Real flowsheets, already solved
Every card is a live, validated simulation — not a mockup. Filter by industry, search by keyword, and open any one straight into your own workspace.
Pumped transfer — where the pump curve meets the system
Water pumped from an atmospheric tank through a control valve and a 150 m line into a tank held at 3 bar — the everyday hydraulic question every plant asks of a centrifugal pump: at what flow does it actually run? The pump is given the way a vendor gives it, as three points off the datasheet curve (45 m at shut-off, 39 m at 2 L/s, 21 m at 4 L/s), not as a flow or an outlet pressure. Design (Steady State). The flow is specified (150 mol/s, about 9.7 m³/h) along with each block's outlet pressure; the solver reports the pump's 1.25 kW shaft duty and the pressure profile down the line. Rating (Run Mode > Pressure Driven). Now nothing specifies the flow. The pump follows its curve, the valve its Cv law, the pipe Darcy-Weisbach friction, and the flow is wherever the pump's head equals what the valve, the line and the 2 bar lift into the receiving tank take away — the operating point on a pump-and-system-curve chart. The valve is sized so that point is the design duty: 150 mol/s, 34 m of head, 4.32 bar at the discharge. Throttle it. Set FCV₁₀₁'s opening to 0.5 on its Input sheet and re-run: the flow falls to 93 mol/s and the discharge pressure rises to 4.98 bar, because a pump delivers more head at less flow and rides back up its curve. That is the behaviour the curve encodes and a fixed pressure rise cannot show.
5 unit ops • NRTL
63 0
View & openTray efficiency: 20 real trays are not 20 stages
Every equilibrium column model quietly assumes each tray reaches equilibrium. Real trays do not, and the gap is not small enough to ignore when you are buying a shell. The same depropaniser is solved twice at 20 trays, a 2.5 reflux ratio and a 50/50 split. As 20 equilibrium stages it puts 99.61% propane overhead. As 20 actual trays on the rate-based model it manages 96.63% — the tower is the same height and the separation is measurably worse, because the model applies an efficiency instead of assuming one. Left to itself it derives that efficiency from the feed's relative volatility and liquid viscosity through the O'Connell correlation; give it a stage_efficiency and it uses yours, and at a punishing Murphree 0.5 the overhead drops further to 95.33%. The practical reading: an equilibrium model sized this column at 20 trays and the tray efficiency spent roughly three points of product purity. Size on equilibrium stages and you will under-build the tower — that is what the efficiency is for. Bounded: this applies a single overall efficiency to every tray, which is a screening treatment. It is not a rigorous rate-based column — for genuine two-film Maxwell-Stefan transfer with per-stage interface composition and energy coupling, use nonequilibrium_distillation instead. O'Connell itself is a correlation fitted to commercial tray data, so it carries that scatter.
4 unit ops • PENG-ROBINSON
174 0
View & openBatch campaign: a batch step inside a continuous flowsheet
Fine chemicals and pharma run batches, but a flowsheet is written in mol/s. These two unit ops reconcile that: they integrate the real batch physics over the batch, then report the products as the RATE that campaign delivers, so a batch step sits inside a continuous flowsheet the way a plant actually runs one. The batch reactor esterifies ethanol with acetic acid to ethyl acetate and water. It reaches 63.2% conversion and stops, and that number is not a specification — it is the equilibrium. The reaction is reversible with an equilibrium constant of 4 (the forward and reverse pre-exponentials are given as 5.0e4 and 1.25e4), and solving K = x(0.10+x)/(0.45-x)^2 = 4 for an equimolar charge carrying 10% water gives an extent of 0.285, which is 63.3% of the acid. Fischer esterification really does plateau there; a model that ran to completion would be describing a different reaction. Turnaround time costs throughput exactly as reaction time does. Two hours of reaction plus one hour of turnaround is a 3-hour cycle and 2922 batches a year, and the reactor is only reacting for 67% of the time it is occupied. That fraction is the number that decides whether you need a second reactor, and it is invisible to a steady-state model. The batch still then takes the crude and recovers the ethyl acetate — the distillate comes off essentially pure, from a pot that started at 28.5 mol% ester — on its own longer 4-hour cycle, so the two units run at genuinely different campaign rates. Bounded: the flowsheet sees each unit's CYCLE AVERAGE, not the transient. A real plant has a surge tank between these two smoothing the pulses, and neither the concentration profile during the batch nor the still's changing overhead composition is visible downstream.
5 unit ops • NRTL
173 0
View & openYield reactor: modelling a reaction nobody has kinetics for
Pyrolysis, gasification and cracking share a problem: the feed is not a set of molecules with a stoichiometry, it is biomass or coal or a heavy residue, and there is no reaction network to write down. What you have instead is a measured product slate from a pilot run. The yield reactor (Aspen's RYield) is the unit op for exactly that case — you give it the slate, it gives you a stream. Here 100 mol/s of feed goes to 823 K and comes out as 112 mol/s of gas at a specified mass yield (31% CO₂, 28% water, 22% methane, 14% ethane, 5% hydrogen), for a 3.66 MW heating duty. The mole count rises and the mass does not: the yields are normalised so the product mass equals the feed mass exactly, which is the entire contract of a yield reactor and the one thing it will not let you get wrong. Be honest about what this is. It has no kinetics, no equilibrium, no residence time and no temperature dependence of the slate — change the outlet temperature and the products do not shift, only the duty does. It is a way to carry a measured yield through a heat and material balance so the rest of the flowsheet is right; it predicts nothing about the reaction itself. If you have a rate law, use kinetic_reactor; if the system reaches equilibrium, use gibbs_reactor. Reach for this one when you have neither, which for solid-feed conversion is most of the time.
3 unit ops • PENG-ROBINSON
171 0
View & openShortcut column sizing (Fenske-Underwood-Gilliland)
Before anyone builds a rigorous column they size it in about a second, and this is that step — Aspen's DSTWU, the calculation every distillation design starts from. State the two keys and how well you want them separated, and the shortcut answers the four questions that set the capital cost. For a depropaniser feed (30% propane, 40% n-butane, 30% n-pentane) at 10 bar recovering 98% of the propane overhead and 98% of the n-butane in the bottoms: Fenske gives 8.34 minimum stages — the count at total reflux, where you spend infinite energy to buy the fewest trays. Underwood gives a minimum reflux of 1.54 — the reflux at infinite stages, the opposite corner. Neither is buildable; the real column lives between them, and Gilliland interpolates: at 1.3x the minimum reflux you need 17.7 theoretical stages, with Kirkbride putting the feed on stage 7.85. That is a little over twice the minimum stage count for 1.3 times the minimum reflux, which is the trade every column is designed on. The energy follows: 1.67 MW reboiler against a 1.35 MW condenser. Bounded, and this is the important part. These are screening numbers, not a design. The shortcut assumes constant relative volatility and constant molal overflow — it cannot see a pinch, a tangent, an azeotrope, or a temperature profile. Its job is to give a rigorous MESH column a starting point that converges, not to replace it. Run the distillation unit op on 18 stages with a feed at 8 to check it.
