# Quantification of Physical Effects in Relation to Potential Loss of Containment in the CCUS Value Chain

Anders Andreasen

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Anders Andreasen. Quantification of Physical Effects in Relation to Potential Loss of Containment in the CCUS Value Chain. 2025. hal-05044974

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2025

# Quantification of Physical Effects in Relation to Potential Loss of Containment in the CCUS Value Chain

Prepared By :

Anders
Andreasen

<figure><img src="image_3.png" /></figure>

<figure><img src="image_4.png" /></figure>
# Summary

This white paper explores potential loss of containment in the CCUS value chain, focusing on CO2's physical effects. It covers release models for liquid and dense phase CO2, running ductile fracture in dense phase CO2 pipelines, and potential BLEVE (Boiling Liquid Expanding Vapor Explosion). Each topic includes theoretical insights and engineering guidance, emphasizing the importance of understanding CO2's unique properties, such as its ability to exist in gas, liquid, and solid states, and its associated risks.

## Main conclusions:

* The Homogeneous Equilibrium Model (HEM) can estimate CO2 release rates with the right discharge coefficient.

* Pipeline design must include prediction of decompression wave speed to assess running ductile fracture risk. The Battelle two-curve method is unreliable; a semi-empirical approach based on experimental data is recommended.

* BLEVE can occur without fire exposure and with non-flammable materials like CO2. The paper provides guidance on potential CO2 BLEVE risk regions and conditions.
<figure><img src="image_5.png" /></figure>

# Table of Contents

<table><tr><td>01</td><td>Abbreviations</td><td>4</td></tr><tr><td>02</td><td>Introduction</td><td>5</td></tr><tr><td>03</td><td>Carbon Dioxide Phase Behavior</td><td>6</td></tr><tr><td>04</td><td>Release Rate Calculation for Liquid and Dense Phase</td><td>8</td></tr><tr><td>4.1</td><td>Theory</td><td>8</td></tr><tr><td>4.2</td><td>Comparison with Experiments</td><td>9</td></tr><tr><td>4.3</td><td>Guidance</td><td>10</td></tr><tr><td>05</td><td>Running Ductile Fracture</td><td>11</td></tr><tr><td>5.1</td><td>Theory</td><td>11</td></tr><tr><td>5.2</td><td>Guidance</td><td>13</td></tr><tr><td>06</td><td>Bleve</td><td>15</td></tr><tr><td>6.1</td><td>Theory</td><td>16</td></tr><tr><td>6.2</td><td>Guidance - ISO 177349 &amp; CO2RISKMAN (DNV)</td><td>17</td></tr><tr><td>6.3</td><td>Additional Reflection</td><td>18</td></tr><tr><td></td><td>References</td><td>19</td></tr></table>
01 Abbreviations

<table><thead><tr><td>Abbreviation</td><td>Description</td></tr></thead><tbody><tr><td>API</td><td>American Petroleum Institute</td></tr><tr><td>BLEVE</td><td>Boiling Liquid Expanding Vapour Explosion</td></tr><tr><td>DNV</td><td>Det Norske Veritas</td></tr><tr><td>EoS</td><td>Equation of State</td></tr><tr><td>HEM</td><td>Homogenous Equilibrium Model</td></tr><tr><td>HNM</td><td>Homogenous Non-Equilibrium Model</td></tr><tr><td>ISO</td><td>International Standard Organisation</td></tr><tr><td>LCO<sub>2</sub></td><td>Liquid CO<sub>2</sub></td></tr><tr><td>LPG</td><td>Liquified Petroleum Gas</td></tr><tr><td>RDF</td><td>Running Ductile Fracture</td></tr><tr><td>SLT</td><td>Superheat Limit Theory/Temperature</td></tr><tr><td>tcPR</td><td>Translated consistent Peng-Robinson (equation of state)</td></tr></tbody></table>
## 02 Introduction

There is a steadily increasing interest in carbon capture, transport, utilization and storage, for now on referred to just as CCUS. This is no longer vague ideas but actually materializing into real projects such as the Longship and Northern Lights project in Norway, the Porthos project in the Netherlands and the Greensand project in Denmark. Carbon capture projects both for carbon capture from waste to energy facilities, biomass fired power plants and cement production are being sanctioned, built and completed. Some examples include the Heidelberg Brevik cement factory in Norway, the Twence waste to energy facility in the Netherlands and the carbon capture projects at Avedøre and Asnæsværket in Denmark.

