Coal Energy Conversion with Aquifer-Based Carbon Sequestration:

An Approach to Electric Power Generation with

Zero Matter Release to the Atmosphere

Final Report

Principal Investigators

Reginald E. Mitchell, Professor, Mechanical Engineering; Christopher Edwards,

Associate Professor, Mechanical Engineering; Scott Fendorf, Professor, Department of

Environmental Earth System Sciences

Post-doctoral and Graduate Student Researchers

Ben Kocar, Adam Berger, Eli Goldstein, John (J.R.) Heberle, BumJick Kim, Andrew

Lee, Paul Mobley, and Rebecca Pass

Abstract

Decarbonization of electricity production is a vital component in meeting stringent

emissions targets aimed at curbing the effects of global climate change. Most projected

pathways toward meeting those targets include a large contribution from carbon capture

and storage. Many capture technologies impose a large energy penalty to separate and

compress carbon dioxide (CO2). Also, injected neat CO2 in a deep saline aquifer is

buoyant compared to the aquifer brine and requires an impermeable seal to prevent it

from escaping the aquifer. The proposed alternative technology processes coal in

supercritical water pumped from a saline aquifer. The entire effluent stream is

sequestered, capturing all carbon and non-mineral coal combustion products in the

process. This stream is denser than the aquifer brine and therefore offers a higher level of

storage security, and could utilize aquifers without suitable structural trapping. This

technology also increases energy security in the United States, allowing for the use of its

coal resources while avoiding atmospheric pollution.

The overall objective of this project was to provide the information needed to enable

development of this technology that takes water from a deep saline aquifer, desalinates

and uses it to process coal at supercritical water (SCW) conditions, and returns the water,

which will contain dissolved CO2 and other combustion products, back to the aquifer.

The sensible energy of the hot combustion products is transferred to the working fluid of

a heat engine for electricity generation. The only system effluents are supercritical water

containing dissolved coal conversion products, and solids, composed of ash and salts that

precipitate from the supercritical water.

A complete architecture employing supercritical water oxidation has been studied,

including a liquid-oxygen-pumped air separation unit and regenerator system that heats

and desalinates the incoming brine. A system-level, thermodynamic analysis indicates an

overall plant thermal efficiency of 36.7%, taking into account all separation and storage

energy penalties. In addition, an exergy analysis gives insights into the least efficient

parts of the proposed system. Results from the model and these analyses elucidate how

the proposed system may be operated as a zero-emission electricity source and the

technical challenges that must be addressed for deployment.

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, was

developed and used to determine properties in the thermodynamic analysis. 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 was 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.

Geochemical models were 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. At the low

concentrations of dissolved trace elements in the brine being returned to the aquifer, no

adverse impacts on aquifer ecology are forecast.

Experimental facilities were built to permit the study of the coal reformation process

in SCW environments as well as to permit the study of combustion of the synthesis gas

produced. The coal reformation study was designed to determine coal conversion rates to

synthesis gas and char over selected ranges of temperature and pressure near the critical

point of water (Tc,H2O = 647 K, Pc,H2O = 221 bar). A 15-meter long reactor that provides

up to 15 minutes of residence time at specified temperature and total pressure was built

for this purpose. A 50 kW facility was also built to demonstrate stable combustion of the

synthesis gas in supercritical water with near-stoichiometric amounts of oxidizer. Stable

combustion has been demonstrated recently, and further combustion studies are

underway.

Wyodak coal, a low-ash, low-sulfur coal, was selected as the base-case coal for the

experimental study of the reformation process in SCW. This is a sub-bituminous coal

from Wyoming. It is typical of the coals from the Powder River Basin that are currently

being used for electric power generation in the United States.

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 capture and storage (CCS) option that may be more

acceptable to the public than other proposed CCS options.

Carbon capture and storage is a class of approaches to allow for the burning of fossil

fuels with minimal emissions of CO2 to the atmosphere. Models projecting technology

paths toward meeting emission targets typically include a large CO2 reduction from CCS

[1]. While enhanced oil recovery and enhanced coal bed methane recovery offer value-

added operation of CO2 storage, deep saline aquifers offer the largest storage potential of

CO2 [2]. The CO2 is less dense than the aquifer brine in an underground reservoir.

Therefore, storage in saline aquifers requires a suitable stratigraphic trap (such as a

caprock) above the reservoir to prevent the CO2 from escaping back to the surface. Over

time the CO2 becomes more secure in the reservoir by residual trapping in the aquifer

pores, dissolving into the brine, and finally forming carbonate minerals.

The purpose of this research is to analyze the feasibility of a plant that bypasses the

need for stratigraphic and residual trapping by injecting the CO2 into the aquifer already

dissolved into the brine. Once dissolved in the brine, it is denser than the surrounding

fluid and will sink to the bottom of the reservoir rather than escape back to the surface.

This type of CCS could allow for decreased monitoring of injectant plumes and less

potential remediation. This dissolved injection could be accomplished using a

conventional separation system on a pulverized coal (PC) power plant [3]. One could

pump the brine to the surface, dissolve the CO2 in it, and pump the brine—and the

captured CO2—back underground. This process would require multiple separation

processes that each carry an energy penalty, reducing overall efficiency. Another method

to achieve CO2 dissolution in aquifer brine, which is the subject of this research,

incorporates a supercritical water oxidation (SCWO) power plant that pumps aquifer

brine to the surface, reforms and combusts coal in supercritical water generated from the

brine, and returns the combustion products in solution to the aquifer. 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. Identified sites across the globe have an estimated capacity to store a minimum of

125 years of CO2 generated in coal-fired power plants at the rate of coal consumption for

electric power generation in 2005 [2]. A case study on a representative aquifer with 100

mD permeability and 12% porosity determined the effluent of a 500 MWe power plant

with a 50 year injection period would cover an area of 63 square miles [3].

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.

There have been few external developments in supercritical water oxidation during

the course of the research project. The most mature experimental SCWO work has been

performed in Switzerland with the goal of use in aqueous organic waste destruction [4].

The SCWO flame would provide an internal heat source for waste destruction processes.

The most recent work from 2007 was interested in minimizing reactor fouling and

plugging from precipitation of salts from the simulated waste fluid [5]. Recently, a low

emission power plant utilizing SCWO combustion of fine coal particles was proposed

and a system model was analyzed with a net efficiency of 28% [6]. This plant

incorporates a SCWO combustor with coal used directly as the fuel—not reformed—and

separates ash and CO2 from the effluent stream after combustion, in contrast to this

research project. Such an architecture does not avoid the separation energy penalties that

this research aims to circumvent, leading to a lower efficiency than the architecture

proposed below.

