Energy Procedia 63 ( 2014 ) 1084 – 1090

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1876-6102 © 2014 The Authors. Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license

(http://creativecommons.org/licenses/by-nc-nd/3.0/).

Peer-review under responsibility of the Organizing Committee of GHGT-12

doi: 10.1016/j.egypro.2014.11.116

GHGT-12

Improving the efficiency of a chilled ammonia CO2 capture plant

through solid formation: a thermodynamic analysis

Matteo Gazzani, Daniel Sutter, Marco Mazzotti*

ETH Zurich, Institute of Process Engineering, Sonneggstrasse 3, 8092 Zurich, Switzerland

Abstract

Post-combustion chemical absorption is regarded as the state-of-the-art commercially-available CO2 capture process. The

adoption of aqueous ammonia as solvent, leading to the so-called Chilled Ammonia capture Process (CAP), has long been

considered one of the most promising alternatives to amine-based for post-combustion carbon capture. This work investigates the

development of a second generation CAP where the capture efficiency is improved by making use of a crystallizer to form solids

in the process. The reference standard CAP and the advanced crystallizer-based CAP are simulated in Aspen using the Extended

UNIQUAC thermodynamic model. The two CAP solutions are compared in term of the different energy penalties introduced

applying the capture process to a conventional Ultra Super Critical (USC) power plant. Thanks to the solid formation, the CAP

with the crystallizer features a lower energy penalization with a decrease of about 10% compared to the total penalty of the

standard CAP.

© 2013 The Authors. Published by Elsevier Ltd.

Selection and peer-review under responsibility of GHGT.

Keywords: Chilled Ammonia Process, CO2 post-combustion capture, CCS, CO2 capture energy penalty

CAP Chilled Ammonia Process

FGD Flue Gas Desulfurizer

LHV Lower Heating Value, MJ/kgfuel

USC Ultra Super Critical

* Corresponding author. Tel.: +41 44 632 2456; fax: +41 44 632 1141.

E-mail address: marco.mazzotti@ipe.mavt.ethz.ch

© 2014 The Authors. Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license

(http://creativecommons.org/licenses/by-nc-nd/3.0/).

Peer-review under responsibility of the Organizing Committee of GHGT-12

Matteo Gazzani et al. / Energy Procedia 63 ( 2014 ) 1084 – 1090 1085

1. Introduction

The power plant flue gas decarbonisation via conventional amine scrubbing is regarded as the state-of-the-art

commercially-available CO2 capture process. This is mainly due to the large number of existing plants for acid gas

treatment and the associated maturity of the technology. On the other hand, the large energy penalty, the solvent

degradation and the difficulty of handling corrosive solutions have prompted the research towards advanced

processes. The use of chilled ammonia as solvent is a promising solution which has long been regarded as a possible

alternative to amines. The main advantages of this technology, also known as Chilled Ammonia capture Process

(CAP), include: i) low cost and large availability of the solvent, ii) chemically stable solution, iii) high stability to

oxygen, iv) regeneration at medium pressure and v) high CO2 carrying capacity.

The performance of the CAP has been assessed in only a few works, without reaching a consistent and thorough

assessment. Darde et al [1], Valenti et al [2] and Versteeg and Rubin [3] report significant thermodynamic

advantages over the conventional amine solutions whilst the contrary is shown in Mathias et al [4]. Recently, Valenti

et al [5] reported substantially equivalent performance between amine and CAP.

The performance evaluation is complex because the chilled ammonia plants entail several interdependent energy

intensive requirements: i) heat for rich solution regeneration, ii) chilling duty for lean solution and flue gas cooling,

and iii) heat requirement for the ammonia wash section. It is well known that the plant operating conditions strongly

affect the chemical behaviour of the process: solid phases, primarily consisting of ammonium carbonate and

bicarbonate, may form in the absorber and in related components. In the existing CAP plants the solid formation has

been strictly avoided due to the complexity of handling solids. On the other hand, solid formation offers different

advantages from a thermodynamic point of view: i) reduction of the mass flow per tonne of CO2 of the rich solution

sent to the stripper, thus resulting in a significant decrease of the heat requirement; ii) reduction of the stripper

dimension; and iii) reduction of the ammonia slip from the absorber. Accordingly, the performance of the next

generation CAP can be improved when exploiting the formation of solids.

This work investigates the possibility of using solids in the CAP with a dedicated solid formation unit and

compares the results of the proposed concept with the standard CAP.

2. Plant layout

2.1. Standard CAP without solid formation

The overall plant layout of the standard CAP is shown in Figure 1. The flue gas exiting the Flue Gas Desulfurizer

(FGD) (at about 40-60 °C, in saturated conditions) enters the direct contact cooling section where most of the water

and residual contaminants are removed along with the gas cooling. The flue gas enters the CO2 capture section at

18°C; the capture scheme is a standard layout with two adsorption/desorption towers, a regenerative heat exchanger

to harness the energy content of the CO2 lean stream and a solution pump to match the different operating pressure

of the absorber/desorber. The treated flue gas exiting the CO2 absorber are sent to the ammonia abatement section

where the NH3 slip is lowered to few ppm. The bulk NH3 removal is carried out in the water wash through a

conventional absorber/desorber process and therefore similar to the cycle for CO2 capture. The flue gas, with a

lower amount of CO2 but containing ammonia, is chilled and enters the bottom of the absorber, where it is contacted

with water flowing from the upper stages of the column [6]. The ammonia content in the flue gas leaving the upper

part of the column is reduced to environmentally acceptable level in a final acid wash. The liquid exiting the bottom

of the absorber is regenerated in the desorption column providing almost pure water back to the absorber. The

ammonia recovered from the flue gas is recycled to the absorber in the CO2 capture island.