For all projects related to CCUS proper evaluation of the hazards related to CO₂ inventory is necessary, both for loss prevention via e.g. inherently safe design, over-pressure protection, protection against low temperature and also for evaluation of consequences in case of loss of containment occurs. In order to do so, an understanding of the physical effects of CO₂ is necessary. CO₂ is special compared to many other chemicals from a risk perspective in that it can be released as gas, liquid and solids in loss of containment scenarios. In addition, CO₂ is toxic, but not flammable.

This white paper addresses the following:

* Estimation of release rate for liquid and dense phase CO₂;

* Evaluation of conditions where a running ductile fracture may occur in a dense phase pipeline and potential catastrophic loss of containment may occur;

* Evaluation of operating conditions where a potential BLEVE may occur in case of a catastrophic loss of containment.

All these topics are highly relevant for the CCUS value chain, and they have all been subject to decades of research, extensive experimental testing, and a vast amount of published scientific literature. Common to all is a deep reliance on advanced thermodynamic analysis and calculations. It is safe to say that in a short article like the present one, full justice cannot be served to the level of detail required in order to fully comprehend the presented phenomena. Nevertheless, the intention is to give an overall understanding and based on the fundamental principles guidance will be given, in line with recognized industry standards from API, ISO and DNV.
03 Carbon Dioxide Phase Behaviour

In order to understand the behaviour of CO₂ it is a good starting point to review the phase diagram as seen in Figure 1. The gaseous state limit follows the sublimation line through the triple point and further along the saturation line up to the critical point. At temperatures higher than this line (or pressures lower) the CO₂ will be in the gaseous state. The liquid state boundary is given by the saturation line and the melting line, both starting from the triple point. Between these two lines and at a temperature below the critical pressure, the CO₂ will be in a liquid state. For pressures above the critical pressure the liquid state is also referred to as dense liquid. For pressures above the critical pressure and concurrent temperature above the critical temperature the CO₂ is in a supercritical state (bounded upwards by the melting line).

<figure><img src="image_6.png" /><figcaption>Figure 1 CO₂ phase diagram. Reproduced from Ben Finney / Mark Jacobs, CCO, via Wikimedia Commons</figcaption></figure>
<figure><img src="image_7.png" /><figcaption>Figure 2 Depressurisation path starting from liquid CO₂ at sub-cooled condition to atmospheric pressure.
© ORS Consulting.</figcaption></figure>

One of the really interesting complications of the CO₂ phase diagram, is the fact that at atmospheric conditions CO₂ can either be in the gaseous, solid or mixed gas/solid conditions, but not in a liquid state. Solid CO₂ is also more popularly referred to as dry ice. Another result of the phase behaviour is also that depressurization of liquid CO₂ will always end as a mixture of gaseous CO₂ and dry ice at -78.5°C. This is illustrated for an isentropic depressurisation in Figure 2 (black line). Sub-cooled liquid at 40 bar will reach the saturation line as pressure is reduced. As the pressure is further reduced the saturation line is followed with gradually increased amount of gaseous CO₂. Once the triple point is reached the further reduction in pressure will result in a gas/dry ice mixture. The amount of dry ice formed depends on the initial conditions, but for liquid CO₂ a 50/50 mixture of gas/dry ice by mass is a good starting point. Starting from a gaseous state may also result in dry ice formation, though the window of initial conditions leading to this is smaller, and generally the amount of dry ice formed will be less by mass.

The formation of dry ice can give a number of complications such as e.g. blockage in downstream piping following a pressure reduction. Thus, the design of a vent/depressurization system must take this into account. When depressuring/emptying process systems (vessels, pipelines etc.), dry ice formation is also a risk to consider, and avoiding pressures below the triple-point is a mitigation measure. Further, in case of release to atmosphere a cloud of dry ice will significantly reduce visibility / line-of-sight.
# 04 Release Rate for Calculation for Liquid and Dense Phase

The estimation of CO₂ release rates in loss of containment scenarios will be the starting point of defining source terms for consequence modelling e.g. toxic CO₂ dispersion modelling. It is therefore important to be able to predict the release rate with some degree of adequacy. In this regard LCO₂ is special compared to many other liquids as LCO₂ may form both gaseous CO₂ and solid CO₂ (dry ice) upon release at atmospheric pressure.