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. Due to the solvation

properties of supercritical water, 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 ultimately separated),

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.

In the following sections, significant results of each Area are discussed.

Area 1: Systems Analysis (Edwards)

System Model

The proposed system model has evolved from its initial concept following the

detailed analysis of the plant. Initially, a single regenerator/desalinator heat exchanger

was employed that preheated the aquifer brine and oxygen with the combustion products

after they exited the main heat exchanger of the heat engine. The two processes are now

split into separate systems, with an active desalinator operated as a two-pressure-level

system. This change avoids difficult practical implications brought on by handling salt

precipitation in a single regenerator/desalinator as the brine becomes supercritical [7].

This also avoids a pinch point that results from the difference between pseudocritical

points in the two streams.

The ASU is modeled as a liquid oxygen (LOX) plant with a high pressure pump.

The LOX pump allows for increasingly high pressures to be achieved with minimal

change in the work investment. The SCWO system is not analyzed in detail here because

the system can be modeled thermodynamically without regard to internal details. The

regenerator/desalinator is discussed in detail below, as well as the heat engine. The ASU

was analyzed completely, but is only discussed briefly in this report.

Aspen Plus 7.1 was used to analyze the system model. A mixed set of property

methods were used throughout the system including the NBS steam tables for water in

the Rankine cycle and after desalination (until entry into the SCWO system). The

electrolyte-NRTL method was used for the aquifer brine through desalination. The Peng-

Robinson property method with Boston-Mathias extrapolations for high temperature

gases was utilized for the remainder of the system.

Figure 1 shows the interactions of the main plant components. The process begins

with the aquifer brine pumped to the surface and entering the regenerator/desalinator.

About 10% of the brine is desalinated, since only that portion is required by the SCWO

system to moderate the combustion process. The remaining brine is needed to keep all of

the combustion products, including CO2, in solution at re-injection. The amount of brine

necessary for dissolution is calculated from published CO2 solubility relations [8]. Neat

oxygen also enters the regenerator/desalinator from an ASU and is preheated. Air

separation is necessary because the solubility of nitrogen in water is much less than that

of CO2, and its inclusion would prohibitively increase the amount of brine needed to

dissolve the combustion products. Some of the desalinated water goes to coal slurry

preparation. The slurry is sent back to the regenerator/desalinator to be preheated before

the oxygen, water, and slurry streams are passed to the SCWO system. The SCWO

system is operated at 250 bar to remain well above the critical pressure (221 bar), thereby

avoiding major fluctuations in properties. A low-sulfur Powder River Basin coal was

chosen for the model [9]. The coal exergy was determined by the methods outlined by

Kotas [10].

The regenerator/desalinator contains a set of staged heat exchangers, designed to best

transfer the enthalpy of the combustion products stream to streams where heating is

required. To understand its operation, it is best to follow both the schematic in Fig. 2 and

the temperature profiles in the various heat exchangers shown in Fig. 3. The products

enter at the top right of Fig. 2 and heat the desalinated water and coal slurry to

supercritical temperature in a regenerator, before they enter the SCWO system. The

products stream temperature profile inflects as it passes over the vapor dome. Therefore,

the desalination process is split into a two-pressure-level system in order to best match

the temperature profile of the products stream in the intermediate pressure (IP) heat

exchanger. Two pressure levels are also necessary to provide enough desalinated water

for the SCWO system without creating a pinch point, and to prevent salts from

precipitating in the brine if too much water is evaporated.

Figure 1: Diagram of a SCWO system combined with a heat engine

powered by the hot combustion products, an air separation unit, a

slurry preparation unit, and a combined regenerator/desalinator.

AQUIFER BRINEINJECTANT

Lift

Pump

Injection

Pump

COAL

Slurry

Prep.

AIRASU

NITROGEN

IP

Pump

LP

Pump

HP

Pumps

Slurry

Pump

Regenerator

IP Desal.

Products

Stream

To SCWO

System

Oxygen

LP Desal.

Q1

Q2

IP HX

LP HX

Brine Econ.

Slurry HX Slurry Regen.

Figure 2: Diagram of the regenerator/desalinator. Q1 and Q2 are transferred

to cooling water during startup and to the HRSG at steady state.

Figure 3: Temperature vs. enthalpy plot of the streams in the regenerator/

desalinator. (Line type denotes the streams in each heat exchanger; arrows

give the direction of each stream)

After the regenerator, the combustion products pass through the IP heat exchanger

before being mixed with the other outlet flows and re-injected. The IP heat exchanger

brings the IP brine stream under the vapor dome, evaporating some water while

concentrating the remaining brine. After flashing in the IP desalinator, the concentrated

brine is drained and used to preheat the low pressure (LP) brine in the LP heat exchanger.

This process brings the LP brine to the saturated-liquid state. That liquid is then

evaporated by the condensation process of the IP desalinator. The LP desalinator flashes

the partially evaporated LP stream, and again returns the concentrated brine to a heat

exchanger to preheat the coal slurry to near-critical temperatures. The condensation

process in the LP desalinator is used to preheat the oxygen stream to near-critical

temperatures. These two preheating processes are not completely matched in capacity to

cool the brine and condense the water vapor, so cooling water or some of the bypassed

aquifer brine is used to provide additional cooling as shown by Q1 and Q2.

The desalinated water streams are then over-pressurized to 450 bar to minimize the

effect of a pinch point in the regenerator from the slopes near the pseudocritical point and

create a better temperature match with the three streams in the regenerator. The pressures

in the desalination process were chosen at 110 bar (IP) and 100 bar (LP) to minimize

exergy destruction in the heat exchangers. The intermediate pressure of 110 bar was

chosen to match temperatures well with the pseudocritical point of the products. This is

about half the critical pressure, which results in a saturation temperature close to the

temperature of the products stream. The composition of the products stream has a lower

critical temperature than pure water, and lowers the temperature of the products stream as

it passes over the vapor dome. There is also a pinch point to consider as the aquifer brine

begins to change phase, chosen to exceed 25 K in temperature difference. With the

desalination pressures close to each other, this allows for a small temperature difference

in the IP desalinator condenser, where the temperature difference is constant as each

stream changes phase. This low temperature difference minimizes exergy destruction in

the process.

Air NitrogenOxygenLOX

Pump

HX 1

LP

HP IP

Turb.

HX 3

HX 3HX 2

High

Press.

Col.

Low

Press.

Col.

Q3

Q3

Figure 4: Diagram of the Air Separation Unit. HX 3 is shown broken up

with both streams in the right block cooled by the stream in the top block.

