This methodology is applicable to any device that requires multilayered PDMS, multiple alignment processes, selective PDMS bonding, and multilayered integration with downstream sensing systems. Along with the detailed step-by-step process, we discuss essential quality assurance checks that can be performed throughout fabrication to assist in troubleshooting and maximizing chip yield.  Specifications tableA programmable microvalve array (PMA) manipulates fluids for critical sample processing operations such as metering, transporting, or combining chemicals or analytes Understanding the mechanism of microvalves and the fluidic resistance of the microvalve network is crucial for PMA design. The microvalves can be modeled as thin elastic membranes, and membrane theory can be utilized to predict the performance of the microvalve. Furthermore, the fluidic resistance of the PMA network affects the dynamic operation of the PMA, hence these theoretical elements must be considered. Once a PMA is designed, it is manufactured via soft lithography. Two master SU-8 molds are used to make PDMS replicas of fluidic and pneumatic layers Additionally, PMAs are frequently integrated with a downstream detection system to create a completely automated platform for biological and chemical sensing. A microfluidic capillary electrophoresis (µCE) is a widely used sensing technique for the highly sensitive detection of amino acids or other organic compounds This method article outlines the complete fabrication process, highlighting the steps important to successful and consistent integration. A detailed discussion of design considerations, mold fabrication, a consistent selective bonding procedure, and alignment considerations for successful integration with a variety of tools are covered in detail. Furthermore, to enhance PMA production capability, examples of quality assurance steps and a plan for addressing potential issues will be discuss along with functional demonstration of PMA. The PMA fabricated by this method can be utilized in a wide range of fields from space exploration t), the pneumatic chamber height (E), and Poisson's ratio (υ).A number of parameters must be considered when building a Programmable Microvalve Array (PMA). The size, number, and arrangement of microvalves are the key parameters to determine the performance of PMA. First, the microvalve size should be determined based on the desired volume and fabrication restrictions. Since the microvalve has a thin elastic membrane, membrane theory can be used to understand the deflection profile by considering the height of the pneumatic chamber. The theoretical deflection of the membrane can be calculated using an analytical model with the membrane theory described in Once the size of the microvalve is determined using all design criteria, the array of microvalves is designed to generate the PMA. In practice, the overall performance of a PMA will be influenced by the fluidic resistance of any pumping path defined by L is the length of the fluidic path, w denotes the channel width, 2 (UV KUB2), followed by a post-exposure bake (5 min at 65 °C then 10 min at 95 °C) before developing the wafer for 4 min and completing a hard bake at 150 °C for 30 min. Prior to use of both molds, the geometry of each fluidic and pneumatic layer is measured with a surface profiler and both molds are inspected visually under a microscope.The fluidic and pneumatic layers are designed in AutoCAD (Autodesk Inc.), converted to photomasks, and fabricated using a standard soft-lithography technique. Briefly, two five-inch silicon wafers (one for the pneumatic mold and one for the fluidic mold) are dehydrated on a hotplate set to 95°C, placed in a spin coater, then SU-8 2050 is poured on top and spun to the desired thickness . The wafers are then removed from the spin coater and placed on a hotplate for the pre-exposure bake (10 min at 65 °C then 16 min at 95 °C). UV exposure is completed using the designed photomasks with a dose of 240 mJ/cmSince the fluidic layer of the PMA is thin , there is a risk of damage to the layer when removing it from the fluidic SU-8 mold. To address this, the fluidic mold is coated with parylene-C to achieve a thickness of less than 1 µm utilizing a Specialty Coating Systems PDS 2010 and a parylene adhesion promoter (Silane A-174) after visual inspection under a microscope. The parylene acts as a surfactant, reducing the amount of pulling force required for detaching the thin PDMS and, contrary to conventional silane coatings, provides visual cues to any damage or wear the coating may have.Pouring begins with the preparation of PDMS (10:1 ratio) and its degassing in a vacuum chamber. The PDMS is then put into the pneumatic bath A and degThe procedure of bonding two layers begins with the removal of the PDMS from the two molds. To prepare the pneumatic PDMS layer a sharp knife was used to cut around the circumference of the molding bath. The layer is removed from the mold using the knife as a lever to peel the PDMS off the mold, and the layer is wrapped in plastic wrap and stored until use C. To preOnce both PDMS replicas are removed from the molds, the pneumatic and fluidic access ports are punched into the pneumatic layer C prior t2 gas, dehydrating on a hot plate set to 45°C, and a final cleaning with tape.For final preparation for the pneumatic and fluidic bonding, the fluidic layer needs to be