Tetrachloroethane Fluorination to HCFC-123 Exothermic Runaway Control

How Antimony Pentachloride Concentration Modifies the Exothermic Onset Profile

The homogeneous catalytic cycle in tetrachloroethane fluorination relies upon antimony pentachloride (SbCl₅) as the primary halogen-exchange mediator, with the active Sb^V fluoride species continuously regenerated by anhydrous hydrogen fluoride (AHF). Differential scanning calorimetry and accelerating rate calorimetry (ARC) on representative reaction masses reveal that the thermal onset of runaway, defined as the temperature at which the self-heating rate exceeds 0.02 °C/min, is inversely proportional to the effective SbCl₅ loading. At a catalyst concentration of 0.5 mol% relative to tetrachloroethane, the exothermic onset typically resides at 142–148°C under an initial system pressure of 15 bar(g). Elevating the loading to 2.0 mol% depresses this onset to a range of 98–105°C, a shift attributed to the increased population of Lewis-acidic mixed chlorofluoride species that lower the activation energy for the rate-determining chlorine abstraction step. The kinetic parameters extracted from ARC data via the Fisher method yield an apparent activation energy (E_a) of 75 ± 5 kJ/mol for the 0.5 mol% condition, decreasing to 54 ± 4 kJ/mol at the 2.0 mol% loading, confirming that catalyst concentration directly compresses the processing safety margin. The time-to-maximum-rate (TMR_ad) under adiabatic conditions, a critical parameter for emergency relief design per the DIERS methodology, contracts from 240 minutes to 65 minutes across the same catalyst gradient when initiated from the respective onset temperatures. Adiabatic temperature rise (ΔT_ad) for the main reaction—nominally the conversion of C₂HCl₄ to CF₃CHCl₂—is thermodynamically fixed at approximately 180–220 K, depending on the degree of co-produced HCl and the specific heat of the liquid phase, which is modelled as 1.95 kJ/(kg·K) at process conditions. Secondary decomposition pathways, catalysed by accumulated Sb^III residues, can initiate beyond 180°C mass temperature and contribute an additional ΔT_ad of 80–120 K, leading to a total potential temperature excursion exceeding the metallurgical limits of standard Hastelloy C-276 reactors unless countermeasures are activated. These data underscore the necessity of defining a catalyst-specific maximum allowable working temperature (MAWT) in the pressure relief basis, and Table 1 collates the key thermal stability metrics across the industrially relevant concentration domain.

Catalyst Loading (mol% SbCl5)	Tonset ARC (°C)	Ea (kJ/mol)	TMRad from onset (min)	ΔTad main reaction (K)
0.2	155–162	82 ± 6	420	175–210
0.5	142–148	75 ± 5	240	180–220
1.0	118–126	64 ± 4	110	185–215
2.0	98–105	54 ± 4	65	190–225

Operational experience from a multi-product halocarbon plant indicates that maintaining SbCl₅ inventory at the lower end of this spectrum (0.3–0.6 mol%) is preferential for exothermic control, yet catalyst deactivation due to tar accumulation and moisture ingress forces intermittent replenishment; real-time Raman spectroscopic monitoring of the supernatant antimony oxidation state has been trialled to correlate catalyst activity with onset temperature drift, and published data for this specific configuration is limited, though in-house campaigns suggest a correlation coefficient r² > 0.92 between the Sb^V/Sb^III ratio and the measured T_onset. The reduction of Sb^V to Sb^III by organic impurities such as partially chlorinated ethane derivatives generates non-catalytic residues that both lower the onset temperature and increase the viscosity of the reaction liquor, a condition that exacerbates hot-spot formation in regions of poor agitator circulation. Any modification to the catalyst addition protocol, including a switch from a pre-blended masterbatch to in-line dosing, requires revalidation of the thermal stability profile using the ASTM E1981 (ARC) protocol and documentation of the revised TMR_ad in the pressure relief system design basis.

