2.1.1. Net electric power selection

A plant’s net electric power P _net (the power it can provide to consumers) is chosen as a baseline parameter for LCOE optimisation as it strongly affects other indicators (specifically, costs). Usually a commercial plant, the net electric power is in range of 50–1200 MW, with the lower limit reserved for simple inexpensive gas-operated turbines and higher power for complex nuclear plants. Exceeding 1200 MW limit is unfavourable as it compromises grid stability, requiring additional costs to ensure grid operation in case of plant failure. To establish the relationship between net electric power and plant costs, it is necessary to understand how the plant net electric power relates to fusion plasma properties.

2.1.2. Fusion reactor indicators

Fusion plant net electric power P _net can be calculated as follows:

(2.1) \begin{equation}P_{\textit{net}}=P_{\textit{gross}}-P_{\textit{aux}}-P_{\textit{heat}}/k,\end{equation}

where P _gross is the total plant generator electric power, P _aux is the plant’s auxiliaries power consumption (pumps, vents and other secondary equipment), P _heat is the power delivered into the plasma for its heating and sustaining plasma temperature, and k is the plasma heating system’s efficiency (usually around 50 %). The ratio P _heat /k is called gross plasma heating power in this article to avoid confusion since plasma heating power P _heat is used to describe the amount of power injected into plasma, while the ratio P _heat /k is used to describe losses in electric power due to operation of plasma heating systems.

Here,P _gross is related to the plant’s heat-to-electricity conversion N coefficient (is usually in the range 35 %–45 % depending on the construction of a turbine island – more on that later); P _tot is the total reactor heat power (produced by fusion reaction, and in case of tritium breeding, fission in blanket):

(2.2) \begin{equation}P_{\textit{gross}}=P_{\textit{tot}}N,\end{equation}

(2.3) \begin{equation}P_{\textit{tot}}=P_{\textit{fus}}+P_{\textit{fis}}.\end{equation}

A plant auxiliaries’ power consumption P _aux is usually a fraction of the plant’s total electric power:

(2.4) \begin{align}P_{\textit{aux}}=P_{\textit{gross}}f,\end{align}

where f is the auxiliaries power fraction (usually 7 % electric power (Farina & Festa Reference Farina and Festa2023)). The plasma heating power can be expressed as a ratio of the plant’s fusion power to the fusion gain Q, indicating how fusion amplifies heating energy, induced by plasma heating systems:

(2.5) \begin{align}P_{\textit{heat}}=P_{\textit{fus}}/Q.\end{align}

For the VOYAGER plant, which uses DD fusion without tritium breeding, the plant total power equals the fusion power (P _fus ). Thus, for the VOYAGER plant, one can express P _net as follows:

(2.6) \begin{equation}P_{\textit{net}}=P_{\textit{fus}}N\!\left[1-f-1/\left(Qk\right)\right]\!.\end{equation}

The latest equation suggests the VOYAGER plant net electric power is strongly dependent on reactor design parameters such as fusion power P _fus and gain Q that can vary greatly. Other parameters (N, f, k) do not affect the net electric power significantly, as they represent already optimised equipment (plasma heaters, turbines and other conventional devices).

2.1.3. Reactor fusion power and fusion gain optimisation

The VOYAGER fusion plant plasma is considered as a diamagnetic bubble of a radius a and a length L in a stationary regime. It confines a purely deuterium plasma of density n. Distribution functions of ions and electrons are both close to Maxwellian with a temperature T. The diamagnetic bubble formation condition is the equality of pressures of the plasma and the vacuum magnetic field B:

(2.7) \begin{equation}\frac{B^{2}}{8\pi }=2nT.\end{equation}

By definition, the magnetic field is close to zero inside the ‘bubble’ and rises up to vacuum one in a thin transition layer. The thickness of this layer is of the order of the ion gyroradius ρ _i (Beklemishev Reference Beklemishev2016):

(2.8) \begin{equation}\rho _{i}=\frac{\textit{v}_{Ti}}{\varOmega _{i}},\quad \textit{v}_{Ti}=\sqrt{\frac{2T_{i}}{m_{i}}},\quad \varOmega _{i}=\frac{eB}{m_{i}c},\end{equation}

where m _i and v _Ti are deuterium mass and thermal velocity, Ω _i is ion cyclotron frequency calculated by the vacuum field B, e is the elementary charge and c is the speed of light. Two types of losses are taken into account. The first is the escape rate of a diamagnetic configuration (Chernoshtanov Reference Chernoshtanov2022):