4 unit ops • PENG-ROBINSON
172 0
View & openIonic liquid breaks the ethanol-water azeotrope
The ethanol-water azeotrope is a wall: at 89.4 mol% ethanol the vapour and the liquid have the SAME composition, relative volatility is 1, and no number of trays gets you past it. This flowsheet walks through it by adding an involatile ionic liquid, [EMIM][BF₄], which binds water preferentially and pulls the two apart. The numbers are the demonstration. NRTL puts the relative volatility at the azeotrope at 0.996 — that is the wall, computed, not asserted. Add the IL and it climbs to 2.10 at 10 mol%, 2.98 at 20%, and 3.66 at 30%. Flash the IL-laden mixture here and the vapour comes off at an ethanol:water ratio of 22.6 against the azeotrope's 8.43 — decisively across. Run the same feed with no IL and there is nothing to separate: the mixture goes straight from all-liquid to all-vapour with no useful two-phase band, which is precisely what an azeotrope means. Why an IL rather than the usual glycol entrainer (see 'anhydrous-ethanol-extractive-distillation'): an ionic liquid has effectively no vapour pressure, so it never contaminates the distillate and it regenerates by flashing rather than by a second column. Bounded, and this is why the example is a flash and not a column: the IL package models the phase behaviour — which is what decides whether an entrainer works — but omits the IL's own pure-component enthalpy, so a rigorous column energy balance is a follow-up. The separation shown here needs only the phase equilibrium. The IL also carries no molar mass in the databank, so mass-basis readouts show a dash; the mole-basis flash is exact.
6 unit ops • IONIC-LIQUID
195 0
View & openAcetone recovery — batch still vs. continuous column
A spent-solvent stream (60 mol% acetone, 40% water) recovered two ways, so the Analysis panel's Batch distillation (Rayleigh) tool has something to be compared against — it had no curated example, and a batch calculation only means anything next to the continuous alternative. The flowsheet is the continuous answer: a shortcut column recovering 98% of the acetone overhead and rejecting 98% of the water, giving 59.8 mol/s of distillate at high purity for 4.4 MW of reboiler duty. The batch answer (Analysis > Batch distillation). Charge 100 mol of the same mixture to a still and boil it down until the pot falls to 10 mol% acetone: you distil 67.4 mol and the average distillate is only 84.2 mol% acetone. That is the Rayleigh result and it is the whole lesson — a simple batch still has one theoretical stage, so the vapour it makes is enriched but never pure, and the composition drifts the entire time. Reaching column-grade purity in a batch needs reflux and cuts (a batch rectifier), or a second pass. Why the comparison is the point. Batch equipment is cheap, flexible, and right for small or campaign volumes; the column is right when the duty is continuous and purity matters. Having both numbers on the same feed is what makes that a decision rather than a preference.
4 unit ops • NRTL
239 0
View & openMethanol column sized by design spec
A methanol-water column sized by design specification rather than by trial and error — Aspen's Design Spec / SimCentral's Adjust, and a capability that had no curated example despite being solver-native. The column is specified the way a shortcut (FUG) column is: light and heavy keys with their recoveries (98% of the methanol overhead, 2% of the water). What is not specified is the reflux ratio. Instead the flowsheet carries a design_specs entry — vary COL.reflux_ratio until the metric shortcut_n_stages equals 14 — and the solver root-finds it, re-solving the whole flowsheet each trial. Open the Design Spec panel to see it, or the solved result: reflux settles at 1.06 for exactly 14.00 stages. Why that is the interesting question. Reflux and stages trade off against each other, and against energy. Left at the initial 1.6, this separation needs only 10.8 stages but 4.35 MW of reboiler duty; pulled down to 1.06 it needs 14 stages and just 3.49 MW — a 20% energy saving bought with three more trays. A design spec lets you state the column you can afford to build and have the solver tell you how to run it, instead of guessing a reflux ratio and reading off whatever height falls out. The Sensitivity tab comes preloaded with that trade-off as a sweep (reflux ratio against reboiler duty) so the curve behind the single design-spec answer is one click away.
4 unit ops • NRTL
239 1
View & openBlending header — one model, three solve modes
A solvent blending and distribution header, built to demonstrate the capability AVEVA sells SimCentral on and Aspen splits across separate products: one flowsheet that solves steady-state, pressure-driven, and dynamically — no re-modelling between them. A water-rich supply and an ethanol-rich additive each pass a control valve into a common header, down a trunk line, and split at a tee to two consumers. 1. Steady-state (Solver menu > Steady-state). You specify the flows (100 and 25 mol/s) and each valve's outlet pressure; the solver reports the blend — 22 mol% ethanol — and the pressure profile down the header. This is the design question: what pressures do I need to move this flow? 2. Pressure-driven (Solver menu > Pressure-driven). Now the feeds specify pressure (9 bar) instead of flow, the consumers are pinned at 2 bar, and each valve's cv becomes a resistance law. Flow is a solved unknown, closed by mass balance against every element's resistance. This is the rating question: what flow do I actually get? The valves here are sized for the design duty, so the answer comes back at the same 125 mol/s and the same 6.0 bar header — the two modes agree because they describe one plant. The tee is where it gets interesting: in pressure-driven mode the 60/40 split is not read from split_fractions (that parameter is ignored) — it is solved from the two consumer valves' Cv against the downstream pressures. Halve HCV_UNIT2's cv and the split moves and the total flow drops; do the same in steady-state mode and nothing budges, because there the split is something you asserted rather than something the network decided. 3. Dynamic (Solver menu > Dynamic). Initialized from the steady-state solution, step the additive feed to 98 mol% ethanol and watch the blend move: the header responds first, then each consumer lags it by its own holdup, all settling at 23.6 mol% — the value a hand mass balance gives, ((100x0.05) + (25x0.98))/125.