When storing or transporting CO₂ over longer distances, this is either done in the liquid phase at 7-15 bara or in the dense liquid phase at a pressure above the critical pressure of CO₂, often even exceeding 100 bar. Thus, release models which works in both the sub-critical liquid as well as the dense liquid region is required.

## 4.1 Theory

For a general estimation of the maximum release rate in case of e.g. a rupture, the maximum flux can be established via the general steady-flow energy balance (neglecting irrelevant terms).

Equation 1

$$h_0 = \frac{v^2}{2} + h_1$$

Equation 2

$$v = \sqrt{2(h_0 - h_1)}$$

Equation 3

$$G = \rho\sqrt{2(h_0 - h_1)}$$

From the steady flow energy balance (Equation 1) the velocity, v, can be expressed (Equation 2) and from the velocity the corresponding mass flux, G, can be found by multiplication with the fluid density, ρ (Equation 3). Above, h₀ is the upstream/stagnation enthalpy of the fluid, and h₁ is the fluid enthalpy at a lower pressure subject to an isentropic expansion. Hence this method is also referred to as the isentropic nozzle flow equation.
The maximum flux, Gmax, is found by performing isentropic expansion varying the pressure from the upstream pressure, Po, to ambient/back pressure and localizing the pressure where the flux is the highest, either above ambient pressure for critical flow or at ambient pressure in case of subcritical flow. Different variants of the above is also included in the API 520 standard [1] for a numerical integration scheme in order to find maximum flux and also a variant of the omega-method applicable to sub-cooled liquid, where the localization of the maximum flux is implicitly included. The presented methods are generally applicable under the assumption of homogeneous equilibrium (HEM).

However, for prediction of CO₂ releases, the applicability of HEM has been widely discussed and investigated. In the following the Homogeneous Non-Equilibrium Model (HNM) proposed by Vianna et al. [2] is also included when comparing model predictions against experimental observations. HNM considers the metastability of CO₂ and vibrational relaxation and their effects on predicted release rates. While the phenomena included in the model seem complex the proposed model is easy to implement. For further information please refer to ref. [2].

## 4.2 Comparison with Experiments

A number of discharge experiments have been compiled from the literature, these include:

* British Petroleum test cases

* Shell test cases

* Health and Safety Executive (HSE) and COOLTRANS test cases

* INERIS test case

The test results have been extracted from ref. [2], [3]. The different flux prediction models referred above are compared to the experimental results in Figure 3 and Figure 4.

<figure><img src="image_8.png" /><figcaption>Figure 3 Comparison between experiments and calculated mass flux. © ORS Consulting.</figcaption></figure>
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Figure 4 Calculated vs experimental mass flux (left) and mass flow (right). For the calculations a discharge coefficient of 0.67 has been applied for HEM and API methods. For the HNM method three separate discharge coefficients are applied.). © ORS Consulting.

As seen from the comparisons, the HEM method presented in the theory section compares very well with the homogeneous non-equilibrium method and also the experimental data is predicted mostly within expected accuracy. In conclusion the HEM method can be applied in order to estimate CO₂ release rates with proper consideration of the applied discharge coefficient.

It is interesting to note that the formation of dry ice does not seem to invalidate the HEM method. This is believed to be due to the fact that the dry ice forms after vena contracta.

## 4.3 Guidance

To summarise, for defining the source term of CO₂ release consequence modelling it is recommended to:

* Apply a HEM calculation methodology;

* Apply a conservative discharge coefficient of 0.7 (70%);

* The LCO₂ will be released as gaseous CO₂ and dry ice. The phase distribution can be determined from a suitable property package.
# 05 Running Ductile Fracture

Compared to natural gas pipelines a running ductile fracture (RDF) is of bigger concern for carbon dioxide pipelines. Sometimes the wall thickness dictated by the design pressure is not enough to ensure that a ductile fracture is arrested. In case of a pipeline fracture the fluid decompresses, but when the CO₂ reaches saturated conditions the decompression speed drops significantly. In case the decompression speed is lower than the speed of a running fracture, the pipeline itself cannot arrest a running fracture. The consequence is that the pipeline basically unzips from the initial fracture in the longitudinal direction of the pipeline. If not arrested the running fracture may extend several hundred meters cf. Figure 5. In order to properly design against a running ductile fracture it is essential to be able to predict the decompression wave speed [4].