1086 Matteo Gazzani et al. / Energy Procedia 63 ( 2014 ) 1084 – 1090

Figure 1: Plant layout of the standard CAP

In order to limit the energy penalty associated with the ammonia slip abatement, the CO2 absorber column is

designed such that the first stages of the column contribute to lower the ammonia vapour pressure. This is achieved

recycling part of the CO2 rich solution from the bottom of the column which is cooled and chilled before re-entering

the column [7]. The combination of high CO2 concentration and low temperature prevents the further CO2 uptake

while favouring the ammonia absorption. A schematic representation of the column is shown in Figure 2.

Figure 2: Details of the absorber configuration [7]

2.2. Advanced CAP with crystallizer for solid formation

From a qualitative point of view the overall plant layout of the crystallizer-based CAP (as introduced in [8] and

shown in Figure 3) is similar to the standard CAP but for the CO2 capture island. Provided that handling solids in a

packed column is not feasible from an engineering point of view, the operating conditions of the plant have been

tuned in order to carefully avoid any solid formation inside the absorber and desorber. The solid formation in the

rich solution is restricted to a crystallizer downstream of the absorber. In order to reduce the mass flow sent to the

stripper as much as possible, all the solid material exiting the crystallizer is separated in a hydrocyclone together

with a part of the liquid solution. The obtained slurry is sent to a solid dissolution reactor and then to a regenerative

heat exchanger; therefore all the solids are dissolved before entering the desorption column. The liquid stream

separated in the hydrocyclone is sent back to the absorber to control the absorber temperature and the ammonia slip.

Matteo Gazzani et al. / Energy Procedia 63 ( 2014 ) 1084 – 1090 1087

The rich solution exiting the absorber must be cooled to 5°C to favour the solid formation. No advanced heat

integration, for example the regenerative heat exchange between the separated slurry and the rich CO2 liquid, has

been considered in this work. The chilling duty for the crystallization has been calculated cooling the solution from

18 to 5°C.

Figure 3: Plant layout of the crystallizer CAP

3. Methodology and approach

The use of an appropriate thermodynamic model that correctly describes the interaction among NH3, CO2 and

H2O in the system and the solid-liquid-vapor equilibria is of a paramount importance for obtaining reliable results in

the simulations of the CAP. Accordingly, this work uses the Extended UNIQUAC model developed by Thomsen [9]

and upgraded by Darde et al [10]. The model considers five different solid phases:

1. Ammonium bicarbonate (BC) NH4HCO3

2. Ammonium carbonate (CB) (NH4)2CO3∙H2O

3. Ammonium sesqui-carbonate (SC) (NH4)2CO3∙2NH4HCO3

4. Ammonium carbamate (CM) NH2COONH4

5. Ice H2O

The solubility data for solids 1-4 used in the parameter fitting procedure are based on Janecke data [11]. The gas-

phase fugacities are calculated with the Soave-Redlich-Kwong equation of state.

Both the CAP solutions with and without crystallizer have been simulated with Aspen Plus©. The vapour-liquid

equilibrium in the absorber/desorber of the CO2 section and the NH3 water wash has been modelled with the

rigorous Aspen RadFraq approach for multistage vapor-liquid systems assigning stagewise Murphree efficiencies

(only in the absorbers) and checking for salt precipitation. The crystallization step is simulated with a continuous

stirred reactor working at thermodynamic equilibrium. The plant simulation is a closed loop with connected streams

between the desorber and absorber section. The convergence is guaranteed by making use of tear streams and

calculating the ammonia and water make-up.

The solid formation along the process is carefully checked by making use of ternary phase diagrams for the CO2-

NH3-H2O system: critical points and the CO2/NH3 absorber profiles are reported on ternary diagrams to control that

no solid formation takes place outside the crystallizer reactor. More details about solid formation are reported in

Sutter et al [12]

The main assumptions for the CAP simulations are reported in Table 1.

1088 Matteo Gazzani et al. / Energy Procedia 63 ( 2014 ) 1084 – 1090

Table 1: Main assumptions for the CAP calculation in Aspen

CAP process specifications

CO2 capture, %

CO2 lean loading (entering the absorber)

CO2 rich loading (exiting the absorber)

CO2 purity before storage

Ammonia slip in the flue gas exiting the CO2 absorber, ppm

Ammonia slip after the water wash, ppm

Flue gas temperature entering absorber, °C

Adsorption/Desorption pressure, bar

90

0.3-0.35

> 0.5

> 0.99

< 8000

200

18

1.01/10

Flue gas composition: CO2 15.6%, N2 66.0%, O2 17.4%, Ar 1.0%

Utilities

Chilling water temperature, °C

Cooling water temperature, °C

Heat exchanger ΔTmin, °C or K

2

15

3

In order to calculate the overall energy penalty of the two CAP solutions, a base case USC plant without carbon

capture has been considered as proposed by EBTF [13]. The power balance of this plant is reported in Table 2.