11 unit ops • NRTL
239 2
View & openEDC pyrolysis furnace: 1,2-dichloroethane to vinyl chloride monomer
The balanced-process route to PVC's monomer: 1,2-dichloroethane (EDC) thermally cracks in a fired furnace to vinyl chloride (VCM) + HCl (C₂H₄Cl₂ → C₂H₃Cl + HCl, exactly mass-balanced), the HCl is stripped overhead in a distillation column (recovered pure — it's recycled to the plant's oxychlorination unit in a real balanced VCM process, not modeled here), and a second column separates VCM product from unconverted EDC (also recycled in a real plant). Unlike Linde's proprietary furnace internals (tube metallurgy, coking-cycle length, radiant-zone geometry), EDC pyrolysis chemistry itself is standard, widely published petrochemical engineering (Kirk-Othmer and equivalent references), not proprietary Linde IP — so this is built directly from the real reaction and typical single-pass conversion (~50-55%, a widely cited industrial range for this process, not a specific plant's exact figure), the same 'illustrative, representative, not literature-pinned to one source' posture used throughout MaximaLabs's cracking-furnace examples.
8 unit ops • PENG-ROBINSON
292 0
View & openOxidative dehydrogenation of ethane to ethylene + acetic acid (EDHOX concept)
Linde's EDHOX technology co-produces ethylene AND acetic acid from ethane and oxygen in one catalytic step (a proprietary mixed-metal catalyst in a multi-tubular salt-cooled reactor), with combined ethylene+acetic-acid selectivity published above 93% and the CO₂ by-product of over-oxidation recovered pure (no nitrogen dilution, since the oxidant is pure O₂, not air). Modeled here as three chained fixed-conversion reactors on real, exactly mass-balanced reactions: the main dehydrogenation (C₂H₆ + 0.5 O₂ → C₂H₄ + H₂O), the acetic-acid co-production path (C₂H₆ + 1.5 O₂ → CH₃COOH + H₂O), and a minor full-oxidation loss path (C₂H₆ + 3.5 O₂ → 2 CO₂ + 3 H₂O) accounting for the un-selective balance — the same 'illustrative, not fitted' posture as the naphtha cracker's furnace, since Linde's catalyst kinetics are proprietary and not published. The per-reaction conversions here are tuned to reproduce the one real published number (>93% combined selectivity: this flowsheet computes ~96%), not an assumed per-pass ethane conversion, which Linde doesn't publish.
10 unit ops • PENG-ROBINSON
292 0
View & openQuench-converter ammonia synthesis loop
The multi-bed quench converter BASF first installed in 1942, which displaced the tube-cooled converters before it, running inside a complete 200 bar synthesis loop. Ammonia synthesis is equilibrium-limited and exothermic, so the catalyst wants to run hot for rate and cool for equilibrium, and no single adiabatic bed can do both. The quench converter's answer is mechanical: split the loop gas, send 55% through the interchanger into bed 1 and inject the other 45% as cold shots between the beds, so each bed heats up adiabatically and each quench pulls it back down. What it computes: the sawtooth, 673 → 766 K in bed 1, quenched to 689 K and out at 760 K, quenched to 712 K and out at 766 K; 12.5 mol% ammonia at the converter exit, in the 12-18% band real quench converters deliver, at 20.9% nitrogen conversion per pass; 44.6 mol/s of 97.8% liquid ammonia off the 250 K separator; and argon and methane held at 3.3% and 4.2% in the loop by a 3% purge, with the recycle closed by a Wegstein tear. The acid test is the counterfactual: one adiabatic bed producing that same 12.51 mol% ammonia exit ends at 834.7 K — 562 C, past the sintering limit of promoted-magnetite catalyst — while the three-bed quench converter delivers the identical duty with a 766 K (493 C) peak.
15 unit ops • PENG-ROBINSON
180 0
View & openJet loop reactor: nitrobenzene to aniline
A circulating loop reactor of the kind BASF and Buss built for strongly exothermic three-phase hydrogenations — here nitrobenzene to aniline (C₆H₅NO₂ + 3 H₂ → C₆H₅NH₂ + 2 H₂O, -443 kJ/mol). Fresh nitrobenzene and hydrogen are injected into a large circulating liquid stream; the reactor itself runs adiabatic and the reaction heat is taken out by a heat exchanger in the EXTERNAL loop, which is the whole architectural point — the exchanger area is set independently of the reactor, and the circulation rate, not a cooling jacket, is what holds the temperature. A gas separator vents the excess hydrogen, 5% of the liquid is taken off as crude aniline, and the rest is pumped back. What it computes: a 19:1 recycle-to-product ratio holds the adiabatic reactor to 429.7 K in / 494.3 K out, a 64.6 K rise, with the loop cooler removing 4.66 MW; overall nitrobenzene conversion is 99.98% (99.70% per pass) and the crude leaves at 34.7% aniline / 64.9% water. Counting the vent as well as the crude, one aniline is made per two water, which is the stoichiometric mass-balance check on the whole loop — the crude by itself runs slightly water-lean because the vent carries water vapour off with the excess hydrogen. The acid test is what happens without the loop: the same feed in one adiabatic once-through reactor is already at 627 K by 30% conversion and runs off the top of the model's temperature range well before full conversion. That gap is the reason the architecture exists.
10 unit ops • PENG-ROBINSON
179 0
View & openAcetic acid recovery — NRTL with vapor-phase dimerization
Concentrating dilute aqueous acetic acid (30 mol%) by distillation, modeled with the NRTL activity package and chemical-theory vapor-phase association (acetic acid dimerizes, 2 A → A2, in the vapor). This is the physics a cubic equation of state gets wrong: the pinned NRTL acetic-acid/water binary (DECHEMA) plus the cited dimerization constants (Nagy et al., Molecules 2020) give the real bubble curve and latent heat, so the close-boiling acid/water pair actually separates and converges where Peng-Robinson does not.
4 unit ops • NRTL-ASSOCIATE
327 0
View & openEthanol-water heat-and-flash (Van Laar)
The simplest separation there is, on the simplest activity model that still gets a non-ideal pair right: a 40 mol% ethanol-water stream heated to 356 K at 1 atm and flashed adiabatically, on the two-constant Van Laar equation — the oldest of the family, temperature-independent, cheap to fit from a single azeotrope point. What it computes: 55% of the feed vaporizes at 356 K, the vapour at 54 mol% ethanol against 22% left in the liquid, on 2.75 MW of preheat. Ethanol-water is the one pair the package carries regressed constants for; other pairs fall back to UNIFAC.