<figure><img src="image_11.png" /><figcaption>Figure 5 Example of running ductile fracture in a buried pipeline. Reproduced from ref. [5] under the creative commons license.</figcaption></figure>

## 5.1 Theory

In the following the fluid is assumed inviscid, in homogeneous equilibrium and the decompression process to be isentropic. The decompression wave speed $w$ can be calculated based on the equations 4, 5 and 6, where $c$ is the speed of sound of the fluid and $v$ is the outflow velocity. In order to calculate the outflow velocity and the speed of sound, a suitable and accurate thermodynamic package is required with accurate fluid description both in the liquid, gaseous and two-phase region is required.
<table><tr><td>Equation 4</td><td>w = c - v</td></tr><tr><td>Equation 5</td><td>v = -∫<sub>p<sub>0</sub></sub><sup>p</sup> (1 / cρ) dp</td></tr><tr><td>Equation 6</td><td>c = √<span style="text-decoration: overline">((dp)/(dρ))</span><sub>s</sub></td></tr></table>

Note: The speed of sound for a two-phase system is not well defined. However, assuming homogenous equilibrium and isentropic conditions it can be calculated from Equation 6. Pressure: p, Density: ρ, Entropy: s.

The calculation is started from the initial conditions (p, T), from which an isentropic path is followed, gradually reducing the pressure. For each step the speed of sound is calculated. Typically, a numerical scheme is required. The resulting decompression wave speed curve for CO₂ can look like the one in Figure 6. Starting from high to low pressure, a sudden drop in decompression wave speed occurs, which coincides with the entry to the two-phase region. To which extent a running ductile fracture will occur, can be assessed by overlaying the pipeline fracture speed curve. The red curve shows a case where the fracture speed exceeds the decompression wave speed and a running fracture may occur. Increasing the pipeline wall thickness or toughness can shift the fracture curve, as exemplified by the green curve, and the situation is safe from running fracture. In simplified terms the high-pressure wave will now travel at a faster speed away from the initiating fracture compared to how fast the fracture will travel, thus relieving the pressure at the tip of the fracture and thereby arresting it. The analysis is known as the Battelle two-curve method.

<figure><img src="image_12.png" /><figcaption>Figure 6 Illustration of the Battelle two-curve method for evaluation of risk of running ductile fracture combining the fluid decompression wave-speed curve with the fracture propagation speed curve. Adapted from [4].</figcaption></figure>
## 5.2 Guidance

In order to properly assess the adequacy of the pipeline design it is pivotal to know the on-set of the sudden drop in fluid decompression wave-speed. The on-set pressure, P, enters as one of the governing parameters in the evaluation method presented in the following. This pressure is also referred to as the plateau pressure if the visual representation is inverted, and the decompression wave speed is depicted on the x-axis and the pressure on the y-axis.

The Battelle two-curve method has been found not to have sufficient reliability and conservatism even with de-rated terms applied and due to this uncertainty the Battelle two-curve method analytical approach is advised against. Instead a semi-empirical approach is suggested based on experimental data [6].

For assessing pipeline design for running ductile fracture, ISO 27913:2024 [7] provides guidance, which is largely based on the semi-empirical based guidance from DNV-RP-F104 [6]. The semi-empirical guidance comes from experience from a large number of tests performed through various programmes.

Based on materials properties, pipeline dimensions and wall thickness combined with the plateau pressure from the decompression waves-speed curve, Figure D.1 and the corresponding formulae can be used to assess if the pipeline design is susceptible to RDF or if arrest of fractures will occur. Figure D.1 is recreated (including experimental data) in Figure 7. If the evaluation is inside the shaded region arrest will not occur, and vice versa if outside the shaded region a fracture will propagate without being arrested i.e. a possible RDF. There is an uncertain region to the left of the shaded region where additional assessment may qualify the design.

<figure><img src="image_13.png" /></figure>
Figure 7 Evaluation of fracture arrest in dense CO₂ stream pipelines according to ISO 27913:2024 and DNV -RP-F104. Experimental data has been extracted from the CO₂ PIPE TRANS, COOLTRANS, CO₂ SAFE-ARREST, SARCO₂ experiments from various literature sources, see also [8]. © ORS Consulting.

In a previous study [4] the decompression wave-speed was compared to experimental data, and it was found that the HEM assumption combined with an accurate equation of state gave good results for pure CO₂. For cases with a significant amount of impurities mostly conservative behaviour was observed i.e. over-prediction of P, but some non-conservative behaviour was also identified e.g. for dense phase CO₂ with a significant fraction of hydrogen.

