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Citation: Cooper-Baldock, Z.;
Perrelle, T.D.L.; Phelps, C.; Russell,
M.; Ryan, L.; Schofield, J.; Nathan,
G.J.; Jafarian, M. Preliminary
Evaluation of Methods for
Continuous Carbon Removal from a
Molten Catalyst Bubbling Methane
Pyrolysis Reactor. Energies 2024,17,
290. https://doi.org/10.3390/
en17020290
Academic Editor: Dmitri A. Bulushev
Received: 6 December 2023
Revised: 24 December 2023
Accepted: 29 December 2023
Published: 6 January 2024
Copyright: © 2024 by the authors.
Licensee MDPI, Basel, Switzerland.
This article is an open access article
distributed under the terms and
conditions of the Creative Commons
Attribution (CC BY) license (https://
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4.0/).
energies
Article
Preliminary Evaluation of Methods for Continuous Carbon
Removal from a Molten Catalyst Bubbling Methane
Pyrolysis Reactor
Zachary Cooper-Baldock 1,2, * , Thomas De La Perrelle 1,2, Callum Phelps 1,2, Millicent Russell 1,2,
Lachlan Ryan 1,2, Joshua Schofield 1,2, Graham J. Nathan 1,3 and Mehdi Jafarian 1,3 ,*
1Centre for Energy Technology, The University of Adelaide, Adelaide, SA 5005, Australia;
graham.nathan@adelaide.edu.au (G.J.N.)
2School of Mechanical Engineering, The University of Adelaide, Adelaide, SA 5005, Australia
3School of Chemical Engineering and Advanced Materials, The University of Adelaide,
Adelaide, SA 5005, Australia
*Correspondence: zachary.cooperbaldock@flinders.edu.au (Z.C.-B.); mehdi.jafarian@adelaide.edu.au (M.J.)
Abstract:
Methane pyrolysis in molten catalyst bubble (MCB) column reactors is an emerging
technology that enables the simultaneous production of hydrogen and solid carbon, together with
a mechanism for separating the two coproducts. In this process, methane is dispersed as bubbles
into a high temperature molten catalyst bath producing hydrogen and low-density carbon, which
floats to the surface of the bath from providing a means for them to be separated. However, the
removal of carbon particulates from a bubbling column reactor is technically challenging due to
the corrosive nature of the molten catalysts, contamination of the product carbon with the molten
catalysts, high temperatures and lack of understanding of the technology options. Four potential
concepts for the removal of carbon particulate from a methane pyrolysis molten metal bubble column
reactor are presented, based on the pneumatic removal of the particles or their overflow from the
reactor. The concepts are evaluated using a cold prototype reactor model. To simulate the operation of
a high-temperature reactor at low temperatures, the dominant dimensionless numbers are identified
and matched between a reference high-temperature reactor and the developed cold prototype using
water, air and hollow glass microsphere particles as the representatives of the molten catalyst, gaseous
phases and solid carbon particulates, respectively. The concepts are tested in the cold prototype. High
rates of particle removal are achieved, but with different tradeoffs. The applicability of each method
together with their advantages and disadvantages are discussed.
Keywords: hydrogen; methane pyrolysis; CO2capture; cold prototype; bubble column
1. Introduction
Hydrogen combustion produces no carbon emissions. Hence, if it can be produced
either from renewable resources or from fossil fuels with CO
2
capture, it can be a carbon
neutral fuel. There are primarily four methods used to produce hydrogen at industrial
scales, including methane reforming (SMR), coal gasification (CG), oil reforming and water
electrolysis [
1
,
2
]. The dominant SMR process produces H
2
at the cost of ~1.5 USD/kgH
2
.
However, it also produces 9–11 of CO
2
per kgH
2
[
3
,
4
]. Integration of carbon capture
to SMR for 90% CO
2
capture also adds some 60% to the costs [
4
]. Nevertheless, it is
still attracting significant investment in countries like the United Kingdom because it is
judged as the technology with the greatest potential to deliver largescale supplies in the
short term [
5
]. The coal gasification route is even more carbon intensive, producing some
18 (kgCO2/kgH2) [6]
. On the other hand, water splitting via electrolysis with renewable
electricity, while enabling net-zero hydrogen production, is even more expensive than SMR
with CCS, despite falling costs [
3
]. Hence, there is a need to invest in the development
Energies 2024,17, 290. https://doi.org/10.3390/en17020290 https://www.mdpi.com/journal/energies
Energies 2024,17, 290 2 of 23
of new technologies with strong potential to yield complete or near-complete mitigation
of CO
2
in the production of low-cost hydrogen. One of the emerging technologies under
development, with strong potential to lower the cost of hydrogen production, is methane
pyrolysis [5], which cracks the methane to produce solid carbon and hydrogen:
CH4(g)→2H2(g)+C(s)(1)
The enthalpy of this reaction, at 75 kJ/molH
2
, is only ~46% of the net enthalpy of SMR
and some 25% that of water electrolysis. Whilst these are different production mechanisms,
the lower specific energy requirement per mole of hydrogen produced implies potential for
lower production costs. In addition, this reaction produces solid carbon as a by-product,
which not only avoids the direct production of CO
2
and hence the cost of sequestration
but it also offers the potential to coproduce valuable carbon coproducts, such as carbon
black, carbon fibers, carbon nanomaterials and cement additives. Even if markets for these
coproducts become saturated, the cost of sequestration of pure solid carbon would be
far less than that of gaseous CO
2
. While the current market is too small to adsorb the
amount of hydrogen anticipated to be needed in the new economy, if it is to be produced
through methane pyrolysis, much larger markets are potentially possible with significant
development, such as the partial replacement of steel and cement with carbon-based
products [
7
,
8
]. Moreover, the full lifecycle emissions of this process depend on the source of
the energy needed to drive the endothermic pyrolysis reactions, the source of the methane,
and on the life of the carbon product, it has potential to be net-zero if the pyrolysis is driven
with renewable energy, or even negative with the use of biomethane instead of natural
gas [
7
]. However, various technical barriers remain to be overcome before this technology
can be widely implemented commercially.
The pyrolysis of methane begins to occur at ~400
◦
C but does not reach completion
without reaching temperatures of more than 1000
◦
C. Moreover, it is a kinetic control reac-
tion. For this reason, there has been significant investment in the development of catalysts
to increase reaction rates. Solid, porous catalysts have been proposed and investigated
vastly [
9
–
11
]. However, they are quickly coked by the carbon byproduct (Equation (1))
and exhibit reduced efficiency. To address this challenge, molten catalysts, e.g., molten
metals [
11
], molten alloys [
10
] and molten salts [
10
] under a gas bubbling flow regime
within a bubble column reactor, have been proposed.
In a methane pyrolysis molten catalyst bubbling (MCB) reactor, methane is dispersed
from submerged nozzles into a molten bath, where it dissociates mostly to H
2
and carbon
(Equation (1)) as the bubbles rise through the bath. The buoyancy force induced by the
substantial difference between the density of the carbon and that of the molten bath floats
the carbon product to the surface of the molten bath, from where it can be removed [
1
].
The methane pyrolysis MCB reactors appearing in the literature operate predominantly
under a semi-batch configuration with a noncontinuous removal of the carbon product.
Nonetheless, continuous operation would be desired in the case of an industrial process to
minimize the capital and operational costs. However, despite the need for the development
of an efficient method for the removal of carbon particulate from a molten metal bubble
column reactor, this has remained a largely uninvestigated aspect of the molten metal
methane pyrolysis process [12].
Recently, von Wald et al. [
13
] proposed a method for carbon removal from a molten
metal bubble column methane pyrolysis reactor that involves entraining the carbon into the
outgoing gas and flowing it through a cyclone to remove the carbon particulate. A math-
ematical model has also been developed based on the expected volumetric gas flow rate
leaving the reactor. However, the concept has not yet been demonstrated. Additionally, a
particle diameter of 0.1 mm and a completely nonporous carbon structure was assumed [
13
].
These could be different from the solid particles generated in a methane pyrolysis reaction,
which has been found to have a range of sizes [
12
] and carbon structures [
13
–
15
]. More
recently, Kudinov et al. [
14
] proposed a conceptual design of a desooting device that uses
Energies 2024,17, 290 3 of 23
a floating sensor to control the catalyst level within the reactor and removes the carbon
particulate using a vacuum device.
The use of mechanical methods such as skimming, similar to that found in the fuming
industry [
14
], or entraining the carbon in the hydrogen gas stream have also been pro-
posed [
11
,
16
]. However, mechanical separation of carbon from the top of the molten bath is
technically challenging due to the corrosive nature of liquid metal, which can exacerbate
the mechanical wear of the moveable metallic components in contact with the molten bath.
Several conceptual designs can also be found in patents [
16
], which are typically based
on physical overflow of the carbon product or entraining carbon particles into a gas flow.
Nevertheless, these reports are mostly conceptual and lack either a systematic theoretical
or experimental evaluation, so that further systematic evaluation is needed both to better
identify their operational challenges and optimize them.
