled water through a 5 micron particle filter and two hydrocarbon
filters in series. The hydrocarbon filters separate out any liquid hydrocarbons that are not
pyrolyzed or oxidized in the system. Typically, unburned hydrocarbons are not present
since the combustion efficiency of supercritical water oxidation is very high. The filtered
water then enters a solenoid three-way valve to send the water to the drain tank during
combustion and also allowing the water to be recycled to the supply tank during warm-up
and cool-down. The drain tank has a centrifugal pump connected to its outlet with a
liquid-level switch to make sure that it does not run dry. The outlet of the pump passes
through a check valve and a manual three-way valve to send the water out the lab drain,
or back through the deionizer and into the supply tank to be used again.
The system for handling high-pressure water is shown in Fig. 23. These components
take filtered water from the low-pressure supply system, pressurize it, and control and
measure the flow rates of the several water streams that enter the high-pressure stack.
After the triplex pump described above, the water passes by a bladder accumulator. Here
an accumulator is employed in its usual role for pulsation dampening only, rather than as
both a reservoir and dampener as in the fuel system. The water then enters a manifold,
where it branches out into six lines. The manifold has a globe valve for each outlet so
that any lines that are not needed may be isolated. Excess flow from the pump leaves
through the manifold back pressure regulator (BPR) and returns to the supply tank. The
BPR is operated by an embedded PID controller that tracks the manifold pressure with a
transducer. This setup allows the BPR to keep a constant, set pressure in the manifold
regardless of changes in flow rate through each output line. This configuration is
intended to decouple the flow control in the six lines.
After the manifold the water lines each have a stepper-driven needle valve, flow rate
sensor, and check valve. Unlike the fuel and oxygen systems, each water line does not
have a solenoid shut-off valve. These are not needed since a small amount of water flow
past a mostly closed metering valve is not a problem in any of the six lines. Furthermore,
the needle valve selected for most of the lines is rated for repeated shut-off by the
manufacturer. After the check valves, two of the lines reach the mixing tees described
earlier for fuel and oxygen. The fuel and oxygen solutions and three more water lines
flow through the heaters and to the combustor. The last line is for dilution cooling water,
which goes directly to the cooling section of the high-pressure stack.
There is a single outlet from the stack at the end of the dilution cooling section. It
runs to a second BPR. This BPR is responsible for holding the desired operating pressure
inside the combustor. Just ahead of it is a pressure transducer for feedback to the BPR
controller and a thermocouple to ensure that the fluid is below the rated maximum
temperature of the BPR (650°F or 617 K). The manual three-way valve after the BPR is
for calibrating the water flow sensors. With the heaters off and the stack cold, cold water
will be run through the pressurized system and out this valve. By measuring the mass of
water output in a known time at several rates, the true mass flow rate will be correlated
with the sensor output.
The heaters, shown assembled in Fig. 24, were designed with operational flexibility
in mind. They must heat at least five separate streams above the critical temperature of
water. A computer model was constructed to predict the necessary length of each line to
heat its stream to 700 K. The model operates under a constant wall temperature
assumption. This assumption is justified by the use of cast aluminum as the medium
between the electric heaters and tubes to be heated. The aluminum has a high thermal
conductivity and low spatial temperature variation has been observed in a similar heater
unit in use for another project in the research group. The model was modified for the use
of variable properties and assumes that the fuel/water line is entirely water. However, the
Figure 23: High-pressure water handling system.
heat transfer model used could only give a rough estimate because it cannot account for
pseudoboiling in the tubes. Pseudoboiling occurs when the tube wall temperature is
above the pseudocritical point (the temperature of maximum specific heat at a given
pressure above the critical pressure) while the bulk of the flow is below the pseudocritical
temperature. At this condition the flow has a low-density supercritical boundary layer
and a high-density liquid core. Therefore, this regime operates very much like film
boiling at subcritical pressures. The supercritical layer does not conduct heat as well into
the core of flow, so deteriorated heat transfer occurs. The computer model was run twice
with different properties to give best- and worst-case scenarios. To work between these
two bounds, it was decided to construct the heaters with a large number of relatively short
tube passes. This design allows the total passage length for each of the five flows to be
varied between experimental runs by connecting together different numbers of tubes.