smoothed, both layers must be cleaned, and practice alignment should be done. First, the plastic film piece is removed from the unfeatured side of the fluidic PDMS. To remove any trapped air, the PDMS layer can be lifted from the edge to the area around the bubble and then slowly placed back down. Several bubbles may exist as a result of particles trapped beneath the PDMS, which must be cleaned using tape before proceeding with the bonding procedure. If bubbles are not effectively eliminated, the fluidic layer may deform unexpectedly during plasma exposure. Finally, particles are removed from the top of the fluidic layer using tape. Dust causes poor bonding and valve failure, thus cleanliness during the bonding stage is crucial. After that, the fluidic layer is covered with a petri dish and placed away during pneumatic cleaning. For the pneumatic PDMS, the cleaning process starts with a rinse in isopropyl alcohol (IPA) and deionized (DI) water to eliminate any remaining materials, followed by drying with N2 flowrate of 7.5 sccm. Both PDMS layers  is used to align multiple microvalves as shown in S layers C and D a2 gas and put on a hot plate set to 45 °C for about 10 min. The PMA is wrapped in plastic wrap and stored until needed.Once the fluidic and pneumatic layers are assembled, the fluidic wells and interfacing holes are punched indicated in Once bonding has finished, the extra PDMS at the perimeter of the fluidic layer is trimmed and the combined pneumatic-fluidic PMA is removed from the lower plastic covering. Next, the function of all microvalves is verified by connecting vacuum and pressure sources to all the pneumatic ports. If the microvalves do not stay actuated under vacuum, there is likely contamination branching two or more pneumatic lines. If a valve does not actuate at all, either the pneumatic mold is damaged, or contamination has blocked the line. Any failure at this stage prompts an investigation of both the PMA and pneumatic mold under a microscope. This identifies if the issue is one-off, i.e. dust contamination, or a more serious issue, i.e. a line broken on the mold. After completing the PMA, the PMA can be bonded to either blank glass or a CE glass chip. However, if the entire surface of the PMA is activated with oxygen plasma, the gates of microvalves will bond to the surface, resulting in sticking issues. A selective bonding strategy is necessary to avoid unwanted bonding of the PMA gate structures to the underlying the glass substrates. In our approach of selective bonding, we utilize a unique microcontact printing technique to passivate the gate After the appropriate deposition period has passed, the stamps are removed from the vacuum chamber and placed in a petri dish covered to prevent contamination. To precisely stamp PFTCS onto the gate of microvalves, the motorized alignment stage is used again. First, the cleaned PMA is placed into the vacuum chuck with the channel side facing down from the chuck. The chuck is then secured to the motorized stage and the stamp is placed on the rotating stage PFTCS up after removing the tape mask. Using the moveable digital microscope and the motorized controls, the PFTCS spots are aligned with the gates and pressed together. Any rotational or translational misalignment can lead to valve sticking so, like bonding, multiple valves are checked before and after pressing the PMA and stamp together to verify alignment. The PMA with the stamp is then placed on a hotplate set to 150 °C with a 3 kg weight to promote the condensation reaction and conformal contact. After 30 min, the weight and stamp are removed from the PMA and placed in the petri dish for storage. The PFTCS stamps can be utilized for confirmation of PFTCS transfer using a high-resolution microscope or optical profilometer. 2SO4:1 H2O2) bath and a DI water bath are prepared with the glass substrate being submerged in the piranha acid. After five minutes, the glass substrate is removed, briefly washed with DI water, and then placed in the water bath for ∼30 s. The glass is then removed from the DI water and N2 gas is used to dry the substrate. After that, the glass substrate is dehydrated at 95°C for 10 min and then allow to cool to room temperature before bonding.The final bonding takes place in a similar way to that of conventional glass-PDMS bonding. To prepare the bonding PMA to a featureless glass substrate, a piranha acid , solenoids , a controller , and vacuum and pressure lines along with a manifold or barbed fittings for interfacing with the chip. General QA is completed throughout the fabrication process. Final QA is performed by setting up the PMA in a testing bed shown in The testing bed is used with valve actuation sequences to move fluid around the chip, demonstrating both the pumping capability and usability in real-world experiments. To evaluate pump performance, a PMA equipped with a 3 × 3 microvalve array and perimeter Input/output microvalves was successively fabricated to transport liquid from an inlet to an outlet indicated in 25 distinct operational states that can be utilized in a number of sequences to perform various fluid processing or delivery functions. Finally, the robustness of the presented approach was evaluated by executing the pumping sequence from To demonstrate the usability of the PMA in real experiments, a mixing sequence is evaluated. 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.


