Automation in Construction 165 (2024) 105571
Contents lists available at ScienceDirect
Automation in Construction
journal homepage: www.elsevier.com/locate/autcon
3D printed sulfur-regolith concrete performance evaluation for waterless
extraterrestrial robotic construction
Ilerioluwa Giwa a , Mary Dempsey b , Michael Fiske c , Ali Kazemian a, d, *
a
Bert S Turner Department of Construction Management, Louisiana State University, Baton Rouge, LA 70803, United States of America
Department of Mechanical and Industrial Engineering, Louisiana State University, Baton Rouge, LA 70803, United States of America
c
Advanced Materials & Manufacturing, Jacobs Space Exploration Group, Huntsville, Alabama, AL 35806, United States of America
d
Division of Electrical and Computer Engineering, Louisiana State University, Baton Rouge, LA 70803, United States of America
b
A R T I C L E I N F O
A B S T R A C T
Keywords:
Planetary robotic construction
Sulfur-regolith concrete
Sulfur concrete
Construction 3D printing
Martian regolith
Vacuum conditions
By leveraging the capabilities of construction 3D printing, building structures in harsh extraterrestrial envi­
ronments is conceivable. Sulfur concrete is a waterless construction material that offers great potential to replace
Portland cement concrete (PCC) in extraterrestrial construction. The shape stability of 3D printed Martian sulfurregolith concrete (SRC) was found to benefit from a lower substrate layer temperature. However, this comes at
the cost of flexural strength, resulting in up to 53% strength loss. The printed SRC specimens demonstrated a
significantly faster strength development rate (gaining about 85% of the ultimate strength after only 12 h)
compared to the printed PCC. The printed SRC specimens also outperformed the PCC specimens in vacuum
conditions at higher temperatures. Furthermore, modifying the SRC materials with Dicyclopentadiene resulted in
up to 44% strength increase and minimized the sublimation rate of the printed specimens in vacuum, especially
at an elevated temperature.
1. Introduction
The possibility of a multi-planetary future for humanity has become
more conceivable following significant investments, advancements, and
innovations in the aerospace sector. NASA’s Artemis program aims to
land astronauts on the Moon in the next few years to establish a sus­
tained human presence and to enable a variety of scientific activities on
the Lunar surface [1]. These missions will also serve as a precursor and
test bed for technologies that will be used to support missions to Mars.
However, micrometeoroid bombardments, ionizing radiation, dusty
terrains, and extreme temperatures pose high risks to the safety of as­
tronauts who will live and work in the challenging space environment.
To promote successful exploration efforts and enable a long-term pres­
ence on the Moon and Mars, supporting infrastructure such as landing
and launch pads, habitats, hangers, research labs, and protective shields
are needed to protect humans and exploratory technological assets in
the harsh Lunar and Martian environments [2–4]. Given these inhospi­
table conditions and the limited available human workforce, conven­
tional construction methods that rely on manual efforts cannot be
adopted for extraterrestrial construction. Furthermore, the limited mo­
tion capabilities of suited astronauts on the Moon is another important
consideration highlighting the necessity of Lunar construction automa­
tion. Taking into account these constraints, extrusion-based Construc­
tion 3D Printing (C3DP) holds significant potential for construction in
such harsh and perilous conditions, owing to its robotic and autonomous
capabilities [5,6].
In the past decade on Earth, extrusion-based C3DP has attracted
significant attention in the construction industry as a promising robotic
technology for realizing freeform structures [7–15]. Complex structures
can be built efficiently using C3DP robots by avoiding the use of form­
works, thereby accruing significant cost savings, accelerating the con­
struction process, increasing job site safety, minimizing waste, and
reducing environmental impacts [7]. While C3DP was originally
conceived for constructing buildings and other infrastructure on Earth,
NASA has been considering this technology for off-world construction.
To explore the possibilities with C3DP, NASA organized a “3D Printed
Habitat Challenge” in 2015–2019, during which a total of $5 M was
awarded to the teams with solutions for planetary C3DP, with a focus on
designing and building shelters utilizing in-situ resources on Mars [16].
In particular, the viability of this robotic technology was demonstrated
by the construction of 1/3-scale habitats during this challenge. In
another effort, a 1700 square-foot habitat (Mars Dune Alpha) was 3D
* Corresponding author at: Bert S Turner Department of Construction Management, Louisiana State University, Baton Rouge, LA 70803, United States of America.
E-mail address: kazemian1@lsu.edu (A. Kazemian).
https://doi.org/10.1016/j.autcon.2024.105571
Received 18 October 2023; Received in revised form 9 June 2024; Accepted 14 June 2024
Available online 18 June 2024
0926-5805/© 2024 Elsevier B.V. All rights are reserved, including those for text and data mining, AI training, and similar technologies.
I. Giwa et al.
Automation in Construction 165 (2024) 105571
printed at NASA Johnson Space Center in 2020 using cementitious
materials to simulate a Mars habitat that will enable research studies on
crews living in the simulated habitat for one-year missions [17]. These
major efforts highlight the anticipated important role of C3DP as a ro­
botic planetary construction technology in future NASA missions.
To realize an efficient planetary C3DP process, in-situ resources on
the Moon and Mars must be used as printing materials to avoid the
prohibitive cost and risks of transporting construction materials from
Earth, which will also considerably reduce the launch mass of the space
vehicle [18]. Due to the relatively low cost and availability of Portland
cement concrete (PCC) on Earth, it is by far the most commonly used
material for traditional construction and C3DP. Beyond Earth, hydraulic
cement-based concrete remains a possible choice of material for space
construction due to the evidence of water on the Moon and Mars. In
2009, the presence of a significant amount of ice water (5.6% of
disturbed regolith mass that reached sunlight) was estimated by the
Lunar Crater Observation and Sensing Satellite (LCROSS) in the
permanently shadowed regions near the Lunar poles [19,20]. However,
the ingredients needed to produce PCC on the Moon and Mars are
complex to obtain [21]. Furthermore, PCC and other alternative Lunar
and Martian binder-based concretes such as Calcium Sulfo-Aluminate
(CSA) and Magnesium Oxy-Sulfate (MOS) require a significant amount
of water in their production that could otherwise be used for life support
or other exploration activities [22,23]. Sulfur concrete (SC), on the other
hand, is a waterless alternative material that seems to be well-suited for
planetary construction. Data collected from remote sensing missions and
returned soil samples from the Moon and Mars to date have verified the
presence of sulfur mineral – commonly in the form of troilite mineral
(Fe–S) [24–26]. The elemental sulfur, which serves as the binder in the
sulfur concrete composite, could be extracted from the sulfate/sulfide
mineral deposits in the Martian or Lunar surface using in-situ resource
utilization (ISRU) technologies [24]. On Earth, sulfur concrete has
shown promising properties such as superior mechanical strength
compared to conventional concrete [27]. Considering its unique prop­
erties and the feasibility of its production on the Moon and Mars using
in-situ resources, sulfur concrete is considered a possible choice of ma­
terial for extraterrestrial construction.
In the existing literature, the properties of sulfur concrete intended
for terrestrial and extraterrestrial construction have only been studied
by testing mold-cast specimens. To the best of the authors’ knowledge,
only one study (conducted in 2016 by Khoshnevis et al. [28]) has
explored the extrusion of sulfur concrete. These researchers studied the
effects of mixture composition and melting temperature on the proper­
ties and deformations of printed sulfur concrete made with natural sand.