Four potential methods are identified here that appear to have the potential to be
applied to the removal of solid carbon particulates from a methane pyrolysis molten catalyst
reactor. These are pneumatic conveyance, venturi vacuums, cyclone flow and an overflow
weir. Pneumatic conveyance, both for fluids and particulate-laden environments, has been
adopted across a large range of industries and applications, e.g., burner feeding [17], feed
lance injection [
18
] and particulate and granule mixing [
19
]. Venturi vacuums and injectors
have also been widely used in many applications, in which a motive fluid is used to draw
another fluid into a different geometry or segment of pipe [
20
]. In recent years, research has
also been conducted into venturi scrubbers and particle-based venturi technologies [
21
].
These studies have reported considerable success in removing particles with a range of
diameters [
21
,
22
]. The use of cyclonic flow for particle separation and mixing is also
widely applied across a range of particle sizes [
23
], which is a specific application of cyclone
separators. However, to the best of the author’s knowledge, no specific reports are available
of the performance of the use of pneumatic conveyance, venturi vacuums and cyclone
flow for the removal of carbon particles under conditions of relevance to a molten catalyst
methane pyrolysis bubble column reactor. Moreover, while the overflow concept has been
proposed [
16
], it has not been thoroughly investigated yet. Hence, the principle objective
of the present investigation is to explore the potential benefits and limitations of these
methods for the removal of carbon particulates from a molten catalyst methane pyrolysis
bubble column reactor.
Experimental evaluation of high-temperature molten catalyst methane pyrolysis re-
actors is both challenging and costly. This is mainly because of the risks associated with
the handling of high-temperature molten catalysts in the presence of combustible gases
(methane and hydrogen). For this reason, a cold model prototype was developed and used
to simulate the behavior of the carbon particles within a high-temperature MCB reactor,
at room temperature and with water, using conditions of similarity. The use of similarity
and dimensionless analysis is well established in the literature to undertake experimental
investigations using alternative fluids with similar dimensionless properties but under
ambient conditions and at much lower costs [
24
]. However, to the best of the author’s
knowledge, no previous cold model of a methane pyrolysis bubble column reactor has
been reported. It should be noted that whilst it is not possible to match all dimensionless
parameters of a molten metal methane pyrolysis bubble column reactor, given that they
scale in different ways, it is possible to match the dominant ones. Therefore, the second
objective of this work is to develop a cold model of the methane pyrolysis MCB reactor and
use it to assess the viability of the abovementioned four identified methods for the removal
of the solid carbon particles.
Contribution
This investigation intends to make two contributions to the current state of the art
with respect to methane pyrolysis. The first contribution is the development of a cold
prototype reactor that simulates the physical interactions and phenomenon of hot molten
metal bubble column reactors, but at ambient temperatures and using readily available
Energies 2024,17, 290 4 of 23
materials. The process to develop this is elucidated below, as well as the physical effect
similarity matching process. The successful development of a physically similar cold
prototype reactor lowers the technology development cost [
24
], while also increasing the
development speed by avoiding the need to address the safety and approvals processes of
operating at high temperatures [9–11] and high pressures [10].
The second contribution is the identification, development and subsequent testing
of mechanisms or methods for carbon removal. This work aims to adapt technologies
from other industries, namely, an overflow weir [
16
], pneumatically assisted overflow
weir [17–19], venturi
ejector [
21
] and cyclone extractor [
23
], for use in methane pyrolysis
reactors. The challenge of carbon removal has limited previous processes to batch opera-
tion [
13
]; thus, the development of a suitable technology for its continuous removal opens
the door to continuous hydrogen production [14].
2. Potential Separation Methods
The four identified methods for the removal of carbon particulate from the molten
catalyst methane pyrolysis reactor are described in Figure 1. Further explanation on each
concept is provided below.
Energies 2024, 17, 290 4 of 23
Contribution
This investigation intends to make two contributions to the current state of the art with
respect to methane pyrolysis. The first contribution is the development of a cold prototype
reactor that simulates the physical interactions and phenomenon of hot molten metal bubble
column reactors, but at ambient temperatures and using readily available materials. The
process to develop this is elucidated below, as well as the physical effect similarity matching
process. The successful development of a physically similar cold prototype reactor lowers
the technology development cost [24], while also increasing the development speed by
avoiding the need to address the safety and approvals processes of operating at high tem-
peratures [9–11] and high pressures [10].
The second contribution is the identification, development and subsequent testing of
mechanisms or methods for carbon removal. This work aims to adapt technologies from
other industries, namely, an overflow weir [16], pneumatically assisted overflow weir [17–
19], venturi ejector [21] and cyclone extractor [23], for use in methane pyrolysis reactors.
The challenge of carbon removal has limited previous processes to batch operation [13];
thus, the development of a suitable technology for its continuous removal opens the door
to continuous hydrogen production [14].
2. Potential Separation Methods
The four identified methods for the removal of carbon particulate from the molten cat-
alyst methane pyrolysis reactor are described in Figure 1. Further explanation on each con-
cept is provided below.
Figure 1. Schematic representation of the four concepts proposed for the removal of carbon partic-
ulates from the molten catalyst methane pyrolysis reactor, namely, (a) overflow weir, (b) pneumatic-
assisted overflow weir, (c) venturi ejector and (d) cyclone extractor.
2.1. Overflow Weir
In this concept, the layer of carbon particles that accumulate on top of the molten bath
overflow from a weir into an adjacent collector, as shown schematically in Figure 1a. This
concept is attractive due to its mechanical simplicity. For example, the height of the weir can
Figure 1. Schematic representation of the four concepts proposed for the removal of carbon particu-
lates from the molten catalyst methane pyrolysis reactor, namely, (
a
) overflow weir, (
b
) pneumatic-
assisted overflow weir, (c) venturi ejector and (d) cyclone extractor.
2.1. Overflow Weir
In this concept, the layer of carbon particles that accumulate on top of the molten bath
overflow from a weir into an adjacent collector, as shown schematically in Figure 1a. This
concept is attractive due to its mechanical simplicity. For example, the height of the weir
can be adjusted to minimize the overflow of the molten catalyst bath. Nevertheless, it may
provide limited control over the particles’ residence time within the reactor headspace.
2.2. Pneumatic-Assisted Overflow Weir
The pneumatic-assisted overflow weir concept utilizes a high-velocity pneumatic gas
jet parallel to the surface of the molten catalyst bath to generate pneumatic conveying of
Energies 2024,17, 290 5 of 23
the floated particles away from the liquid surface, as shown schematically in Figure 1b. It
should be noted that the conveyed particles may require additional separation later using
a filtration system, such as a typical cyclone separator, in the case of molten metal and
particulate being extracted together. A heat exchanger would also be needed to recover the
enthalpy of the conveying gas, depending on the process configuration. That is because
passing an inert gas through the reactor could lead to increasing its temperature, which, in
return, would affect the thermal load of the reactor. Nevertheless, detailed assessment of
these aspects is beyond the scope of the present investigation.
2.3. Cyclone Extractor
The cyclone extractor, as shown in Figure 1c, employs a swirling flow that is exerted
using gas jet nozzles, above the fluid surface of the reactor column, to generate a high-
velocity flow field, transporting the particles from the reactor head space once they become
entrained in the jet flow. This concept requires the same heat exchanger proposed for the
overflow weir, as the injected gas temperature may increase, affecting thermal load.
2.4. Venturi Ejector
The venturi ejector system employs the venturi effect with a high-velocity venturi jet
to produce a low static pressure region adjacent the reactor headspace, which is used to
suck particles from the reactor into the venturi connecting arm, through the suction port,
before they are exhausted from the reactor column. This concept is schematically shown in
Figure 1d.
3. Methodology
This section details the development of a reference high-temperature reactor, cold pro-
totype and dimensionless matching to ensure the concept’s physical effects are applicable
to the hot MCB methane pyrolysis reactor designs. Subsequently, the experimental arrange-
ment, operating conditions and specific separation devices are detailed and discussed.
3.1. Reference High-Temperature Reactor, Cold Prototype and Dimensionless Matching
To test the effectiveness of the proposed removal concepts, a test reactor was required
to be independently developed. This development consisted of several key steps. The
first was the synthesis of the available literature pertaining to methane pyrolysis in MCB
reactors, with a focus on technologies that could be adapted to pilot- or large scale [
24
].
This was used to develop a theoretical hot reference model reactor. The details of the hot
reference reactor are provided in Table 1. After the hot model was determined, this could
then be used to generate a cold prototype reactor with similar physical characteristics,
at ambient temperatures. The shared characteristics between the hot reference and cold
prototype are of great importance, as they allow for dimensionless matching between the
proposed concepts to ensure the physical phenomenon will be seen in a hot reactor.
To develop an accurate hot reference reactor, the composition, geometry and operating
conditions must be known from the existing literature. These values are presented in
Table 1. The model hot reactor has a 26–73% nickel–bismuth ratio [
9
] and a density of
9500 kg/m
3
[
11
,
25
,
26
]. In addition to the composition of the reactor, its physical geometry
must be theorized. It was seen that a diameter of 1.2 m, height of 1.0 m and operating
environment of 15 bar at 1000
◦
C could be developed if available reactors were scaled
in size [
11
]. The modelled hydrogen, methane and carbon production rates were set to
0.01022 kg/(s.m
3
), 0.08140 kg/(s.m
3
) and 0.06092 kg/(s.m
3
), respectively, to correlate to this
scaling [
26
,
28
]. For the produced carbon byproduct, the modelled density was 1800 kg/m
3
with a diameter of 56 microns, based on the existing literature [
25
]. For the hydrogen
produced, it was also assumed to have the properties as detailed in Table 1[25].