Objective: The purpose of this study was to analyze the feasibility of repairing a ruptured intervertebral disc using a patch secured to the inner surface of the annulus fibrosus (AF). Different material properties and geometries for the patch were evaluated.Methods: Using finite element analysis, this study created a large box-shaped rupture in the posterior-lateral region of the AF and then repaired it with a circular and square inner patch. The elastic modulus of the patches ranged from 1 to 50 MPa to determine the effect on the nucleus pulposus (NP) pressure, vertical displacement, disc bulge, AF stress, segmental range of motion (ROM), patch stress, and suture stress. The results were compared against the intact spine to determine the most suitable shape and properties for the repair patch.Results: The intervertebral height and ROM of the repaired lumbar spine was similar to the intact spine and was independent of the patch material properties and geometry. The patches with a modulus of 2–3 MPa resulted in an NP pressure and AF stresses closest to the healthy disc, and produced minimal contact pressure on the cleft surfaces and minimal stress on the suture and patch of all models. Circular patches caused lower NP pressure, AF stress and patch stress than the square patch, but also caused greater stress on the suture.Conclusion: A circular patch with an elastic modulus of 2–3 MPa secured to the inner region of the ruptured annulus fibrosus was able to immediately close the rupture and maintain an NP pressure and AF stress similar to the intact intervertebral disc. This patch had the lowest risk of complications and produced the greatest restorative effect of all patches simulated in this study. Intervertebral discs (IVDs) cushion compression forces on the spine whilst providing considerable flexibility. Disc herniation manifests as a protrusion of the soft inner nucleus pulposus (NP) through the outer annulus fibrosus (AF), which can ultimately compromise the structural integrity of the disc. The protrusion can also compress nerves in the region, with patients often complaining of pain and numbness in the back and lower extremities. Discectomy is an effective treatment to relieve the symptoms of herniated discs, but there is a risk of re-herniation due to original annular rupture .Studies have reported a reoccurrence of symptomatic herniated discs in 7%–18% of patients within 2 years of surgery , and thein vitro and cellular experiments. However, implant dislocation is still a concern due to the difficulty of fixation  analysis allows for one input factor such as the force or moment to be altered while keeping other factors constant. This is a convenient and inexpensive method of simulating different physiological conditions and has been widely used for assessing lumbar implants .This pilot study uses a FEM of the lumbar spine to simulate the repair of a large box-shaped AF rupture using an AF patch placed on the inner surface of the annulus. The material properties and geometry of the patch were altered to determine the most stable combination to improve the primary stability and strength of the repaired tissue.The intervertebral height of the model in this study was 9.90 mm, the vertebral body height of L4 was 31.06 mm, and the vertebral body height of L5 was 28.56 mm, all of which are within the ranges reported in literature . Mesh coin vitro experimental data , and the official website shows that the material is ethylene terephthalate. The interface between the patch and surrounding tissue was simulated with surface-to-surface contact elements. The friction coefficient between the patch and surrounding tissue was 0.1  is difficult, and the smaller the size of the patch, the easier the operation. At the same time, the patch must be able to completely plug a 5 mm × 5 mm square rupture. Round and square patches are more suitable for plugging square ruptures, which have a greater effective area utilization. Other shapes, such as triangles and rectangles, require a larger volume of patch to repair the same size rupture. Considering the technical feasibility, this study repaired the AF rupture using a square patch and circular patch secured to the inner surface of the and recorded the pressure on the nucleus pulposus, vertical displacement, disc bulge, contact pressure on the cleft surfaces (CPCS), vertebral ROM and the maximum stress on the AF, patch and suture. The size of the disc bulge was defined as the maximum radial displacement of AF after compression. The CPCS is the maximum stress on the walls of the four sides of the AF rupture. A square patch was chosen because this was the design used in previous studies, which allows for better comparison , and boxAs shown in The ROM of the herniated lumbar FEM was comparable to experimental results from Mimura et al. when placed under a torque of 10 Nm and axial compression force of 400 N . The gooThe maximum AF stress for the intact model was 0.464 MPa, but was exceeded by both repair models. As shown in The CPCS and maximum suture stress were at their lowest when the Young’s modulus of the patch was 2 MPa, with minimum values 0.587 and 6.574 MPa, respectively. The maximum stress on the patch increased with the Young’s modulus . The NP To assess how the geometry of the patch may affect the success of the repair, the modulus of the patch was set as 2 MPa and an axial compression force of 400 N and torque of 3 Nm were applied to the upper surface of the L4.When the lumbar segment was placed under an axial compression force of 400 N, the circular patch showed a lower NP pressure, CPCS and patch stress than a square patch. However, the suture stress was greater. The degree of degeneration had a considerable effect on the NP pressure when using a square patch, but the circular patch was able to maintain a relatively constant pressure regardless of the severity of the herniation . In flexThere was no significant difference in the ROM or compressive stiffness of the lumbar spine after repairing the ruptured annulus with either a square patch or circular patch. However, the NP pressure was significantly less in all motions when using a circular patch .As a novel technique, annular repair is attracting more attention as an effective method of treating a ruptured disc while retaining the physiological environment. This study aimed to assess the feasibility of placing a patch on the inner surface of the AF to seal a ruptured disc using a validated finite element model. This technique is designed to reduce the loss of NP tissue and prevent re-leakage. The results showed that a circular patch with a Young’s modulus of 2–3 MPa applied to the inner surface of the AF region could immediately close the rupture and maintain a low NP pressure and restore the lumbar stability and stress distribution similar to the intact model.A meta-analysis by Miller et al. reported that the risk of re-herniation and reoperation after lumbar discectomy was approximately threefold higher in patients with large versus small annular ruptures . There iTo determine the most suitable material properties for the patch, different Young’s moduli were evaluated (up to a maximum of 50 MPa ) in termThe loss of intervertebral height after repairing with the inner patch was less than that reported for tissue-engineered discs (0.34 vs. 0.48 mm) . MaintaiThis study also considered the stress distribution on the repair patch and assessed the risk of failure. The patch must be able to withstand the maximum stress conveyed from the NP and inner region of the AF under different physiological loading conditions. The results predicted a maximum stress on the patch of 1.84 MPa, so the strength of the patch should exceed 1.84 MPa. The polypropylene sutures have an ultimate tensile strength of 23.15 ± 2.87 MPa . The maxFrom an in-vivo animal study, Chiang et al. showed fewer degenerative changes occur in repaired disks than unrepaired disks . This cuin vitro, animal models, are needed for product development to determine more details. Second, this study did not consider the implantation method for the patch, which is currently under investigation by the authors. Future studies may introduce custom surgical instrumentation, once it has first been confirmed that the patching technique is an effective method for repairing herniated IVDs. Third, the results of this study may only be applicable to cases where the AF rupture is located around the posterolateral aspect of the AF. This is the most common site of injury in the lumbar spine and so potentially is applicable in a large proportion of cases. Finally, the AF has a limited healing potential, and so whether the repair could generate clinically significant improvements over alternative methods requires further large-scale and comparative clinical studies.There are some limitations to this study. First, this was a preliminarily assessment of the reliability of AF repair with a medial patch which focused on the most suitable material properties for the device. Further studies, such as This study verified the mechanical feasibility of repairing large annular ruptures with an inner patch and investigated how the material properties and geometry of the inner patch can influence the results. The results suggested that the AF inner patch should have a modulus in the range of 2–3 MPa and a strength greater than 1.84 MPa. In addition, the circular patch was better able to restore the properties of the IVD than the square patch. An inner patch with these characteristics can restore disc height and stress distribution. This study provides a basis for the design and development of related products in the future.


The well-known principle of time–temperature superposition (TTS) is of prime interest for polymers close to their glass transition. First demonstrated in the range of linear viscoelasticity, it has been more recently extended to large deformations in tension. However, shear tests were not yet addressed. The present study depicted TTS in shearing conditions and compared it to results in tensile conditions both for low and high strains for a polymethylmethacrylate (PMMA) of different molar masses. The main objectives were to enlighten the relevance of the principle of time–temperature superposition for shearing at high strain and to discuss the way shift factors should be determined. It was suggested that shift factors could be dependent on compressibility, which should be taken into account when addressing various types of complex mechanical loadings. Modeling and understanding the mechanical behavior of polymers close to their glass transition are still interesting open problems because those materials can be widely used in that domain either for processing amorphous polymers or for end use as semi-crystalline materials.The glass transition in polymers is not a phase transition. It is a gradual change in the mobility of the polymer chains that are entangled or tangled when the temperature drops. However, this phenomenon is not associated with a critical change in the structure of the polymer. Nevertheless, it is of key importance for polymers because it delimits their fields of application. Studying the behavior of these materials around the glass transition is therefore of great interest. Furthermore, despite many results, some questions remain open.On the scale of physics, deforming a material means modifying the organization of the atoms that constitute it. The stress defined by the mechanical engineer results from the energy variations involved in these modifications. In the case of polymers, these modifications are of two natures, and their contributions highly depend on the temperature. The interactions between atoms or groups of unbound atoms define the first contribution. The second one is torsion potentials involved in changes in the local conformation of backbone and side groups. When the latter is facilitated by temperature, conformational changes in the backbone become cooperative and introduce a change in the entropy of the system. This is the α transition of the polymer associated with the glass transition.Eyring  proposedThe jump frequency of the change is classically controlled by an activation energy  depends on the temperature according to a law for which there is no theoretical form. One generally uses the empirical Vogel, Fulcher, and