5 unit ops • VAN-LAAR
92 0
View & openBeer column: 8% ethanol to 53% (Wilson)
The first column of a distillery on the Wilson equation — the oldest local-composition model, fine for a fully miscible pair like ethanol-water and unable by construction to represent a liquid-liquid split, which is why it is offered for this pair and not as a default. An 8 mol% fermenter beer, 12 stages with the feed near the top (stage 3) as a beer column is run, reflux 1, 15% distillate: 53 mol% ethanol overhead and a stillage bottoms with no ethanol left, 1.37 MW on the reboiler for 100 mol/s of beer. Ethanol-water is the one binary the package carries regressed Wilson parameters for; everything else falls back to UNIFAC.
4 unit ops • WILSON
93 0
View & openEthanol-water rectifier (UNIQUAC)
The ethanol-water rectifier on UNIQUAC — Abrams and Prausnitz's local-composition model with its surface and volume parameters, here on the one binary the package carries regressed parameters for (every other pair falls back to UNIFAC, which the package description says). 30 mol% feed, 20 stages, reflux 2.5, 40% distillate: 75 mol% ethanol overhead and ethanol-free water in the bottoms at 5.7 MW. The cut is kept at 0.40 on purpose: at 0.32 the spec asked for a 94% distillate past the 89.4% azeotrope and the column reported exactly that, and at 0.36 it pinched a hair short of it (balance 5e-3). Compare with the same column on NRTL and Wilson to see how much three regressed models agree on one well-measured pair — a few tenths of a percent in the distillate.
4 unit ops • UNIQUAC
93 0
View & openToluene removal from 1-butanol solvent (modified UNIFAC, Dortmund)
A recovered 1-butanol solvent contaminated with 5 mol% toluene, redistilled to send the toluene overhead: 24 stages, reflux 3, an 8% distillate. There is no fitted toluene/1-butanol binary, so this runs on modified UNIFAC (Dortmund) — the variant that cuts infinite-dilution error from 42% to 16% in this tree, and a dilute contaminant in a solvent is exactly the infinite-dilution regime. Dortmund puts the toluene/butanol azeotrope at 67 mol% toluene (original UNIFAC: 70%), which is why the overhead comes out at 43% toluene rather than pure, and the bottoms at 1.7% toluene — two-thirds of the contaminant removed per pass. The 3-point azeotrope difference between the two variants is the size of the model choice on this pair.
4 unit ops • UNIFAC-DORTMUND
92 0
View & openn-Hexane / 1-butanol column with no fitted binary (UNIFAC)
A binary the databank has no regressed NRTL parameters for, so the column runs on predictive original UNIFAC from the two molecules' groups alone: 50/50 n-hexane and 1-butanol, 20 stages, reflux 2, 45% distillate. UNIFAC predicts the minimum-boiling azeotrope at 97 mol% hexane and 341.7 K, and the column lands its distillate exactly there (96.7%) with an 88% butanol bottoms — the azeotrope, not the stage count, is what caps the overhead purity, and that is a prediction from group contributions with no data on this pair behind it. Read it as a screening result: original UNIFAC's mean error against fitted binaries in this tree is 0.08 in ln gamma, and an alkane/alcohol pair is in its well-behaved range.
4 unit ops • UNIFAC
93 0
View & openAcetone-water column at 3 atm (PR + Wong-Sandler)
An acetone recovery column at 3 atm on Peng-Robinson with the Wong-Sandler mixing rule — the one Gᴱ rule that keeps the quadratic composition dependence of the second virial coefficient, so it stays thermodynamically consistent at high pressure. 20 mol% acetone feed, 14 stages, reflux 2, 25% distillate: 80 mol% acetone overhead and acetone-free water in the bottoms, 2.65 MW on the reboiler. The cut is set at 0.25 rather than the feed's exact acetone fraction of 0.20, which pinned the column on a razor edge (temperature residual 2e-3, balance 3%); at 0.25 it closes to zero. Bound: the rule is driven by the low-pressure acetone/water NRTL binary and, as records, NeqSim's Wong-Sandler kij is deliberately not wired in, so the second-virial cross term uses the van der Waals kij.
4 unit ops • PR-WONG-SANDLER
94 0
View & openMethanol-water column at 10 atm (PR + Huron-Vidal)
A methanol-water column at 10 atm on Peng-Robinson with the original Huron-Vidal (1979) infinite-pressure Gᴱ mixing rule, driven by the fitted methanol/water NRTL binary. 30 mol% feed, 16 stages, reflux 2, 30% distillate: 96.2 mol% methanol overhead, 1.6% in the bottoms, 3.4 MW on the reboiler at 452 K. The point is the pressure: at 10 atm a gamma-phi activity model is at the edge of its ~10 bar envelope, and a cubic with a Gᴱ rule carries no such ceiling. Bound, as the package's own description says: Huron-Vidal wants activity parameters regressed at infinite pressure, and these are the low-pressure DECHEMA ones, so treat the split as the rule's reading of that binary rather than a validated high-pressure number (the SRK + NeqSim card is the one with parameters fitted for the rule).
4 unit ops • PR-HURON-VIDAL
93 0
View & openPolystyrene devolatilization, two solvents (Sanchez-Lacombe)
A solution-polymerized polystyrene (100 kg/mol) carried in toluene and benzene, flashed at 480 K and 0.2 atm to strip the solvent — the multi-solvent case the Flory-Huggins package cannot take (one polymer, one solvent) and the reason the Sanchez-Lacombe lattice-fluid equation of state is in the tree. Characteristic parameters are the published ones: polystyrene T 735 K, P 357 MPa, rho 1105 kg/m3; toluene 543 K, 402 MPa, 966 kg/m3; benzene 523 K, 444 MPa, 994 kg/m3 (Sanchez & Lacombe, J. Phys. Chem. 80 (1976) 2352 and Macromolecules 11 (1978) 1145), with the segment number r from M P/(R T rho). What it computes: 94% of the feed moles leave overhead as solvent vapour and the melt retains 0.16 wt% residual solvent — read it on mass, not moles, since a 100 kg/mol chain is one mole against 0.09 kg/mol solvents. No binary k_ij (none is published for this pair); a single equilibrium stage, no diffusion-limited devolatilization kinetics.