(2.9) \begin{equation}\tau _{\parallel }=\eta R_{\textit{v}}\tau _{\textit{gd}}\frac{a}{\rho _{i}},\quad \tau _{\textit{gd}}=\frac{RL}{\textit{v}_{\textit{Ti}}},\end{equation}

where τ _gd is a gas-dynamic confinement time, enhanced by the factor of vortex section $R_{\textit{v}}$ , and R is the vacuum mirror ratio. The expression contains the product $\eta \cdot R_{\textit{v}}$ with the factor of order of unity $\eta \sim 1$ . Therefore, product $\eta \cdot R_{\textit{v}}$ is replaced by $R_{\textit{v}}$ for simplicity. The mirror field $B_{m}$ is fixed at 20 T as an already achieved value for mirrors (Baburin et al. Reference Baburin, Zapretilina, Hitrov, Kolomentseva, Mednikov, Moizykh, Ovsyannikov, Rodin and Shcherbakov2024). Now, the mirror ratio is rewritten as R = $B_{m}$ /B. Equation (2.9) does not depend on the vacuum magnetic field, thus:

(2.10) \begin{equation}\tau _{\parallel }=R_{\textit{v}}\frac{aL}{\textit{v}_{\textit{Ti}}^{2}}\varOmega _{\textit{mi}},\end{equation}

where $\Omega_{mi}$ is the ion cyclotron frequency calculated by the mirror field $B_{m}$ . The current of axial losses does not depend on the device length, while the power of axial losses is defined by the average energies of escaping particles:

(2.11) \begin{equation}J_{\parallel }=\frac{nV}{\tau _{\parallel }}=\frac{\pi a}{R_{\textit{v}}\varOmega _{mi}}n{\textit{v}}_{Ti}^{2},\quad P_{\parallel }=\kappa _{\parallel }\textit{TJ}_{\parallel },\end{equation}

where κ_‖ is assumed to be constant κ_‖ = 7. The second type is radiation losses for electrons:

(2.12) \begin{equation}P_{r}=P_{\textit{Br}}+\alpha _{\textit{cy}}P_{\textit{cy}}, \quad P_{\textit{Br}}=\pi a^{2}\textit{LC}_{\textit{Br}}n^{2}\sqrt{T},\quad P_{\textit{cy}}=2\pi a\rho _{\textit{i}}\textit{LC}_{\textit{cy}}B^{2}nT,\end{equation}

where P _Br is Bremsstrahlung radiation generating in the whole plasma volume, P _cy is cyclotron radiation generating in the transition layer, and coefficients C _Br and C _cy are taken from Huba (Reference Huba2013) and Peakman & Lindley (Reference Peakman and Lindley2023). The fraction of cyclotron radiation absorbed by the wall is denoted by α _cy . The following calculation assumes that only α _cy = 30 % of cyclotron radiation is lost; the rest of the power is reflected by the wall and returned to the plasma. Radial losses from the transition layer ( $J_{\bot}$ , $P_{\bot}$ ) driven by Coulomb collisions were also analysed. It is found to produce 5 %–6 % of power and up to 40 % of particle losses.

Figure 2.

Dependence of plasma parameters (density n, ion temperature T _i , electron temperature T _e , vacuum magnetic field B and inverse fusion gain 10/Q) versus vortex section factor ( $R_{\textit{v}}$ ).

A purely deuterium–deuterium ‘primary’ fusion produces rather low power; to make the fusion reactor compact, the VOYAGER design uses ‘secondary’ fusion reactions with by-products (tritium and helium-3) to contribute to the overall fusion power, with an assumption that all the tritium produced in primary reactions is returned back into the plasma for the secondary reaction:

(2.13) \begin{equation}D+D\rightarrow \left\{\begin{array}{c@{\quad}c} 50\,{\% }\colon & p+T\\[3pt] 50\,{\% }\colon & n+{}^{3}\textit{He} \end{array}\right\},\quad D+T\rightarrow n+{}^{4}\textit{He},\quad D+{}^{3}\textit{He}\rightarrow p+{}^{4}\textit{He}.\end{equation}

The energy gain factor is

(2.14) \begin{equation}Q=\frac{P_{\textit{fus}}}{P_{\textit{heat}}},\quad P_{\textit{heat}}=P_{\parallel }+P_{\bot }+P_{\textit{r}}-P_{\textit{p}},\end{equation}

where P _fus is the total fusion power, P _p is the part of P _fus released in charged products and spent for plasma heating, and P _heat is the total external heat power introduced to the plasma.