The total power output of the heaters was chosen based on the enthalpy required to
preheat enough water, oxygen, and fuel to fire the combustor at rates up to 50 kW and
with mixed mean outlet temperatures from 1200 to 1800 K. About 40 ft. of electric
Figure 24: Assembled heater units. Aluminum filler is visible through
the casting holes in the shells. At the end of each unit, six cartridge
heating elements and their wires are visible, along with inner and outer
circles of Inconel 625 fluid tubes.
resistance cartridges of the chosen model are required to provide the desired 72 kW of
heating. For ease of fabrication, transport, and use, it was decided to use a larger number
of shorter cartridges. This choice works well with the choice of many short fluid tubes.
Due to the large numbers of heaters and tubes, two identical heater units were designed
and constructed. This geometry provides even heat transfer and convenient fluid tube
connections. Each unit is independently controlled. This arrangement allows for the
possibility of operating streams at different temperatures to simulate various operating
conditions in later experiments.
Each unit is comprised of a steel shell with machined plates welded to each end. The
end plates have holes for six 6 kW electric cartridges, 40 in. long, and twelve ¼” OD
Inconel 625 fluid tubes. The heaters are inserted into thin-walled steel sleeves. After
insertion of the tubes and heaters, the shells were filled with cast aluminum. The
aluminum transfers heat from the heater cartridges to the fluid tubes. The tubes are 60 in.
long—20 in. longer than the heaters—to allow for tubes to be shortened when worn
connections need to be cut and replaced. (The tubes are not replaceable since they are
cast in the aluminum.) The tube ends were offset axially so they could be placed close
together without interference between adjacent fittings.
Calcium silicate was selected as the insulation for the heaters. Two inch thick
insulation wraps around the heaters along with 2 in. thick end caps drilled for the tubes to
pass through. Any gaps in the hard insulation will be filled with fiberglass insulation.
The expected heat loss to the environment from the heaters is less than 2%, based on a
finite element model of the heaters using a free convection heat transfer coefficient at the
outer insulation boundary.
A cut-away view of the high-pressure stack is shown in Fig. 25. The stack has three
major sections: the combustor, an elbow, and the dilution cooler. These sections are
connected with Grayloc-style flanges (in blue), which eliminates the need for welding
traditional flanges to the 2 in. thick vessel walls. They also allow the two flanges to be
placed close together without interference. The body of each section is machined from
solid, forged Inconel 625. All fluid, thermocouple, and pressure fittings are cone-and-
thread high-pressure fittings to ensure proper sealing at elevated temperatures.
The combustor (the lower, vertical section in Fig. 25) has two functional zones. The
lower portion is the header, with inlet ports for the fuel/water solution, oxygen, and
water. These streams flow up through three coaxial passages formed by the outer vessel
and two nozzles. Nozzles of varying diameter may be used to change the relationships
between mass flow rates, fluid velocities, and Reynolds numbers at the burner. Swirlers
may also be inserted to affect flame stabilization. Near the center of the combustor are
two opposed sapphire windows for observation of the flame. In this drawing the nozzles
end just above the bottom of the window so that flame attachment at the burner can be
observed. Shorter nozzles could be substituted to observe flame lengths longer than the
window aperture. Above the windows there is a liner (shown in red). This perforated
metal liner functions like that in a gas turbine combustor, allowing relatively cool fluid
(supercritical water in this case) to enter from the outer annular passage, shielding the
liner and the pressure wall from the hot combustion products in the core flow. This liner
extends beyond the combustor, through the flanged joint and into the elbow.