Energies 2024,17, 290 6 of 23
Table 1.
Key dimensions and operating conditions of the developed reference methane pyrolysis
reactor, together with the physical properties of the molten bath, gaseous phases and carbon particle.
Liquid Metal Catalyst Properties (Bubbling Media)
Parameter Value Reference
Composition (CL) 26–73% nickel–bismuth [11,25,26]
Density ( ρL)9500 kg/m3[11,25,26]
Reactor geometry and operating conditions
Diameter (DR) 1.2 m [11]
Height (hR) 1.0 m [11]
Pressure (PR) 15 bar [11]
Temperature (TR) 1000 ◦C [11]
Conversion efficiency ( ηR)90% [11]
Hydrogen production rate (rH)0.01022 kg/(s.m3)[27]
Methane input rate (rM)0.08140 kg/(s.m3)[27]
Carbon production rate (rC)0.06092 kg/(s.m3)[27]
Carbon properties
Density (ρC)1800 kg/m3[27]
Diameter ( DC)56 micron [27]
Hydrogen properties
Pressure (PH)15 bar [25]
Density (ρH)0.4669 kg/m3[25]
Viscosity (µH)16.9 µPa/s [25]
After this analysis had been conducted and the reference hot model developed, di-
mensionless matching could be undertaken. A dimensionless match is required between
the conceptual removal methods, cold test reactor and the hot reference model. To di-
mensionlessly match the removal methods, their dominating physical relationships were
determined. These are presented in Table 2, with further explanation provided in Table 3.
For the pneumatically assisted overflow weir, these are the particle Stokes and Froude
number and the Reynolds number. This is the same for the cyclone extractor. The venturi
ejector needs to have a matched particle Stokes number.
Table 2. Identified dominating dimensionless group for each removal method.
Removal Method Identified Dimensionless Groups
Pneumatic assisted overflow weir Particle Stokes number (Stk)
Particle Froude number (Frp)
Reynolds number (Re)
Cyclone extractor Particle Stokes number (Stk)
Particle Froude number (Frp)
Gas Froude number (Frg)
Reynolds number (Re)
Venturi ejector Particle Stokes number (Stk)
Reynolds number (Re)
Energies 2024,17, 290 7 of 23
Table 3.
Summary of dimensionless similarity parameters used to match key phenomena in the cold
prototype experiment to the developed reference methane pyrolysis reactor.
Dimensionless Similarity Parameters
Variable Formula Physical Relevance Interaction Represented
Particle Froude number Frp=vp
((1−s)gD)1
2
Particle inertia to gravitational force
Influence of gravity on particles
Particle Stokes number Stk =ρpd2
pvg
18µD
Characteristic particle response
time to flow time
Particle response to large scale flow
Gas Froude number Frg=vg
((1−s)gD)1
2
Flow inertia to gravitational force Buoyancy effects
Reynolds number Re =ρgvgD
µ
Ratio of inertial forces to
viscous forces
Flow speed and
configuration effects
Density ratio (particle to
liquid bath) s=ρp
ρlParticle–liquid buoyance ratio Particle–liquid buoyancy effects
To match the dimensionless groups, as detailed in Table 2, their relationships to the
physical properties, environmental conditions and operating parameters must also be
known. These relationships are provided in Table 3, which details their specific interac-
tions, physical relevance and associated variables. The particle Froude number
Frp
is
calculated based on the mean particle velocity (
vp
), the ratio of particle to fluid densities (s),
acceleration due to gravity (g) and the diameter of the reactor column (D). For the particle
Stokes number
(Stk)
the particle density (
ρp
, particle diameter (
dp
, mean velocity of the
gas (
vg
), gas dynamic viscosity (
µ
) and diameter of the reactor column (D) were used. The
gas Froude number is determined in a similar manner to the particle Froude number, but
using the velocity of the gas
vg
). The Reynolds number is given by the gas density
ρg
,
the gas velocity
vg
, reactor diameter (D) and the dynamic viscosity of the gas
(µ)
. The
density ratio (s) is given by the density of the particles
ρp
divided by that of the liquid
phase (ρl).
Once the relationships of each concept and the physical interactions were understood,
a physically similar cold prototype reactor could be developed. This reactor was scaled
proportionally to ensure that it provided a dimensionless match to the hot reference reactor
for each concept. To model the molten metal phase of the hot reactor, water was selected,
while air was selected for the gas phase. These two fluids were matched to the hot reference
reactor using the similarity parameters in Table 3. The carbon particulate product was
replaced with 7019-grade Q-Cell hollow glass microspheres. The microspheres and water
were selected, as they have similar dimensionless density ratios and diameters to those
of the modelled liquid metal and carbon phases. The selected microspheres have an
effective density of 0.19 kg/m
3
. Consequently, they float on the water surface within
the column headspace, mimicking the behavior of carbon particulate in MCB methane
pyrolysis reactors. Moreover, the particles’ density-to-water density ratio is similar to that
of the molten catalyst to the carbon product Table 1.
In a high-temperature methane pyrolysis bubble column reactor, methane dissociates
to H
2
and carbon as the bubbles rise through the molten bath column. To better simulate this
phenomenon, the 7019-grade Q-Cell hollow microspheres were introduced at the bottom
of the bubble column. In so doing, the particles were entrained into the inlet airstream
prior to being injected into the bottom of the column. The particles had a mean diameter
of 80
µ
m with a maximum size of 150
µ
m. In addition, the diameter, height, pressure and
temperature were scaled to ensure a similar physical domain to the hot prototype. This is
detailed in Table 4. The subscript used to denote the phase of the hot reference liquid is L,
C for carbon and H for hydrogen. Then, to ensure the selected parameters and subsequent
separation concepts are applicable to the hot reference model, their operating conditions
are scaled to match Table 1’s hot model, in accordance with the relationships described in
Energies 2024,17, 290 8 of 23
Table 3. These experimental conditions and scaled up hot model conditions are provided
in Table 5.
Table 4.
Key dimensions and operating conditions of the developed cold prototype methane pyrol-
ysis reactor, together with the physical properties of the pseudo molten bath, gaseous phases and
carbon particle.
Bubbling Media
Parameter Value Reference
Composition (CL) 100% water n/a—selected
Density ( ρL)1000 kg/m3n/a—selected
Reactor geometry and operating conditions
Diameter (DR) 0.09 m n/a—selected
Height (hR) 1.00 m n/a—selected
Pressure (PR) 1 bar [10,11,25–27]
Temperature (TR) 25 ◦C [10,11,25–27]
Modelled conversion efficiency ( ηR)90% [10,11,25–27]
Air bubble flow rate (modelled hydrogen) (rH)0.01836 kg/(s.m3)[25]
Microsphere flow rate (modelled carbon particle
buildup) (rH)0.06092 kg/(s.m3)[25]
Glass microsphere properties
Density ( ρC)190 kg/m3[25]
Diameter (DC) 80 micron [25]
Air properties
Pressure (PH)1 bar [25]
Table 5.
Experimental removal methods for geometric and flow configuration data in addition to
assumed geometric constraints for the developed hot model, to be used for dimensionless matching,
where additional hot model geometric data are required.
Removal Method Parameter Cold Prototype Condition Theoretical Reference Condition
Pneumatic-assisted
overflow weir
Slit height 0.003 m 0.008 m
Gas velocity (at slit) 21.9 m/s 17.8 m/s
Cyclone extractor Cyclone nozzle diameter 0.05 m 1.10 m
Gas velocity (nozzle) 27.7 m/s 51.2 m/s
Venturi ejector
Volumetric flow rate 308 SLPM 1000 SLPM
Gas viscosity 1.81 Pa.s 16.9 Pa.s
Throat diameter 0.0053 m 0.005 m
Figure 2provides a visual depiction of how close the dimensionless match is between
the cold prototype (blue) and the hot reference model (yellow). For the venturi, the particle
Stokes number is higher in the hot model than the cold prototype. For the cyclone extractor,
the Froude number provides good similarity, but this is not true of the particle Stokes
number, which is lower. For the pneumatically assisted overflow weir, the Stokes, Reynolds
and gas Froude number provide good similarity. This is not true of the particle Froude
number for the weir however, which is approximately half. The density ratio, between the
cold prototype and hot model, is an exact match, which should negate any differences in
density effects.
Energies 2024,17, 290 9 of 23
Energies 2024, 17, 290 9 of 23
Stokes number is higher in the hot model than the cold prototype. For the cyclone extractor,
the Froude number provides good similarity, but this is not true of the particle Stokes num-
ber, which is lower. For the pneumatically assisted overflow weir, the Stokes, Reynolds and
gas Froude number provide good similarity. This is not true of the particle Froude number
for the weir however, which is approximately half. The density ratio, between the cold pro-
totype and hot model, is an exact match, which should negate any differences in density
effects.