4 unit ops • SANCHEZ-LACOMBE
94 0
View & openEthanol-water column at 6 atm (SRK + MHV1)
The ethanol-water column everyone knows, run at 6 atm on the Gᴱ-mixing-rule package: Soave-Redlich-Kwong whose mixture energy parameter comes from NRTL's excess Gibbs energy through the MHV1 rule (Michelsen 1990) instead of a kij — the package built for polar, hydrogen-bonding mixtures at pressure, where a γ-φ activity model is past its ~10 bar ceiling and a plain cubic with van der Waals mixing has no hydrogen bonding at all. A 10 mol% ethanol feed through 16 stages at a reflux ratio of 2.5 gives a 66.7 mol% distillate and an ethanol-free bottoms, 2.6 MW on the reboiler. Read it against NRTL, which at 6 atm is still inside its own range: the two packages put the pressure-shifted azeotrope in different places — 0.88 mol fraction ethanol here, 0.84 on NRTL at 6 atm (0.97 against 0.89 at 1 atm) — so this card shows where the MHV1 rule sits, not a number to prefer over the fitted binary at low pressure. The cut is kept well short of the azeotrope on purpose: at a distillate ratio of 0.12 the profile pinched against it and the column reported exactly that. Each solve is about a minute, because every stage K-value is a Gᴱ mixing-rule evaluation.
4 unit ops • SRK-MHV1
92 0
View & openNitric acid: the Ostwald process (NH3 -> NO -> NO2 -> HNO3)
The three-stage industrial route to nitric acid, each stage a genuinely different reaction regime. First a catalytic burner oxidizes ammonia in air over a Pt-Rh gauze (4 NH₃ + 5 O₂ → 4 NO + 6 H₂O) — this is a KINETICALLY selective step, not an equilibrium one: thermodynamics actually favours N₂, and only the short contact time on the hot catalyst steers it to nitric oxide, so it's modeled as a fixed-conversion reactor (a Gibbs-minimization reactor here would wrongly predict N₂). The gas is cooled and the nitric oxide is oxidized to nitrogen dioxide (2 NO + O₂ → 2 NO₂), an exothermic step favoured by the low temperature. Finally the NO₂ is absorbed in water to form nitric acid (3 NO₂ + H₂O → 2 HNO₃ + NO) — modeled as NO₂ + water capture into the acid liquor. Cited heats of reaction (NH₃ oxidation -226.3 kJ/mol NH₃; NO oxidation -57.0 kJ/mol NO).
9 unit ops • PENG-ROBINSON
268 0
View & openSulfuric acid: catalytic SO2 oxidation (contact process)
The heart of a sulfuric-acid plant: roaster/burner gas (SO₂ in excess air) is oxidized to SO₃ over a vanadium-pentoxide catalyst bed, then the SO₃ is absorbed to make acid. The converter is a Gibbs-minimization reactor (Aspen RGibbs equivalent) that finds the SO₂ ⇌ SO₃ equilibrium from first-principles formation energies — no assumed conversion — so it captures the real design tension of the contact process: the reaction SO₂ + ½O2 → SO₃ is exothermic, so a lower bed temperature drives the equilibrium toward SO₃ (higher conversion) while a higher temperature is needed for catalyst activity. At 700 K this converts ~99% of the SO₂; raise the sensitivity temperature and watch the equilibrium conversion fall — the reason real plants use multiple catalyst beds with interstage cooling. The SO₃-rich gas then goes to an absorption tower (modeled as SO₃ capture into the acid product, leaving a spent tail gas of N₂ + excess O₂). Formation-energy source: CRC Handbook / NIST-JANAF (ΔHf°/ΔGf° for SO₂ and SO₃).
5 unit ops • PENG-ROBINSON
284 0
View & openIntegrated styrene chain: benzene to styrene monomer
The full two-step styrene chain connected as one flowsheet rather than two standalone process snippets: benzene alkylated with ethylene to ethylbenzene, purified in a recovery column, then fed directly to the dehydrogenation reactor that makes styrene monomer. The unreacted ethylbenzene the dehydrogenation column recovers is reported as its own product stream rather than recycled back onto the alkylation feed (an honest simplification — closing that loop needs a torn-recycle edge back into ALKCOL, which the plant does but this flowsheet does not).
12 unit ops • PENG-ROBINSON
326 0
View & openStyrene monomer via ethylbenzene dehydrogenation
Ethylbenzene dehydrogenates over an equilibrium/conversion reactor to styrene monomer plus hydrogen — the endothermic reaction that supplies the world's polystyrene/SBR-rubber feedstock. A downstream column recovers unreacted ethylbenzene overhead for recycle while polymer-grade styrene leaves the bottoms. Vasudevan design, Ind. Eng. Chem. Res. 2009, 48, 10941 (Figure 15.1).
7 unit ops • PENG-ROBINSON
326 0
View & openEthanolamines plant: MEA / DEA / TEA from ethylene oxide + ammonia
The nitrogen analogue of the ethylene-glycol chain, and a major gas-treating-solvent process in its own right. Ethylene oxide reacts with ammonia through the same kind of CONSECUTIVE addition chain the glycols follow — EO + NH₃ → monoethanolamine (MEA), EO + MEA → diethanolamine (DEA), EO + DEA → triethanolamine (TEA), all atom-balanced and keyed on the shrinking EO pool. A large ammonia excess pushes selectivity toward MEA (the ~80/13/5 MEA/DEA/TEA slate a high NH₃:EO ratio makes, amine-side mirror of how a high water:EO ratio favours MEG). The separation strips the excess ammonia (recovered for recycle) and the reaction water, then splits the amines by boiling point (MEA 170 C < DEA 269 C < TEA 335 C) into ~99.8% MEA, high-purity DEA, and a TEA bottoms cut. HONEST SCOPE: the reaction chemistry is rigorous stoichiometry (real atom balances, realistic selectivity from the fixed conversions), and the amines are characterized as pseudo-components (Tc/Pc/omega from the open-data databank — MEA/DEA/TEA have no CoolProp entry) so they flash under Peng-Robinson. The purification uses spec-based component-split separators (the Aspen 'Sep'-block technique), NOT rigorous vacuum columns — the ethanolamines are wide-boiling with narrow adjacent-amine relative volatilities, the same wide-boiling-MESH limit measured for the glycol columns. The recovered ammonia is shown as an open recycle stream (an honest simplification, like the parent EO example).