Therefore, there is very limited knowledge of the performance of 3D
printed sulfur concrete, and there is no information available in the
existing literature on the properties of 3D printed sulfur-regolith con­
crete (SRC) made with high-fidelity regolith simulants. The main goal of
this study is to characterize the fresh-state properties, mechanical
properties, and microstructure of 3D printed SRC specimens intended
for extraterrestrial construction. To assess the suitability of SRC as a
resilient and sustainable construction material, the impact of modifying
SRC with Dicyclopentadiene (DCPD) and the recyclability prospect of
printed SRC specimens was explored. Furthermore, the behavior and
performance of 3D printed SRC in comparison to its PCC counterpart
were studied in space-representative conditions (under vacuum condi­
tions at different temperatures). The findings from this study will pro­
vide insights into the feasibility of extrusion-based C3DP using sulfurregolith concrete as a viable technology for space construction and to
support NASA’s missions in the near future.
desulfurizing crude oil and gas was found in construction as well. Early
studies investigating the properties of sulfur concrete have demon­
strated that sulfur concrete has unique properties such as superior me­
chanical strength, rapid hardening, and resilience in acidic or saline
environments [27,29–31]. As such, sulfur concrete has been extensively
used in the production of corrosion-resistant sewer pipes, roadblocks
and sidewalks, foundations, railway ties and sleepers, bridge decks, and
acid or sewage tanks [32]. Sulfur used in construction applications is
typically modified with organic or inorganic additives such as bitumen,
styrene, or dicyclopentadiene (DCPD). The purpose of modifying the
sulfur binder is to enhance the properties and durability of the sulfur
concrete material by primarily preventing the inherent crystallization of
sulfur crystals [27]. The S8 sulfur exists in three main polymorphs
(orthorhombic (α) sulfur, monoclinic (β) sulfur, and plastic (γ) sulfur).
These polymorphs show different properties such as color, density,
melting point, and stability, and undergo allotropic transformation at
different temperatures. The α- sulfur is the most stable allotrope at
ambient conditions and is obtained by cooling molten sulfur below
96 ◦ C. The α‑sulfur consists of puckered rings of eight sulfur atoms ar­
ranged in a rhombic lattice. The β ‑sulfur is characterized by needle-like
crystals and is found to be only stable between 96 ◦ C and its melting
point (119 ◦ C). The γ‑sulfur which has a helical structure consists of
various intramolecular allotropes and is obtained by rapidly cooling
liquid sulfur melted above 150 ◦ C. The γ‑sulfur is the densest and least
stable of the three allotropes. In the context of sulfur concrete, when
pure sulfur concrete is rapidly quenched from a melting temperature
above 159 ◦ C during production, it mainly consists of an equilibrium
mixture of metastable polymeric (μ) sulfur (possessing long molecular
chains) and cycloocta‑sulfur that acts as the plasticizer [33,34]. Within a
few hours, the polymeric sulfur concrete which possesses plastic prop­
erties becomes brittle as it rapidly depolymerizes into the orthorhombic
species when cooled at ambient conditions [35]. On the other hand, at
temperatures below 159 ◦ C, the sulfur concrete principally contains
cycloocta sulfur species that rapidly solidify and crystallize into the
metastable monoclinic (β) species. It then rapidly transforms into a
composite that contains the denser orthorhombic (α) species at tem­
peratures below 96 ◦ C [34,36]. As a result of the allotropic trans­
formation of the sulfur crystals between the different polymorphs, there
is a volumetric change (~7% decrease in volume) as the sulfur crystal­
lizes from the monoclinic to the denser orthorhombic polymorph
[27,37]. As such, shrinkage and internal stresses can occur within the
composite and lead to the formation of cracks and defects that can
significantly affect the strength and durability. Hence, to address this
issue, chemical modifiers have been proposed and tested to plasticize
sulfur by cross-linking the sulfur bi-radicals with some organic co­
monomers at either low or high (termed inverse vulcanization) reaction
temperatures [34]. Although the modification of sulfur using such
organic additives can improve the structural properties and durability,
the modifiers (such as DCPD) are expensive and emit an offensive odor
[38].
2.1. Properties of sulfur concrete for terrestrial applications
An important aspect of sulfur concrete production is its workability
and fresh-state properties, which have not been investigated extensively.
Gwon et al. explored the rheological properties (yield stress and plastic
viscosity) of the DCDP-modified sulfur composite. These researchers
reported a considerable increase in these rheological properties as the
temperature was raised from 120 ◦ C to 140 ◦ C, and when a larger
amount of micro-fillers (a blend of Portland cement and fly ash) were
added to the composite [39]. In the only published study on the prop­
erties of 3D-printed sulfur concrete, Khoshnevis et al. reported that the
printing temperature (140 ◦ C and 150 ◦ C) and sulfur content (30% and
35% by weight) are two critical parameters that impact the fresh
properties, printability, and microstructure of printed sulfur concrete
[28]. The researchers revealed that extruding sulfur concrete at 150 ◦ C
2. Background
Since prehistoric times, sulfur has been a commonly used commodity
in medicine, textiles, metalwork, and military hardware. Following
World War I, the alternative use of sulfur obtained as a byproduct of
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I. Giwa et al.
Automation in Construction 165 (2024) 105571
yielded a notable improvement in pore distribution throughout the
sample compared to at 140 ◦ C. Additionally, higher sulfur content (35%)
resulted in a reduced layer quality when compared to layers extruded at
a lower sulfur content (30%). Similarly, previous studies have explored
the mechanical properties of mold-cast sulfur concrete made with nat­
ural mineral aggregates. Gregor and Hackl revealed that the aggregate
particle size distribution and aggregate type (angular or spherical) have
significant impacts on the mechanical strength of unmodified sulfur
concrete [40]. Furthermore, the same study reported that the
compressive and flexural strength was increased by 51% and 33%,
respectively, when the sulfur binder was modified with DCPD (3% of
sulfur weight). Dugarte et al. considered the effect of sulfur content
(20–50% by weight) on the mechanical properties of DCPD-modified
sulfur concrete (5% DCPD loading). These researchers found that the
maximum compressive strength (27 MPa) was achieved at 30% sulfur
content, and 73% of the ultimate strength was attained within the first
24 h [41]. Regarding the reusability of sulfur concrete, Gulzar et al.
reported a 30% improvement in the compressive strength of sulfur
polymer concrete when it was remelted, remixed, and remolded after 48
h. However, a drastic reduction in strength was reported on the second
attempt at recycling the material, mainly due to the loss of the sulfur
binder [42].
well understood. Therefore, the properties and performance of DCPDmodified SRC are also investigated in this study under representative
space conditions.
Based on the extensive literature presented, most of these studies
have only investigated the performance of mold-cast sulfur concrete in
the context of terrestrial and extraterrestrial applications. However,
since no study has focused on the evaluation of 3D printed sulfur con­
crete, there is a knowledge gap on the behavior and properties of
additively manufactured sulfur concrete for earth and beyond.
3. Methodology
The primary objective of this study is to investigate the fresh and
hardened state properties of 3D printed SRC materials. Furthermore, the
performance of the printed SRC specimens was compared to the per­
formance of similarly printed PCC specimens under near-vacuum con­
ditions and temperature variations. The conventional properties of the
printing mixtures were first characterized by studying the fresh prop­
erties and the mechanical strength of mold-cast specimens. Next, the
printability of the sulfur-regolith mixtures was studied by evaluating the
effects of the two main printing process parameters, namely, printing
temperature and interlayer time gap, on the shape stability of the SRC
extrudates. Two levels of printing temperature (PT) - 130 ◦ C and 150 ◦ C
[±5 ◦ C] - and three levels of the interlayer time gaps (I) - 45 s, 120 s, and
180 s - were selected. These parameters are selected given the impact of
temperature on the rheology (viscosity) of sulfur. Hence, the tempera­
ture levels were selected to remain above the melting point of sulfur
(about 120 ◦ C) but below the ring-opening polymerization (ROP) floor
temperature of sulfur (where the properties of sulfur concrete such as
the viscosity drastically change and may become too stiff to extrude).
Furthermore, since the deposited material temperature will change with
time, it is also important to understand how the layer temperature
(given by the interlayer time gap) could affect the buildability and
interlayer bond strength. The interlayer delay values were selected to
account for the rapid solidification of sulfur at room temperature.
Additionally, the layer cooling process was accelerated by using an air
circulation fan to ensure uniform heat loss across the printed specimens
during the interlayer delay period. The accelerated cooling was only
performed during the extended interlayer time delay periods beyond 45
s. Consequently, the effects of these process parameters on the shape
stability and mechanical properties were studied, resulting in the se­
lection of the optimal parameters used for the next round of testing. In
addition, the effect of DCPD modification and recyclability of SRC were
also studied.