Tammann (VFT) law ,3,4, whiA, This expression can be rewritten as the Williams, Landel, and Ferry (WLF) law :(2) τα=τTherefore, the behavior of polymers depends on both temperature and time. In the case of a mechanical loading at non-zero, monotonic, or oscillatory velocity, the behavior depends on the temperature and the strain rate (E) as:In addition to the shift factor  under forced oscillation, which is one of the key tools used in this analysis, can be used to determine the shift factors as a function of temperature. During this test, the sample is subjected to an oscillatory force or displacement . The complementary quantity (displacement or force) is measured.In the range in which the behavior remains linear (low deformation), the response has the same form and the same frequency as the stress. For a material with pure elastic behavior, strain and stress remain in phase. If the behavior of the material is viscoelastic, a phase shift (δ) results from the “dissipated” energy at each cycle due to the anelasticity.Considering the geometry of the loading and assuming that the strain and the stress are uniform, two quantities . This is the equivalent of characterizing the double dependence on time and temperature of the elastic modulus or of quantities associated with linear viscoelasticity as described above.However, TTS can also apply to large deformations . This was recently illustrated thanks to well-instrumented tensile tests of amorphous ,7,8 or sOn the other hand, the shear tests were not really addressed with regard to solid polymers (seemingly because tension was much more easily controlled). However, this type of loading path presents a major difference in tension because it is isochoric . In parallel, the glass transition is clearly influenced not only by the rigidity of the chains but also by the free volume (and therefore the density) of the material. It is thus interesting to investigate the differences between these two modes of solicitation.Firstly, to enlighten the relevance of the time–temperature superposition principle for shear tests at high strain.Secondly, to discuss the role of volume change due to mechanical loading.Consequently, the main objectives of this study were:The first of these goals was pursued using shear tests at high strain conducted under well-controlled conditions and dealt with promoting local measurements and constant strain rates in isothermal conditions during the tests. Indeed, despite their widespread use, nominal quantities  would not have allowed a rigorous characterization of the state of deformation and stress. In the same manner, very often only the crosshead velocity is kept constant, which implies that the zone of observation would have experienced a variable strain rate. This paper describes the way the tests were conducted to overcome those limitations. The results for two PMMAs of different molar masses will be discussed as a proof of how TTS applies in high-strain shear tests in comparison with tensile conditions. In that part of the study, TTS was calibrated using shear DMA tests.In a second step, the shift factors (Equation (2)) deduced from shear tests were compared to those obtained in tensile conditions. This was conducted using DMA analyses of amorphous polymers (PMMAs of different molar masses and different glass transition temperatures and a polyetherimide (PEI)) and on a semi-crystalline polyetherketoneketone (PEKK). The effect of the volume change due to mechanical loading was pointed out. We concluded that the shift factors should be revisited to account for that phenomenon.As stated previously, the database gathered using the above protocol aimed to enlighten the relevance of the time–temperature superposition principle for shear tests (including high strains) and to discuss the role of volume change due to mechanical loading. The description of the application of the TTS in shear tests that included a high strain, which was innovative,Its comparison with the tensile loadings , which The influence of the material’s compressibility on shift factors, which represented a perspective for further improvements.To progress gradually from the first of these conclusions to the last one, we broke down our analyses into three parts:Altogether, three PMMAs were used that were representative of the wide range of accessible PMMAs. They differed in their glass transition temperatures and in their rubbery behaviors. The α transition temperature  made it possible to illustrate that the TTS was applied in the shear as it was applied in the tension [® curve-fitting option. To achieve this, frequencies scans were performed every 10 °C. The WLF parameters that were identified for a reference temperature of 130 °C are gathered in Shift factors were estimated for the three materials in the tension and torsion using the ExcelUsing the shift factors extracted from the torsion DMA tests  was related to the equivalent strain rate at For a given equivalent strain rate at the same reference temperature, the type of macroscopic behavior  was the same in the tension and in the shearing . However, ifferent .Even if the differences appear small, two sets of experiments for PMMA 93 using two injection-molding campaigns and two sets of machines were conducted to highlight the order of magnitude of the differences between the shearing and tension. In addition, it can be seen in The same general trends were observed for other polymers such as PEI and PEKK , but theIn the case of amorphous polymers,  aTTrefT . From a  aTTrefT . From a more physical point of view, the same effect would have been obtained if a density correction (Equation (3)) had been applied to account for a higher density in the torsion than in the tension. It was then tempting to relate this to the free volume that could be increased in the tension while it remained constant under shearing, which is known to be an isochoric phenomenon.