Figure 2. Comparison of dimensionless numbers for each of the removal mechanisms for the devel-
oped hot model reactor (denoted in orange) and the cold prototype reactor (denoted in blue).
3.2. Experimental Arrangement and Operating Conditions
Figure 3 shows the experimental arrangement of the cold prototype test reactor. To the
right of Figure 3, air enters the system, passing a flow valve, before entering a pressurized
vessel. In the pressurized vessel, the air mixes with the hollow microsphere particles. This
mixture simulates the hydrogen and carbon formation of the hot reference reactor, as out-
lined in Table 4. Prior to entering the column, this mixture passes a laser particle sensor. This
sensor is used to check that the air and particle mixture enters the reactor consistently during
testing. After passing this sensor, the mixture enters the base of the reactor column through
a nozzle, before bubbling to the top of the column, where it is separated from the fluid in-
terface by the removal concept.
Figure 3 also details the separation concepts connection. In the top right of Figure 3,
pressurized air is passed through a rotameter and control value before entering the separa-
tion concept. This allows for the airflow rate, into the concepts, to be controlled for testing
purposes. Connected to the separation concept is the removal pathway. This pathway,
shown to the left of Figure 3, is used to allow for the separated particles to exit the system,
where they are filtered, collected and weighed.
Figure 2.
Comparison of dimensionless numbers for each of the removal mechanisms for the
developed hot model reactor (denoted in orange) and the cold prototype reactor (denoted in blue).
3.2. Experimental Arrangement and Operating Conditions
Figure 3shows the experimental arrangement of the cold prototype test reactor. To the
right of Figure 3, air enters the system, passing a flow valve, before entering a pressurized
vessel. In the pressurized vessel, the air mixes with the hollow microsphere particles.
This mixture simulates the hydrogen and carbon formation of the hot reference reactor, as
outlined in Table 4. Prior to entering the column, this mixture passes a laser particle sensor.
This sensor is used to check that the air and particle mixture enters the reactor consistently
during testing. After passing this sensor, the mixture enters the base of the reactor column
through a nozzle, before bubbling to the top of the column, where it is separated from the
fluid interface by the removal concept.
Figure 3also details the separation concepts connection. In the top right of Figure 3,
pressurized air is passed through a rotameter and control value before entering the separa-
tion concept. This allows for the airflow rate, into the concepts, to be controlled for testing
purposes. Connected to the separation concept is the removal pathway. This pathway,
shown to the left of Figure 3, is used to allow for the separated particles to exit the system,
where they are filtered, collected and weighed.
Figure 4shows the experimental apparatus of Figure 3, as used for testing. The red
square (Figure 3) indicates where the separation technology is attached to the column. The
position of the flow valves, pressurized air, particle sensors and filtration and collection
system is also indicated.
Schematic representation of the developed cold prototypes of the overflow weir, pneumatic-
assisted overflow weir, cyclone extractor and venturi ejector concepts used in this investigation
are shown in Figures 5A, 6A, 7A and 8A, respectively. The cross-sections of these systems
together with their main dimensions are also shown in Figures 5B, 6B, 7B and 8B.
Energies 2024,17, 290 10 of 23
Energies 2024, 17, 290 10 of 23
Figure 3. Process flow diagram (PFD) of the cold prototype reactor configuration. Pressurized air is
passed through a flow meter and used to convey the hollow microsphere particles into the nozzle
at the base of the test column, where they are bubbled by the air vertically, through the fluid column,
towards the separation technology test area. Pressurized air is fed through a rotameter into the sep-
aration technology to remove the microspheres from the separation zone and convey them into the
particle collection area, which is placed on a set of scales. Laser particle sensors are employed on the
inlet and outlet to ensure particles are entering and exiting the system.
Figure 4 shows the experimental apparatus of Figure 3, as used for testing. The red
square (Figure 3) indicates where the separation technology is attached to the column. The
position of the flow valves, pressurized air, particle sensors and filtration and collection sys-
tem is also indicated.
Figure 4. Experimental configuration of the cold prototype reactor configuration (as detailed in Figure
3). Rotameter not shown, as this was located adjacent to the test rig, attached to the test facility’s air
lines. The rotameter was used for variable airflow control (Figure 3) and controlled the flow rate of air
Figure 3. Process flow diagram (PFD) of the cold prototype reactor configuration. Pressurized air is
passed through a flow meter and used to convey the hollow microsphere particles into the nozzle at
the base of the test column, where they are bubbled by the air vertically, through the fluid column,
towards the separation technology test area. Pressurized air is fed through a rotameter into the
separation technology to remove the microspheres from the separation zone and convey them into
the particle collection area, which is placed on a set of scales. Laser particle sensors are employed on
the inlet and outlet to ensure particles are entering and exiting the system.
Energies 2024, 17, 290 10 of 23
Figure 3. Process flow diagram (PFD) of the cold prototype reactor configuration. Pressurized air is
passed through a flow meter and used to convey the hollow microsphere particles into the nozzle
at the base of the test column, where they are bubbled by the air vertically, through the fluid column,
towards the separation technology test area. Pressurized air is fed through a rotameter into the sep-
aration technology to remove the microspheres from the separation zone and convey them into the
particle collection area, which is placed on a set of scales. Laser particle sensors are employed on the
inlet and outlet to ensure particles are entering and exiting the system.
Figure 4 shows the experimental apparatus of Figure 3, as used for testing. The red
square (Figure 3) indicates where the separation technology is attached to the column. The
position of the flow valves, pressurized air, particle sensors and filtration and collection sys-
tem is also indicated.
Figure 4. Experimental configuration of the cold prototype reactor configuration (as detailed in Figure
3). Rotameter not shown, as this was located adjacent to the test rig, attached to the test facility’s air
lines. The rotameter was used for variable airflow control (Figure 3) and controlled the flow rate of air
Figure 4.
Experimental configuration of the cold prototype reactor configuration (as detailed in
Figure 3). Rotameter not shown, as this was located adjacent to the test rig, attached to the test
facility’s air lines. The rotameter was used for variable airflow control (Figure 3) and controlled the
flow rate of air into the separation technology. Red box is shown to indicate where the separation
technology (Figure 3) is placed during testing.
Energies 2024,17, 290 11 of 23
Energies 2024, 17, 290 11 of 23
into the separation technology. Red box is shown to indicate where the separation technology (Figure
3) is placed during testing.
Schematic representation of the developed cold prototypes of the overflow weir, pneu-
matic-assisted overflow weir, cyclone extractor and venturi ejector concepts used in this in-
vestigation are shown in Figures 5A, 6A, 7A and 8A, respectively. The cross-sections of these
systems together with their main dimensions are also shown in Figures 5B, 6B, 7B and 8B.
Figure 5. Schematic diagram of (A) three-dimensional and (B) cross-section of the overflow weir con-
figuration. The key dimensions of the system are also shown.
Figure 6. Schematic diagram of (A) three-dimensional and (B) cross-section of the pneumatic-assisted
overflow weir configuration. The key dimensions of the system are also shown.
Figure 7. Schematic diagram of (A) three-dimensional and (B) cross-section of the cyclone extractor
configuration. The key dimensions of the system are also shown.
Figure 5.
Schematic diagram of (
A
) three-dimensional and (
B
) cross-section of the overflow weir
configuration. The key dimensions of the system are also shown.
Energies 2024, 17, 290 11 of 23
into the separation technology. Red box is shown to indicate where the separation technology (Figure
3) is placed during testing.
Schematic representation of the developed cold prototypes of the overflow weir, pneu-
matic-assisted overflow weir, cyclone extractor and venturi ejector concepts used in this in-
vestigation are shown in Figures 5A, 6A, 7A and 8A, respectively. The cross-sections of these
systems together with their main dimensions are also shown in Figures 5B, 6B, 7B and 8B.
Figure 5. Schematic diagram of (A) three-dimensional and (B) cross-section of the overflow weir con-
figuration. The key dimensions of the system are also shown.
Figure 6. Schematic diagram of (A) three-dimensional and (B) cross-section of the pneumatic-assisted
overflow weir configuration. The key dimensions of the system are also shown.
Figure 7. Schematic diagram of (A) three-dimensional and (B) cross-section of the cyclone extractor
configuration. The key dimensions of the system are also shown.
Figure 6.
Schematic diagram of (
A
) three-dimensional and (
B
) cross-section of the pneumatic-assisted
overflow weir configuration. The key dimensions of the system are also shown.
Energies 2024, 17, 290 11 of 23
into the separation technology. Red box is shown to indicate where the separation technology (Figure
3) is placed during testing.
Schematic representation of the developed cold prototypes of the overflow weir, pneu-
matic-assisted overflow weir, cyclone extractor and venturi ejector concepts used in this in-
vestigation are shown in Figures 5A, 6A, 7A and 8A, respectively. The cross-sections of these
systems together with their main dimensions are also shown in Figures 5B, 6B, 7B and 8B.
Figure 5. Schematic diagram of (A) three-dimensional and (B) cross-section of the overflow weir con-
figuration. The key dimensions of the system are also shown.