2.2. Properties of sulfur concrete for extraterrestrial applications
Considering the favorable properties of sulfur concrete in terrestrial
applications, it also has great potential for planetary construction in
locations where the surface temperature is safely below temperatures at
which sulfur can become unstable (>110 ◦ C). For example, sulfur con­
crete can be used in the permanently shadowed region (PSR) of the
Moon, where the surface temperature can be as low as -180 ◦ C, or in
most locations on Mars where the maximum daily temperature is about
20 ◦ C [22]. In several efforts to assess the suitability of sulfur concrete
for planetary applications, previous research studies have reported the
performance of mold-cast sulfur-regolith concrete made with Lunar and
Martian regolith simulants. Omar et al. (1993) reported compressive and
tensile strength values of 34 MPa and 4 MPa, respectively, for SRC
materials made with JSC-1 Lunar simulant and an optimum sulfur
content of 35% [43]. Wan et al. reported superior compressive strength
performance of SRC made with JSC-1A regolith simulant (>50 MPa at
50% sulfur content) compared to sulfur concrete made with natural sand
(25–28.3 MPa at 15–35% sulfur content) [44]. A study by Toutanji et al.
reported that the compressive strength of JSC-1 Lunar sulfur-regolith
concrete made with 35% sulfur content was about 28% higher than
JSC-1 Lunar Portland cement concrete with a 0.42 water-to-cement ratio
[45]. Grugel et al. also found that the average compressive strength of
the samples exposed to Lunar temperature cycles (− 191 ◦ C and 20 ◦ C in
80 cycles) was five times lower than the control samples (unexposed)
due to the weakened bonding between the sulfur and aggregate particles
[46]. A primary concern with using sulfur concrete for planetary con­
struction is sulfur sublimation under vacuum conditions. Due to weak
intermolecular forces, sulfur can sublimate at temperatures and pres­
sures below its relatively low triple point, resulting in the transitioning
of solid sulfur molecules with higher vapor pressure directly to gas
[47–49]. Based on numerical models and experimental data collected at
a vacuum pressure of 5 × 10− 7 Torr at room temperature for 60 days,
Grugel and Toutanji estimated a time duration of 6.5 years and 4.4 years
for 10 mm thick sulfur concrete specimens made with 30% and 43%
sulfur content to sublimate, respectively [47]. Rahim et al. also reported
up to 48% reduction in the compressive strength of the unmodified
Martian SRC cured in NaOH solution for up to 7 days [50]. This finding
reveals some potential durability concerns for unmodified sulfur con­
crete considering the basic nature of Martian soils. Currently, there is
limited data available on the sulfur crystallization process under nonstandard atmospheric and environmental conditions in space, and the
performance of SRC intended for space construction in this context is not
3.1. Materials and mixture proportions
In this study, the SRC materials were prepared using a commercially
available elemental sulfur powder (99.9% purity) as the binder, and
Mars Global Simulant (MGS-1) as the fine aggregate. MGS-1 is a highfidelity analog that is representative of the basaltic soil found on the
Martian Gale Crater (manufactured by Exolith Lab) [51]. The specific
gravity and mean particle size of the regolith simulant were determined
to be 3.085 g/cm3 and 90 μm, respectively. Table 1 presents the
chemical composition of the MGS-1 regolith simulant. Based on pre­
liminary trials, a sulfur content of 33% (by weight) was selected to
develop printing materials with the MGS-1 regolith simulant. Per ACI
548.2R, the unmodified mixture was prepared by first melting the sulfur
powder at the designed testing temperature and preheating the regolith
simulant to ~170 ◦ C in the oven for 2 h [52]. The heated ingredients
were then mixed for two minutes in a three-stage mixing sequence
(Fig. 1). Each mixing stage is followed by placing the mixture in the oven
at the testing temperature for three minutes. A total mixing duration of
15 min was adopted for each batch. The modified SRC mixtures (MSRC)
were prepared by using 10% DCPD (% sulfur weight) to modify the
sulfur before the regolith simulant addition. The DCPD (supplied by
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Automation in Construction 165 (2024) 105571
Table 1
Chemical composition of the MGS-1 regolith simulant [51].
Oxides
Si02
Ti02
Al203
FeO
MnO
MgO
CaO
Na2O
K2O
P2O5
LOI
Wt. (%)
42.9
0.6
12.8
11.2
0.1
14.6
7.4
1.5
0.6
0.1
5.5
Fig. 1. Plain and modified SRC preparation process.
Millipore Sigma) was mixed with molten sulfur at 140 ◦ C under a fume
extraction system, and continuously stirred for 2 h. The modified sulfur
is then mixed with the preheated regolith following the three-stage
mixing sequence described for the unmodified mixture. To evaluate
the performance of the SRC in comparison to conventional concrete,
PCC samples were also 3D printed and tested. The PCC mixtures were
prepared using Type I/II Portland cement (ASTM C150) as the binder
and silica sand (maximum particle size of 1 mm) as the fine aggregate. A
commercially available limestone (LS) powder – HuberCrete Preferred
grade - with an average particle size of 6 μm and specific gravity of 2.7
g/cm3 was used to replace the cement (15% cement replacement) in the
matrix to reduce the cement content. A water/binder ratio of 0.40 was
selected and a high-range water-reducing admixture (HRWRA) - ADVA
CAST 585 - was used to achieve the required flowability for a successful
print. The cement, sand, and LS powder were first dry mixed for 3 min
before water and the necessary amount of HRWRA was added and mixed
for an additional 7 min. Both printing materials (SRC and PCC) were
designed to have similar binder volumes. Hence, the total volume of the
molten sulfur in the SRC mixture as well as the binder paste of the PCC
(consisting of Portland cement, LS, and water) was maintained at 0.42
m3/m3. The SRC and PCC mixture proportions are presented in Table 2.
were printed using a printing system designated for cementitious ma­
terials. The PCC printing system has a build envelope of 2.1 m (length) ×
0.9 m (height) and a closed-loop extrusion system (Fig. 2b). Both
printing systems use a rectangular nozzle with 40 mm (width) by 20 mm
(height) dimensions and side trowels to improve the surface quality of
the extrudate. The printing process for the SRC and PCC materials was
carried out at a constant printing speed of 5 mm/s and 40 mm/s,
respectively, while the extrusion rate was adjusted in both cases to
achieve the desired layer dimensions. Considering the differences in the
printing systems used for both printing materials, the optimal print
parameters were selected based on the preliminary investigations. For
both experiment groups, the printing materials were first prepared and
mixed (SRC or PCC), and then were manually fed into the extrusion
system (after reaching the printing temperature in the case of SRC). A
fume extraction system as well as personal protective equipment were
also utilized during the high temperature printing experiments.
3.3. Fresh-state properties and shape stability tests
The conventional properties of the SRC and PCC were evaluated by
measuring the wet density, and flowability. The flowability of the
mixtures was measured using a flow table per ASTM C1437 and a slump
test using a mini-slump cone (half the standard slump cone size
described in ASTM C143). To prevent the material from rapid solidifi­
cation upon contact, the testing procedure for the SRC was modified by
heating the surface of the flow table, flow mold, and mini-slump cone to
120 ± 10 ◦ C before starting each experiment. In the flow table test, the
material was placed inside the mold and the table was jolted 15 times in
10 s. The spread diameter of the material following the flow table test
and the slump in height following the mini-slump test were measured
accordingly. To assess the printability of the SRC mixtures, a shape
stability test similar to the test proposed by [53] was adopted to study
the effect of the printing process on the ability of the unmodified SRC
mixture to support its weight and the weight of the successive layers
(Fig. 3a). The shape stability test involves printing a two-layer specimen
3.2. Printing systems
Two custom-made extrusion-based printing systems were used in this
study for the SRC and PCC materials. The SRC specimens were printed
using a gantry-type printing system developed for high-temperature
materials (Fig. 2a). The printing system has a build envelope of 1.2 m
(length) × 0.3 m (height), and the printhead was implemented with a
close-loop heating control mechanism. Thermocouples connected to a
PID temperature controller were installed on the extruder to maintain
the desired temperature levels. The high-temperature printing system
utilizes a screw-type extrusion system with a printhead composed of an
insulated aluminum extrusion tube, a steel auger, insertion heaters, and
a closed-loop DC servo motor. On the other hand, the PCC specimens
Table 2
Mixture Proportions for the SRC, MSRC, and PCC printing materials.