−1 at 130 °C up to a strain of 0.5. This was validated thanks to measurements of the volume variation during tension as described above.One additional clue was the PMMA 93 that exhibited a tensile behavior close to incompressibility between an equivalent strain rate of 0.0001 to 5 sTo limit the range of the strain rate for which the polymer would be incompressible in the conditions in which we identified the shift factors, we analyzed the master curves. The low deformation simplified the analysis, and we could assume that an apparent Poisson’s ratio could be defined. If the behavior of a material is linear elastic, it should exhibit a Poisson coefficient  and torsion (G′ and G″) master curves at 130 °C for PMMA 93 are plotted. The estimations of G′ and G″ from E′ and E″ and Equation (9) is superimposed in grey.It was possible to validate this observation in the DMA conditions shown in −4 to 1 s−1 at 130 °C). Due to TTS itself, this corresponded to the range of temperatures “surrounding” the α transition at 1 s−1; i.e., 112 °C for PMMA 93 and 132 °C for PMMA 3500.At this stage, we could assume that the behavior of PMMA 93 was close to incompressibility in the range of the equivalent strain rate for which the polymer exhibited a viscoelastic or rubbery behavior  was provided as casted 4 mm-thick plates. The others of lower mass  were provided as 4 mm-thick plates that were injection molded. The dimensions of the plates were 300 × 300 mm and 100 × 100 mm, respectively. For the following analysis, the samples were processed at various dimensions using a 3-axis cutting machine.The tensile DMA tests were performed using 30 mm × 4 mm × 1 mm rectangular samples. The effective gauge length was 5 mm. The maximum strain was 0.001 and the heating rate was 1 °C/min for temperature-scanning tests. From a technical point of view, it was difficult to reach a flowing zone in tension. Conversely, by combining oscillatory rheology and torsion DMA, a wider range of temperatures was explored in shearing. Shear loading in the solid state resulted from torsion of the rectangular samples (30 mm × 4 mm × 1 mm), while the parallel discs were used for the flowing state (discs with a diameter of 25 mm and a 1 mm thickness).® PEI and KEPSTAN 7002®PEKK from Arkema  were gathered on 2 mm-thick injection-molded plates. The tensile samples were then 20 mm × 4 mm × 2 mm, whereas the torsion ones were 35 mm × 10 mm × 2 mm. The maximum strain was 0.004.For comparison, some results for ULTEM 1010The different geometries are depicted in High-strain tests consisted of tensile tests aimed at exploring compressibility behavior and shearing tests aimed at exploring TTS in the latter conditions. In both cases, digital image correlation (DIC) was used to assess local strains and stresses.Since all tests were performed above the glass transition temperature, the materials were considered as renewed by the pre-heating (10 min). Experiments were then performed without any additional pre-conditioning.(a) Tensile testsDumbbell-shaped samples were tooled from the 4 mm-thick PMMA sheets b in the A random pattern was sprayed on the surfaces (face and thickness) of the samples. A high-sampling-rate video camera was used to record the tests. The sample thickness was recorded as well to assess the strain in the third (thickness) direction. This was made possible by using a prism that reflected the image of the thickness to the front camera b.The strain field was assessed via the deformation of the pattern using image-correlation software. Hence, the local strains in the 3 main directions were known. Hencky’s strains were used. Finally, the volumetric strain during the test was calculated using the following equation :(10)ΔVVo(b) Shearing testsAn Iosipescu configuration  with a sSimilar to tensile loading tests, an exponential crosshead velocity was imposed to promote a local true strain rate as constant as possible. DIC allowed measurement of the true strains on the front face.Prior using this set-up, a validation was conducted to assure that the local strain was mainly under shear conditions, and a value of 25% was chosen as a maximum strain to guarantee “pure” local shear with no sliding. We observed that the deformation was quasi-homogeneous in the central zone c. HoweveAB is the length between the notches and h is the thickness . The same equivalent strain rate led to the same type of behavior  in the tension and under shearing and to the same behavior for each of the given loading geometries. The use of equivalent strain rates at α transition temperature allowed us to define a type of universal scale of strain rates for comparing different materials.Complementary to previous studies performed under tension, one of the main conclusions of this paper was that the time–temperature superposition principle applied in the region of the glass transition of polymers regardless of the loading (even at high strain). This suggested that kinetics of elementary processes are equivalent at any level of strain and are only based upon small changes in local conformation. Differences between high and low strains would then be due to the number of active elementary processes that are accumulated during the history of loading.However, differences existed between shift factors extracted from tension and those extracted from torsion tests. This could be observed for various polymers in addition to PMMA.These differences could be due to the fact that the shearing was a loading condition that kept the volume unchanged, whereas the volume could increase during tension. It was then suggested that shift factors could be sensitive to the interrelation between volume changes due to mechanical loading and the free volume in the polymer.This also suggested that they should be corrected to account for compressibility in the behavior, which can depend on the temperature and strain rate.A solution could be to reintroduce the density correction (Equation (3)) and to express the dependence of the density not only on the temperature but also on the deformation. The first invariant or volume variation would be the parameter to use.One could then define a usual shift  factor and a second shift depending on compressibility of the behavior in an manner inspired by the density correction (Equations (3) and (4)).Obviously, additional studies should be performed to validate this idea using various conditions of loading: compression, biaxial, etc.


