Supplementary material for
Thermodynamic equilibrium analysis of entrained flow gasification of spent pulping liquors
Erik Furusjö*, Yawer Jafri
Luleå University of Technology, Division of Energy Sciences, SE-971 87 Luleå, Sweden
* Corresponding author; e-mail ; phone +46-920-492545
1.Feedstock heat of formation
Solving the gasifier energy balance requires the enthalpy of all streams entering and leaving the process. For the oxygen, nitrogen and gasification products enthalpy estimation is straightforward since the chemical composition is known. However, for the spent pulping liquors, which are complex mixtures of organic and inorganic species it is not possible to use tabulated data for heat of formation. Instead their enthalpies of formation are calculated using Hess’ law from higher heating values (HHV) determined by bomb calorimetry and assumptions about the bomb combustion products as described earlier[1]. When this approach is used it is very important to know exactly which heating value is measured, i.e. what are the final combustion products in the bomb calorimeter? Spent pulping liquors of the type studied in this work are characterized by high sodium and sulfur contents, which are typically 20% and 5% on a dry weight basis. In addition, there is often potassium present up to 20% of the sodium concentration. In the context of this discussion, there is no major difference between the behavior of sodium and potassium. Hence, below we use the notation sodium to represent both alkali metals.
In the presence of high partial pressures of CO2 in a bomb calorimeter, sodium and potassium will tend to form carbonate. During normal combustion fuel sulfur typically forms SO2, but during HHV determination in the bomb calorimeter fuel sulfur is oxidized to SO3. In addition, water is added to create a saturated atmosphere in which the formed SO3 reacts to form sulfuric acid H2SO4. H2SO4 (pKA=-3) and HSO4- (pKA=2) are both stronger acids than H2CO3 (pKA,apparent=6) and HCO3- (pKA=10). Hence, the equilibriums in Eq 1 are driven strongly to the rightmost side and sulfur in the bomb combustion products is found as Na2SO4(aq). However, if the fuel composition is such that there is excess H2SO4, i.e. when fuel molar ratio S/(Na+K)2>1, the reaction will stop at NaHSO4,as illustrated in Eq.2 for a case with S/Na2=2.
Na2CO3(aq) + H2SO4(aq) ↔ NaHSO4(aq)+ NaHCO3(aq) ↔ Na2SO4(aq) + CO2(g) + H2O Eq 1
Na2CO3(aq) + 2 H2SO4(aq) ↔ 2 NaHSO4(aq)+ CO2(g) + H2O Eq 2
The conclusion that all fuel sulfur ends up as HSO4- or SO42- is supported by the fact that the standard method used for the determination of BL sulfur (SCAN-N 35:96 Annex B) is based on determination of sulfate in the bomb calorimeter combustion products by ion chromatography. A similar method is used for solid biomass (ISO 16994:2015 Method A). It can be noted that a so-called sulfur correction is sometimes used to correct measured fuel heating values for the fact that SO3/H2SO4 is formed in the bomb calorimeter instead of SO2, but that is not the case for the heating values used in this work.
For BL, which has a molar ratio S/(Na+K)2<1, the bomb combustion products are, based on the chemical equilibriums discussed above, assumed to be CO2, H2O, Na2SO4, K2SO4, Na2CO3, K2CO3, NaCl, KCl and N2. All sulfur is assumed to form sulfate salts while the remainder of the sodium and potassium is assumed to form carbonate salts. The partition between Na and K salts is made simply based on fuel Na/K ratio.
For STL, which has a molar ratio S/(Na+K)2>1, it is not possible for all sulfur to form sulfate salts as noted above. In this case, the combustion products are assumed to be CO2, H2O, Na2SO4, K2SO4, NaHSO4, KHSO4, NaCl, KCl and N2. Hence, sodium and potassium are assumed to form chloride, hydrogen sulfate and sulfate to balance all Na, K and S. All carbon forms CO2 in this case.
It can be noted that the presence of a rather large amount of water in the calorimetric bomb after combustion (added water and combustion water) will most likely mean that the formed Na and K salts discussed above will partly be present in hydrated and/or dissolved form. In a previous study for black liquor[1], it was concluded that accounting for such effects had an effect of 2.5% compared to the heating value. We have disregarded such effects in the present work.
The above mentioned approach to estimate the inorganic combustion products is also used to calculate the stoichiometric oxygen demand. BL and STL heat capacities are based on the correlations presented by Zaman et al. [2]. Experimentally determined heat losses are included in the energy balances for the pilot scale process as described below.