Mixture ID
Sulfur
(kg/m3)
Regolith1 (kg/m3)
DCPD (kg/m3)
Cement (kg/m3)
Sand1 (kg/m3)
Limestone Powder
(kg/m3)
Water (kg/m3)
HRWRA2
(%)
SRC
MSRC
PCC
872
799
–
1744
1596
–
–
80
–
–
–
510
–
–
1464
–
–
76
–
–
234
–
–
0.36
1
2
Saturated surface dried (SSD).
Percentage of cement by mass.
4
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Automation in Construction 165 (2024) 105571
Fig. 2. (a) High-temperature sulfur concrete printing system (b) Cementitious material printing system.
Fig. 3. (a) Shape stability test (b) Temperature measurements using an infrared camera during the shape stability test.
at the designated interlayer time gap and measuring the change in the
height (if any) of the bottom layer due to plastic deformation. During
this test, a maximum dimensional error of 10% of the nozzle dimensions
was allowed for each printed layer to ensure consistency. Temperature
data of the printed layers was collected during the shape stability test
using a FLIR infrared camera (Fig. 3b). Preliminary tests confirmed a
maximum temperature difference of under 8 ◦ C between the FLIR
camera measurements (layer surface) versus the internal layer temper­
atures (measured by a thermometer with its probe inserted inside the
same layer). As such, and considering the short interlayer delays
selected in this study, infrared camera measurements were used for layer
temperature measurements.
DC motor is connected to the four-sided vane (diameter and height =
101.6 mm) that shears the mixture in a 3.5-l stainless steel container
(container diameter and height = 178 mm). To maintain the testing
temperature, flexible heaters were wrapped around the container in
addition to a hotplate which was placed under the measurement
container. The rheological properties were measured by collecting data
on the electrical power consumption of the driving motor during the
test. Based on these power consumption data, the torque measurements
were estimated and used to calculate the rheological properties of the
material (Eq. 1). Two test protocols, namely, the constant shear rate test
(CSRT) and flow curve test (FCT), were followed to determine the yield
stress and plastic viscosity, respectively. In the CSRT, the material was
sheared at a constant rotational velocity of 2 RPM for a test duration of
120 s (Fig. 4b). The peak torque (to initiate flow) was measured and
reported. Furthermore, the dynamic yield stress and plastic viscosity
were determined using the FCT protocol (Fig. 4c). First, the material was
pre-sheared for 20 s to maintain a consistent shear history for all the
tested mixtures. The test was performed at a descending shear rate (from
30 RPM to 18 RPM) to obtain the up-curve used to determine the
Bingham parameters (dynamic yield stress and plastic viscosity). It
should be mentioned that the developed rheometer was initially
3.4. Rheological properties using custom-made rheometer
A custom-made coaxial cylinder-type rheometer with a four-sided
vane was developed for measuring the yield stress and plastic viscos­
ity of the SRC printing material (Fig. 4a). The principal components of
the developed SRC rheometer comprise a brushed DC motor with an
encoder, a Roboclaw used for closed-loop motor control, and a Rasp­
berry Pi 4 single-board computer for processing and data collection. The
5
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Automation in Construction 165 (2024) 105571
Fig. 4. (a) Custom-made rheometer for sulfur concrete (b) Constant shear rate test protocol (c) Flow curve test protocol.
calibrated and verified using a standard viscosity fluid of known vis­
cosity (30,000 cP) and several concrete mixtures with different flow
properties (rheological properties as measured by the ICAR plus com­
mercial PCC rheometer) [54]. The fundamental parameters (shear
stress) were determined from the relative parameters (torque) using the
Reiner-Riwlin equation. The rheological properties of the SRC mixtures
were measured at different temperatures ranging from 125 ◦ C – 165 ◦ C.
In addition, the rheological properties of the PCC printing mixture were
also determined.
τGM = (IM − INL )Kt Nη
gearing efficiency (%).
3.5. Hardened-state properties tests
The dry density and compressive strength of the mold-cast SRC and
PCC specimens were determined using 50 mm cubes per ASTM C109.
The compressive strength test was performed at a loading rate of 0.25
MPa/s (Fig. 5a). A three-point bending test was performed on the cast
and 3D printed specimens (tested in the X or Z directions, in-plane)
following ASTM C293 and at a displacement rate of 0.2 mm/min
(Fig. 5b). The prism specimens used for the flexural strength test
measured 40 mm in width, 40 mm in height, and 160 mm in length. The
cast and printed SRC specimens were stored in dry room conditions
(20 ◦ C [±◦ 5 ◦ C] and RH >65%), while the PCC specimens were cured at
(1)
where τGM is the torque generated by the gear motor (oz-in), IM is the
motor current (A), INL is the motor no-load current constant (A), Kt is the
motor torque constant (oz-in/A), and N is the gear ratio, and η is the
Fig. 5. Mechanical performance tests (a) Compression test (mold-cast SRC specimen) (b) Three-point bending test of printed SRC specimen under different
loading directions.
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dry room conditions or in a water tank at room temperature. Printed
specimens that were fabricated during the shape stability test were also
tested after 24 h considering the rapid strength development of sulfur
concrete. A vacuum chamber with an ultimate vacuum pressure of 0.07
Torr (when unloaded) was utilized to measure the effect of vacuum
condition and temperature variations – 20 ◦ C and 55 ◦ C [±5 ◦ C] - on the
weight loss and mechanical strength of the printed specimen. The
weight loss was measured after 3 days and 7 days, respectively, using a
digital scale with a 0.01 g resolution. Similarly, the effect of these
conditions on the flexural strength of the printed specimens was
measured after 3 days and 7 days. The printed SRC and PCC specimens
were placed in the vacuum chamber 4 h and 8 h, respectively, after
fabrication to allow initial strength development. To study the changes
in the microstructural morphology of the pure sulfur and SRC specimens
due to DCPD modification and sample aging, scanning electron micro­
scopy (SEM) was also used to image the samples using a JEOL JSM6610LV SEM device with an Energy Dispersive X-ray (EDX) detector
used for compositional analysis. The SEM characterization was carried
out on fragmented samples sputter coated with platinum at an acceler­
ating voltage of up to 30 kV.
steady rise in the measured SRC yield stress from 145 ◦ C to 165 ◦ C. On
the other hand, the plastic viscosity of the SRC decreased from 125 ◦ C to
145 ◦ C, and then began to increase when the temperature was beyond
145 ◦ C until 165 ◦ C. The decreasing plastic viscosity trend from 125 ◦ C
to 155 ◦ C is consistent with the workability results presented before.
Furthermore, the observed trend is similar to the published data on the
viscosity of plain molten sulfur, in which the viscosity decreased from
0.3 to 0.07 Pa.s as the temperature increased from 120 ◦ C to 159 ◦ C,
before the viscosity rapidly increased to about 0.932 Pa.s at 190 ◦ C
[55,56]. Based on the obtained SRC viscosity data and comparing them
with the existing literature on plain sulfur, it is concluded that the
addition of regolith particles to molten sulfur increases its plastic vis­
cosity by up to 2 orders of magnitude. Moreover, the increase in yield
stress and plastic viscosity at temperatures around 150 ◦ C and higher
can be attributed to the formation of cross-linked catena‑sulfur chains
due to the ring-opening polymerization reactions of sulfur bi-radicals
when heated above 150 ◦ C. For comparison purposes, the static and
dynamic yield stress of the PCC mixture was measured as 870 Pa and
440 Pa, respectively, while the plastic viscosity was measured as 34 Pa.s.
The higher yield stress and plastic viscosity values of the PCC material
compared to SRC material at temperatures between 125 ◦ C to 155 ◦ C,
are consistent with the lower flowability and slump values reported
earlier.