A LOX ASU was analyzed based on available designs, and further optimized for this

system [11, 12]. The architecture used is displayed in Fig. 4. The system includes three

pressure levels of air compression. One stream is pressurized to that of the high pressure

(HP) column, and cooled in a series of heat exchangers using the separated nitrogen and

oxygen streams. Another is over-pressurized, cooled, and then expanded across a throttle

for Joule-Thomson cooling. The final air stream is over-pressurized, cooled and then

sent through a turbine to drive the IP compressor and in doing so obtain additional

cooling. These streams are then sent throughout the system with a HP column containing

a condenser that acts as a reboiler for the LP column. Streams entering the LP column

are subcooled in heat exchanger (HX) 3 by the nitrogen product.

The SCWO system is modeled as a single combustion process because the outlet state

of the system only depends on the streams entering the system. However, one could

imagine a system similar to a gasifier, where the coal slurry enters a reformer with some

oxygen and supercritical water. The reformer breaks down the coal into a synthetic fuel,

the solids are separated, and the remaining oxygen and water are injected in a combustor

at near stoichiometric proportions (since both the fuel and oxygen are valuable

resources). The combustion products then transfer heat to a heat engine, since the

products need to remain pressurized for re-injection. Finally, the products are cooled in

the regenerator/desalinator heat exchangers and injected back into the saline aquifer.

Main Heat

Exchanger

Brayton Cycle

Rankine Cycle

Work

Work

Heat Rejection

from Condenser

COMBUSTION

PRODUCTS

TO REGEN/

DESAL

Figure 5: Heat engine layout for the plant, including a helium Brayton

cycle, bottomed by a Rankine cycle.

The heat engine, depicted in Fig. 5, consists of a helium Brayton cycle heated by the

combustion products in the main heat exchanger, and bottomed by a conventional (water-

based) Rankine cycle. Helium was chosen as the working fluid because of its high

thermal conductivity and specific heat ratio. Both are important to aid heat transfer with

the combustion products. The helium flow rate is chosen to match the capacity of the

products stream in the main heat exchanger. To reduce creep stresses at this high

temperature, the pressures of the streams entering the heat exchanger are matched. The

Rankine cycle pressure is specified to ensure that the quality of the water exiting the

turbine is greater than 0.88. Maintaining a pinch point difference in the HRSG of 17 K

dictates the water flow rate in the Rankine cycle.

There are a number of parameters that are variable within the heat engine and

regenerator/desalinator. These systems are highly coupled and therefore their parameters

are difficult to optimize to find the maximum plant efficiency. For example, the Brayton

pressure ratio determines the temperature of the helium entering the main heat exchanger

after compression. This temperature sets the total heat transferred, for a specific

effectiveness, in the main heat exchanger. Meanwhile, a pinch point is possible near the

low-temperature side of the main heat exchanger, due to the concavity in the products

stream temperature profile as it nears the pseudocritical point. In addition, the products

stream must have a high enough temperature to heat the slurry and water to supercritical

temperature in the regenerator. However, as the slurry and water temperature exiting the

regenerator increases, additional desalinated water is required to moderate the SCWO

system. This, in turn, raises the products stream heat transfer in the IP heat exchanger,

increasing the chance of a pinch point within the regenerator and IP heat exchanger.

These competing factors, among others, must be optimized to find the most efficient

operating conditions for the system.

Model Results

The LOX ASU was found to have a work requirement of 1.4 MJ/kg-O2, including

pressurization to the system combustion pressure of 250 bar. Figure 6 shows the oxygen

work intensity of the LOX ASU and a range for a state-of-the-art GOX ASUs with gas

compression to varying pressures. The gaseous work requirement was found from

published data at 1.7 bar [12] and was compressed further with a polytropic efficiency of

90%, in four intercooled stages. The number of intercooled stages could make the

gaseous system a more, or less, work-intensive option compared to the liquid-pumped

ASU. The oxygen leaving the ASU at full pressure from a LOX/pumped plant is near

ambient temperature, while a GOX stream could be at a much higher temperature from

the compression process, if not well intercooled. This hot oxygen could be advantageous

to reduce preheating loads or eliminate the need for preheating in the

regenerator/desalinator. However, the low heat capacity of the oxygen requires less than

14% of the heating load from the LP desalinator condenser as seen in Fig. 3. In addition,

decreased levels of oxygen preheating would coincide with increased oxygen work

intensity as intercooling is reduced. Therefore, the differences in the systems are quite

small, with both gaseous and liquid ASUs appearing as good options for this plant.

0 100 200 300 400 500

0.8

1

1.2

1.4

1.6

Exit Pressure, bar

Oxygen Inte

nsity, M

J/k

g O

2

Liquid Pumped ASU

Compressed Gas IPCC ASU-High

Compressed Gas IPCC ASU-Low

Figure 6: Plot of the oxygen intensity of the modeled liquid pumped ASU

versus the high and low range of published data [12] for a gaseous ASU

with intercooled compression. The dashed line is the nominal operating

pressure of the SCWO plant.

The temperature vs. products-specific enthalpy plot shown in Fig. 3 displays the

functioning of the regenerator/ desalinator. The IP desalinator and regenerator are

designed to minimize temperature differences while cooling the products stream

(decreasing irreversibility). The rest of the heat exchangers are heated by the cooling of

the concentrated brines and condensation of water vapor, shown underneath the IP

desalinator in Fig. 3, as described in the Model section. The flow direction of each

stream in the heat exchangers is shown by the arrows on the plot. A pinch point of 25 K

was found in the IP desalinator, and 11 K in the regenerator.

After an exergy analysis—investigating the irreversibility in each plant component —

it was found that large exergy destruction occurs in the LP desalinator and brine

economizer. This destruction is because the extra cooling loads Q1 and Q2, shown in

Fig. 2, remove enthalpy from a valuable resource at high temperature and provide no

useful service. Given that the LP desalinator saturation temperature is above that of the

water in the HRSG, it is possible to heat a portion of the HRSG evaporator while

condensing steam in the LP desalinator. The brine economizer can also transfer heat to

the HRSG economizer. This heat integration, shown in Fig. 7, may be difficult to

accomplish during startup procedures in a practical installation. Therefore, the LP

desalinator should be setup to be cooled by bypassed aquifer brine during startup

procedures, while using the heat integration during steady state operation for maximum

efficiency.

Main Heat

Exchanger

Brayton Cycle

Rankine Cycle

Work

Work

Heat Rejection

from Condenser

COMBUSTION

PRODUCTS

TO REGEN/

DESAL

Q1 Q2

Figure 7: Heat engine schematic of the plant at steady state conditions,

including heat input to the economizer and boiler from the

regenerator/desalinator as depicted by Q1 and Q2.

The power balance for a 500 MW plant is shown in Table 1. An overall efficiency of

36.7% was found. This efficiency includes all energy penalties for oxygen separation

and pressurization of the waste stream to the injection pressure.