For the gas phase, comparison between model results and experimental data is straightforward. The inorganic smelt phase is, however, dissolved to form GL. In this work, inorganic components in GL are compared to smelt phase predictions. The components are sulfur, carbonate and total inorganic carbon (TIC). TIC is the sum of carbon present as carbonate and hydrogen carbonate. Hydrogen carbonate is not present in the smelt leaving the gasifier but is formed in the quench as discussed below. Sodium and potassium are not compared, since both model and experimental data predicts these to be completely found in GL.
2.Comparison of thermodynamic models
In order to evaluate the importance of the more advanced TEM implemented in FactSage compared to SIMGAS, three OPs each from Jafri et al. (BL) [1] and Furusjö et al. (STL) [3] were modeled using both tools. Here, we only compare the outputs of the models to each other; the agreement with experiments is discussed in the main text.
2.1Method
In FactSage, inorganic melt solution models specially developed for black liquor [4] were used to reflect the importance of inorganic products, due to the very high ash content of spent pulping liquors. The solution phases FTpulp-MeltA, FTpulp-ORTA, FTpulp-NKCA, FTpulp-NKCB and FTpulp-Gsrt of the FactSage FTpulp database were used in addition to the FactPS database in agreement with previous work [5] on TECs of BLG.
In comparison, SIMGAS uses a simpler TEM, including only ideal mixtures of components for both the gas and the inorganic smelt phases. Included components are listed in Table 1 of the main paper.; this selection is based on species for which significant concentrations were found during numerous runs with varying process conditions. The pure component data for gas components and solid carbon are taken from the NIST Chemistry WebBook while data for pure inorganic smelt components are based on Lindberg [4], which is the same source as used in the FactSage FTpulp data (but excluding interaction parameters). In SIMGAS, Gibbs energy minimization is accomplished using an active-set method in order to be able to include both linear and non-linear constraints.
2.2Results
The difference in predicted temperatures between the two models is 6.8±0.2 °C and 3.2±0.7 °C for the BL (Jafri OP3-5) and STL (Furusjö OP 7-9) gasification cases, respectively. Considering that the predicted temperatures are 1100-1200 °C, this is a fairly small difference. The predicted main products are compared in Figure 1, showing only small differences for both gas phase and smelt phase products. The component that shows the largest relative difference is K2S for which the predicted flow is 2-3 times higher for the FactSage model. It is not unexpected that the major difference appears in the composition of the inorganic smelt phase, since the thermodynamic description of this phase is the most important difference between the models.
If describing specific inorganic smelt properties is the aim of the simulation this difference between the models can be significant. Indeed, smelt properties can be important in entrained flow gasification to describe flowability of slag and interactions with containment materials as studied for solid biomass gasification[6, 7]. However, if prediction of overall performance and process efficiency is the major aim then differences for K2S, which represents less than 1% of the smelt phase, are not significant. A separate study (not shown) shows that the effect of non-ideality of the gas phase is approximately 0.7 °C on the predicted temperature.
The conclusion from the comparison between the two thermodynamic models is that the simpler implementation in SIMGAS is sufficient to describe the process for the purposes of this study. Hence, the remainder of the work described in this paper is based on the SIMGAS model. The primary motivation for using this TEM is the ease with which constraints can be implemented.
Figure 1. Predicted main products from the gasification process for gas (a) and smelt (b) phases. Dark bars represent SIMGAS and light bars FactSage for each component. Simulated cases are OPs 3-5 from Jafri et al. [1] and OPs 7-9 from Furusjö et al. [3].
3.References
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2. Zaman AA, Tavares SA, Fricke AL (1996) Studies on the Heat Capacity of Slash Pine Kraft Black Liquors: Effects of Temperature and Solids Concentrations. J Chem Eng Data 41:266–271. doi: 10.1021/je9501245
3. Furusjö E, Stare R, Landälv I, Löwnertz P (2014) Pilot Scale Gasification of Spent Cooking Liquor from Sodium Sulfite Based Delignification. Energy & Fuels 28:7517–7526. doi: 10.1021/ef501753h
4. Lindberg D (2007) Thermochemistry and melting properties of alkali salt mixtures in black liquor conversion processes. Åbo Akademi University
5. Wiinikka H, Johansson A-C, Wennebro J, et al (2015) Evaluation of black liquor gasification intended for synthetic fuel or power production. Fuel Process Technol 139:216–225. doi: 10.1016/j.fuproc.2015.06.050
6. Carlsson P, Ma C, Molinder R, et al (2014) Slag Formation during Oxygen-Blown Entrained-Flow Gasification of Stem Wood. Energy and Fuels 28:6941–6952. doi: 10.1021/ef501496q
7. Ma C, Backman R, Öhman M (2015) Thermochemical Equilibrium Study of Slag Formation during Pressurized Entrained-Flow Gasification of Woody Biomass. Energy and Fuels 29:4399–4406. doi: 10.1021/acs.energyfuels.5b00889
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