The fractional clearance of proteins as measured in healthy human subjects increases 10,000–100,000- fold when studied in nephrotic patients. This remarkable increase cannot be accounted for by extracellular biophysical mechanisms centered at the glomerular filtration barrier. Rather, it is the nephron and its combination of filtration and cellular uptake that can provide a plausible explanation of these fractional clearance changes. The nephron has two regions that critically determine the level proteinuria/albuminuria. Glomerular filtration of plasma proteins is primarily a size selective event that is basically unchanged in acquired and genetic kidney disease. The glomerular concepts of ‘charge selectivity’ and of ‘large pores’, previously used to explain proteinuria, are now recognized to be flawed and non-existent. Filtered proteins then encounter downstream two protein receptors of the Park and Maack type associated with the proximal tubular cell. The high capacity receptor is thought to retrieve the majority of filtered proteins and return them to the blood supply. Inhibition/saturation of this pathway in kidney disease may create the nephrotic condition and hypoproteinemia/hypoalbuminemia. Inhibitors of this pathway (possibly podocyte derived) are still to be identified. A relatively small proportion of the filtered protein is directed towards a high affinity, low capacity receptor that guides the protein to undergo lysosomal degradation. Proteinuria in normoproteinemic states is derived by inhibition of this pathway, such as in diabetes. The combination of glomerular sieving, and the degradation and retrieval pathways can quantitatively account for the changes in fractional clearance of proteins in the nephrotic condition. Finally, the general retrieval of filtered protein by the proximal tubular cell focuses on the teleological importance of this cell as this retrieval represents the third pillar of retrieval that this cell participates in . Richard Bright in 1827 was one of the first to recognize that the appearance of protein in the urine was specifically due to kidney disease. What is remarkable that almost 200 years later a debate still continues as to whether proteinuria is a glomerular centric problem or a nephron problem.New England Journal of Medicine adopt a glomerular centric approach based on biophysical models of Haraldsson and Deen and others  accurately estimated the predicted glomerular flux of albumin governed by size selectivity alone; the calculated GSC was 0.04. This meant that this pathway was relatively high capacity as compared to the degradation pathway. Similar results were obtained using ex situ kidney perfusion.The first and only demonstration so far of circulating plasma albumin being filtered and then returned to the renal vein  was published in 1999 . Retrievdegraded . This reThe existence of the PTC retrieval pathway was confirmed 14 years later using a transgenic mouse with podocyte-specific expression of doxycycline-inducible tagged murine albumin .The nephron then has two opportunities to govern the urinary excretion of proteins; one at the glomerular filtration level and the other further downstream in terms of PTC uptake. Based on the discussion below it is likely that glomerular filtration may be basically invariant for acquired and genetic kidney disease. With this in mind, we would suggest that one aspect that limits urinary protein excretion is PTC processing. Data in the literature is indicative that the process is biphasic which ensures that proteinuria/albuminuria is not a continuum. There are two major lines of evidence that support this; albumin uptake by perfused tubules and protein fractional clearances in diseased states.−1) and a low affinity, high capacity site (Michaelis constant = 1.2 mg ml−1)  . These dmg ml−1) . This ba−5–10−6 to 10−4 can be quantitatively related to the JALBmax and Km of the low capacity pathway o change .dd-line to the n-line has been interpreted in our laboratory as being governed by the retrieval pathway  that suggested that albumin was essentially impermeable to the GFB.The best test for charge selectivity is through direct analysis of albumin interaction with glycosaminoglycans (GAGs). This interaction can be of two types. An equilibrium interaction is one where excluded volume is relevant; this will be important in estimating the ability of albumin to enter the GAG-containing charged regions of the GFB. The second interaction is a relative mobility interaction; this will be important in determining transport within the GAG-containing charged regions of the GFB.In the 1930s and 1940s there was the realization by a number of polymer physicists like Flory, Huggins and Staudinger (see Kurata ) that eqCa is the mass concentration of the solute a, Mna is the number average molar mass of the solute a and A2aOS and A3aOS are the osmotic second and third virial coefficients. The second virial coefficient is an average of all the binary interactions for all solute molecules and will generally dominate quantitatively for polymer concentrations up to 50 mg ml−1. For a mix of polymer solutes a and b