Figure 8 shows the exergy distribution of the incoming coal for the SCWO plant

under startup and at steady state conditions. During startup, roughly 25% of the fuel

exergy leaves as net work out and the rest is destroyed in the various sections of the plant.

A large amount of exergy destruction occurs in the regenerator/desalinator since the heat

exchangers are inherently irreversible due to temperature differences between the

streams. After heat integration is incorporated, the exergy distribution, shown by the

green bars on the right in Fig. 8, display an 11% decrease in exergy destruction in the

regenerator/desalinator. Almost all of that exergy is transferred into increased work

output. The integrated system has an efficiency of 36.7% on an LHV basis. The brine

economizer is able to provide 30% of the heating in the economizer and 56% of the

heating in the evaporator. This extra heating permits a larger water mass flow rate in the

Rankine cycle, providing a higher net work output from the plant. In addition, a closer

temperature match exists with the helium in the superheater as a result.

Table 1: Work flow of power plant under steady state conditions.

Component Power (MW)

Brayton Cycle

Compressor -314.4

Turbine 527.1

Net 212.7

Rankine Cycle

Condensate Pump -0.03

Feed Pump -2.69

Turbine 461.6

Net 458.8

ASU -146.1

Water Pumps -25.5

Overall Plant 500.0

Firing Rate (LHV basis) 1364.1

Efficiency ( % LHV) 36.7%

Figure 8: Clustered bar plot comparing the plant under startup (left) and

steady state (right) conditions with heat integration, leading to increased

efficiency.

This SCWO plant is compared to other coal technologies with CO2 capture in Fig. 9.

The SCWO plant is a more efficient option than any of the three currently available

technologies using coal, including subcritical and supercritical PC with post-combustion

capture, and IGCC with pre-combustion capture [13]. These plants only capture 90% of

the CO2, and do not capture other toxins such as mercury. The SCWO plant captures

100% of the CO2 from the plant with zero emissions to the atmosphere, including trace

species. The SCWO injectant is also the most secure given that it is pre-equilibrated with

the aquifer brine.

Figure 9: Overall efficiency of current coal technologies with 90% CO2

capture and a SCWO plant with 100% capture and equilibrated.

Further increases in efficiency could follow from the insights gained from the exergy

analysis. Currently, the largest exergy destruction occurs due to irreversibilities during

combustion. The important variables that affect exergy destruction from combustion in a

SCWO system are the reactant and product temperatures. Figure 10 displays the

percentage of incoming fuel exergy destroyed by combustion against these two

parameters. This contour plot illustrates the Extreme States Principle as outlined by Teh

[14]. Teh found that combustion exergy destruction could be lowered by going to high

energy states, at high pressures and temperatures, before combustion. In the case of

SCWO systems, exergy destruction from combustion has a weak dependence on

pressure. This causes reactant and product temperatures to be the dominant factors.

Figure 10 shows the current system, marked by a circle on the plot, operating at a

maximum product temperature of 1600 K. Material constraints in the main heat

exchanger limit the product temperature in this case. The dashed arrow in Fig. 10 shows

that a reactant temperature increase of roughly 75 K at a fixed product temperature could

lead to a 5% reduction in exergy destruction, while another 200 K increase would be

necessary to achieve another 3% reduction. The heat integration effects reveal that most

of this reduction could translate into a similar increase in net efficiency. Future analysis

will attempt to achieve such changes while balancing the optimum operation of the plant

as outlined in the System Model section. Future research will also include a sensitivity

analysis for the important parameters of the system.

Conclusions

The modeled system has furthered understanding of the SCWO power plant initially

proposed under this project. Either a liquid-pumped or gaseous ASU was found to be

appropriate for the modeled SCWO plant. Heat integration was found to substantially

increase the overall efficiency of the plant. The efficiency of the plant was found to be

36.7 %, higher than current low-emission coal technologies. A path towards further

efficiency improvements through combustion exergy destruction was also identified.

Figure 10: Combustion exergy destruction (% of fuel exergy) as a function

of reactant and product temperature at 250 bar. The present system state is

marked with a circle. The x denotes the state that would lower exergy

destruction by 5% if the reactant temperature is increased by 75 K.

The analysis of the full SCWO system shows that it is a potential option for the

efficient use of coal in electricity generation with zero emissions to the atmosphere. The

SCWO technique has the added advantage of injecting a pre-equilibrated mixture of

aquifer brine and combustion products, including CO2. This mixture is denser than the

original aquifer brine, allowing for a more secure storage strategy than neat CO2

injection. This type of injection may allow for decreased monitoring and liability costs

for CO2 storage in saline aquifers. It may also lead to an increased number of aquifers

available for CCS, if it is found that a suitable caprock is not required for effective long-

term storage. Increasing the potential number of storage aquifers suitable for CO2 storage

could reduce costs of establishing that an aquifer will retain the majority of the injected

CO2, lowering costs associated with permitting. Injection sites could also be located

closer to CO2 sources, reducing transportation costs. Reducing the transmission, storage,

and monitoring costs all could have significant impacts on reducing greenhouse gas

emissions, however, the largest cost of CO2 abatement is from the capture process. This

technology could have a large impact on the cost of CO2 capture, since separation

processes are minimized as well as their associated energy penalties, resulting in an

efficient power plant.

Publications

1. Heberle, J.R., & Edwards, C. (2009). Coal energy conversion with carbon

sequestration via combustion in supercritical saline aquifer water. International

Journal of Greenhouse Gas Control, 3 (5), 568-576.

2. Mobley, P.D., Pass, R.Z., Edwards, C.F. Exergy Analysis of Coal Energy

Conversion With Carbon Sequestration via Combustion in Supercritical Saline

Aquifer Water. 2011 Energy Sustainability Conference. Accepted for publication.

Contacts (Systems Analysis)

Christopher F. Edwards: cfe@stanford.edu

Paul D. Mobley: pdmobley@stanford.edu

Rebecca Z. Pass: rzpass@stanford.edu

References (Systems Analysis)

1. EPRI, 2007. The Power to Reduce CO2 Emissions.

2. Benson, S. C. (2005). Underground geological storage. In B. D. Metz, IPCC

Special Report on Carbon Dioxide Capture and Storage (pp. 195-276).

Cambridge: Cambridge University Press.

3. Burton, M. B. (2007). Eliminating buoyant migration of sequestered CO2 through

surface dissolution: implementation costs and technical challenges. Society of

Petroleum Engineers Annual Technical Conference and Exhibition. Anaheim, CA,

USA.

4. 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.

5. Prikopsky, K., Wellig, B., & von Rohr, R. (2007). SCWO of salt containing

artificial wastewater using a transpiring-wall reactor: Experimental results.