The elbow section turns the flow and holds a third sapphire window to provide an
axial view of the flame in the combustor below. The window design (visible in this
view) is identical to that of the two windows in the combustor. The bolted flange holding
the window allows the window seals to be preloaded so they maintain a seal at the target
operating conditions of 700 K and 250 bar. There is an inlet near the window for
injection of 700 K water to prevent impingement of very hot (1800 K or more)
combustion products on the window.
Figure 25: Section view of high pressure stack design, showing fluid
ports, burner, axial viewing window, and liners.
The final section of the high-pressure stack is the dilution cooler. This section houses
another perforated liner and fluid inlets. Here, high-pressure, ambient temperature water
is mixed with the combustion products to cool them below the stack BPR maximum
temperature of 617 K.
The high-pressure stack is shown installed in the lab in Fig. 26. The vessel is
insulated from the black supporting stand by Macor ceramic parts. The stand is bolted to
an optical table so that optical diagnostics may be added easily.
Construction Progress
At the beginning of this project, the allocated lab space did not have adequate utilities
to run the experiment. Domestic cold water supply, lab drain, and process cooling water
lines have been installed. A 200 A, 480 V, 3 phase service was installed for the electric
heaters and triplex water pump motor. All necessary utilities upgrades are now complete.
All of the systems described in the preceding section have been installed. All
pressure and flow rate sensors have been calibrated. The feed systems have been run at
ambient temperature and nominal operating pressure long enough that pressure and flow
control are well understood.
Several attempts have been made at heating the system while pressurized in
preparation for flame experiments. All of these tests were stopped after a few hours of
heating when steam was observed leaking from the high-pressure stack. It was
determined that the metal c-seals between the sections of the vessel were defective. The
manufacturer has provided a replacement set of seals. These new seals are being
installed and testing will resume. Thus far, temperatures of 450 K have been reached
before leaks occurred. As soon as the vessel can be heated to and held at 700 K without
Figure 26: The high-pressure stack shown without and with
insulation.
the appearance of steam leaks, injection of fuel and oxygen will begin in attempts to
produce an autoignited flame.
Publications 1. Heberle, J.R., Edwards, C.F. Coal Energy Conversion with Carbon Sequestration via Combustion
in Supercritical Saline Aquifer Water. International Journal of Greenhouse Gas Control.
Accepted for publication.
References (Supercritical Combustor Development) 1. Wellig, B., 2003. Transpiring-wall reactor for supercritical water oxidation. Doctoral thesis no.
15,038, Swiss Federal Institute of Technology (ETH) Zurich. Available http://www.e-
collection.ethz.ch.
Contacts (Supercritical Combustor Development) Christopher F. Edwards: cfe@stanford.edu
J.R. Heberle: jrh@stanford.edu
Paul D. Mobley: pdmobley@stanford.edu
Area 4: Aquifer Interactions. (Fendorf) Fate of Trace Elements in Aquifer Brine Solutions.
Within this task, we assess the fate of potentially toxic trace elements in CO2-
equilibrated brine solutions under elevated pressure and temperature. To accomplish this
objective, geochemical models, including EQ3/6, PHREEQC, and MIN3P (1-3), were
used to simulate post coal-combustion geochemistry of brine solutions delivered to deep
aquifers.
Bulk and Trace Element Composition of Simulated Brine and Coal
Multiple factors will dictate the geochemistry of trace elements within CO2-
equilibrated brines, including 1) the original composition of the brine solution, 2) the
composition of the combusted coal, 3) the ratio of combusted coal:brine solution, 4)
aquifer redox status, and 5) trace element removal
during the coal gasification/combustion process.
To establish a base-case scenario of trace element
speciation and fate within a deep aquifer, factors
1-4 (above) are explored with geochemical
simulations; factor 5 is excluded until trace
elements are measured in solids (ash) removed
during the combustion process (a future task).