17 unit ops • PENG-ROBINSON
325 2
View & openEthylene glycol plant: closed water loop + multi-effect evaporator dehydration
The water-integrated evolution of the fiber-grade MEG plant. Two changes turn the once-through EO/glycol chain into a real, water-economical process: (1) the recovered process water is RECYCLED back to the hydration reactors through a purge splitter. A pure recycle is inventory-singular (reactors are keyed on the shrinking EO pool, so water consumption is fixed regardless of how much water circulates), so an 8% purge pins the loop and makes it well-posed, cutting fresh water makeup from 10 to ~1.75 mol/s (a >80% reduction). (2) The bulk dehydration is done by a genuine TWO-EFFECT EVAPORATOR TRAIN (forward-feed, real steam economy: effect 1's low-temperature vapor is the heating steam for effect 2) doing rigorous (P,H)-flash water removal, rather than a single spec-based split. The recycle converges through the solver's Wegstein tearing (~15 outer passes) and MEG still comes out fiber-grade (>=99.9%). HONEST SCOPE: the reaction chemistry is rigorous stoichiometry (real atom balances, ~90/9/1 selectivity) and the evaporators are real energy-balanced flash effects — but flash evaporation CANNOT reach glycol dryness without slipping glycol into the overhead vapor (MEG has a real vapor pressure at 90 C), so the evaporators run cool and only pre-concentrate. An overhead knockout returns the ~1-2% slipped glycol to the product (no yield loss), and a final vacuum-refining polish (still a spec-based split, representing the refluxed dehydration column a flash cannot replicate) removes the last water. This mirrors a real MEG plant's multi-effect-evaporator + vacuum-refining dehydration section. The product columns remain spec-based (a converged 99.9% fiber-grade MEG column is not tractable under Peng-Robinson here — the MEG/DEG relative volatility is too narrow for the wide-boiling MESH path).
32 unit ops • PENG-ROBINSON
330 4
View & openEthylene glycol plant: fiber-grade MEG + DEG/TEG byproducts
Extends the EO/glycol chain all the way to separated products. Ethylene + O₂ make ethylene oxide over a silver catalyst (with the competing total-combustion side reaction), EO condenses out, and then hydrates through the real CONSECUTIVE glycol reactions — EO + H₂O → MEG, EO + MEG → DEG, EO + DEG → TEG (all atom-balanced, keyed on the shrinking EO pool) — giving the industrial ~90/9/1 mono-/di-/tri-ethylene-glycol selectivity that a high water:EO ratio produces. The purification train recovers the water for recycle and splits the glycols into fiber-grade MEG (>=99.9%), DEG, and TEG products. HONEST SCOPE: the reaction chemistry is rigorous stoichiometry (real atom balances, realistic selectivity set by the fixed conversions); the PURIFICATION is modelled with spec-based component-split separators to the known product purities (the Aspen 'Sep'-block technique for a well-understood separation section), NOT rigorous vacuum distillation columns — a converged 99.9%-fiber-grade MEG column is not tractable in this solver under Peng-Robinson (the MEG/DEG relative volatility is too narrow for the wide-boiling MESH path; a real plant uses large multi-effect evaporators + vacuum columns). The water-recycle stream is left open (not looped back) — an honest simplification, like the parent EO example's ethylene recycle.
23 unit ops • PENG-ROBINSON
329 0
View & openEthylene oxide synthesis through to glycol
The full EO/MEG chain in one flowsheet, upstream of the standalone hydration example: ethylene and oxygen react over a silver catalyst to ethylene oxide (low per-pass ethylene conversion, ~10%, is realistic — high conversion pushes the competing total-combustion side reaction, which this model represents as a second reactor consuming a fixed share of the same ethylene at ~80% EO selectivity). A cooled flash condenses EO (and reaction water) from the unreacted ethylene/oxygen/CO₂ vented for combustion-side purge; the condensed EO then hydrates with fresh water to MEG exactly as in the standalone hydration example. No ethylene/O₂ recycle loop (an honest simplification — real plants recycle unreacted ethylene at high ratio).
17 unit ops • PENG-ROBINSON
326 1
View & openEthylene oxide hydration to mono-ethylene glycol (MEG)
Ethylene oxide reacts uncatalyzed with excess water at 200 C to mono-ethylene glycol (MEG, antifreeze/PET feedstock); the excess-water dilution that suppresses the over-reaction to di-/tri-ethylene glycol byproducts is reflected in the large water excess on the feed, and a downstream column concentrates MEG in the bottoms while excess water leaves overhead for recycle. Kinetics basis: Ind. Eng. Chem. Res. 2009, 48, 10840.
10 unit ops • PENG-ROBINSON
325 0
View & openMethanol synthesis (syngas loop)
Low-pressure syngas-to-methanol loop (CO + 2 H₂ <-> CH₃OH, equilibrium reactor at 80 bar/510 K). The high-pressure separator's crude liquid still carries several mol% dissolved CO₂/H₂/CO from the reactor loop, so a letdown valve plus a second flash (mirroring the HDA casebook's own SEP-LETDOWN-SEP2 degassing pattern) strips the bulk of it before the atmospheric column — dissolved permanent gas at that scale can otherwise mislead a bubble-point search onto a spurious low-temperature root nowhere near methanol/water's real ~340 K bubble point, so keeping the column's own feed genuinely light on non-condensables is the honest fix, not a thermo-solver workaround. From the ChemSep casebook (Methanol_iecr49p6150).
11 unit ops • PENG-ROBINSON
325 0
View & openDME synthesis (methanol dehydration)
Catalytic methanol dehydration (2 CH₃OH → (CH₃)2O + H₂O, 82% conversion) followed by two atmospheric-pressure columns: DMECOL rejects dimethyl ether overhead (~88% pure — the ternary's relative volatility at this pressure caps a single column's overhead purity there) via the inside-out method (a genuinely wide-boiling ternary — DME boils at -24 C, water at 100 C), then MEOHCOL recovers unconverted methanol (99.9% pure) from the water byproduct (99.3% pure). From the ChemSep casebook (DME_ie101583j).