Figure 6. Schematic diagram of (A) three-dimensional and (B) cross-section of the pneumatic-assisted
overflow weir configuration. The key dimensions of the system are also shown.
Figure 7. Schematic diagram of (A) three-dimensional and (B) cross-section of the cyclone extractor
configuration. The key dimensions of the system are also shown.
Figure 7.
Schematic diagram of (
A
) three-dimensional and (
B
) cross-section of the cyclone extractor
configuration. The key dimensions of the system are also shown.
Energies 2024,17, 290 12 of 23
Energies 2024, 17, 290 12 of 23
Figure 8. Schematic diagram of (A) three-dimensional and (B) cross-section of the venturi ejector con-
figuration. The key dimensions of the system are also shown.
4. Experimental Procedure
Each of the particulate removal devices (shown in Figures 5–8) were installed on the
top of the cold prototype bubble column reactor (indicated in red in Figure 3) for testing
purposes. During the experiments, the volumetric flow rates of the inlet air to the column
were monitored and each removal mechanism was tested across a range of inlet gas flow
rates (𝑄) to determine its effectiveness in removal of the particles from the reactor. The
flow rate of the outlet air (𝑄) is equal to the sum of the flow rates of two airstreams
(𝑄 and 𝑄). The first airstream is injected into the column to carry the particles
into the system (𝑄) and simulate the bubbling regime of the reactor, while the second
the airstream (𝑄) is injected to pneumatically remove the particles. The flow rate of
the air injected into the reactor column, 𝑄, was fixed at 14.5 standard liters per minute
(SLPM) to ensure the matching of the identified dimensionless groups between the hot and
cold prototype reactors (listed in Tables 3 and 4). However, the flow rate of the injected gas
employed to remove the particles pneumatically was changed to assess the performance of
each concept. In each experiment, a fixed airflow rate was set for a period of 10 min, while
the mass of particles removed from the system was continuously measured and recorded.
This was then used to generate the mass removal rate in mass per time, e.g., g/min. The
pneumatic airflow speeds into the removal concepts were also area averaged to yield super-
ficial gas velocity. On this basis, the gas fraction (Equation (2)) shows the pneumatic airflow
rate through the removal mechanism against that of the bubble column flow rate.
𝑋=𝑅𝑒𝑚𝑜𝑣𝑎𝑙 𝑑𝑒𝑣𝑖𝑐𝑒
𝑓
𝑙𝑜𝑤 𝑟𝑎𝑡𝑒
𝐵𝑢𝑏𝑏𝑙𝑒 𝑐𝑜𝑙𝑢𝑚𝑛
𝑓
𝑙𝑜𝑤 𝑟𝑎𝑡𝑒 =𝑄
𝑄 (2)
The particle fraction (Equation (3)) was defined as the mass of particles removed from
the system (𝑚,) divided by the mass of particles entering (𝑚,) the cold prototype sys-
tem.
𝑋=𝑃𝑎𝑟𝑡𝑖𝑐𝑙𝑒𝑠 𝑟𝑒𝑚𝑜𝑣𝑒𝑑 (𝑚𝑎𝑠𝑠)
𝑃𝑎𝑟𝑡𝑖𝑐𝑙𝑒𝑠 𝑎𝑑𝑑𝑒𝑑 (𝑚𝑎𝑠𝑠) =𝑚,
𝑚, (3)
The saltation velocity (Equation (4)) is the velocity that marks the transition point be-
tween settling or particles being lifted and conveyed within a system. Hence, it was also
used to compare the removal mechanisms against one another. This is of relevance, as many
of the concepts are intrinsically reliant upon pneumatic conveyance. The saltation velocity
(𝑣) was estimated (Equation (4)) using the Rizk correlation [23], as follows:
𝑀
𝜌∙𝑈 ∙
𝐴
= 1
10(
.)𝑣
𝑔∙𝐷(.) (4)
Figure 8.
Schematic diagram of (
A
) three-dimensional and (
B
) cross-section of the venturi ejector
configuration. The key dimensions of the system are also shown.
4. Experimental Procedure
Each of the particulate removal devices (shown in Figures 5–8) were installed on the
top of the cold prototype bubble column reactor (indicated in red in Figure 3) for testing
purposes. During the experiments, the volumetric flow rates of the inlet air to the column
were monitored and each removal mechanism was tested across a range of inlet gas flow
rates (
Qconcept
) to determine its effectiveness in removal of the particles from the reactor.
The flow rate of the outlet air (
Qtotal )
is equal to the sum of the flow rates of two airstreams
(
Qreactor
and
Qconcept
). The first airstream is injected into the column to carry the particles
into the system (
Qreactor
) and simulate the bubbling regime of the reactor, while the second
the airstream (
Qconcept
) is injected to pneumatically remove the particles. The flow rate
of the air injected into the reactor column,
Qreactor
, was fixed at 14.5 standard liters per
minute (SLPM) to ensure the matching of the identified dimensionless groups between
the hot and cold prototype reactors (listed in Tables 3and 4). However, the flow rate of
the injected gas employed to remove the particles pneumatically was changed to assess
the performance of each concept. In each experiment, a fixed airflow rate was set for a
period of 10 min, while the mass of particles removed from the system was continuously
measured and recorded. This was then used to generate the mass removal rate in mass
per time, e.g., g/min. The pneumatic airflow speeds into the removal concepts were also
area averaged to yield superficial gas velocity. On this basis, the gas fraction (Equation
(2)) shows the pneumatic airflow rate through the removal mechanism against that of the
bubble column flow rate.
Xg=Remo val devi ce f l ow ra te
Bubb le co lumn f lo w r ate =Qconcept
Qreactor (2)
The particle fraction (Equation (3)) was defined as the mass of particles removed from
the system (
mp,out
) divided by the mass of particles entering (
mp,in
) the cold
prototype system
.
Xp=Particles removed (mass)
Particles added (mass)=mp,out
mp,in
(3)
The saltation velocity (Equation (4)) is the velocity that marks the transition point
between settling or particles being lifted and conveyed within a system. Hence, it was
also used to compare the removal mechanisms against one another. This is of relevance, as
many of the concepts are intrinsically reliant upon pneumatic conveyance. The saltation
velocity (vSA LT) was estimated (Equation (4)) using the Rizk correlation [23], as follows:
.
Ms
ρg·Uslt ·A=1
10(1440dp+1.96)"vSALT
pg·D#(1100d+2.5)
(4)
Energies 2024,17, 290 13 of 23
Here,
.
Ms
is the particle flow rate,
ρg
is the gas density, Ais the face area of the
particulate bed,
dp
is particle diamter, gis gravitational acceleration and Dis a characteristic
length, which is assumed to be the diameter in each developed concept.
5. Error Analysis
Experimental uncertainties exist due to the measured masses of added and removed
particles, water contained in the extracted material, measured volumetric flow rates of
inlet gas (both bubbled through the column and used in pneumatic methods) and system
losses. The particles were seen to clump in the presence of hydrostatic forces, resulting
in some water being removed from the cold prototype reactor during testing, the exact
amount of which was difficult to accurately measure. In doing so, a sample of removed
particles was taken and weighed and then allowed to dry, after which it was weighed again.
It was determined that ~22% of the initial weight was water. It was also found that up
to 10% of the microspheres remained in the system tubing, as shown in
Figures 3and 4
,
during experimental testing. This effect is accounted for as detailed in Table 6. The EK-610i
precision scales used to weigh the particulate mass have an uncertainty of
±
0.01 g. The
particle feeder system also had an uncertainty of
±
0.1 g when operating at a speed of
1.2 g/min
. The experiment duration has an uncertainty of
±
1 s, to account for the time
delay between the shutoff of the removal system and the final fluctuations in the recorded
mass. Table 6presents the uncertainties calculated for this investigation.
Table 6.
Variables and associated uncertainties in the operation and measurement of data from the
cold prototype reactor.
Symbol Parameter Measurement Component Uncertainty
mtotal,out Total measured mass removed EK-610i precision scales ±0.01 g
mp,out Particle mass removed Calculated Equation (7)
mp,in Particle mass added Particle feeder ±0.10 g
mwater Additional mass extracted from the system (liquid) Removal and filter system ±22%
mtubin g Mass not extracted from the reactor (tubing losses) Removal and filter system ±10%
Qconcept Volumetric flow rate into removal device Rotameter ±8%
Qreactor
Volumetric flow rate into the bottom of the reactor column
ALICAT flow meter ±0.7%
In order to reliably determine the particle removal fraction of each separation method
(Equation (3)), several important masses and uncertainties are recorded, as outlined in
Table 6.
mtotal,out
denotes the total mass removed from the system. The equation for
this is given (Equation (5)). This total mass removed consists of the removed particle
mass (
mp,out
), the removed liquid mass (
mwater
) and the losses of mass, where particulate
has remained trapped in the tubing (
mtubing
). This equation can be rearranged to yield
(Equation (6)), which denotes the particle mass removed from the system (
mp,out
), which is
used to calculate the particle removal fraction (Equation (3)). The uncertainty in the total
mass (∆mtotal,out) is also given (Equation (7)).
mtotal,out =hmp,out +mwater −mtubingi±∆mtotal,out (5)
mp,out =hmtotal,out −mwater +mtubingi±∆mtotal,out (6)
∆mtotal,out =q∆mp,in2+∆mwater2+∆mtubin g2(7)
There is also uncertainty associated with the measurements taken from the pneumatic
air system and its volumetric flow rates. These uncertainties were present in the rotameter
and ALICAT flow meters (Figure 4). The rotameter and ALICAT flow meters preformed
separate tasks. The rotameter was responsible for determining the airflow rate into the
removal system (
Qconcept
), whereas the ALICAT flow meter was used to measure the flow
Energies 2024,17, 290 14 of 23
rate of air into the base of the reactor (
Qreactor
). The rotameter had an uncertainty of
±
8%
and the ALICAT flow meter had an uncertainty of 0.1 SLPM when operating at the fixed
speed of 14.5 SLPM (approximately
±
0.69%). These uncertainties are factored into the
volumetric flow rates, for both the removal concept flow rate (Equation (8)) and the reactor
column flow rate (Equation (9)).