For evaluation, it is strictly required to ensure proper validation of the equation of state used to establish the decompression wave-speed curve. In case of presence of impurities, we suggest that as a minimum the equation of state should be able to predict the following with adequate accuracy:

* Bubble point and dew point curves for the binary pair CO₂/X at various conditions, where X is the impurity. This applies to all impurities present in a significant fraction.

* Preferably also prediction of bubble and dew point lines for ternary pairs of CO₂/X/Y, but acknowledging that experimental data may not be readily available for all combinations.

* Prediction of density as a function of pressure and temperature over a wide range of conditions and covering all the states gaseous, liquid (subcritical), two-phase, dense phase (liquid) and supercritical (in case this is also a possible operating scenario).
## 06 Bleve

Boiling liquid expanding vapor explosion (BLEVE) is often associated with fires exposing separators and other vessels containing pressurized liquid hydrocarbons in the Oil & Gas Industry, LPG tanks and similar and known as very violent events causing both pressure waves, heat radiation (fire ball) and flying fragments all of which needs to be considered in risk assessments. However, a BLEVE can occur even without fire exposure and can also occur for non-flammable materials like CO₂.

A BLEVE without ignition may occur when a vessel storing pressure liquefied gas such as e.g. propane, LPG, chlorine, or ammonia is subject to accidental rapid depressurization e.g. due to a catastrophic loss of containment. When the pressurized liquid is depressurized very rapidly it may result in the liquid evaporating homogeneously throughout the whole fluid creating a shockwave.

To which extent pressurized liquid CO₂ is also prone to potential BLEVE has been widely debated and investigated in the literature. In the following the fundamental requirements for a BLEVE to occur is briefly discussed followed by guidance on regions of potential risk of CO₂ BLEVE. As a potential CO₂ BLEVE will have much more violent initial consequences than a large continuous CO₂ release it is important to be able to predict when a CO₂ BLEVE is possible. When a CO₂ BLEVE is possible the potential consequences can be estimated by similar methods as applied for BLEVEs in general e.g. reference is made to methods of the Yellow Book [9].

The following text is cited verbatim from ISO 17349:

> "Boiling Liquid Expanding Vapour Explosion (BLEVE) is an explosion resulting from the failure of a container of liquid at a temperature significantly above its boiling point at normal atmospheric pressure. This rare but extremely catastrophic event can occur when a vessel containing liquid above its normal boiling point fails. It is noteworthy that the BLEVE does not cause the tank rupture. The BLEVE results from the sudden opening of the vessel which stems from flawed materials, fatigue, corrosion, poor manufacture, thermal stresses, pressure stresses, reduction in material strength due to high wall temperatures or external factors. The hazards generated by BLEVE include shock overpressures, high velocity expanding vapour and flashing liquid, fragmentation of the container, release of projectiles and release of the contained fluids".

Both ISO 17349 [10] and DNV CO2RISKMAN [11] contains descriptions on CO₂ BLEVE and has also been a source of inspiration for this white paper.
## 6.1 Theory

In order to describe the basics, Figure 8 is used as a guide. The Figure shows an isothermal decompression path ABCD starting from the fluid in the dense liquid state (A). During expansion the pressure drops dramatically until the bubble point is reached (B), from which the further expansion occurs at unchanged pressure (for a pure substance) until all liquid is evaporated at the dew point (C), and the expansion continues at decreasing pressure (D).

<figure><img src="image_14.png" /><figcaption>Figure 8 Pressure-volume relation. The figure has been re-created from ref. [10].</figcaption></figure>

In case the depressurization is very fast, there is insufficient time for heat ingress to evaporate the liquid along the saturation line BC, and the pressure follows the line BS, which is a metastable state. If the reduction in pressure is not too fast, at any point along line BS the fluid can return to line BC via violent boiling. If not, and the spinodal point (S) is reached. The solid line from the critical point through S is the liquid spinodal. The definition of the spinodal point i.e. any point on the spinodal is given by Equation 7.

Equation 7

$$ \left(\frac{\partial p}{\partial v}\right)_T = 0 $$

In case the spinodal point is reached then phase separation must occur according to thermodynamics and separation into liquid and gas will rapidly occur. The density variations develop spontaneously into liquid and gas regions. This occurs homogeneously throughout the whole liquid [10].
## 6.2 Guidance - ISO 177349 & CO2RISKMAN (DNV)

For a BLEVE to occur, the fluid must be on the spinodal curve between 1 atm (assuming release to ambient) and the critical point where the curve ends as per Figure 8. When a catastrophic pressure vessel failure happens very fast, there is not sufficient time for heat to be absorbed by the fluid.