8 unit ops • PENG-ROBINSON
330 0
View & openDilute acetic acid recovery by extractive distillation
Recovering acetic acid from a dilute aqueous stream (e.g. cellulose acetate or terephthalic acid plant wastewater) using diisobutyl ketone (DIBK) as an extractive entrainer. Because acetic acid and water form no azeotrope but have a relative volatility that flattens toward 1 as the mixture concentrates, straight distillation needs an impractically tall column and heavy reboil duty; the entrainer raises water's relative volatility so the extractive column takes water overhead in far fewer stages, and a second column then strips the entrainer from the acid bottoms for recycle. Modeled with NRTL, and the property method is the whole point here rather than a detail: an entrainer works by changing liquid non-ideality, and a cubic equation of state with van der Waals mixing has no term for it — no binary interaction parameter ships for acetic-acid/water or DIBK/water, so Peng-Robinson runs this flowsheet at kij = 0 and represents none of the effect the column depends on. It does not converge either. NRTL carries the DECHEMA-fitted acetic-acid/water binary and reaches the DIBK pairs through UNIFAC.
7 unit ops • NRTL
329 1
View & openAcetic acid via methanol carbonylation (Cativa process)
Methanol carbonylated with CO to acetic acid over the iridium-based Cativa catalyst (BP Chemicals), which runs at higher selectivity and lower water content than the older Monsanto rhodium process. The homogeneous catalyst itself isn't a flowing component in this model — only the carbonylation stoichiometry and downstream light-ends recovery are represented. Unreacted CO is flashed off and a column strips residual methanol from the acetic acid product.
9 unit ops • PENG-ROBINSON
332 0
View & openSite water reuse: RO + EDI polish + MD to near-ZLD
The full hybrid a large chemical site builds when it is reusing its own effluent rather than buying river water: reverse osmosis does the bulk desalting, electrodeionisation polishes the permeate to boiler-feed quality, and the RO concentrate is warmed by waste heat and pushed to salt by membrane distillation and a crystalliser. It is the companion to industrial-water-reuse-ro-md-zld, which stops at reuse-grade water; the delta here is the polishing leg, and it is a genuinely different technology rather than a tighter membrane. Why EDI and not more RO. RO is pressure-driven and pushes water, so its energy scales with the volume produced. Electrodialysis drags ions out under a field and the water never crosses, so its energy scales with the salt removed — which is why it is hopeless on seawater and excellent on an RO permeate that is already 99% desalted. This model shows that rather than asserting it: at the same stack and current the moles removed are fixed by Faraday's law regardless of how salty the feed is. The reason the polisher is resin-filled is the interesting part. Plain electrodialysis has a limiting current density proportional to the diluate concentration, so as the water approaches pure the boundary layer runs out of ions to deliver and the salt flux plateaus — ED cannot reach ultrapure at any voltage, and the voltage runs away as the water stops conducting. Packing the diluate compartment with ion-exchange resin fixes both: the resin carries the current, and the water splitting that occurs above the classical limiting current — a fault in a solution-filled stack, causing a pH excursion and scale — is exactly what regenerates the resin in place. That is why an EDI polisher needs no chemical regeneration, and why the same operating point means opposite things in the two stacks.
11 unit ops • PENG-ROBINSON
169 0
View & openIndustrial water reuse: RO + membrane distillation to near-ZLD
A chemical-site water-reuse train of the kind large industrial water users build when the intake is salinising and freshwater is getting scarce: reverse osmosis makes reuse-grade water, the RO concentrate is warmed by low-grade waste heat and pushed further by membrane distillation, and the MD concentrate is crystallised to salt. Overall water recovery comes out at 94.9%. The point of the flowsheet is why MD is there at all. RO is pressure-driven, so its ceiling is osmotic pressure — and this feed hits that wall hard: at 25 bar the same 75% recovery is infeasible, and the model says so with a number rather than a shrug (osmotic pressure 27.2 bar against a 25 bar feed, a structured SPEC_THERMODYNAMICALLY_IMPOSSIBLE error, which is why the feed here runs at 45 bar). Membrane distillation is driven by a vapour-pressure difference instead, so only vapour crosses and there is no osmotic ceiling at all: it takes the RO concentrate from 2.0 mol% to 9.7 mol% salt, recovering 81% of the water RO had to leave behind, and its distillate is solvent-only. The waste-heat coupling is not decoration. At the RO concentrate's own 25 C the MD flux is about 1 kg/m2/h and the module is pointless; warmed to 60 C it runs at 21.5 kg/m2/h, inside the 10-50 kg/m2/h band real DCMD modules achieve. That is the entire commercial case for MD — it is a heat-driven process, so it is only cheap where low-grade heat is already being vented.
9 unit ops • PENG-ROBINSON
170 0
View & openPhosphoric acid concentration: microwave-then-steam hybrid (29 → 42 → 54% P2O5)
The transition path the concept proposes: microwaves take the acid from 29 to 42% P₂O₅ in a first vacuum flash (29.1 MW absorbed), a pump returns the 42% acid to atmospheric pressure, and the existing steam exchanger and second flash finish it to 54.0 wt% — 23.4 t/h of steam (0.59 t/t) plus 1,120 kWh of electricity per tonne of P₂O₅ at 65% magnetron efficiency, the same 64.0 t/h of water removed in total. The split is not free: two-thirds of the latent duty moved from a co-product steam to purchased electricity, and nothing about the second stage got easier. What a hybrid does buy is a pilot that can be run beside the existing evaporator without betting the line on it, and the freedom to put the microwave stage where it has a case — the viscous, fouling end — rather than on the dilute feed. Compare phosphoric-acid-concentration-steam and -microwave. Model scope, stated up front. The acid runs on the brine package, which carries phosphoric acid as a non-volatile molecular solute with its CRC liquid heat capacity (145 J/mol/K) — the streams show the real H₃PO₄/water compositions and the acid's sensible heat is in the balance. What that package does NOT yet carry is the acid's non-ideal water activity: it runs Raoult's law, so the boiling-point elevation is the colligative one, right in sign but about a third of the real value (+9 K at 54% P₂O₅ here; the real acid runs 20-30 K hotter, and 85 wt% acid boils at 158 °C at atmospheric pressure). The Pitzer fit that closes it is published (Bakher & Kaddami 2018, Braz. J. Chem. Eng. 35(3) 1153, open access) and named in the package as the parameters to transcribe. No heat of dilution is carried. Both omissions move the acid temperatures, not the water removed nor the energy-carrier comparison, which is what the study turns on. The 320 K feed stands in for a hotter plant feed for the same reason.