Qconcept =Qconce pt ±∆Qconcept =Qconce pt ±8% (8)
Qreactor=Qreactor ±∆Qreactor =Qreactor ±0.70% (9)
This evaluation, for both mass fraction and gas fraction, was applied to all tests
conducted using the cold prototype reactor and is reported as the error bars shown in
the results.
6. Results and Discussion
Figure 9presents the performance of each removal method in the cold prototype test
apparatus. Depicted is the particle removal fraction (
XP
) and the gas fraction (
Xg
). It can
be seen from Figure 9that the performance of each concept is different and linked to the
gas fraction at each instance. The removal performance of the pneumatically assisted weir
was seen to increase sharply as the gas fraction increased between 10 and 20. This reached
a peak removal rate of 55%, in which the assisted weir was able to remove 55% of all
particulates entering the system. After this, the performance of the assisted weir stagnated,
declining as the gas fraction was increased beyond 30.
Energies 2024, 17, 290 15 of 23
Figure 9. Comparison of particle removal fraction (𝑋) as a function of the gas fraction (𝑋) for the
four removal devices tested in the cold prototype reactor. The axes are non-dimensionalized in ac-
cordance with the methodology outlined in Section 4.
Figure 10. Performance data for the pneumatically assisted weir. Airflow rate (𝑄) in SLPM is
shown vs. the particle mass removed (𝑚,) in grams for each of the 10 min trials conducted at a
fixed speed (circles).
Figure 9.
Comparison of particle removal fraction (
XP
) as a function of the gas fraction (
Xg
) for
the four removal devices tested in the cold prototype reactor. The axes are non-dimensionalized in
accordance with the methodology outlined in Section 4.
The cyclone extractor, depicted in Figure 9, is seen to only remove particulate when
the gas fraction is higher than 17. Beyond a gas fraction of 17, the cyclone extractor begins
to remove particulate, becoming exponential between gas fractions of 30 and 38. The
maximum removal rate of the cyclone extractor was 28%.
The venturi ejector was seen to begin removing particulate at the lowest gas fraction
of 6. The removal performance of the ejector was approximately linear and could be seen to
Energies 2024,17, 290 15 of 23
constantly increase to a maximum removal rate of 54% at a gas fraction of 22. The overflow
concept is also included in Figure 9for completeness. This removal system was passive
and did not feature a separate airflow rate through the concept itself. It is for this reason
that its gas fraction remains at 0, as the numerator of
Xg
is 0, whilst there is still a non-zero
denominator due to the gas entering at the base of the reactor column. This concept was
seen to remove the particulate mass from the system over the 10 min trial. However, at the
later stages of each test, the overflow device contained large amounts of the reactor column
fluid, in addition to the particles. Which is extremely undesirable.
Figures 10–12 detail the specific, non-dimensionalized performance of each of the
concepts. These plots present the actual mass removed, in grams, on the y-axis and the
flow rate into the concept, in standard liters per minute, on the x-axis. These plots offer a
comparison of the actual mass of particulate removed from the system in each 10 min trial.
For the pneumatically assisted overflow weir, at peak removal, 23.83 g is removed from
the system. This is significantly higher than the 12.15 g that are removed from the cyclone
extractor, but close to that of the venturi ejector, which was able to remove 23.05 g.
Energies 2024, 17, 290 15 of 23
Figure 9. Comparison of particle removal fraction (𝑋) as a function of the gas fraction (𝑋) for the
four removal devices tested in the cold prototype reactor. The axes are non-dimensionalized in ac-
cordance with the methodology outlined in Section 4.
Figure 10. Performance data for the pneumatically assisted weir. Airflow rate (𝑄) in SLPM is
shown vs. the particle mass removed (𝑚,) in grams for each of the 10 min trials conducted at a
fixed speed (circles).
Figure 10.
Performance data for the pneumatically assisted weir. Airflow rate (
Qconcept
) in SLPM is
shown vs. the particle mass removed (
mp,out
) in grams for each of the 10 min trials conducted at a
fixed speed (circles).
Another important aspect is the airflow rate required to reach the maximum particulate
removal. For the pneumatically assisted weir, this occurs at 316 SLPM, 553 SLPM for the
cyclone extractor and 308 SLPM for the venturi ejector. The lower relative airflow rates of
the pneumatically assisted weir and venturi ejector highlight their suitability for industrial
applications. This is because, in an industrial device, the net energy requirement for the
removal of a given mass impacts the final production cost, which includes the required
flow rate, cooling of the hot gas phase and compression before reinjection to a particle
removal device. Relatively low airflow rates, but high extracted mass, such as those seen
in the assisted weir and venturi ejector (Figures 10 and 12), incur lower parasitic energy
requirements than the cyclone extractor, which requires much higher flow speeds.
Furthermore, when discussing suitability of each method to the real world, the physi-
cal behavior beyond just the extracted mass must be assessed. Figure 13 shows the three
pneumatic concepts during testing. Figure 14 shows the overflow concept. For the pneu-
matic concepts, Figure 13A is the assisted overflow weir, 13B the cyclone extractor and 13C
Energies 2024,17, 290 16 of 23
the venturi ejector. The white masses shown in Figure 13 are the microsphere particles used
to represent the carbon particulate. This particulate was seen to build up and accumulate
in different manners, depending on the separation concept. Further analysis of this is
provided in Figures 14–16.
Energies 2024, 17, 290 16 of 23
Figure 11. Performance data for the cyclone extractor. Airflow rate (𝑄) in SLPM is shown vs. the
particle mass removed (𝑚,) in grams for each of the 10 min trials conducted at a fixed speed (cir-
cles).
Figure 12. Performance data for the venturi ejector. Airflow rate (𝑄) in SLPM is shown vs. the
particle mass removed (𝑚,) in grams for each of the 10 min trials conducted at a fixed speed (cir-
cles).
Furthermore, when discussing suitability of each method to the real world, the physical
behavior beyond just the extracted mass must be assessed. Figure 13 shows the three pneu-
matic concepts during testing. Figure 14 shows the overflow concept. For the pneumatic
concepts, Figure 13A is the assisted overflow weir, 13B the cyclone extractor and 13C the
venturi ejector. The white masses shown in Figure 13 are the microsphere particles used to
represent the carbon particulate. This particulate was seen to build up and accumulate in
Figure 11.
Performance data for the cyclone extractor. Airflow rate (
Qconcept
) in SLPM is shown
vs. the particle mass removed (
mp,out
) in grams for each of the 10 min trials conducted at a fixed
speed (circles).
Energies 2024, 17, 290 16 of 23
Figure 11. Performance data for the cyclone extractor. Airflow rate (𝑄) in SLPM is shown vs. the
particle mass removed (𝑚,) in grams for each of the 10 min trials conducted at a fixed speed (cir-
cles).
Figure 12. Performance data for the venturi ejector. Airflow rate (𝑄) in SLPM is shown vs. the
particle mass removed (𝑚,) in grams for each of the 10 min trials conducted at a fixed speed (cir-
cles).
Furthermore, when discussing suitability of each method to the real world, the physical
behavior beyond just the extracted mass must be assessed. Figure 13 shows the three pneu-
matic concepts during testing. Figure 14 shows the overflow concept. For the pneumatic
concepts, Figure 13A is the assisted overflow weir, 13B the cyclone extractor and 13C the
venturi ejector. The white masses shown in Figure 13 are the microsphere particles used to
represent the carbon particulate. This particulate was seen to build up and accumulate in
Figure 12.
Performance data for the venturi ejector. Airflow rate (
Qconcept
) in SLPM is shown
vs. the particle mass removed (
mp,out
) in grams for each of the 10 min trials conducted at a fixed
speed (circles).
Energies 2024,17, 290 17 of 23
Energies 2024, 17, 290 17 of 23
different manners, depending on the separation concept. Further analysis of this is provided
in Figures 14–16.
Figure 13. Operational images of the pneumatically assisted overflow weir (A), the cyclone extractor
(B) and the venturi ejector (C).