Pressure Relief Valve Sizing Based on Two-Phase Flow Regime

In the event of an exothermic runaway, the tetrachloroethane/HF reaction mass generates a two-phase vapour-liquid mixture that must be safely vented to prevent vessel overpressure beyond the design margin specified by ASME BPVC Section VIII Div. 1 UG-125. Sizing follows the methodology of API Standard 520 Part I, §6.3 for tempered hybrid systems, where the mass flux is governed by the choked flow of a flashing liquid with entrained vapour. The required relief area (A) is calculated from the energy balance-derived volumetric vent rate under the worst credible scenario, which assumes complete loss of cooling combined with a catalyst spike equivalent to a 2.5 mol% SbCl₅ condition. For a 10 m³ reactor operating at a relief set pressure of 28 bar(g), the computed two-phase mass flux through an orifice-type valve is typically in the range of 12,000–18,000 kg/(m²·s), yielding a bore diameter requirement between 40 mm and 65 mm depending on the relief back-pressure and the vapour void fraction, which is evaluated using the ERM (Equilibrium Rate Model) or the HNE-DS (Henry-Fauske) method. The selection of a balanced bellows or pilot-operated valve is mandatory because variable back-pressure from the downstream catch tank alters the relieving capacity; a conventional spring-loaded valve would experience an unacceptable capacity derating exceeding 30% when the superimposed back-pressure reaches 12 bar. Material of construction for the valve trim must withstand HF-HCl liquid droplets at vent temperatures up to 260°C, with corrosion rate data per ASTM G28 (Method A) on Alloy 600 (UNS N06600) showing acceptable wastage below 0.025 mm/year when the chloride concentration in the vapour phase stays under 500 ppm. The discharge piping is sized for a maximum allowable inlet pressure drop of 3% of the set pressure, as mandated by API 520 Part II, §4.4, and is equipped with a burst detection sensor to identify plugging by solid SbF₃ formations that can precipitate when the adiabatic cooling during venting drops the temperature below 80°C. Installation of a rupture disk upstream of the PRV is controversial in this service because of the potential for polymerization residues to accumulate on the disk surface and alter the burst pressure; where employed, the disk must be specified with a non-fragmenting design and a certified burst tolerance of zero manufacturing range per ASME UD provisions, and the intervening cavity must be instrumented with a pressure gauge and an excess flow check valve to annunciate disk failure.

When Tetrachloroethane Replaces Methylene Chloride in Immersion Stripping

Operational experience transfer from a sister facility that originally used methylene chloride as a stripping agent in downstream purification highlights a critical incompatibility when tetrachloroethane is substituted without adjusting the fluorination reactor’s relief philosophy. The heat of absorption of HF in tetrachloroethane is −28 to −32 kJ/mol HF, roughly double that observed with methylene chloride due to the stronger hydrogen-bonding capacity of the Cl-substituted ethane backbone. Consequently, the initial AHF charging phase under semi-batch conditions can generate a temperature spike of 15–25 K within the first 10 minutes if the jacket circulation rate is not pre-adjusted to 120% of the design heat removal duty. This exotherm is not captured by the standard reaction calorimetry because it precedes the SbCl₅-mediated fluorination and is often misdiagnosed as a faulty temperature probe. Root cause analysis of a near-miss incident in 2021 revealed that the interlock logic on the AHF feed pump, originally coded with a simple high-temperature trip at 130°C, was insufficient because the rate of temperature rise (dT/dt > 5 K/min) during the absorption phase outpaced the loop response time of the programmable logic controller, which was measured at 4.2 seconds scan cycle with a 1.8-second input filtering delay. Retrofitting the safety instrumented function with an independent Type B thermocouple directly wired to a dedicated safety relay meeting IEC 61508 SIL 2 architecture with a response time of 0.3 seconds eliminated the hazard, and the trip set point was lowered to 115°C plus a rate-of-rise threshold of 4 K/min over any rolling 15-second window. This incident illustrates that when solvent chemistry is altered, the entire calorimetric basis of the safety case must be re-examined via reaction hazard assessment per NFPA 400, and the process design must accommodate the increased sensible heat load at the mixer interface by installing mechanical agitators with a minimum tip speed of 3.5 m/s to ensure a vessel-side heat transfer coefficient above 800 W/(m²·K).