4. Results and discussion
4.1. Characterization of the printing materials (SRC and PCC)
4.1.2. Hardened-state properties (mold-cast specimens)
Fig. 7 shows the 7-day compressive and flexural strength of the SRC
and PCC mold-cast specimens. The SRC specimens prepared at 155 ◦ C
showed comparable compressive and flexural strength values to speci­
mens prepared at 130 ◦ C. For PCC, it was observed that the compressive
and flexural strength values of the water cured mold-cast specimens
were 57% and 29% higher, respectively, compared to the strength
values of the air-dried specimens. Given the widely established rapid
strength development of sulfur concrete, where it achieves its ultimate
strength within a few days, the 7-day strength was regarded as the ul­
timate strength, and no measurements were conducted at 28 days. On
the other hand, the 28-day compressive and flexural strength of the PCC
specimens subjected to water curing were determined to be 34.1 MPa
and 6.4 MPa, respectively, while the strength values for air-dried PCC
specimens were found to be 28.9 MPa and 5.82 MPa, respectively.
Interestingly, these obtained data show that even after 28 days of water
curing, the strength of the mold-cast PCC specimens did not reach the
same levels as those achieved after 7 days by the mold-cast SRC
specimens.
4.1.1. Fresh-state properties
The average wet density values of the SRC and PCC printing mate­
rials were determined to be 2.310 g/cm3 and 2.096 g/cm3, respectively,
which show a higher density value of the SRC compared to PCC given
the similar binder volume of the two materials. The workability test
results show an increase in the flowability and slump of the SRC with an
increase in material temperature. The average flow diameters of the SRC
material at 130 ◦ C and 150 ◦ C were 172.5 mm and 194.5 mm, respec­
tively. Similarly, the average slump values at 130 ◦ C and 150 ◦ C were
measured as 66 mm and 98.5 mm, respectively. The increase in flow­
ability and slump with an increase in the material temperature (130 ◦ C
to 150 ◦ C) reveals the temperature-dependent behavior of the molten
sulfur within the SRC matrix. On the other hand, the flow diameter and
slump of the PCC were measured as 165 mm and 30 mm, respectively.
The printable PCC material exhibits lower flowability and slump values
compared to the printable SRC material due to the differences in their
composition and rheology. Fig. 6(a-b) presents the yield stress and
plastic viscosity values based on the high-temperature rheological
measurements performed on the SRC at different temperatures. At
temperatures between 125 ◦ C and 155 ◦ C, there were no significant
changes in the yield stress (static and dynamic). However, there was a
Fig. 6. Measured rheological properties of SRC materials at different temperatures (a) Static and dynamic yield stress (b) Plastic viscosity.
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Fig. 7. Mechanical properties of the mold-cast specimens (a) 7-days compressive strength (b) 7-days flexural strength.
4.2. Performance of printed SRC specimens
in the plastic state with minimal strength to sustain the loads associated
with the deposition of the next layer (nozzle extrusion pressure and the
weight of the next layer). The relatively high wet density of SRC mate­
rials (2.310 g/cm3) - which translates into heavier layers - also con­
tributes to larger plastic deformations in this scenario. In contrast,
longer interlayer delays (120 s and 180 s) allow the bottom layer to
solidify as the layer temperature decreases, forming a stable layer with
sufficient strength to support successive layers without any visible
deformation. Furthermore, the deformations measured for layers prin­
ted at 130 ◦ C were less than those measured at 150 ◦ C (about 20% less
deformation), which is associated with the lower bottom layer temper­
atures before the next layer was printed.
4.2.1. Effects of process parameters on shape stability
Fig. 8 shows the cross-sectional profiles of the printed SRC specimens
following the shape stability test. Based on the results, it was observed
that layer deformations solely occurred at the shortest interlayer time
gap (45 s) for both printing temperatures (130 ◦ C and 150 ◦ C). With a 45
s interlayer delay, layers printed at 130 ◦ C and 150 ◦ C exhibited about
10% and 23% deformation in layer height, respectively. In contrast, no
visible layer deformation was observed at the longer interlayer time
gaps (120 s and 180 s) at both temperatures. Considering the sulfur
solidification process, the observed deformations are attributed to the
bottom layer temperature, before the deposition of the next layer.
Table 3 presents the average temperature values measured for the bot­
tom layers using an infrared FLIR camera during the shape stability test.
Based on these measurements, with an interlayer time gap of 45 s, the
bottom layer temperature was near or higher than the melting point of
sulfur for both printing temperatures. Hence, the bottom layer was still
4.2.2. Effects of process parameters on flexural strength s
Subsequently, the shape stability specimens from the previous stage
were tested to determine the effect of the printing process parameters on
flexural strength. Based on the results presented in Fig. 9a, specimens
printed at 150 ◦ C showed up to 19% higher flexural strength compared
Fig. 8. Cross-sectional profiles of printed SRC specimens following the shape stability test.
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Table 3
Average bottom layer temperatures during the shape stability test.
Sample ID
Bottom layer temp. ( C)
◦
PT130I45s
PT130I120s
PT130I180s
PT150I45s
PT150I120s
PT150I180s
120.7 [1.6]
102.2 [1.1]
94.2[0.9]
132.2 [2.6]
105.2 [1.1]
102.1[3.7]
Fig. 9. Flexural strength results (a) shape stability 2-layer specimens tested in the X-direction (in-plane) (b) 8-layer specimens tested in the Z-direction (in-plane).
to the ones printed at 130 ◦ C with the same interlayer delay. In addition,
SRC specimens printed with longer interlayer time gaps (120 s and 180
s) showed up to 23% lower flexural strength compared to specimens
printed with a shorter interlayer delay (45 s), at both printing temper­
atures. The observed reduction in flexural strength for the longer
interlayer delays can be attributed to the formation of cold joints arising
from poor bonding between the newly deposited layer (with a higher
temperature) and the preceding layer with a lower temperature.
Considering the data presented in Table 3, specimens printed with
longer interlayer delays have lower bottom layer temperatures and have
already transitioned to the solid state. This results in a weaker interlayer
bond compared to the scenario with both layers in the plastic state (45 s
interlayer delay). Hence, the results demonstrate that higher layer
temperature is beneficial for interface adhesion. The results also show
that increasing the printing temperature from 130 ◦ C to 150 ◦ C lessens
the negative impact of the longer interlayer time gaps (120 s and 180 s)
on flexural strength, owing to the elevated temperature of the newly
printed layers which enhances interlayer adhesion. Interestingly, for
specimens printed at 130 ◦ C and with longer interlayer delays, the
interlayer interface was visible (which indicates a higher interface
porosity and weaker bonding), while no interface could be observed for
specimens printed at 150 ◦ C (Fig. 8). This observation provides addi­
tional confirmation of the beneficial effects of the elevated printing
temperature on interlayer adhesion.
To further study the effects of printing temperature and interlayer
delay on interlayer adhesion in 3D printed SRC elements and eliminate
the potential impact of layer deformations on accurate strength mea­
surements, 8-layer SRC specimens (40 mm × 40 mm × 160 mm) with no
deformations were printed and tested under flexure (Z-direction inplane) by loading the specimen parallel to the interlayer interfaces.
Similar printing temperatures (130 ◦ C and 150 ◦ C) and interlayer delays
corresponding to substrate layer temperatures of 55 ◦ C and 85 ◦ C
[±5 ◦ C] were considered. In other words, each new layer of the 8-layer
specimen was only deposited when the measured temperature of the
preceding layer reached the target substrate temperature value. These
layer temperature levels (interlayer delays) were selected based on
practical considerations and the logistics required to print larger speci­
mens. Fig. 9b presents the flexural strength results of the 8-layer SRC
specimens. Similar to the results obtained for the two-layer specimens
tested in the X-direction, Z-direction specimens printed at 150 ◦ C
showed higher flexural strength values compared to 130 ◦ C specimens.
Furthermore, the flexural strength of the specimens printed with inter­
layer delays equivalent to an 85 ◦ C substrate temperature was 113% and
41% higher at 130 ◦ C and 150 ◦ C, respectively, compared to specimens
printed with a 55 ◦ C substrate temperature. These results confirm the
importance of the printing temperature and interlayer delay on the
quality of interlayer adhesion which in turn, affects the structural and
anisotropic properties of 3D printed structures. Based on this finding, a
surface heat treatment approach - similar to [57] - could be envisioned
for the SRC printing process to improve interlayer adhesion by reheating
the previous layer before the new layer is deposited. Unlike PCC 3D
printing where time is the critical factor due to the ongoing timedependent chemical reactions, in the case of SRC 3D printing, the
layer temperatures mainly determine the properties of the printed
element. Future research is needed to better evaluate and quantify the
implications of a layer reheating approach on the SRC 3D printed ele­
ments. Based on these findings, an optimal printing temperature of
150 ◦ C and an interlayer time gap of 120 s was determined and selected
for further investigation on the SRC properties.