Thermodynamically speaking, during failure the entropy of the fluid is constant. For a BLEVE to occur the initial entropy must be higher than the entropy of the spinodal point at atmospheric pressure [11]. Using this information a region of initial conditions for pure CO₂ that could give a spinodal state can be established i.e. where the entropy requirement for ending up at the spinodal point with instantaneous phase separation as a consequence cf. Figure 9, see also [12]. The region of initial conditions that may lead to the spinodal state in an isentropic expansion is illustrated by the grey region in the Figure.

<figure><img src="image_15.png" /><figcaption>Figure 9 Domain (grey) which may lead to a spinodal state upon instantaneous depressurization of CO₂. Thermodynamic calculations have been made using the translated and consistent Peng-Robinson EoS, tcPR [13] via thermopack [14]. © ORS Consulting.</figcaption></figure>

As seen from Figure 9 the lowest pressure and temperature which may theoretically lead to a CO₂ BLEVE is 291.5 K (18.4°C) and 56 bara with the chosen EoS. In the case of presence of other components these should ideally be included in the thermodynamic analysis. For liquefied CO₂ the concentration of impurities is very low, as e.g. seen from the CO₂ specification for the Northern Lights project [15]. Hence, pure CO₂ may be used as a reasonable proxy for thermodynamic analysis.

For storage of liquid CO₂ under normal condition (7-15 bara), the pressure and temperature are on the saturation curve and quite far from where a spinodal state may be reached.
## 6.3 Additional Reflection

A variant of the above also referred to as the super-heat limit theory (SLT) [16], [17] defines a minimum temperature at which BLEVE can occur i.e. the superheat limit temperature (also SLT). The theoretical SLT can be considered as the spinodal at atmospheric conditions, or it can be determined experimentally, where a value of 267 K has been reported [18]. This apparently extends the region as visualised in Figure 9. Consulting the literature on CO₂ BLEVE, the evaluation criteria provided by the SLT and in the previous section may be thermodynamically sound and correct in theory. However, it does not change the fact that BLEVE's have been reported to occur also at lower initial temperatures (and pressures) as shown in the shaded region in Figure 9 and also lower than the SLT [17], [19]. Experiments suggest that the overpressure created following a vessel rupture gradually decreases as the initial temperature is decreased. While the blast pressure is highest when the initial temperature is above the SLT, it does not entirely disappear even when the initial temperature is below the SLT [19]. Some refer the overpressure / blast generated below the SLT as a cold BLEVE (as opposed to hot BLEVE when above the SLT). Clearly, depressurisation below the SLT results in a meta-stable state, where violent boiling will occur, yet if this should be classified as a BLEVE is not as simple as that.

It is also important to recall that for a BLEVE to occur in the first place a catastrophic failure of the vessel holding the CO₂ must take place. As also stated by ISO 17349, the BLEVE does not cause the rupture - it may be a result of the rupture. For flammable inventories, or inventories which are stored nearby flammable substances an external fire e.g. a pool fire or a jet fire may cause weakening of the CO₂ storage vessel and cause both high pressure and rupture due to thermal weakening of the vessel. Often for CO₂ storage vessels, these may be far from flammable substances, and a catastrophic rupture would be due to integrity failure of the vessel e.g. corrosion, poor welding etc. or an external force.
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## About the Author

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Anders Andreasen has worked extensively on projects spanning the entire CCUS value chain, including carbon capture, CO2 compression, dehydration, purification, liquefaction, pipeline transport, ship-loading, injection, and utilization. While doing so, he has served both national and international clients. Supporting client projects has included utilizing existing open-source tools or developing new tools for tackling problems such as CAPEX estimation, compressor system optimization, running ductile fracture, and CO2 solubility in brines. Mr. Andreasen has also contributed to both the Greensand (https://pilot.greensandfuture.com/en) and INNO-CCUS (https://inno-ccus.dk/) research and demonstration projects. His work within CCUS is also showcased in numerous scientific papers.

Anders Andreasen holds both an M.Sc. and Ph.D. degree in Chemical Engineering from Aalborg University and the Technical University of Denmark, respectively.

Anders Andreasen

Principal Consultant

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