The listed equations were used to calculate fusion reactor parameters using BINP SB RAS proprietary calculation software, starting with fusion power P _fus = 3 GW, and length and radius of diamagnetic bubble of 100 and 1 m, respectively (see figure 2). Notice the fusion gain of an open trap based reactor rises together with a vortex section factor, with ignition occurring at $R_{\textit{v}}$ = 3800 (P _heat → 0 and Q → ∞), see table 1.

Table 1.

Parameters of reactor with ignited plasma.

The fusion power P _fus , fusion gain Q and vortex section factor $R_{\textit{v}}$ are related to each other. For example, fusion gain Q rapidly rises with vortex section factor $R_{\textit{v}}$ at a fixed fusion power P _fus , as shown in figure 2. Understanding the relationship among these parameters is useful for reactor optimisation. Let us analytically consider the mode with fixed Q neglecting radial diffusion and cyclotron radiation for simplicity. The power balance equation (2.14) in this case gives

(2.15) \begin{equation}P_{\parallel }=\left(\alpha -\frac{1}{Q}\right)P_{\textit{fus}}-P_{\textit{Br}},\end{equation}

where α is a fraction of fusion power that heats plasma. Fusion power and Bremsstrahlung power are both proportional to reactor length and density square as well as the right side of (2.15):

(2.16) \begin{equation}P_{\textit{fus}},P_{\textit{Br}}\sim Ln^{2}.\end{equation}

Parallel losses are defined by the density n and vortex section factor $R_{\textit{v}}$ :

(2.17) \begin{equation}P_{\parallel }\sim \frac{n}{R_{\textit{v}}}.\end{equation}

Combining (2.15)–(2.17), one gets

(2.18) \begin{equation}R_{\textit{v}}\sim \frac{1}{\sqrt{P_{\textit{fus}}L}}.\end{equation}

Fusion power is proportional to VOYAGER reactor length since the heat dissipation via reactor walls is limited (see § 2.3.1 for more information), thus:

(2.19) \begin{equation}R_{\textit{v}}\sim \frac{1}{P_{\textit{fus}}}.\end{equation}

Thus, at a given fusion gain Q, the vortex section factor is inversely proportional to the reactor fusion power. It is reasonable to keep the vortex section factor $R_{\textit{v}}$ and fusion power P _fus as low as possible to minimise plant complexity and cost, while keeping net electric power P _net high to provide revenue, so balancing for Q, R _v , P _fus and P _net is necessary to minimise LCOE. One can calculate several options for balancing listed in table 2. This table suggests the following.

1. (i) Options A and B suggest using low P _fus (1000 MW) reactors to minimise the cost for reactor assembly. At this power level, high $R_{\textit{v}}$ is required to maintain reasonable Q. Options C and D use high P _fus (3000 MW) to minimise $R_{\textit{v}}$ . Notice designing a reactor with P _fus higher than 3000 MW is not reasonable as it results in P _net exceeding 1200 MW and thus is difficult to use in real-grid scenarios.

2. (ii) Options A and C feature finite Q and thus require development of new, continuously operating heating systems (more on them next section), making cost predictions difficult. Thus, the cost and LCOE estimations for these set-ups were not calculated.

3. (iii) Options B and D have infinite Q due to sufficiently high $R_{\textit{v}}$ and P _fus , enabling us to design and estimate costs for VOYAGER components, specifically, plasma heating and other systems. Notice option B produces three times less net electric power, thus providing much lower revenue and requiring significant cost savings in both CAPEX and OPEX to achieve similar LCOE. Thus, option D was selected for further description in the article, though a qualitative comparison between options B and D is presented in § 3.

Table 2.

Options for balancing fusion power and vortex section factor in VOYAGER plant.