4.2.3. Effects of DCPD modification
Since DCPD was added to the MSRC printing material, it exhibited a
higher level of flowability during the initial printing trials due to the
higher liquid-to-solid ratio compared to the SRC printing material with
the same regolith/sulfur ratio for a similar volume of material. As such, a
reduced printing temperature of 120 ◦ C [±5 ◦ C] was necessary for
successfully fabricating the MSRC specimens with minimal de­
formations. The 7-day compressive and flexural strength of the moldcast MSRC specimens were measured as 58.5 MPa and 9.52 MPa,
respectively. Compared to the SRC mold-cast counterparts, the sulfur
modification using DCPD led to up to 44% and 34% increase in
compressive and flexural strength, respectively. Furthermore, the
average 7-day flexural strength of the printed MSRC specimens (13.3
MPa) was about 41% higher than the printed SRC counterparts. The
significant strength enhancement can be attributed to the free-radical
chain-growth polymerization reaction between the sulfur and the
DCPD comonomer resulting in the formation of polysulfides shown to
have a higher molecular weight [35,58]. To understand the underlying
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cause of the strength improvement achieved through sulfur modification
with DCPD, SEM imaging was performed. Fig. 10(a-d) shows the SEM
images of both the plain sulfur and SRC samples (with and without
DCPD modification) at 1-day age. The SEM images of the plain sulfur
samples were obtained to provide insight into the sulfur crystal
morphology. Based on the results, a crystalline structure corresponding
to orthorhombic sulfur crystals was observed in the unmodified sample
(Fig. 10a). On the other hand, a smooth surfaced crystal structure with
no phase separation as seen in the unmodified sample was observed for
the modified sample (Fig. 10b). This could be attributed to the successful
copolymerization of the monomers where the DCPD comonomer is
cross-linked with the sulfur biradicals. The observed microstructural
morphology of the modified and unmodified samples is also consistent
with results presented in other studies [55,59,60]. Ghumman et al. re­
ported that in contrast to the crystalline structure observed in unmodi­
fied sulfur, modified sulfur exhibits an amorphous morphology due to its
cross-linking structures [60]. On the other hand, the SEM imaging of the
SRC samples provides more insight into the microstructural interaction
between the sulfur binder and the regolith aggregate. The SEM image of
the SRC material reveals regions of poor bonding between the crystalline
sulfur crystals and the regolith particles (Fig. 10c). In contrast, the MSRC
samples show regions of better bonding between the regolith particle
and the smooth surfaced bulk sulfur due to the DCPD modification
(Fig. 10d). The improved bonding between the modified sulfur and
regolith particles could potentially be due to the amorphous morphology
of the modified sulfur crystals. Hence, the increase in strength of the
printed MSRC compared to the SRC can be attributed to the improve­
ment in the engineering properties of the modified sulfur and the
microstructural packing which, in turn, leads to a lower porosity. Gan­
non et al. also attributed the differences in the microstructure of un­
modified and modified sulfur sample (based on a proprietary additive)
to the improved bonding or bridging between the modified sulfur and fly
ash which was used as the fine aggregate [61].
4.2.4. Recyclability of the printed SRC
To evaluate the feasibility of recycling 3D printed SRC, the original
SRC and MSRC specimens (control) were first tested after 7 days. Upon
completion of the flexural test, the specimens were immediately recy­
cled by reheating them in the oven and then reprinting new specimens
without any compositional modification. The printed recycled unmod­
ified specimens (SRC_R) and modified specimens (MSRC_R) were then
tested after 7 days to determine their flexural strength. Based on visual
inspections throughout the remelting process and by measuring the flow
table diameter of the recycled material, no significant changes were
observed in the flowability and workability of the SRC and MSRC ma­
terials. Next, the recycled materials were successfully reprinted using
similar extrusion process parameters as the original materials. Results
presented in Fig. 11a confirm that there are no significant changes in the
flexural strength of the recycled SRC and MSRC specimens compared to
the original samples. Interestingly, in two other studies on modified
mold-cast sulfur concrete, up to 20% strength improvement was re­
ported after recycling [42,44]. These researchers attributed the strength
improvement to the lower amount of entrapped air in the recycled moldcast specimens due to the additional compaction. In the case of the 3D
printed specimens in this study, a more consistent compaction is antic­
ipated for both the original and recycled materials due to a mechanized
extrusion system. Hence, these preliminary results highlight the feasi­
bility of recycling 3D printed sulfur concrete without strength loss,
enabling a cradle-to-cradle sustainability approach since the recycled
sulfur concrete can be reused for other construction purposes at the end
of its life cycle. Furthermore, considering the energy-intensive processes
needed for the extraction of raw materials (such as elemental sulfur),
this would lead to significant savings and conservation of the limited
planetary resources.
4.3. Performance comparison between printed SRC and PCC
4.3.1. Strength development
Fig. 11b shows the flexural strength of the SRC and PCC specimens at
Fig. 10. SEM images of unmodified and modified samples (a) Unmodified plain sulfur (b) Modified plain sulfur (c) SRC (d) MSRC.
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Fig. 11. (a) Flexural strength of recycled and reprinted SRC and MSRC specimens (b) Strength development of 3D printed SRC and PCC specimens over time.
Fig. 12. SEM images at different ages (a) unmodified plain S8 at 1-day (b) unmodified plain S8 at 7 days (c) unmodified plain S8 at 7 days (higher magnification) (d)
modified plain S8 at 1-day (e) modified plain S8 at 7 days (f) SRC at 1-day (g) SRC at 7-days (h) MSRC at 1-day (i) MSRC at 7-days.
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different ages. The average flexural strength of the SRC and MSRC
specimens after 12 h was 8 MPa and 10 MPa, respectively. The SRC and
MSRC specimens then reached an ultimate strength value of 9.4 MPa
and 13.3 MPa, respectively, after 7 days. This shows that the printed SRC
and MSRC specimens could obtain about 75–85% of their ultimate
strength after only 12 h. These findings are also consistent with previous
studies that investigated the strength development of sulfur concrete
made with natural aggregates [38,62]. In the case of the PCC, the tested
specimens reached a maximum flexural strength value of 1.1 MPa after
12 h, which then later increased to 5.6 MPa after 7 days. Moreover, the
28-day flexural strength of the printed PCC was measured as 5.8 MPa
indicating a marginal improvement from the 7-day values. Hence, these
findings reveal the rapid strength development of the printed SRC (as
well as a superior ultimate strength) in comparison to the printed PCC
which only obtained about 19% of its ultimate strength after 12 h.
Considering the existing literature, it is well established that the
mechanical strength development of cement-based materials is due to
the formation of hydration products as a result of chemical hydration
reactions that progress over time. In the case of SRC and MSRC mate­
rials, SEM imaging was performed to better understand their micro­
structure. The microstructural morphology of the plain sulfur and SRC
samples (unmodified and modified) were studied after 1 day and 7 days,
respectively. For the plain sulfur sample, the SEM image reveals crys­
talline features within the bulk sulfur after 1-day. However, these fea­
tures were not visible after 7 days - Fig. 12(a-b). The bulk sulfur
observed after 1 day appears to have developed into macro-sized crystals
visible at higher magnification after 7 days (Fig. 12c). The smooth sur­
faced bulk sulfur due to DCPD modification was observed after 1 day and
7 days - Fig. 12(d-e). This indicates an effective modification of sulfur
with DCPD since no notable changes were observed in the microstruc­
ture within this period. However, Fig. 12e shows some sulfur fragments
which could indicate the presence of some unreacted sulfur or depoly­
merized sulfur after 7 days. For the SRC samples, the SEM image
revealed the presence of regolith particles that are unbounded to the
bulk sulfur after 1 day (Fig. 12f). Moreover, after 7 days, the unmodified
bulk sulfur has developed into macro-crystals that grow to fill the voids
around the regolith particles (Fig. 12g). Therefore, the microstructure
after 7 days appeared to be more dense and well-packed compared to
what was observed after 1 day. Furthermore, the SEM images of the
MSRC sample revealed that the modified bulk sulfur completely en­
capsulates the regolith particles after 1 day. After 7 days, a unique
microstructure was observed indicating that the modified bulk sulfur
has also developed into macro-crystals with better angularity that are
well distributed and bound to the regolith particles. As a result, a refined
microstructure with a dense and well-compact microstructure due to
better packing between the modified sulfur crystals and regolith parti­
cles was observed after 7 days. This indicates that modified sulfur (with
a potentially amorphous structure with better angularity) forms a better
bonding with the regolith particles compared to unmodified sulfur with
a crystalline structure. A study by Bright et al. also noted the differences
in the crystallinity of the modified sulfur crystals in comparison to un­
modified sulfur due to their polysulfide fractions which are higher in the
former [63]. Hence, the obtained results show that improvement in the
microstructure of the SRC and MSRC from 1 day to 7 days could be the
reason for the flexural strength improvement within this period. In
addition, the EDS analysis performed following the SEM imaging suggest
no significant compositional changes in the SRC and MSRC materials
over time. This finding is also consistent with the results presented in
previous studies [55,64].