The composition of brine solutions may vary
considerably, and are dictated by several factors
including i) the composition of porewater at the
time of burial, ii) the net effect(s) of diagenesis on
porewater composition as influenced by ambient
solids and other proximal porewaters, and iii) physical transport of materials to and from
a the system via advective flow and solute mixing (4). While different combinations of
these factors may result and yield brines with disparate chemistries, sedimentary brines
typically possess salinities greater than 100 g/L and chloride concentrations greater than
10 g/L (4-6). Alkalinity may also vary considerably and range from 10 to 1000 mg/L.
Table 1. Aqueous Constituents in
Simulated Brine
Constitutent mmol/kg H2O
Sodium 1000
Chloride 1000
Lithium 3
Potassium 4
Magnesium 30
Calcium 80
Bromide 2
Barium 2.5
Alkalinity 10
pH 7.1
For modeling purposes, a circumneutral brine solution was assumed as possessing ~1000
mmol/kg chloride and 10 mmol/kg alkalinity, along with multiple trace salts (Table 1).
Despite substantial variation, generalities can be made regarding the quantity of
potentially toxic trace elements in sub-bituminous coal. For example, trace quantities of
lead and cadmium, and to a lesser extent, arsenic and mercury, are often found. The high
coal throughput for modern power plants, however,
results in large quantities of these trace elements
accruing in various combustion products, including
fly ash, slag, and gaseous emissions (7, 8). In our
geochemical models, a suite of trace elements are
stipulated at concentrations representative of those
found within coal combusted at conventional plants
(refs 7, 8; Table 2). Furthermore, coal used in the
simulations is assumed to possess ~67 moles of
carbon/kg dry weight, which will dictate the
quantity of CO2 equilibrated with brine.
Simulating the fate of trace elements in CO2-
equilibrated brine solutions
The geochemical modeling software package
PHREEQC in conjunction with PITZER and
WATEQ databases were used to simulate trace
elements post-combustion in brine solutions (2).
PHREEQC was written to simulate the equilibrium
chemistry of aqueous solutions interacting with
minerals and mineral surfaces and gas dissolution/exosolution, and using the PITZER
database allows for calculating the activity of charged aqueous species under high ionic
strength conditions (e.g. > 1M) representative of saline brines. For geochemical species
absent from the PITZER database, the WATEQ database was used with an extended
Debye-Huckle formulation to calculate activity coefficients.
To demonstrate potential trace elements concentrations within post-combustion brine
solutions, stoichiometric concentrations of coal constituents were added incrementally to
one kg of brine, thus representing various ratios of coal to brine within the combustion
process. Reaction constituents were then free to react with the aqueous milieu, and a
variety of geochemical parameters were calculated, including the formation of aqueous
species (inclusive of ion pairs), solid phases (based on saturation indices), and gases.
Intuitively, as trace element mixture is added to one kg of brine, aqueous concentrations
of all trace elements increase (Fig. 27). However, when solid phases are allowed to
precipitate, the activity of multiple trace elements is repressed. For example, when
anoxic, low redox potential (Eh) aquifer conditions are specified, sulfate reduction is
thermodynamically favored, leading to the precipitation of trace element-bearing sulfide
minerals such as realgar, covellite, cinnabar, and galena (Fig. 28).
Table 2. Solid phase constituents
in simulated coal combustion.
Constituent mmol/kg (dw)
Carbon 66666
Arsenic 0.6
Cadmium 4.7
Chromium 0.8
Copper 0.4
Mercury 0.08
Nickel 0.5
Lead 1.5
Antimony 0.04
Selenium 0.05
Zinc 0.7
Calcium 350
Magnesium 50
Iron 1000
Titanium 50
Potassium 216
Sodium 600
Thus, anoxic conditions may lower trace element concentrations (Fig. 29) owing to the
sink provided by sulfidic minerals (9).