Journal of Supercritical Fluids , 40, 246-257.

6. Donatini, F. (2009). Supercritical water oxidation of coal in power plants with

low CO2 emissions. Energy , 34, 2144-2150.

7. Armellini, F. T. (1994). Precipitation of sodium chloride and sodium sulfate in

water from sub- to supercritical conditions: 150 to 500 C, 100 to 300 bar. The

Journal of Supercritical Fluids , 7 (3), 147-158.

8. Hangx, S. (2005). Behavior of the CO2-H2O System and Preliminary

Mineralisation Model and Experiments. Utrecht University, Department of Earth

Sciences.

9. Argonne Premium Coal Sample. (n.d.). Retrieved August 10, 2010, from Argonne

National Laboratory: http://www.anl.gov/PCS/report/part2.html

10. Kotas, J. (1985). The Exergy Method of Thermal Plant Analysis.

11. Howard, H. (2008). Patent No. 7,437,890. US.

12. Thambimuthu, K. S. (2005). Capture of CO2. In B. Metz, IPCC Special Report on

Carbon Dioxide Capture and Storage (pp. 105-178). Cambridge: Cambridge

University Press.

13. DOE/NETL. (2010). Cost and Performance Baseline for Fossil Energy Plants,

Volume 1: Bituminous Coal and Natural Gas to Electricity.

14. Teh, K.-Y. (2007). Thermodynamics of Efficient, Simple-Cycle Combustion

Engines. PhD Thesis, Stanford University, Palo Alto, CA.

Area 2: Supercritical coal reforming (Mitchell)

The proximate and ultimate analyses of Wyodak coal, the sub-bituminous coal from

Wyoming selected for study, are presented in Table 1. On a dry basis, the coal has a

lower heating value of 28.2 MJ/kg. Samples of the coal were ground and size classified

to obtain particles in the size range 75 – 106 m for use in the SCW reforming tests.

Table 2: Proximate and Ultimate Analyses of Wyodak Coal

Proximate Analysis Ultimate Analysis

Property Wt-% Property Wt-%

Fixed carbon 35.06 Carbon 51.42

Volatile matter 33.06 Hydrogen 4.16*

Moisture 26.30 Nitrogen 0.69

Ash 5.58 Sulfur 0.32

Chlorine < 0.01

Oxygen 11.53#,*

Ash 5.58

*excludes moisture #determined by difference

When coal particles are heated, volatile matter (organic coal decomposition products)

is released and char particles are produced. The volatile matter consists of evolved gases

(CH4, CO, CO2, H2, H2O, and light hydrocarbons (C2H2, C2H4, etc.)), and tars. The coal

decomposition products classified as tars are organic compounds having molecular

weights greater than benzene (78 kg/kmol) and which condense at low temperatures. The

tars released from the coal (the primary tars) are oxygenated compounds. These are

converted to secondary tars (phenolic compounds and olefins) and tertiary tars

(polycyclic aromatic hydrocarbons, PAH) in steam gasification environments. The

secondary tars are also converted to PAH in such environments. In SCW environments,

the tars and PAH hydrolyze rapidly, yielding CO, CO2, H2, and H2O, primarily, with

smaller amounts of CH4 and other light hydrocarbons. These species are all completely

miscible in supercritical water.

The char particles remaining subsequent to coal devolatilization react with water in

SCW environments at a rate somewhat slower than the rate that the tars are converted to

dissolved gases. Consequently, the size of the reformation reactor is governed by the

conversion time of the coal char in SCW environments. Our experimental efforts are

devoted towards providing this information.

Autothermal Reformer Operation

Calculations using the thermodynamic model developed in Area 1 activities indicate

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. 11 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).

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 Fig. 12, 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.

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 11: Required oxygen for autothermal operation of the

reformer at 647.3K with preheating of the reformer feed water.

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 12: Heating values of synthesis fuel at 647.3K

with water preheating.

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.

Shown in Fig. 13 is a comparison between predicted synthesis gas compositions

assuming ideal gas behavior and real gas behavior at 290 atm and 647.3 K, a supercritical

water environment. The elemental composition of Wyodak coal, shown in Table 2, was

used to determine the atom populations. The scale of the vertical axis is logarithmic.

Noticeable differences are significant. In particular, when real gas behavior is taken into

account, the predicted synthesis gas has less CO and H2 and more CH4 than the synthesis

gas predicted assuming ideal gas behavior.

Figure 13: Predicted equilibrium evolved gas composition at 290 atm, 647.3 K.

Atom populations were determined from the elemental analysis for Wyodak coal,

presented in Table 2.

Coal Char Reforming

No gasification model has been developed that permits the accurate prediction of

particle mass loss rates under supercritical water conditions. Based on previous work, we

have developed the heterogeneous reaction mechanism presented in Table 3. In the

mechanism, water dissociatively chemisorbs onto free sites (Cf) on the carbonaceous

surfaces yielding adsorbed OH and H (C(OH) and C(H), respectively). Adsorbed O

(C(O)) is also formed as C(OH) dissociates on the surface. The adsorbed species are

involved in surface reactions that result in the release of H2 and CO, primarily, with

smaller amounts of CH2, CH and COH (which rearranges to HCO) also released. Once

in the gas phase, the CH and CH2 are involved in homogeneous reactions that result in

the formation of methane. In order to accurately account for the changes in the gas phase

composition due to chemical reaction, a detailed homogeneous reaction mechanism is

required. In order to avoid this complexity, we assume that the global reactions shown in

Table 4 are kinetically limited, and that the H2/O2 system of reactions is equilibrated.

In our approach, the Arrhenius parameters for the reaction rate coefficients of the

heterogeneous reactions in Tables 3 and for the rate-limiting homogeneous reactions in

Table 4 are determined by adjusting the parameters to provide agreement with data

obtained in experiments performed in this project. The rate coefficients for the reverse

reactions of the rate-limiting homogeneous reactions are determined from the equilibrium

constants and the forward rate coefficients. Equilibrium constants for the homogeneous

reactions are determined from data provided in the JANAF Thermochemical Tables and

fugacity coefficients to account for non-ideal gas behavior.

The intrinsic chemical reactivity of the carbonaceous material is defined as the rate

that a carbon atom is removed from the bulk. Based on the heterogeneous reaction

mechanism shown in Table 3, carbon reactivity (Ri,C, in kg/m2-s) is given by

RiC k4 C(H) 2 k6 C(O) k7 C(H) k8 C(OH)

where the brackets denote the concentration of the species enclosed.