4.3.2. Performance under vacuum conditions and temperature variations
Fig. 13(a-b) presents the weight loss in the printed SRC and PCC
specimens, respectively, when exposed to near-vacuum conditions. The
SRC specimens exposed to vacuum conditions at room temperature
(25 ◦ C) showed negligible weight loss (0.010% and 0.016% of original
weight) after 3 days and 7 days, respectively. However, when the vac­
uum chamber temperature was elevated to 55 ◦ C, the weight loss
significantly increased by an order of magnitude to 0.82% and 0.92%,
after 3 days and 7 days, respectively. Our findings based on testing SRC
are consistent with previous studies in which sulfur-based materials
were subjected to vacuum conditions at different temperatures [47,65].
Tucker et al. also reported that the comparative ratio of the rate of plain
sulfur sublimation measured at 24–26 ◦ C and 40–45 ◦ C temperature was
about 1:80 [66]. The increase in sublimation rate as the temperature
Fig. 13. Performance of printed specimens in vacuum conditions (a) SRC and MRSC weight loss over time (b) PCC weight loss over time (c) Surface quality of the
SRC25◦ C specimen (d) Surface quality of the SRC55◦ C + Vac specimen (e) Surface quality of the MSRC55◦ C + Vac specimen.
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increases is attributed to the increase in the vapor pressure of the
orthorhombic sulfur crystals considering the triple point phase of sulfur
[67]. In contrast, the MSRC specimens showed a significantly lower
weight loss even at the elevated temperature (55 ◦ C). This demonstrates
up to 85% decrease in weight loss when compared to the weight loss
value reported for the SRC specimens after 7 days. This result shows that
modifying SRC with DCPD can potentially help reduce the rate of sub­
limation due to the dominant polysulfide species with relatively lower
vapor pressure compared to the principal orthorhombic species in the
SRC specimens [68]. Fig. 13 (c-e) presents the surface quality of the
control and vacuum-exposed printed specimens, observed using an op­
tical microscope (75× magnification). For the printed SRC specimen
exposed to near-vacuum conditions and elevated temperature, the effect
of sublimation was visible through the exposed regolith particles not
coated with sulfur on the surface of the specimens (Fig. 13d). In contrast,
the printed MSRC sample showed no exposed regolith particles
(Fig. 13e). For PCC, the measured weight loss due to the evaporation of
free moisture in the printed PCC specimens are presented in Fig. 13b.
The moisture loss in the printed PCC specimens exposed to vacuum
conditions at room temperature was about 32% and 25% higher than in
printed specimens air-dried in ambient conditions after 3 days and 7
days, respectively. The effect was even more pronounced when the
temperature increased from 25 ◦ C to 55 ◦ C. The measured weight loss
increased by about 102% and 95% after 3 days and 7 days, respectively,
when compared to the weight loss of the air-dried specimens stored in
ambient conditions.
Fig. 14(a-b) shows the 3-day and 7-day flexural strength of the
printed SRC and PCC specimens exposed to vacuum conditions at the
two different temperature levels. SRC specimens exposed to vacuum
conditions at room temperature (25 ◦ C) showed no significant changes
in flexural strength after 3 and 7 days, when compared to SRC specimens
stored in ambient conditions (control SRC specimens). However, when
the temperature was elevated to 55 ◦ C, the flexural strength of the SRC
specimens reduced by 12% and 7%, after 3 days and 7 days, respec­
tively, compared to the printed control specimens. Similarly, the MSRC
specimens subjected to vacuum conditions at 55 ◦ C also exhibited up to
17% reduction in flexural strength after 3 days and 7 days, when
compared to the modified control specimens. The decline in the flexural
strength data observed in both the SRC and MSRC specimens exposed to
vacuum conditions and elevated temperatures can be attributed to the
increased porosity within the composite due to sulfur sublimation as
observed in the previous stage. Another likely reason for the reduced
flexural strength is the negative impact of vacuum conditions on the
microstructural evolution of SRC and MSRC materials. Interestingly, it
was observed that the flexural strength of the SRC and MSRC specimens
still increased from 3 days to 7 days under vacuum conditions. On the
other hand, the printed PCC samples exposed to vacuum conditions at
room temperature showed a 5% and 30% reduction in flexural strength
at 3 days and 7 days, respectively, compared to the printed control
specimens stored in ambient conditions. Similarly, the printed PCC
specimens exposed to vacuum at the elevated temperature (55 ◦ C)
showed about 78% lower flexural strength, compared to the control
specimens. The notable strength loss can be attributed to elevated
porosity levels within the PCC matrix, arising from the depletion of free
water needed for the formation of hydration products that typically
reduce matrix porosity in cementitious materials. In contrast to the re­
sults obtained for the SRC and MSRC specimens, the PCC flexural
strength decreased from 3 days to 7 days. This suggests severe impacts of
near-vacuum conditions on the ongoing hydration reactions due to
water loss, as well as a possible microcracking and deterioration of the
internal microstructure of the cementitious material exposed to vacuum
conditions at early ages before reaching higher levels of mechanical
strength. Based on these results, it is concluded that the negative effects
of vacuum on the properties and flexural strength of the printed PCC
specimens, compared to the printed SRC specimens, are considerably
more pronounced.
As a side investigation and to explore the operational aspects of
Lunar and Martian construction using SRC and PCC, the behavior of
freshly mixed SRC and PCC materials in vacuum conditions was also
studied. When the freshly prepared SRC and PCC samples were exposed
to vacuum conditions, they became unstable, showing a visible increase
in the apparent volume (foaming), along with the formation of visible
air pockets on the surface of the samples upon removal from the vacuum
chamber. To determine the ideal initial time delay before exposing the
printed specimens to vacuum conditions, the specimens were stored in
ambient conditions for 0–6 h before exposure to near-vacuum conditions
(room temperature). The exposed printed specimens were then visually
inspected for any physical defect or damage. Upon exposure to vacuum,
the 4-h printed PCC specimen imploded into fragments (Fig. 15) while
the printed SRC specimen showed no visible damage. This observation
can be attributed to the SRC reaching sufficient strength levels (due to its
rapid strength development) to withstand the internal stresses caused by
exposure to vacuum. It was observed that although the PCC reached its
initial setting, it did not develop sufficient strength to withstand the
vacuum conditions until after the 6-h initial period in regular atmo­
spheric pressure. On the other hand, the printed SRC could be placed in
vacuum conditions as early as 15–30 min after mixing, without any
visible damage or fracture appearance. These findings, in general,
highlight the superior resistance of the SRC materials to near vacuum
conditions, compared to PCC-based printing materials. Moreover, the
quick strength development of sulfur concrete can enable swift con­
struction and deployment of 3D printed infrastructure on the Moon and
Mars to support manned missions.
Fig. 14. Flexural strength of the printed samples in vacuum and temperature variations (a) SRC and MSRC (b) PCC.
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Fig. 15. A 4-h air-dried printed PCC specimen before and after exposure to near-vacuum conditions.