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
SI
-5
0
5
10
15
20
Realgar (AsS)
CdS
Covellite (CuS)
Cinnabar (HgS)
Galena (PbS)
Eskolaite (Cr2O3)
Anoxic Mineral Saturation Indicies
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
SI
-15
-10
-5
0
5
10
15
20
25
Delafossite (CuFeO2)
Calomel (HgCl)
Pyromorphite (Pb5(Cl(PO4)3)
Chromite (Fe,Mg)Cr2O4
Oxic Mineral Saturation Indicies
Mineral Saturation Indicies
Coal:Brine Ratio Coal:Brine RatioReaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
SI
-5
0
5
10
15
20
Realgar (AsS)
CdS
Covellite (CuS)
Cinnabar (HgS)
Galena (PbS)
Eskolaite (Cr2O3)
Anoxic Mineral Saturation Indicies
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
SI
-15
-10
-5
0
5
10
15
20
25
Delafossite (CuFeO2)
Calomel (HgCl)
Pyromorphite (Pb5(Cl(PO4)3)
Chromite (Fe,Mg)Cr2O4
Oxic Mineral Saturation Indicies
Mineral Saturation Indicies
Coal:Brine Ratio Coal:Brine Ratio
Figure 28: Mineral saturation indices for trace element-bearing minerals under anoxic (left)
and oxic (right) conditions (values exceeding 0 suggest precipitation). Arrows represent the
nominal coal:brine used with the reactor design presented in this report.
Figure 27: PHREEQC simulations of aqueous trace elements
after coal combustion and equilibration with brine. “Coal:Brine
Ratio” represents fractions of coal (kg) and associated trace
elements used for equilibration with 1 kg solution.
Precipitation of solid phases not allowed.
Mineral Precipitation Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
No Solid Phases
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
As Cr Cu Hg Pb Zn Cd
Oxic Brine/Effluent Equilibration
Coal:Brine Ratio
Mineral Precipitation Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
No Solid Phases
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
As Cr Cu Hg Pb Zn Cd
Oxic Brine/Effluent Equilibration
Coal:Brine Ratio
Minerals may also precipitate under oxic conditions and are capable of scavenging
trace elements (e.g., delafossite, Cu-bearing phase, and chromite, a Cr-bearing solid)
(Fig. 30). Additinally, brines that become oxygenated will favor the precipitation of
Fe(III) bearing minerals such as goethite (FeOOH), which act as strong adsorbents
(scavengers) of various trace elements (10). Alternatively, Fe-oxides such as goethite
may form in-situ if aerated solutions are injected into anoxic aquifers that have an
abundance of Fe(II). Either of these scenarios will result in a decrease in trace element
concentrations due to sequestration reactions on the iron oxide surfaces (Fig. 31). In
particular, arsenic is predicted to adsorb strongly to Fe(III) (hydr)oxides, resulting in a
dramatic drop in its aqueous concentration.
Figure 29: Aqueous concentrations of trace elements as a function of
coal:brine under anoxic conditions without solid phase precipitation (left) and
with trace element-bearing solid phase precipitation allowed (right).
Mineral Precipitation Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
No Solid Phases
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
As Cr Cu Hg Pb Zn Cd
Anoxic Brine/Effluent Equilibration
Coal:Brine Ratio Coal:Brine Ratio
Mineral Precipitation Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
No Solid Phases
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
As Cr Cu Hg Pb Zn Cd
Anoxic Brine/Effluent Equilibration
Coal:Brine Ratio Coal:Brine Ratio
Figure 30: Aqueous trace element concentrations as a function of coal:brine
under oxic conditions without solid phase precipitation (left) and with trace
element-bearing solid phase precipitation allowed (right).