Table 3: Reduced heterogeneous reaction mechanism for char reactivity to H2O

2Cf + H2O <==> C(OH) + C(H) (1)

C(OH) + Cf <==> C(O) + C(H) (2)

C(H) + C(H) <==> H2 + 2Cf (3)

C(H) + C(H) ==> CH2 + Cf (4)

C(H) + C(OH) ==> H2 + C(O) + Cf (5)

C(O) + Cb ==> CO + Cf (6)

C(H) + Cb ==> CH + Cf (7)

C(OH) + Cb ==> COH + Cf (8)

Table 4: Rate-limiting global and equilibrium homogenous reactions

involved in methane formation and HCO destruction

Rate-limiting global reactions

CH + 2 H2O <==> CH4 + O + OH (9)

CH2 + H2O <==> CH4 + O (10)

CHO + H2O <==> CO2 + H2 + H (11)

Equilibrated reactions

H + OH + M <==> H2O + M (12)

H + H + M <==> H2 + M (13)

O + O + M <==> O2 + M (14)

If the gas phase species reach a state of chemical equilibrium, the above mechanism

will predict the expected behavior at both low and high pressures. Shown in Fig. 14 are

predicted synthesis gas compositions for a gasifier operated under autothermal, SCW

conditions at 290 atm and 647.3 K, with and without preheat. Atom populations were

based on the composition of Wyodak coal. Compared to the synthesis gas produced in a

gasifier operated under typical autothermal, Integrated Gasification Combined Cycle

(IGCC) gasification conditions (29 atm, 1700 K yielding primarily CO and H2), the

synthesis gas of the SCW gasifier has a considerably higher methane fraction with little

CO and H2.

Figure 14: Predicted equilibrium evolved gas compositions. SCW gasification at

290 atm, 647.3 K. IGCC gasification at 29 atm, 1700 K. Atom populations were

determined from the elemental analysis for Wyodak coal, presented in Table 2.

Experimental Facility

An experimental facility was designed, built and assembled that permits the

acquisition of data needed to determine the rate coefficients for reactions in Tables 3 and

4 that are 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 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

SCW (no preheating) SCW (preheating)

NH3. A schematic of the experimental facility is shown below in Fig. 14. A picture is

shown in Fig, 15.

Figure 14: Experimental setup for coal gasification under supercritical conditions.

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 a high-pressure

air tank. The air to drive this pump is regulated by a FESTO air controller to control the

water pressure. The water supplied to the reactor is preheated using three Chromalox

MaxiZone cartridge heaters surrounded by aluminum powder and calcium silicate

insulation. The water pump that supplies water to the slurry is manufactured by Haskel,

and the pump-driving air is controlled by another FESTO air controller.

A high-pressure piston-accumulator purchased from High Pressure Equipment

Company is used to deliver the coal-water slurry to the reactor. The slurry is loaded on

one side of the piston-accumulator by a peristaltic pump before the start of an

experiment. Pure high-pressure water from the Haskel pump delivers the loaded slurry

to the reactor by moving the piston from the pure water side toward the slurry side of the

accumulator. For slurry loadings at low pressures, a Vector peristaltic slurry pump is

used instead of the Haskel pump.

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 required flow rates of pressurized water. The fluctuations in the

flow due to these pump strokes are dampened by using pulse dampening accumulators

between the water pumps and the pressure transducers. The accumulators contain a

nitrogen-filled bladder inside a steel housing. The water pressure fluctuations are

dampened by expanding and contracting the bladders. The bladder accumulators were

purchased from HyDac.

Figure 15: A picture of the experimental facility.

The oxygen that is supplied to the reactor is pressurized using a SC Hydraulic gas

booster that 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 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.

The coil-shaped 15-meter long reactor (see Fig. 16) is made of Inconel-625 to endure

high-pressure, high-temperature environments. The reactor is submerged in a bath filled

with sand (pulverized aluminum oxide) and sixteen Watlow Firerod cartridge heaters.

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.

The flow leaving the flow reactor is rapidly quenched in a cool-down tube in order to

prevent further reactions. To decrease the cool-down time, the tube diameter at the

reactor exit is decreased in order to increase the velocity of the flow and improve the

overall heat transfer by convection. Cooling is provided by submerging the cool-down

tube in a chamber filled with flowing cold water. Calculations indicate that the 1-meter

long cool-down tube is sufficient to freeze all chemical reactions in less than 20 seconds.

With this arrangement, the coal reaction time is essentially the residence time of the

particles in the 15 meter coiled reactor.

Figure 16: The 15-m coiled tubular reactor.

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 entire system pressure is controlled by a Tescom back-pressure-regulator. A

Marsh Bellofram air controller supplies sufficient air to the back-pressure-regulator to

maintain system pressures up to 374 atm.

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.

Air supplies to drive the water pumps, the oxygen booster, and the HIPCO valves are

controlled by ASCO solenoid valves, which are operated by OPTO22 modules on a data

acquisition box. The data acquisition/system control box consists of a data acquisition

board (DaqScan 2001) combined 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

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).

Continuing Research (Supercritical water reforming)

Coal conversion tests in supercritical water environments are underway. Extents of

coal conversion are being measured at selected residence times in various SCW

environments. The compositions of the synthesis gases produced are being measured as

well for each of the environments. The data will be used to determine rate coefficients

for the reactions in Tables 3 and 4. In order to help define the rates of some of the

reaction pathways, reactivity tests are also being performed in our pressurized

thermogravimetric analyzer.

In March 2010, the high-pressure compressor that supplies the high pressure air to the

pumps and solenoid valves broke, preventing the SCW facility to be operated.

Diagnosing the reason for the failure and deciding whether or not the compressor was

repairable, securing funds to either repair or replace the compressor, purchasing a new

compressor and installing it took a considerable amount of time. It was not until

February 2011 that the SCW facility was operational, permitting experiments to continue.

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. BumJick 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 above.

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 discharge [1]. 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.

Fig. 17 is a simplified schematic showing 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. Several changes have been made in the configuration of these

systems since the last annual report.

Figure 17: Diagram of key components of the supercritical water 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 18: Liquid fuel delivery system.

The fuel system is shown in Fig. 18. Prior to running the combustor, a low-pressure,

air-powered pump (not shown) transfers fuel into a bladder accumulator in the laboratory.

This accumulator was 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 the fuel

mass fraction entering the combustor can be varied.

Figure 19: Oxygen delivery system.