2.2.1. Fusion chamber dimensions

Fusion power radiates from a plasma into a reactor wall and into expander sections in the form of neutrons and gamma-radiation. A fusion chamber wall should withstand such power, which limits the minimum chamber inner surface. Taking into account complete burn of tritium produced by the DD reaction (due to the tritium recycling system described later), 68 % of the total fusion energy (2002 MW) is released as fast neutrons (table 1). They immediately leave the fusion chamber after synthesis, passing through the fusion chamber first wall and are captured by a blanket. The acceptable neutron load on the first wall is estimated in the literature as 2.5–5 MW m⁻² (Merola Reference Merola2009; Kuang et al. Reference Kuang2018; Lion et al. Reference Lion2025). For the VOYAGER option D with the highest expected fusion power (3000 MW) and assuming conservative 3 MW m⁻² neutron load, the inner surface should be no less than 690 m², translating to a 100 m long reactor central section with 1.1 m first wall radius. Those dimensions are large enough to accommodate plasma with dimensions proposed in § 2.2.2 and thus were used for subsequent cost estimations for reactor assembly.

2.2.2. Radiation protection

A significant portion of the plant energy is emitted as fast neutrons in the central section of the reactor requiring radiation protection for HTS coils. The rest of the energy is produced as fast ions that are absorbed by plasma receivers in expander sections with simpler radiation protection and heat dissipation design. The authors expect central section reactor components (wall, shielding, blanket and magnets) replacement every 25 years (maintenance period τ). Let us estimate the volume for protection as it influences magnet inner diameter, and overall system complexity and cost.

A blanket in the central section is placed between the plasma chamber and HTS magnets (figure 3). The blanket provides radiation protection for HTS magnets and collects heat for consequent conversion into electricity. The blanket consists of a tungsten first wall, vacuum vessel, liquid metal coolant (lead) and moderator (boron carbide) with a total 1.1 m thickness.

Figure 3.

Central section cross-section. 0 cm signifies inner winding of HTS magnet, 218 cm signifies reactor axis centre.

The HTS magnet should not be irradiated above the maximum HTS neutron fluence (Fischer et al. Reference Fischer, Prokopec, Emhofer and Eisterer2018) before maintenance (HTS magnet replacement). Neutron flux attenuation ( $\unicode{x03BC}$ ) during the maintenance period is calculated with neutron fluence ( $\varPhi _{\textit{HTS}}$ ) and values in the table 3 using (2.20)–(2.23).

Table 3.

Parameters for radiation protection.

Neutron flux attenuation ( $\unicode{x03BC}$ ) can be achieved with just 0.4 m of boron carbide (Thomas, Paluszek & Cohen Reference Thomas, Paluszek and Cohen2019) that can be located inside the 1.1 m thick blanket section, and thus enables blanket and magnet replacement within 25 years or 3 times during the 80-year expected plant lifetime.

(2.20) \begin{equation}E_{\textit{tot}}=P_{\textit{fus}}\tau,\end{equation}

(2.21) \begin{equation}N_{\textit{DD}}=\frac{E_{\textit{tot}}}{(E_{\textit{DD}}+E_{\textit{DT}})},\end{equation}

(2.22) \begin{equation}\varPhi _{\textit{HTS}}=\frac{\textit{MN}_{\textit{DD}}}{S_{\textit{magnet}}},\end{equation}

(2.23) \begin{equation}\mu =\frac{\varPhi _{\textit{HTS}}}{\varPhi _{\textit{max}}}.\end{equation}

2.2.3. Fuel handling system: injection, removal of fusion products and separation

Deuterium fuel enriched with by-product tritium is injected into the reactor via pellets similar as Baylor et al. (Reference Baylor2009) to keep the tritium concentration at 5 % in the plasma. The key feature of open traps is the natural channel for removing impurities and reaction products. VOYAGER is a ‘flow’ system, so the accumulation of impurities, including helium, heavy (from the wall) and light (such as deuterium) elements, is not expected. All substances from the plasma escape through the mirror section rather equally – the retention times for all elements are similar. The deuterium flow required to maintain balance is approximately 800 eq. A and fusion reactions give approximately 2–3 eq. A of helium (table 1); the ratio of their densities will be approximately less than 1 %. With such density ratios, even a low degree of purification would not affect plasma ignition (Z _eff ). Several approaches are suitable for separation: membrane diffusion (Willms Reference Willms2016) or gas centrifuge. Notice that separation is used not only for removal of non-reactive isotopes (specifically, helium) but also to acquire tritium for re-injection into plasma for increased fusion power (see § 2.2.2).