Mineral Precipitation Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
No Solid Phases
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
As Cr Cu Hg Pb Zn Cd
Oxic Brine/Effluent Equilibration
Coal:Brine Ratio Coal:Brine Ratio
Mineral Precipitation Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
No Solid Phases
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
As Cr Cu Hg Pb Zn Cd
Oxic Brine/Effluent Equilibration
Coal:Brine Ratio Coal:Brine Ratio
Summary
High concentrations of problematic trace elements present within CO2-equilibrated brine
solutions are likely partially mitigated by secondary precipitation/adsorption reactions
occurring within the aquifer. The extent of these reactions will depend on the redox
status (aeration) of the delivered brine, as well as the redox status and mineralogical
makeup of the receiving aquifer. Future work will involve bench-scale testing of coal
combustion product equilibration with artificial brines in representative aquifer materials,
such as calcite or pyrite-bearing sandstones or shales.
References (Aquifer Interactions) 1. Mayer, K. U.; Frind, E. O.; Blowes, D. W., Multicomponent reactive transport modeling in
variably saturated porous media using a generalized formulation for kinetically controlled
reactions. Water Resour. Res 2002, 38, (9), 1174.
2. Parkhurst, D. L.; Appelo, C. A. J., User's guide to PHREEQC (version 2)-a computer program for
speciation, reaction-path, 1D-transport, and inverse geochemial calculations. U.S. Geol. Surv.
Water Resour. Inv. Rep. 1999, 99, 99-4259.
3. Wolery, T. J. EQ3/6, A Software Package for Geochemical Modeling of Aqueous Systems:
Package Overview and Installation Guide (Version 7.0); UCRL-MA-110662-PT-I; Lawrence
Livermore National Laboratory, Livermore, California: 1992.
4. Hanor, J. S., Physical and Chemical Controls on the Composition of Waters in Sedimentary
Basins. Marine and Petroleum Geology 1994, 11, (1), 31-45.
5. Martel, A. T.; Gibling, M. R.; Nguyen, M., Brines in the Carboniferous Sydney Coalfield, Atlantic
Canada. Applied Geochemistry 2001, 16, (1), 35-55.
6. Fontes, J. C.; Matray, J. M., Geochemistry and Origin of Formation Brines from the Paris Basin,
France .1. Brines Associated with Triassic Salts. Chemical Geology 1993, 109, (1-4), 149-175.
7. Klein, D. H.; Andren, A. W.; Carter, J. A.; Emery, J. F.; Feldman, C.; Fulkerson, W.; Lyon, W. S.;
Ogle, J. C.; Talmi, Y.; Vanhook, R. I.; Bolton, N., Pathways of 37 Trace-Elements through Coal-
Fired Power-Plant. Environmental Science & Technology 1975, 9, (10), 973-979.
Figure 31: Aqueous concentrations of trace element contaminants
as a function of coal:brine under aerated brine conditions without
(left) and with adsorption to the iron oxide mineral goethite (right).
No Sorption Reactions
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-8
1e-7
1e-6
1e-5
1e-4
1e-3
As Cr Cu Hg Pb Zn Cd
Sorption to Goethite Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-8
1e-7
1e-6
1e-5
1e-4
1e-3
Reaction Progress
Constituent Sorption to Precipitated Goethite
8. Querol, X.; Fernandezturiel, J. L.; Lopezsoler, A., Trace-Elements in Coal and Their Behavior
During Combustion in a Large Power-Station. Fuel 1995, 74, (3), 331-343.
9. Huertadiaz, M. A.; Morse, J. W., Pyritization of Trace-Metals in Anoxic Marine-Sediments.
Geochimica Et Cosmochimica Acta 1992, 56, (7), 2681-2702.
10. Brown, G. E.; Parks, G. A., Sorption of trace elements on mineral surfaces: Modern perspectives
from spectroscopic studies, and comments on sorption in the marine environment. International
Geology Review 2001, 43, (11), 963-1073.
Contacts (Aquifer Interactions)
Scott Fendorf: sfendorf@stanford.edu
Ben Kocar: kocar@stanford.edu
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