Figure 19 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 just after it can deliver oxygen at an approximately constant pressure while the

piston in the booster cycles. Next are valves for manually isolating and bleeding

different parts of the system. A stepper-driven metering valve is used for remote control

of the flow rate, similar to the fuel system. The tube downstream of these valves is fitted

with a water 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. The remaining components are similar in function to those in

the fuel system. There is a remote solenoid valve for positive oxygen shut-off. The flow

meter consists of a needle valve in parallel with a differential pressure transducer. This

arrangement is similar in operation to a laminar flow element. The needle valve is set to

create an appropriate pressure drop, but is not adjusted during the experiment. A

thermocouple is located just upstream of the flow meter so that the latter’s readings may

be corrected for temperature variations. A gauge pressure transducer after the flow meter

arrangement serves a similar purpose. 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 by changing the relative

density and viscosity of the fuel/water and oxygen/water streams.

The low-pressure water system can be seen in Fig. 20. 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

rotameter (shown as FI for flow indicator) 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. 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 isolation 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. 21. 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-level switch that controls the motor driving the triplex plunger pump to

prevent it from running dry.

The outlet stream from the high-pressure apparatus re-enters the low pressure system

near the top left of Fig. 20 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 an average of about 16°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 to the drain tank. This tank

has a centrifugal pump connected to its outlet with a liquid-level switch to make sure that

Figure 20: Low-pressure water handling system.

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. 21. 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. At

the outlet of the triplex pump described above is a relief valve set to 5000 psi. After the

pump, the water passes by a bladder accumulator and through a manual needle valve.

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 accumulator and valve

together act as a low-pass RC filter to attenuate high-frequency pressure waves. The cut-

off of the filter can be tuned by adjusting the valve. The water then enters a manifold,

where it branches out into seven 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 decouples

the flow control in the output lines from each other.

Figure 21: High-pressure water handling system.

After the manifold, the water lines each have a stepper-driven needle valve, spur-gear

flow meter (except one), 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 two lines

are for dilution cooling water. One goes directly to the cooling section of the high-

pressure stack. The other joins the line from the stack before the BPR.

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. (This temperature limit is discussed below.) 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 is run through the pressurized system and out through this

valve. By measuring the mass of water output in a known time at several rates, the true

mass flow rate is correlated with the sensor output. The response is quite linear even at

low flow rates, despite the fact that this type of flow meter is intended for use with fluids

more viscous than water.

The BPR downstream of the combustor was initially rated to 650°F (617 K) by the

manufacturer. This is the highest temperature rated model produced by Tescom. The

difficulty in operating this assembly at high temperatures lies with the seals. After two

identical failures of a spring-energized, fluorocarbon radial c-seal, the failing seal was

replaced with an EPDM o-ring. This material is only rated to 250°F (394 K). To meet

this lower maximum inlet temperature, more dilution cooling was needed. With a

reasonable pressure drop from the water manifold to the stack BPR, the original 1/4 in.

dilution cooling line was not sufficient, even with the metering valve fully open. A new

1/2 in. line was installed to satisfy the increased demand. This new line is the right-most

water line leading from the manifold in Fig. 21. Presently it does not have a spur gear

flow meter as do the other lines. The flow rate is simply adjusted up or down depending

on the temperature of the mixed stream flowing into the BPR.

The heaters, shown assembled in Fig. 22, 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

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 in

series.

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

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 by an embedded process controller.

This arrangement allows for the possibility of operating streams at different temperatures

to simulate various operating conditions in later experiments.

Figure 22: Photograph of the 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.

Each unit is comprised of a steel shell with machined header 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 were filled with fiberglass insulation. The

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. 23. 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 Inconel 625 forgings. 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. 23) 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. (The viewing axis is normal

to the page, so only the inside of the far window is visible in this section.) 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 a

darker blue-gray than the main body). 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. The pattern of 3/64 in. holes for this purpose is

shown. 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. (This inlet is not in the section plane; the

intersection of the port with the main passage is circled.)

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.

The high-pressure stack is shown installed in the lab in Fig. 24. 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.

Figure 23: Section view of high pressure stack design, showing fluid

ports, burner, axial viewing window, and liners. Ports are labeled with

their present use: TC=thermocouple, P=pressure transducer, N.C.=no

connection (these ports are plugged).

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

seals are Inconel 718 covered with 0.001 in. silver electroplate, with a thin gold layer in

between to improve adherence of the silver. The first set of seals received had pits in the

silver layer. With most seals from this batch, there was at least one and usually several

pits on the seal line that were too large to be filled by yielding silver during installation.

The result was leaks that began small, but grew noticeably over several minutes as the

escaping water and steam eroded the leak channel. The manufacturer provided a

replacement set of seals with a smoother silver layer. These seals have operated

satisfactorily in the header section below the burner and at the windows. However, they

have continued to be problematic at the two clamped joints.

In response to the continued leaks, other sealing methods were sought. C-seals

operate by using fluid pressure to press the seal against the gland. Operating on the

theory that increased pressure could result in a good seal, a ring seal was made having a

solid cross section with a raised point on each side (Fig. 25). This seal had a diameter

and height such that it would fit in place of a c-seal. The height of this seal at its inner

diameter matched the size of the c-seal gland (0.075 in.), but the raised points were taller

than the gland height such that they were crushed during installation. It was thought that

having pressure on the seal surfaces high enough to yield stainless steel or Inconel would

provide a reliable seal. In fact, one such seal in each of stainless steel and Inconel did

Figure 24: The high-pressure stack shown without and with insulation.

seal at room temperature and during heat-up of the pressure vessel. However, they both

leaked dramatically during cool-down. It is believed that this behavior is due to thermal

expansion. During heat-up, the ring in contact with preheated water is likely warmer than

the surrounding vessel. Hence, thermal expansion keeps a positive load on the sealing

surfaces. During cool-down, the opposite thermal gradient is present, contracting the ring

faster than the vessel. This action unloads the seal and creates a large leak path. Once

these seals leaked, they would not re-seal. Either the ring had yielded to a size smaller

than the gland, or the gland had been enlarged by the warmer ring during heat-up.

The solid ring seals demonstrated that a seal could be made by applying sufficient

stress to mating surfaces. They failed due to a lack of resilience in the assembly.

Although the clamps themselves behave elastically, once the mating surfaces of the two

parts of the vessel are touching, there is nothing except fluid pressure to act on the seal.

A seal placed between the large, formerly-mating surfaces, rather than in the gland, could

remain loaded by the clamps at all times since it would hold the vessel sections apart.

However, such a seal would have to be quite thin to not have an adverse affect on the

operation of the clamps. This is because the clamps are designed to interfere with the

flanges, and too much interference will cause them to not load properly, or will prevent

assembly entirely.

After an attempt with aluminum foil, gasket seals were fabricated from 0.008 in.

fully-annealed copper sheet. One such seal is shown in Fig. 26, ready for installation.