4.4. Discussion
construction, ISRU-based modifiers would have to be sought to avoid the
prohibitive costs of shipping raw materials such as sulfur modifiers from
Earth. Other solutions such as employing micro-fillers to physically
control the volumetric shrinkage of the microstructure due to the allo­
tropic transformation of the unmodified sulfur concrete have also been
discussed [32]. Based on preliminary investigations performed by the
authors using sulfur concrete made with silica sand and limestone
powder micro-fillers, such fissures were not observed in printed and cast
specimens made out of these materials.
The outcomes of vacuum testing at two different temperatures
revealed that the printed SRC specimens generally outperform their PCC
counterparts. The strength development of the SRC specimens continued
even in vacuum conditions, albeit to a somewhat limited degree. On the
other hand, the impact of vacuum and temperature on the printed PCC
specimens led to a substantial strength loss after 7 days and hindered the
overall strength development from 3 days to 7 days. This is evidenced by
a significant loss of free moisture from the printed PCC specimens due to
outgassing in vacuum conditions. For planetary C3DP operations on the
Moon and Mars, a temporary pressurized enclosing structure may be
required to mitigate the instability issues of printing materials in vac­
uum conditions during the 3D printing process. However, at early ages
following the printing process, the sulfur-based printing materials seem
to have a significantly superior resistance to vacuum conditions in
comparison to PCC-based printing materials. Another interesting finding
was that, in addition to enhancing the mechanical performance, modi­
fying the SRC material with DCPD significantly reduced the sublimation
rates as well, even at elevated temperatures which are known to accel­
erate the sulfur sublimation process. These discoveries further
strengthen the case for utilizing modified SRC as a printing material for
planetary C3DP in extreme environments. Finally, the results from this
study highlight the possibility of remelting and repurposing printed SRC
materials without strength loss, which would significantly increase its
value as a sustainable construction material on Earth, Moon, or Mars.
Based on the SRC performance data presented earlier, it is evident
that the previous layer (substrate) temperature - determined by factors
such as the extrusion temperature, interlayer time gap, and ambient
temperature - can influence the shape stability and flexural strength of
the printed SRC specimens. Our findings from the shape stability test
suggest that the extent of plastic deformations becomes more pro­
nounced as the material extrusion temperature is increased from 130 ◦ C
to 150 ◦ C. This observation is attributed to the decrease in the measured
plastic viscosity and the increase in the material flowability as the
extrusion temperature increases within this range. With a longer inter­
layer time gap leading to the substrate temperature dropping to values
below the sulfur melting temperature, plastic deformations are pre­
vented due to the rapid solidification and structurization of the printed
SRC layers. However, a tradeoff in the flexural strength was observed
due to the poor interlayer adhesion between the substrate layer with a
lower temperature and the newly deposited molten layer. This outcome
was further substantiated by the similar trend observed in 8-layer tested
specimens in which the specimens were loaded along the interface lines.
Interestingly, increasing the extrusion temperature (from 130 ◦ C to
150 ◦ C) can help improve adhesion between the new layer and the
substrate with a lower temperature. Therefore, a layer reheating
approach (increasing the temperature of the substrate layer right before
new layer deposition) may be useful in ensuring strong interlayer
adhesion and preventing the formation of cold joints. The 7-day me­
chanical strength of the mold-cast SRC specimens exhibited notably
higher compressive and flexural strength values when compared to the
28-day strength of the PCC specimens. Similarly, the printed SRC ma­
terials showed superior flexural strength values compared to the PCC
specimens. However, several fissures or cracks were observed in the
microstructure of the mold-cast and printed SRC specimens (Fig. 8).
These fissures are attributed to the volumetric shrinkage of the sulfur in
the matrix as it solidifies and undergoes subsequent allotropic trans­
formation. Such observed defects could have also discounted the
measured mechanical properties reported in this study. From the exist­
ing literature, chemical sulfur modifiers such as DCPD are typically
employed to prevent this issue and the related possible long-term
durability concerns. Introducing DCPD to modify the sulfur proved
effective in mitigating the severity of these defects, although it did not
completely resolve the issue. As such, a higher dosage of the DCPD
(>10% by sulfur weight) or the use of different modifiers might be
required to completely prevent sulfur crystals from depolymerizing back
to the orthorhombic form, which culminates into this defect. Moreover,
the mechanical strength of the printed and mold-cast SRC specimens was
significantly enhanced by incorporating 10% DCPD. For space
5. Conclusions
This paper provides crucial data on the feasibility of robotic C3DP
using Martian and Lunar resources. The space resilience and reusability
of 3D printed SRC materials, and the impacts of printing process pa­
rameters (such as extrusion temperature) on the properties of waterless
SRC extrudates, are quantitatively studied for the first time to deliver
new insights into waterless planetary C3DP. The following are the main
conclusions and findings of this paper:
• The printable Martian SRC and PCC materials developed using
similar binder volumes exhibit distinct properties. The printable SRC
14
I. Giwa et al.
Automation in Construction 165 (2024) 105571
materials demonstrated higher flowability supported by the
measured lower yield stress and plastic viscosity values, compared to
the printable PCC. These findings reveal differences in rheological
requirements for printable sulfur concrete in comparison to the
widely used cementitious printing materials.
CRediT authorship contribution statement
Ilerioluwa Giwa: Writing – original draft, Methodology, Investiga­
tion, Data curation. Mary Dempsey: Investigation, Data curation.
Michael Fiske: Writing – review & editing, Supervision, Conceptuali­
zation. Ali Kazemian: Writing – review & editing, Supervision, Meth­
odology, Funding acquisition, Data curation, Conceptualization.
• The Martian SRC material was successfully 3D printed, indicating the
feasibility of using sulfur-regolith concrete for waterless off-world
additive construction. Based on the quantitative data obtained in
this study, the geometrical accuracy and flexural strength of the
printed SRC elements are influenced by the extrusion temperature
and substrate layer temperature. Up to 53% strength loss was
observed in the 8-layer SRC specimens (tested in the Z-direction) as a
result of lower substrate temperatures during the printing process.
These findings highlight the importance of process optimization
while taking into account both the process and ambient temperature
parameters - especially in extreme environments - in C3DP of SRC
materials.
• The SRC specimens exhibited higher compressive and flexural
strength compared to the PCC specimens (in both mold-cast and 3D
printed specimens). Interestingly, the 3D printed Martian SRC
specimens were remelted and reprinted without the need for
compositional modifications, maintaining similar strength levels to
the original SRC specimens. This finding highlights another advan­
tage of sulfur concrete as a sustainable construction material for
planetary applications. This enables new possibilities for reusing
demolished 3D printed structures at the end of their service life,
which is significant due to the anticipated limited resources and the
required energy for the extraction and production of new raw ma­
terials using in-situ planetary resources.
• The Martian SRC specimens exhibited rapid strength development,
gaining about 85% of their ultimate strength after only 12 h, while
the PCC specimens only gained about 19% of their ultimate (28-day)
strength within the same period. Furthermore, subjecting the SRC
and PCC specimens to vacuum conditions at different temperatures
revealed that the negative effects were more pronounced in the PCC
specimens (up to 30% strength loss) than in the printed SRC speci­
mens (up to 7% strength loss) after 7 days of exposure. Considering
the extreme conditions in space, 3D printed infrastructure made with
SRC could offer better resiliency than those fabricated with PCC.
• Modifying the Martian SRC printing material with 10% DCPD proved
effective in enhancing its compressive strength (up to 44% increase)
and flexural strength (up to 34% increase). SEM images also revealed
a dense and refined microstructure in the modified SRC samples
compared to the microstructure of the unmodified samples.
Furthermore, the modification of the printing material with DCPD
significantly reduced the sulfur sublimation rate (by an order of
magnitude). Based on these preliminary results, modifying SRC
presents a viable strategy for improving the resiliency of the 3D
printed SRC materials.
Declaration of competing interest
The authors declare that they have no known competing financial
interests or personal relationships that could have appeared to influence
the work reported in this paper.
Data availability
The data that support the findings of this study are included in the
paper.
Acknowledgments
This work is funded by the Louisiana Board of Regents (RCS: LEQSF
(2022-25)-RD-A-11). The authors also acknowledge the support from
NASA Marshal Space Flight Center (MSFC) and Huber Engineered
Materials.
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