2.2.4. Magnet system design

The VOYAGER magnet system comprises four main components: central section, transition section, vortex section and mirror section (figure 1). General specifications for each section are given in table 4. The central section is designed modularly: it comprises several solenoids covering its whole length with spacing for diagnostics, piping, plasma heating, etc. Solenoids provide a sufficient magnetic field (5 T in normal operation, see figure 2), while providing enough space for radiation protection blanket. Modular design provides ease of assembly and maintenance. Vortex and mirror sections are designed with a 12 T magnet field, with a total length of 12 m each. The transition section is necessary for combining central and vortex sections, ensuring smooth field transition between central and mirror sections.

Table 4.

Magnet system specification.

Magnet system size, material usage and cost calculations are based on coil strength, HTS current density and magnet field modelling developed by SuperOx LLC. Strength estimations use the free-turn model taking into account the internal radius of the magnet (the most stressed part). The maximum load on the HTS cable does not exceed 800–1100 MPa, which is lower than the load values in similar projects (Dunn Reference Dunn2024) and was demonstrated by Moyzykh (Reference Moyzykh2023). The magnetic field is designed using the current density in the range of 100–200 A mm⁻² (achievable for HTS (Mounet Reference Mounet2022)).

Figure 4.

Dependence of plasma parameters versus the relative deuterium density for the second startup stage.

2.2.5. Fusion plasma startup and plasma heating equipment

During power production, the VOYAGER reactor operates in ignited state (see § 2.2.2 for details) and does not require plasma heating, essentially due to having high enough fusion power P _fus and vortex section factor $R_{\textit{v}}$ . However, during startup of the reactor fusion power starts from zero, requiring a dedicated startup procedure divided into two stages.

1. (i) The first stage assumes gas-dynamic plasma confinement, high external heating power and significant tritium content (provided externally). The magnetic field is reduced in the central reactor section, while in the other sections (vortex, mirror, expander), it remains normal. At the first stage, the plasma pressure increases to achieve ‘diamagnetic bubble’ configuration and ignition.

2. (ii) The second stage assumes the transition to a pure deuterium reactor. The portion of tritium in the plasma gradually decreases, along with a decrease in the reaction rate, which is compensated by an increase in density, temperature and magnetic field.

Let us consider the second stage first. It is modelled in the same way as in § 2.2.2, but external sources of tritium are taken into account. Plasma is already ignited and the goal is to decrease tritium content so the reactor can self-sustain its tritium content. To reduce the tritium content, the fusion power increases so that the plasma remains at the border of the ignition mode. The dependence of plasma parameters on the relative density of deuterium is shown in figure 4. The right edge corresponds to a pure deuterium reactor with a plasma tritium content of 0.9 %. This is the final configuration of the second stage. The initial configuration is selected based on the following considerations. On the one hand, the plasma parameters should be noticeably lower than the final ones. On the other hand, the magnetic field must remain large enough for the Larmor radius of fusion products to be acceptable. Thus, the initial configuration of the second stage was selected at 93 % D + 7 % T plasma, see table 5.

Table 5.

Initial configuration of the second startup stage.

Figure 5.

Dependence of plasma parameters on the relative pressure for the first startup stage.

As the reactor consumes external tritium during this stage, it is necessary to estimate its content to include the tritium cost in reactor OPEX maintenance. The duration of the startup second stage is limited by the magnet field ramp-up of the central section, taking 45 minutes, corresponding to up to 15 g of tritium consumption.

Now, let us consider first start up stage. Initially, the plasma at the first start-up stage has an open trap gas-dynamic plasma confinement (without a ‘diamagnetic bubble’). Plasma parameters are estimated similar to § 2.2.2, but with the following three main changes.

1. (i) The absence of a ‘bubble’ cancels the condition (2.7). Instead, the relative plasma pressure $\beta_{\bot}$ is introduced as a parameter:

   (2.24) \begin{equation}\beta _{\bot }=\frac{16nT}{B^{2}}, \quad B_{\textit{pl}}=\sqrt{1-\beta _{\bot }}B,\end{equation}
   where B is the vacuum magnetic field and B _pl is the reduced magnetic field in the plasma. The value of the vacuum field is chosen the same as at the beginning of the second stage: B = 1 T, see table 5.