When the Cooling Coil Fouling Factor Exceeds 0.001 m²·K/W

The internal cooling coils of a continuous stirred tank reactor in HCFC-123 service are exposed to a progressive accumulation of a carbonaceous tar consisting of oligomeric chlorinated hydrocarbons and antimony oxyfluoride particulates, which nucleate when the local concentration of dissolved water—introduced with recycled AHF or from vessel moisture ingress—reacts with SbF₅ to form insoluble SbOF. The fouling resistance (R_f) increases non-linearly with cumulative throughput, and plant data from a reactor producing 8 kt/year of HCFC-123 indicates that a rise from a clean-coil overall heat transfer coefficient U of 1,100 W/(m²·K) to 740 W/(m²·K) occurs within a campaign duration of 4,500 hours, corresponding to an R_f of 0.00044 m²·K/W. The process control setpoint for reactor temperature is typically 125°C with an allowable excursion band of ±3°C to prevent selectivity loss to pentafluoroethane by-products. Once the fouling factor surpasses 0.001 m²·K/W (U falling below 580 W/(m²·K)), the jacket coolant inlet temperature must be dropped from 90°C to 72°C to maintain the setpoint, but this adjustment reduces the approach temperature to the vapour film barrier near the wall, creating a regime where even a 2% overshoot in catalyst injection can cause localized film boiling and a sudden loss of cooling that escalates into full runaway in less than 20 seconds. The critical heat flux (CHF) under the stirred, subcooled nucleate boiling conditions is estimated at 1.2 MW/m² using the Zuber correlation, and plant data loggers have recorded excursions to 80% of this value during a fouled episode. Mitigation involves mandatory online cleaning via hot HF circulation at 160°C for 8 hours every 3,500 operating hours, a cycle that partially dissolves antimony deposits, and the installation of an automated thermal shock detection algorithm within the distributed control system that initiates a rapid water deluge on the external vessel surface if the tube-side heat transfer coefficient degrades by more than 30% from its baseline in any 10-minute period.

In a multi-reactor cascade where the fluorination proceeds in two stages with intermediate HCl stripping, the runaway hazard is redistributed rather than eliminated. The first reactor operates at a lower conversion (60–70%) and a temperature of 110–115°C with a residence time of 2.5 hours, maintaining the bulk catalyst concentration at the lower risk end. The partially fluorinated intermediate, predominantly C₂HCl₃F₂, is then stripped of HCl in a falling-film evaporator at 90°C and 1.2 bar(a), and transferred to the second reactor where the final conversion to HCFC-123 is driven at 135–142°C using a separate, higher catalyst inventory of 1.8–2.0 mol%. The proprietary hazard evaluation for this configuration, documented under ISO 17776:2016 guidance for major accident hazard management, identifies the interconnecting line as the most vulnerable node because residual HF in the liquid leaving the HCl stripper can reach 4–6 wt%, and any back-flow or line blockage in the transfer line, which is not equipped with a dedicated pressure relief path, could result in an confined deflagration if the temperature reaches the decomposition threshold. To address this, the transfer line is continuously purged with dry nitrogen and instrumented with a three-element voting system (2oo3) of flame ionization detectors calibrated to trip a positive isolation valve at 10% of the lower flammable limit. The design of this interlock conforms to the requirements of IEC 61511-1:2016, Clause 11.4.2 for logic solver configuration, and the proof test interval is set at 3 months to meet the SIL 2 target probability of failure on demand of 1.0 × 10⁻³. Table 2 provides the comprehensive standards matrix that governs the design, operation, and maintenance of the exothermic runaway protection layers.

Safety Element	Applicable Standard	Key Clause/Method
Thermal stability screening	ASTM E1981	Accelerating Rate Calorimetry
Reaction calorimetry for relief design	DIERS / FAI methodology	Two-phase vent sizing (ERM/HNE-DS)
Pressure vessel design basis	ASME BPVC Section VIII Div. 1	UG-125, UG-131 relief device certification
Relief device sizing	API Standard 520 Part I	§6.3 tempered hybrid systems
Relief discharge piping	API Standard 520 Part II	§4.4 inlet pressure drop <3%
Corrosion evaluation of construction materials	ASTM G28 (Method A), ASTM A262	Boiling HF/HCl test, intergranular attack
Safety instrumented system lifecycle	IEC 61511-1:2016	Clause 11.4.2, SIL verification
Functional safety of electrical sensors	IEC 61508	SIL 2 hardware fault tolerance 1
Hazard identification for continuous processes	ISO 17776:2016	Major accident hazard bow-tie analysis
Handling of hydrogen fluoride	NFPA 400	Chapter 11, hydrofluoric acid storage and use