Table 2: Power balance of the considered USC plant without CO2 capture as in [13]

Net power output, MW

Fuel input, MW

Net LHV efficiency, %

Flue gas mass flow, kg/s

Specific emissions, gCO2/kWhel

758.6

1676.6

45.2

740

772

The energy penalties introduced with the CAP process are due to the need of thermal and electric energy in five

main processes: i) thermal energy for the CO2 capture reboiler, ii) electric energy for the chilling duty, iii) thermal

energy for the NH3 wash reboiler, iv) electric energy for the CO2 compression and v) electric energy for the pumps.

The penalty associated with the use of steam in the plant reboilers has been calculated considering the decrease in

the steam turbine net power. The resulting difference in the power required to handle the condenser duty has been as

well considered. The electric energy associated with the chilling duty has been computed using the Coefficient of

Performance (COP) of the cooling cycle.

%12達竪辿狸狸辿樽巽 噺 探坦奪脱探狸"奪脱脱奪達担奪樽奪嘆巽湛"辰奪鱈叩樽辰 噺 濯冬塘套嶋嶋套投塔茸冬搭悼東梼刀棟棟搭梼 (1)

The COP has been calculated considering the ideal Carnot COP derived from the temperature of the evaporator

and condenser in the inverse Rankine cycle and a second-law efficiency as following:

%12達竪辿狸狸辿樽巽 噺 %12大叩嘆樽誰担 "と瀧瀧 噺 鐸刀湯倒鐸冬搭投凍貸鐸刀湯倒 "と瀧瀧 (2)

Where: Teva = 273 K, Tcond = 298 K and ηII = 0.6.

The energy required in the CO2 compression has been simulated in Aspen Plus© as reported in [14].

Matteo Gazzani et al. / Energy Procedia 63 ( 2014 ) 1084 – 1090 1089

4. Results and Discussion

The results in terms of energy penalties of the standard CAP and the crystallizer CAP are shown in Figure 4.

Compared to the state-of-the-art CAP without solid formation, the mass flow sent to the stripper is reduced by about

40% thanks to the higher CO2 loading, thus decreasing the heat required for regenerating the rich solution. The

penalization on the steam turbine power output reduces accordingly. In the standard CAP the decrease in the steam

turbine power due to the reboiler duty accounts for 70% of the total energy penalty whilst in the crystallizer CAP it

decreases to about 51.1% (based on the total energy penalty of the standard case). On the other hand, in order to

trigger the precipitation in the crystallizer, the entire solution exiting the absorber has to be cooled thus entailing a

large energy consumption because of the chilling cycle. The associated energy penalty increases from 3.9 to 19.7

percentage points. The pump and waste heat management play a secondary role; anyhow, the crystallizer CAP

allows reducing significantly the pump power thanks to the smaller circulating flow.

When the desorber pressure is constant in the two CAP solutions, no differences arise in the CO2 compression

section. However, it is worth mentioning that an increase in the desorber pressure of the crystallizer CAP does not

suffer of a higher pump power thanks to the reduced solvent flow rate. The additional design cost would also be

limited thanks to the smaller desorber size.

Another important feature of the crystallizer CAP is the reduction in the consumption of the ammonia slip

process. In fact, the CAP with solid formation can further reduce the ammonia slip in the top part of the absorber

thanks to the lower temperature and higher flow rate of the pump around. Accordingly, the consumption decreases

from 10 to 5.7 percentage points for the standard and crystallizer CAP respectively.

The resulting overall energy penalization is reduced by about 10% adopting a second generation CAP with the

crystallizer. Moreover, the chilling duty can be further reduced by improving the heat integration within the whole

plant.

Figure 4: Comparison of the energy penalties between the standard and crystallizer CAP; bar length is normalized based on the standard CAP

penalty = 100, whilst numbers within the bar refer to the total penalty of each layout.

5. Conclusions

This work discussed the development of an advanced CAP layout where a crystallizer for solid formation is

adopted to reduce the capture energy penalization. Two plant layouts were simulated in Aspen Plus with the

Extended UNIQUAC model and compared using a reference USC power plant without CO2 capture. Compared to

1090 Matteo Gazzani et al. / Energy Procedia 63 ( 2014 ) 1084 – 1090

the standard CAP, the advanced solution has shown a reduction in the energy penalization of about 10%. Further

energy saving can be pursued increasing the crystallizer heat integration with other components.

Nevertheless, in order to further develop this concept more work has to be done: i) the solid formation kinetics

have to be investigated, ii) the crystallizer design has be optimized and iii) the full process energy requirements have

to be optimized by manipulating the plant operating variables.

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