The inner diameter is slightly smaller than the outer diameter of the gland, but is larger

than the combustor and cooler liners so as not to block the fluid flow in the annulus

outside the liner. Three locating tabs that do reach the outer diameter of the liner keep

the gasket centered during assembly. The outer diameter of the gasket is large enough

that the gaskets are strong enough to be easy to handle, but small enough that a pressure

Figure 25: Design of solid cross-section metal seal rings.

exceeding the yield stress of annealed copper is generated by the clamps.. Based on

analysis of the clamped joint and impressions on pressure-sensitive film, a pressure of

about 30,000 psi is applied to the gasket. Removed gaskets show complete impressions

of machining marks from the adjacent Inconel surfaces. This design of copper gasket has

proven successful in sealing the vessel during both heat-up and cool-down.

Fuel Choice

Even with the combustor sealed, a few more questions remained to be answered

before an attempt at flame ignition. Chief among these was whether or not the n-heptane

fuel and water would actually form a supercritical solution while flowing through the

heaters. It was important to verify this mixing since stable, steady operation of the

combustor is predicated on a stable composition and state of the reactant streams. To this

end, the stack BPR was placed near the heaters and a temporary line was run from the

fuel/water exit of the heaters to the BPR, along with a temporary dilution cooling line.

With the heater on, fuel and water flow was turned on at rates similar to what would be

used in initial experiments (20-25% mass fraction fuel, 5-10 kW firing rate). The flow

coming out of the BPR had separate phases of water and n-heptane. There was an

unpleasant, ―cooked‖ smell, indicating that some of the fuel had pyrolyzed or hydrolyzed,

but based on the volume of fuel at the outlet, most of the fuel passed through the system

in its original form. Since the fuel and water flow through tubes with an inner diameter

of 0.120 in., it is likely that there was not a sufficient interface area for the water and n-

heptane to mix thoroughly during the residence time in the heaters.

Due to the importance of fuel/water miscibility and the uncertainty of extent of

mixing, the decision was made to conduct initial experiments with a fuel that is miscible

even at ambient conditions. The fuel system was drained of n-heptane, cleaned, and

refilled with methanol. The change to methanol also required replacement of o-rings in

the solenoid shut-off valve. Additionally, a small diameter nozzle was installed in the

Figure 26: Photograph of copper foil gasket before installation

fuel/water mixing tee so that the fuel would be introduced to the water as a turbulent jet

to promote mixing.

Flame Experiments

With functioning seals and the fuel switched to methanol, the apparatus was reset for

ignition experiments. This experiment took place in December 2010. The windows were

not in place at this stage as a precaution, so indication of reaction was by temperature

data only. Figure 27 shows the methanol and oxygen mass flows and the downstream

temperature trace (measured by ―Combustor TC‖ in Fig. 23) during the first simultaneous

injection of reactants. The temperature rise indicates that exothermic reaction has

occurred. Note that the temperatures shown are subcritical since the measurement

location is about 12 in. downstream of the burner nozzle; significant cooling of the flow

occurred by this point since the vessel walls had not been completely heated to a steady-

state condition. Without optical access, it cannot yet be said that a hydrothermal flame

was produced, but it is sure that fuel and oxygen have reacted spontaneously when

brought in contact in supercritical water.

To be confident that reaction occurred, the effect of another source of the temperature

rise was tested. When reactant injection begins, the total mass flow also increases. Since

the flow is cooling a significant amount before reaching the downstream thermocouple,

Figure 27: Reactant flow rate and downstream temperature histories

during first injections of fuel and oxygen. The temperature rise after

introduction of fuel indicates that exothermic reaction has occurred.

an increased flow speed could also increase the observed temperature. After the reactant

flow was stopped (at time 181 min.), water flow continued until the temperature fell to

the pre-ignition temperature (460-470 K). Next, the water flows through the fuel/water

and oxygen/water lines were increased slightly to simulate the added mass flow from fuel

and oxygen. The effect on the temperature is much smaller than when actually injecting

fuel and oxygen; the small peak centered around 200 mins. is the result. Next, fuel and

oxygen were injected a second time to see if the first observed temperature rise was

repeatable. The temperature trace shows that exothermic reaction occurred again.

The second set of ignition experiments was conducted in January 2011. The reactant

flow and downstream temperature plots are shown in Fig. 27. This time, there is not an

obvious temperature rise accompanying the flow of methanol and oxygen except for the

second and fifth injections, indicating that reaction did not occur the other three times.

The primary parameters not shown in Figs. 27 and 28 are the temperatures of the streams

at the point of injection. These temperatures were lower during the second set of

experiments by approximately 15-20 K. Although this is a small difference compared to

the absolute temperatures, it is a large difference compared to the critical temperature.

The decreased temperatures prevailing during this set of experiments are evidently just

Figure 28: Reactant flow rate and downstream temperature histories

during the second set of flame experiments. The lack of clear

temperature increases as in Fig. 21 may indicate that the combustor was

operating just below the autoignition temperature.

below the autoignition temperature for the particular concentration of fuel that was

injected. Flow rates and stream inlet temperatures were varied throughout the second set

of experiments and only in two cases were the flows heated enough to produce a reaction.

These results indicate that the apparatus can be operated both above and below the

autoignition temperatures for various mixtures. This characteristic will allow the

mapping of autoignition and flame stability limits.

Conclusions

Experiments will transition to utilizing the sapphire windows for optical access once

the combustion process is better understood. The sapphire windows require more

controlled operation of the experiment to ensure their structural integrity, since they are

very susceptible to thermal stresses. Initial experimental results of the supercritical water

combustor demonstrate the feasibility of the use of SCWO for power generation. Further

studies of burner designs and operability limits are necessary for development of the

proposed plant.

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)

This research area involved assessing 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 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 was excluded

from study since the 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.

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 5).

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 6). 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.

Table 6. Solid phase constituents in simulated coal combustion

Constituent mmol/kg (dw) Constituent mmol/kg (dw)

Carbon 66666 Antimony 0.04

Arsenic 0.6 Selenium 0.05

Cadmium 4.7 Zinc 0.7

Chromium 0.8 Calcium 350

Copper 0.4 Magnesium 50

Mercury 0.08 Iron 1000

Nickel 0.5 Titanium 50

Lead 1.5 Potassium 216

Sodium 600

Table 5. Aqueous Constituents in

Simulated Brine

Constituent 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

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. 29). 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. 30).

Thus, anoxic conditions may lower trace element concentrations (Fig. 31) owing to the

sink provided by sulfidic minerals (9).

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

Figure 29: 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.

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 30: 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 31: 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

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. 32). 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. 33). In

particular, arsenic is predicted to adsorb strongly to Fe(III) (hydr)oxides, resulting in a

dramatic drop in its aqueous concentration.

Figure 33: 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

Figure 32: 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.

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