2. (ii) The axial confinement time instead of (2.9) becomes gas-dynamic, taking into account beta and the vortex sections factor:

   (2.25) \begin{equation}\tau _{\parallel }=\frac{R_{\textit{v}}}{\sqrt{1-\beta _{\bot }}}\tau _{\textit{gd}}.\end{equation}

3. (iii) The plasma radius is no longer constant. Instead, the magnetic flux Φ = na ² B _pl is assumed to be conserved. The magnitude of the magnetic flux Φ is determined by the magnetic flux of the transition layer at the beginning of the second start-up stage. The term for the cyclotron loss power is also different from (2.11). However, since this power turned out to be small, the exact expression is not presented in the text.

The results of the evaluation of plasma parameters for the first start-up stage are shown in figure 5. The highest relative pressure value of $\beta_{\bot} = 99.5\,\%$ corresponds to the conditions of the ‘diamagnetic bubble’ formation with up to 70 MW plasma heating power required. Notice that modern studies demonstrate the formation of a ‘diamagnetic bubble’ in milliseconds (Roche et al. Reference Roche2025), indicating the first start up stage of the VOYAGER plant should occur in the sub-second format.

Let us consider how to setup 70 MW, 1-second plasma heating. At present, the Budker Institute has completed the development of a 100 keV, 1-second impulse – class injectors (Belchenko et al. Reference Belchenko2018; Shikhovtsev et al. Reference Shikhovtsev2024) for use in magnetic plasma confinement facilities. A batch of similar injectors can be used in the VOYAGER plant to achieve the necessary power. The injectors have a deuterium ion beam with a current of 75 A, the equilibrium yield of deuterium atoms at an energy of 100 keV is 0.44 and the power of the fast deuterium neutral beam of 3.3 MW. The capture coefficient of the beam injected into the inner region of the plasma is estimated as 0.92. As a result, the power injected into the plasma from one deuterium neutral beam is 3.0 MW. So to achieve a 70 MW heating startup power, 24 deuterium neutral beams are necessary – a number which was used in the later economic calculations.

The injectors are planned at both ends of the central section with an injection angle in the vicinity of 45°. Shutters separate injectors from fusion radiation after startup. Notice that the injectors are purely deuterium (tritium is injected separately via mixed D and T pellets) so as not to adapt existing technology to new beam atoms.

After initial heating of the enriched fuel reaches the ignition state, the second stage begins, featuring an absence of external heating, gradual increase in fusion power and reduction of tritium content to self-sustainable volume (roughly 1 %), reaching full fusion power.

2.2.6. Conversion to electric power

Most fusion power is released into the central section blanket (88 %, including neutrons, Bremsstrahlung and cyclotron radiation, radial diffusion from the transition layer), while the rest is absorbed in expander sections (12 %, diamagnetic trap losses). This energy is converted to electric power similarly to conventional liquid metal coolant nuclear power plants (Patel Reference Patel2016), which are also suitable for fusion (Rapisarda et al. Reference Rapisarda, Fernández-Berceruelo, García, García, Garcinuño, González, Moreno, Palermo, Urgorri and Ibarra2021).

The conversion system consists of two separated circuits: primary liquid metal circuit – for cooling the irradiated components (first wall, limiters, blanket, expanders etc.); and secondary water circuit – to power the turbines, with heat exchanger between the circuits. Magnetic field effects on the liquid metal flow are eliminated by placing coolant piping along the reactor (horizontal direction) to avoid coolant flow in the gradient magnetic field.