11 unit ops • BRINE
96 0
View & openPhosphoric acid concentration: microwave heating (29 → 54% P2O5)
The same duty with the steam exchanger replaced by volumetric microwave heating: 43.4 MW absorbed in the acid, then the same vacuum flash to 54.0 wt% P₂O₅ and the same 64.0 t/h of water removed — the microwave does not remove water, it supplies the latent heat. At a 65% wall-plug magnetron efficiency (915 MHz industrial units reach ~85%) that is 66.7 MW of electricity, 1,670 kWh per tonne of P₂O₅ (1,280 at 85%). The comparison that decides it is the energy carrier, not the heating mechanism: at Jorf Lasfar the baseline's 1.79 t/t of steam is a co-product of the sulfuric-acid plants, and the site turns that steam into electricity at roughly 30%. Sending it through a turbine, a magnetron and back into the acid delivers about a fifth of the heat the exchanger delivers directly (0.30 × 0.65 ≈ 0.2), so on this site the microwave route needs roughly five times the primary steam energy of the baseline. Where microwaves do earn their place is not here but at the superphosphoric end (>70% P₂O₅, viscous, fouling, where the recent literature targets them), and an electric route for the 29-54 step should be compared against mechanical vapour recompression, which reuses the latent heat at a coefficient of performance of 10-30. Sweep the absorbed duty to see the product concentration and the electricity follow it. Model scope, stated up front. The acid runs on the brine package, which carries phosphoric acid as a non-volatile molecular solute with its CRC liquid heat capacity (145 J/mol/K) — the streams show the real H₃PO₄/water compositions and the acid's sensible heat is in the balance. What that package does NOT yet carry is the acid's non-ideal water activity: it runs Raoult's law, so the boiling-point elevation is the colligative one, right in sign but about a third of the real value (+9 K at 54% P₂O₅ here; the real acid runs 20-30 K hotter, and 85 wt% acid boils at 158 °C at atmospheric pressure). The Pitzer fit that closes it is published (Bakher & Kaddami 2018, Braz. J. Chem. Eng. 35(3) 1153, open access) and named in the package as the parameters to transcribe. No heat of dilution is carried. Both omissions move the acid temperatures, not the water removed nor the energy-carrier comparison, which is what the study turns on. The 320 K feed stands in for a hotter plant feed for the same reason. Microwave penetration depth (about a centimetre in a conductive acid at 2.45 GHz) and the HF attack on quartz windows are equipment questions this model does not carry.
5 unit ops • BRINE
96 0
View & openPhosphoric acid concentration: steam evaporator baseline (29 → 54% P2O5)
The conventional wet-process step, at the scale of one Jorf Lasfar concentration line: 138 t/h of 29% P₂O₅ acid, heated by 3.5 bar LP steam in a graphite exchanger and flashed under vacuum (0.12 bar) to 54% P₂O₅, evaporating 64.0 t/h of water. The steam feed is tuned so the concentrate lands on 54.0 wt% P₂O₅ (the user variable p2o5_wt_product), which takes 71.7 t/h of steam — 1.79 t steam per tonne of P₂O₅, 43.4 MW of condensing duty. That is the single-effect ideal; the published plant benchmark of ~2.8 t/t carries exchanger approach, losses and off-design operation on top. The point of the number here is what it costs the site: at Jorf Lasfar this steam is a co-product of the exothermic sulfuric-acid plants, which is the fact any electric alternative has to beat. Compare the two sibling examples, phosphoric-acid-concentration-microwave and -hybrid. Model scope, stated up front. The acid runs on the brine package, which carries phosphoric acid as a non-volatile molecular solute with its CRC liquid heat capacity (145 J/mol/K) — the streams show the real H₃PO₄/water compositions and the acid's sensible heat is in the balance. What that package does NOT yet carry is the acid's non-ideal water activity: it runs Raoult's law, so the boiling-point elevation is the colligative one, right in sign but about a third of the real value (+9 K at 54% P₂O₅ here; the real acid runs 20-30 K hotter, and 85 wt% acid boils at 158 °C at atmospheric pressure). The Pitzer fit that closes it is published (Bakher & Kaddami 2018, Braz. J. Chem. Eng. 35(3) 1153, open access) and named in the package as the parameters to transcribe. No heat of dilution is carried. Both omissions move the acid temperatures, not the water removed nor the energy-carrier comparison, which is what the study turns on. The 320 K feed stands in for a hotter plant feed for the same reason. The acid heated at 1 atm here partly boils in the exchanger (the model's colligative boiling point at 1 atm); the real acid is kept liquid under static head and flashes in the chamber.
7 unit ops • BRINE
96 0
View & openPressure-swing ethanol dehydration (Gᴱ mixing rule)
Ethanol–water is the classic azeotrope, and pressure-swing distillation breaks it without an entrainer: the azeotrope moves with pressure, so a low-pressure column and a high-pressure column pass each other's azeotropic distillate and each recovers a pure product. The whole process only works if the property package tracks that shift — which is exactly where a conventional package choice falls between two chairs. This flowsheet runs the high-pressure column at 15 bar on pr-mhv1: Peng-Robinson with an MHV1 excess-Gibbs mixing rule, so the cubic equation of state gets its attraction parameter from NRTL's excess Gibbs energy instead of from a single binary interaction constant. Switch the thermo package (Solver menu) and compare the predicted azeotrope: | package | 1 atm | 15 bar | valid at 15 bar? | |---|---|---|---| | NRTL | 0.891 | 0.802 | no — γ-φ is a low-pressure formulation (~10 bar) | | Peng-Robinson (kij) | 0.586 | 0.613 | yes, but a kij cannot represent this azeotrope | | pr-mhv1 | 0.949 | 0.798 | yes | (mole fraction ethanol; the repo's DECHEMA-validated 1 atm anchor is 0.894.) At 15 bar pr-mhv1 lands within 0.005 of NRTL while remaining a genuine equation of state, whereas plain Peng-Robinson is off by ~0.19 and puts the azeotrope in the wrong place entirely. Selecting nrtl here also trips the applicability guard, which warns that the activity model is past its pressure ceiling and names the fix. The flowsheet demonstrates the mechanism on itself. Drop the column pressure to 1 atm and re-run, changing nothing else: the solve fails with SPEC_THERMODYNAMICALLY_IMPOSSIBLE, because at atmospheric pressure the requested bottoms purity sits beyond the azeotrope and no column can reach it. At 15 bar the same specification converges and the bottoms leaves at x_EtOH ≈ 0.924 — past the atmospheric azeotrope of 0.894, which is precisely the composition an atmospheric column cannot cross.
5 unit ops • PR-MHV1
264 0
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