Mechanically, the operation of the overflow concept was the simplest. This can be seen
in Figure 14. The particulate level would build over the duration of the 10 min trial. When
this level was sufficiently high, the particles would overflow the sides of column and into
the extraction system. Whilst this worked well initially, resulting in predominantly particu-
late separation, the bubbling of the column resulted in water being extracted. The combina-
tion of particles and bubbling fluid caused the particles to clump, as can be seen at B in
Figure 14. Operationally, this would present significant challenges in a MCB methane py-
rolysis reactor, as the liquid metal may be extracted with any carbon particulate, complicat-
ing the refinement of the carbon product.
Interesting air and particle dynamics were observed during the operation of the pneu-
matically assisted overflow weir. Figure 15 displays images of both the overhead view (left)
and side view (right) of the weir. Particle formation and clumping were seen to occur on the
roof of the weir duct at A and also above the pneumatic air inlet at B. Shown in the righthand
side of Figure 15 is the turbulent disturbance to the particle fluid interface. The surface of
this was affected by the pneumatic air passing over it. This caused the surface to rise, closest
to the inlet, and drop on the opposite side of the reactor particle fluid interface. This phe-
nomenon would be undesirable if the technology was deployed in an upscaled, functioning
MCB methane pyrolysis reactor. This could correlate to reduced removal efficiency for
larger diameter columns, as the effect will have a large surface to act on.
During the operation of the cyclone extractor, depicted in Figure 16, several observa-
tions were made. The degree of particle entrainment varies as the vertical height of the col-
umn is increased. Furthest from the fluid particle interface, at A, there is less particles than
lower down at B and C. This is likely due to two factors: The first is the disbursement of the
pneumatic airstream. The nozzles are located at C, which results in a high-speed, strongly
swirling flow that separates as it moves vertically up the column. This separation then re-
sults in a vertical and swirl speed decrease as it transitions from C to B and then from B to
A. The second factor is the entrainment of the particles. Some particles do not remain suffi-
ciently entrained in the flow. This causes them to separate and fall back down to the particle
interface. This was seen to occur at B and C, where the flow is chaotic and turbulent. This
Figure 13.
Operational images of the pneumatically assisted overflow weir (
A
), the cyclone extractor
(B) and the venturi ejector (C).
Energies 2024, 17, 290 18 of 23
phenomenon indicates that scaling the separation technology to industrial scale—by in-
creasing the extractor height—may result in a greater degree of difficulty entraining these
particles in the flow, decreasing separation efficiency. However, it was noted that as the
particle–fluid level increased, as more particulate entered the test reactor, the separation be-
tween the nozzles and interface reduced. This reduction corresponded to a greater degree
of particles becoming entrained in the airstream.
Figure 14. Operational images of the overflow separation concept. The overflow edge indicated (A)
and the particle–fluid bed indicated (B).
Figure 15. Operational image of the pneumatically assisted overflow weir. Left image is above the
removal concept and the right image is a photo taken through the side wall. (A) shows particle splatter
formation on the roof and (B) shows splatter formation on the side walls, above the inlet slit (bottom
right).
Figure 14.
Operational images of the overflow separation concept. The overflow edge indicated
(A) and the particle–fluid bed indicated (B).
Mechanically, the operation of the overflow concept was the simplest. This can be
seen in Figure 14. The particulate level would build over the duration of the 10 min
trial. When this level was sufficiently high, the particles would overflow the sides of
column and into the extraction system. Whilst this worked well initially, resulting in
predominantly particulate separation, the bubbling of the column resulted in water being
extracted. The combination of particles and bubbling fluid caused the particles to clump,
as can be seen at B in Figure 14. Operationally, this would present significant challenges in
a MCB methane pyrolysis reactor, as the liquid metal may be extracted with any carbon
particulate, complicating the refinement of the carbon product.
Energies 2024,17, 290 18 of 23
Energies 2024, 17, 290 18 of 23
phenomenon indicates that scaling the separation technology to industrial scale—by in-
creasing the extractor height—may result in a greater degree of difficulty entraining these
particles in the flow, decreasing separation efficiency. However, it was noted that as the
particle–fluid level increased, as more particulate entered the test reactor, the separation be-
tween the nozzles and interface reduced. This reduction corresponded to a greater degree
of particles becoming entrained in the airstream.
Figure 14. Operational images of the overflow separation concept. The overflow edge indicated (A)
and the particle–fluid bed indicated (B).
Figure 15. Operational image of the pneumatically assisted overflow weir. Left image is above the
removal concept and the right image is a photo taken through the side wall. (A) shows particle splatter
formation on the roof and (B) shows splatter formation on the side walls, above the inlet slit (bottom
right).
Figure 15.
Operational image of the pneumatically assisted overflow weir. Left image is above the
removal concept and the right image is a photo taken through the side wall. (
A
) shows particle
splatter formation on the roof and (
B
) shows splatter formation on the side walls, above the inlet slit
(bottom right).
Energies 2024, 17, 290 19 of 23
The operation of the venturi ejector, shown in Figure 17, displayed the smallest de-
gree of turbulence and particle disturbance. The ejector or suction port, located at A, was
seen to relatively cleanly extract the particles present on the fluid interface at B. Unlike the
pneumatically assisted weir and cyclone extractor, there was no turbulent flow regime
present in the test reactor. This can be seen by the lack of particulate splaer present in
Figure 17. During operation, the particle bed height increased as the microspheres built
up in the system. Upon reaching the level of the suction port (Figure 16A), the particles
were efficiently removed from the system. It was also seen that at higher flow speed, and
subsequent higher venturi pressures, the particle-fluid level height would be drawn up
closer to the suction port at A. This effect scaled proportionally with flow speed into the
venturi, as detailed in Figures 9 and 12. The relative ease of operation, and limited turbu-
lence indicates that the venturi ejector concept should be investigated further.
Figure 16. Operational image of the cyclone extractor geometry. (A) indicates the upper cyclone
assembly, (B) the middle and (C) the lower assembly. Particulate buildup is shown to occur at (C).
The final comparison between the removal methods is specific to those that utilize
pneumatic air. As discussed in Section 4, the saltation velocity (Table 7) is the point that
marks the transition between particles at rest, and those that are conveyed within a sys-
tem. This is calculated according to (4), using the density of the pneumatic air (1.20 kg/m3),
area of the reactor cross fluid interface (m2), mean diameter of the microsphere particles
(80 micron), gravity (9.81 m/s2) and the diameter of the reactor (0.09 m). This can be used
to determine the expected speed at which the particles will be conveyed. As can be seen
from Table 7, the concepts were unable to begin conveying the particles at the expected
speed. These differences can be aributed to the fact that each pneumatic concept did not
apply the effect directly at the fluid level, but at a distance above it, due to the transient
bubbling of the column. Decreasing the separation distance between the concept’s actua-
tion point (where the air is applied) and the fluid level may decrease the difference be-
tween the recorded and expected saltation velocity.
Figure 16.
Operational image of the cyclone extractor geometry. (
A
) indicates the upper cyclone
assembly, (B) the middle and (C) the lower assembly. Particulate buildup is shown to occur at (C).
Interesting air and particle dynamics were observed during the operation of the
pneumatically assisted overflow weir. Figure 15 displays images of both the overhead view
(left) and side view (right) of the weir. Particle formation and clumping were seen to occur
on the roof of the weir duct at A and also above the pneumatic air inlet at B. Shown in
the righthand side of Figure 15 is the turbulent disturbance to the particle fluid interface.
The surface of this was affected by the pneumatic air passing over it. This caused the
surface to rise, closest to the inlet, and drop on the opposite side of the reactor particle
fluid interface. This phenomenon would be undesirable if the technology was deployed in
Energies 2024,17, 290 19 of 23
an upscaled, functioning MCB methane pyrolysis reactor. This could correlate to reduced
removal efficiency for larger diameter columns, as the effect will have a large surface to
act on.
During the operation of the cyclone extractor, depicted in Figure 16, several observa-
tions were made. The degree of particle entrainment varies as the vertical height of the
column is increased. Furthest from the fluid particle interface, at A, there is less particles
than lower down at B and C. This is likely due to two factors: The first is the disbursement
of the pneumatic airstream. The nozzles are located at C, which results in a high-speed,
strongly swirling flow that separates as it moves vertically up the column. This separation
then results in a vertical and swirl speed decrease as it transitions from C to B and then from
B to A. The second factor is the entrainment of the particles. Some particles do not remain
sufficiently entrained in the flow. This causes them to separate and fall back down to the
particle interface. This was seen to occur at B and C, where the flow is chaotic and turbulent.
This phenomenon indicates that scaling the separation technology to industrial scale—by
increasing the extractor height—may result in a greater degree of difficulty entraining these
particles in the flow, decreasing separation efficiency. However, it was noted that as the
particle–fluid level increased, as more particulate entered the test reactor, the separation
between the nozzles and interface reduced. This reduction corresponded to a greater degree
of particles becoming entrained in the airstream.