Due to the high neutron flux central section, the blanket contains a neutron moderator (boron carbide, see § 2.4) to absorb neutrons, which generates heat. This heat is collected with cooling pipes running through the moderator by liquid metal coolant, comprising a primary cooling circuit, operating at 540 °C and near-atmospheric pressure. The estimated coolant flow rate is 2400 kg s⁻¹, proposed by Candido et al. (Reference Candido2026). A possible coolant piping material can be Inconel 718 (Kuang et al. Reference Kuang2018) with a lifetime of 25 years given a moderately low operating temperature. The second circuit operates in subcritical conditions (7 MPa, 326 °C) or super-critical conditions (25 MPa, 530 °C). For such operational temperatures, one can expect an efficiency N in the range from 36 % (subcritical conditions (ROSATOM Group Reference Roche2025)) up to 46 % (for supercritical conditions (Kirillov et al. Reference Kirillov, Pomet’ko, Smirnov, Grabezhnaia, Pioro, Duffey and Khartabil2005)). The lower margin is similar to older nuclear power plants; the higher one to modern designs. Thus, with 3000 MW fusion power capacity, the gross electric power P _gross is between 1080 and 1380 MWe (heat to electricity conversion efficiency 36 %–46 %). The reactor’s self-consumption is estimated at 7 % of the gross power (Meekal & Majid Reference Meekal and Majid2024), and consequently, net electric power is expected in the 1004–1283 MW range.

Table 6.

Construction cost estimation.

2.3.1. Construction cost estimation

Construction costs were determined using a method similar to that for other fusion plant concepts described by Baylor et al. (Reference Baylor2009), Kuang et al. (Reference Kuang2018), Farina & Festa (Reference Farina and Festa2023), Shikhovtsev et al. (Reference Shikhovtsev2024) and The MANTA Collaboration et al. (2024). The expense list, along with their definition and approach for calculation for the VOYAGER concept, is given in table 6. The costs were calculated in the range to account for contingency and to demonstrate sensitivity to cost variation of different expenses. The range between min capital cost and max capital cost can be used as an estimate for subsequent sensitivity analysis.

The plant construction time is expected to be 5–8 years, similar to nuclear plants (Lovering, Yip & Nordhaus Reference Lovering, Yip and Nordhaus2016). However, since a fusion plant does not store significant volumes of fissile material and thus requires less effort for certification and contingency reasons, it is safe to assume the construction time will be at the minimum range, close to 5 years.

Table 7.

Material cost (stock exchange prices on January 2025).

2.3.2. Operation (maintenance) cost estimation

Plant maintenance consists of the following operations.

1. (i) Waste disposal covers removal and processing of plant fissile materials, which was studied by Kuang et al. (Reference Kuang2018) and Farina & Festa (Reference Farina and Festa2023) to be in the range of 2.7–3.3 MUSD annually. Since the VOYAGER plant does not operate with large quantities of fissile material (no tritium breeding), waste disposal cost can be expected to be as low as 10 % of a tritium breeding reactor, giving a 0.3 MUSD annual waste disposal cost.

2. (ii) Maintenance – studied by Farina & Festa (Reference Farina and Festa2023) as well, gives 61.4 MUSD for high-power plant (EUROfusion News 2022) and 46.6 for a lower power (Kuang et al. Reference Kuang2018). Since the high-power plant (EUROfusion News 2022) has a similar power to the VOYAGER plant, 61.4 MUSD was used as an estimation.

3. (iii) Components replacement was calculated directly, with the total component replacement cost esteemed over the plant lifetime (80 years) and averaged annually:

   1. (a) first wall and blanket replacement due to radiation erosion – 231 MUSD every 25 years, replacement time – 3 months;

   2. (b) magnet system replacement due to radiation damage to superconductor material – 322 MUSD every 25 years, replacement time – 3 months;

   3. (c) tritium for start-up after magnet and blanket replacement – 15 g (authors expert guess), 30 000 USD g⁻¹ (The MANTA Collaboration et al. 2024); total for one startup – 0.45 MUSD.

4. (iv) Decommissioning: 3 % of operation cost (Farina & Festa Reference Farina and Festa2023).

Costs for operation were combined to provide the resulting value, being 85 MUSD annually (see § 3, table 9). Since this value is the lowest in the range of similar commercial plant projects (Kuang et al. Reference Kuang2018, EUROfusion News 2022), it may be necessary to assume the resulting plant may be in the range between maximum and minimum operating costs (85–155 MUSD) as a risk factor for sensitivity analysis.

Maintenance downtime for fusion plants is studied by Farina & Festa (Reference Farina and Festa2023), providing a capacity factor between 71 % and 88 %, which is in a similar range to contemporary nuclear power plants (Nuclear Energy Institute 2024) and thus feels like a suitable range for the VOYAGER plant concept as well.

Table 8.

Comparison between several plant projects.