The operation of the venturi ejector, shown in Figure 17, displayed the smallest degree
of turbulence and particle disturbance. The ejector or suction port, located at A, was seen
to relatively cleanly extract the particles present on the fluid interface at B. Unlike the
pneumatically assisted weir and cyclone extractor, there was no turbulent flow regime
present in the test reactor. This can be seen by the lack of particulate splatter present in
Figure 17. During operation, the particle bed height increased as the microspheres built
up in the system. Upon reaching the level of the suction port (Figure 16A), the particles
were efficiently removed from the system. It was also seen that at higher flow speed,
and subsequent higher venturi pressures, the particle-fluid level height would be drawn
up closer to the suction port at A. This effect scaled proportionally with flow speed into
the venturi, as detailed in Figures 9and 12. The relative ease of operation, and limited
turbulence indicates that the venturi ejector concept should be investigated further.
Energies 2024, 17, 290 20 of 23
Figure 17. Operational image of the venturi ejector. (A) indicates the location of the suction port. (B)
shows the clumping present in the fluid interface.
Table 7. Pneumatic concept’s recorded and expected saltation velocities.
Removal Method Theoretical Saltation Velocity Observed Saltation Velocity
Pneumatically assisted
overflow weir 0.97 m/s 3.49 m/s
Cyclone extractor 1.94 m/s 4.23 m/s
Venturi ejector 0.97 m/s 3.49 m/s
7. Conclusions
Experimental testing was undertaken to assess the performance of the four distinct
separation technologies—overflow, pneumatically assisted overflow weir, cyclone sepa-
rator and venturi ejector—for the removal of solid particles from a MCB methane pyroly-
sis reactor, using a cold prototype column. A dimensionally similar cold prototype reactor
was developed to simulate a functioning hot reactor. The cold prototype was developed
based on the existing literature and the important physical effects that would be present
at high temperature. To ensure the cold prototype simulated the physical effects accu-
rately, its particle Stokes number, particle Froude number, Reynolds number, gas Froude
number and density ratio were matched. The close physical similarity, as depicted in Fig-
ure 2, indicates a robust testing model, that should provide similar physical interactions
to a hot reactor. Subsequently, this model was then used for experimental testing that was
able to determine the removal rate of particulate from the cold prototype reactor, in addi-
tion to the flow rates required to do so. All concepts tested were able to remove particulate
from the cold prototype reactor to differing degrees. The overflow achieved a maximum
particle removal rate of 100%, pneumatically assisted overflow weir 55%, the cyclone ex-
tractor 28% and the venturi ejector 54%. For all concepts other than the overflow, some
particulate was retained in the reactor, accumulating predominantly in the extraction pip-
ing and reactor overhead. The primary findings from these experiments are:
a. While the overflow separation mechanism can achieve full particle removal (100%),
the design is limited by the particle fluid interactions and the extracted mass was
contaminated with a large volume of water. If this technology were adapted to a
Figure 17.
Operational image of the venturi ejector. (
A
) indicates the location of the suction port.
(B) shows the clumping present in the fluid interface.
Energies 2024,17, 290 20 of 23
The final comparison between the removal methods is specific to those that utilize
pneumatic air. As discussed in Section 4, the saltation velocity (Table 7) is the point that
marks the transition between particles at rest, and those that are conveyed within a system.
This is calculated according to (4), using the density of the pneumatic air (1.20 kg/m
3
),
area of the reactor cross fluid interface (m
2
), mean diameter of the microsphere particles
(
80 micron
), gravity (9.81 m/s
2
) and the diameter of the reactor (0.09 m). This can be used
to determine the expected speed at which the particles will be conveyed. As can be seen
from Table 7, the concepts were unable to begin conveying the particles at the expected
speed. These differences can be attributed to the fact that each pneumatic concept did not
apply the effect directly at the fluid level, but at a distance above it, due to the transient
bubbling of the column. Decreasing the separation distance between the concept’s actuation
point (where the air is applied) and the fluid level may decrease the difference between the
recorded and expected saltation velocity.
Table 7. Pneumatic concept’s recorded and expected saltation velocities.
Removal Method Theoretical Saltation Velocity Observed Saltation Velocity
Pneumatically assisted
overflow weir 0.97 m/s 3.49 m/s
Cyclone extractor 1.94 m/s 4.23 m/s
Venturi ejector 0.97 m/s 3.49 m/s
7. Conclusions
Experimental testing was undertaken to assess the performance of the four distinct
separation technologies—overflow, pneumatically assisted overflow weir, cyclone separator
and venturi ejector—for the removal of solid particles from a MCB methane pyrolysis
reactor, using a cold prototype column. A dimensionally similar cold prototype reactor
was developed to simulate a functioning hot reactor. The cold prototype was developed
based on the existing literature and the important physical effects that would be present at
high temperature. To ensure the cold prototype simulated the physical effects accurately,
its particle Stokes number, particle Froude number, Reynolds number, gas Froude number
and density ratio were matched. The close physical similarity, as depicted in Figure 2,
indicates a robust testing model, that should provide similar physical interactions to a hot
reactor. Subsequently, this model was then used for experimental testing that was able to
determine the removal rate of particulate from the cold prototype reactor, in addition to the
flow rates required to do so. All concepts tested were able to remove particulate from the
cold prototype reactor to differing degrees. The overflow achieved a maximum particle
removal rate of 100%, pneumatically assisted overflow weir 55%, the cyclone extractor 28%
and the venturi ejector 54%. For all concepts other than the overflow, some particulate was
retained in the reactor, accumulating predominantly in the extraction piping and reactor
overhead. The primary findings from these experiments are:
a.
While the overflow separation mechanism can achieve full particle removal (100%),
the design is limited by the particle fluid interactions and the extracted mass was
contaminated with a large volume of water. If this technology were adapted to a
functioning MCB methane pyrolysis reactor, it could result in some of the liquid metal
being extracted. Further research into mechanical skimming or the use of filtration
systems should be conducted. However, it should be noted that the corrosive nature
of the molten metal bubble column materials may hinder prolonged use in industrial
environments, where such a mechanism may be detrimentally corroded.
b.
The pneumatically assisted overflow weird was able to induce separation at lower
gas flow speeds up to a removal rate of 55%. At high flow speeds, turbulent effects
became present, with the injected air displacing the particle–fluid interface. This may
present challenges if the technology is scaled up for MCB reactors, as it is unknown
what effect increasing the diameter of the reactor will have.
Energies 2024,17, 290 21 of 23
c.
The cyclone extractor can remove particulate, with relatively lower quantities of liq-
uid than the overflow extraction. However, the effectiveness of the particle separation
varies with the height of the column and the height of the nozzles from the particle
bed. Moreover, the cyclone mechanism is more effective at high gas flow rates,
which may hinder its feasibility in industrial operations, due to the impracticality of
implementing high flow rates in a sealed MCB methane pyrolysis reactor.
d.
The venturi ejector was effective in removing particulate but was seen to draw up
the fluid bed height, decreasing suction port separation. When the bed height was
increased, some fluid was drawn into the suction port, contaminating the extracted
particulate. This mechanism was successful enough to warrant further analysis,
displaying a linear trend in particle removal for increases in flow rate. Further testing
should be conducted into adaptions with more than one suction port.
It should be noted that the impact of additional constraints, such as surface tension
and wettability, were not explored for the cold prototype or the reference hot reactor. This
would be a natural extension of the work and would allow for their respective impacts to
be accounted for in further iterations of the removal designs. Further refinement of the
presented concepts numerically, computationally or experimentally would be extremely
beneficial to the wider state of the art.
Author Contributions:
Conceptualization, Z.C.-B., T.D.L.P., C.P., M.R., L.R., J.S., G.J.N. and M.J.;
methodology, Z.C.-B., T.D.L.P., C.P., M.R., L.R. and J.S.; validation, Z.C.-B., T.D.L.P., C.P., M.R., L.R.
and J.S.; formal analysis, Z.C.-B., T.D.L.P., C.P., M.R., L.R. and J.S.; writing—original draft preparation,
Z.C.-B., G.J.N. and M.J.; writing—review and editing, Z.C.-B., G.J.N. and M.J.; supervision, G.J.N.
and M.J. The first six listed authors had an equal contribution to the planning, setup, execution and
analysis of the experiments contained within this article. All authors have read and agreed to the
published version of the manuscript.
Funding:
This research was supported by the Future Fuels Cooperative Research Center (FFCRC)
under Project ID 81125509.
Data Availability Statement: Data are contained within this article.
Conflicts of Interest: The authors declare no conflicts of interest.
Abbreviations
Nomenclature used within the manuscript body. Described below are the symbols used to
denote each term and the definition or meaning of each term. Listed in order
of appearance.
Symbol Definition
CO2Carbon dioxide
SMR Steam methane reforming
CG Coal gasification
H2Hydrogen
CH4Methane
C Composition
ρDensity
D Diameter
h Height
P Pressure
T Temperature
ηConversion efficiency
rProduction rate
SLPM Standard liters per minute
FrpParticle Froude number
Energies 2024,17, 290 22 of 23
Stk Particle Stokes number
FrgGas Froude number
Re Reynolds number
sDensity ratio
XgGas fraction
XpParticle fraction
.
MsParticle flow rate
g Gravity
vSALT Saltation velocity
mMass
∆mMass uncertainty
QGas flow rate
∆QGas flow rate uncertainty
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