the osmotic pressure can be described asA*ab is the interaction coefficient between a and b. The second virial coefficient can be described in modelling terms as steric exclusion of a compound from part of a solution by the presence of another identical molecule that leads to an increase in its concentration in the remaining space. Similarly, the interaction coefficient will describe the excluded volume between different molecules, in this case polymers a and b. The magnitude of the exclusion is intimately related to the size and shape of the interacting molecules.Here Therefore, osmotic pressure measurements are an excellent way of understanding the equilibrium interactions of anionic GAG chains with themselves and with other polymers such as albumin. The second virial coefficient of GAGs has been measured extensively in the 1960s and 70s by Cleland, Laurent, Matthews, Ogston and Preston . The int−1, albumin transport may be reduced by 50%   providedin vivo labeling studies of the glomerular basement membrane for laminin  with the same molecular radius . MicropuApparent charge selectivity was measured in the cold isolated kidney perfusion system (cIPK) through the difference in the fractional clearance of albumin and Ficoll with the same hydrodynamic radius of 36 Å . In thos−8 m2 s−1volt−1  which may be inhibited in nephrotic states.14C-labeled proteins, albumin (radius: 36 Å) and IgG (radius: 55 Å), in control and in PAN nephrotic Sprague-Dawley rats. The plasma elimination rates were then matched to the urinary excretion of these labeled materials. In control rats albumin and IgG plasma retention rates were identical and far enhanced compared with the retention rates of inert transport markers of equivalent hydrodynamic radius; their elimination rate corresponded to the elimination of a 75 Å radius Ficoll. In PAN, they were eliminated as ∼36 and ∼55 Å radii Ficoll, respectively, equivalent to their hydrodynamic radii. In contrast, there was no comparable increase in the elimination rate of Ficoll in PAN. The total plasma clearance of Ficoll in control and PAN rats and the urinary clearance in PAN rats were essentially the same for all radii. On the other hand, the urinary clearance of >45 Å radii Ficoll in controls was considerably lower with increasing radii, demonstrating a postfiltration cellular uptake in controls, which, when inhibited in nephrotic states, would account for apparent large pore formation.In order to address this we examined  plasma eRelevant to understanding these clearances is the notable observation of Rippe et al.  that in The fact that in the nephrotic condition albumin is cleared as a 36 Å molecule is supported by the finding that the loss of plasma albumin to give rise to hypoalbuminemia is quantitatively accounted for by the appearance of albumin in the urine .The failure of the old Nephrology concept of GFB impermeability to albumin in healthy controls and the ‘development of filtration of nephrotic levels of albumin in nephrotic states’ is no more exemplified than in a recent publication by Christensen’s group . CubilinPerhaps the single most influential experiment that has formed the basis of thinking about glomerular ultrafiltration is from the original studies on micropuncture , performConfidence in the micropuncture technology was enhanced by Brenner’s group  who demoThe micropuncture studies of Tojo and Endou  exemplifOur interpretation of these data is that the earliest sample taken by Tojo and Endou probably reflected the true GSC of albumin and that over time albumin was being continually removed from the stationary filtrate by both the retrieval pathway and the degradation pathway in the PTC.125I-labelled protein. After infusion, a whole body vascular washout was started with heparinized horse serum via the jugular vein at a rate of 20 ml min−1 for approximately 8 min. The trichloroacetic acid (TCA) precipitable urine and cortical tissue (not the inner medulla) were included in radioactivity measurements. The overall problems with the glomerular uptake technique is that it does not include the non-TCA-precipitable radioactivity in the urine derived from renal degradation of filtered material, does not include the material that is bound to the surface of the PTCs that could be exchanged with unlabelled albumin during the washout, fails to prove the assumption that there was no washout of retrieved albumin in the PTCs during the 8 min washout period and is unable to account for retrieved albumin that had already been returned to the blood supply during the 8 min infusion.Rippe and coworkers  used a gOnce it was realized that charge selectivity was not a factor in governing the transglomerular transport of albumin there was a compelling need to measure the albumin GSC directly. We expected the GSC to be high yet when the first two photon measurements of the albumin GSC of 0.03 were published  there waThese studies do provide what was originally expected when cell-mediated effects are not involved and knowing that charge selectivity is not a factor; the GSC of 36 Å radius Ficoll in healthy control humans = GSC of 36 Å radius Ficoll in nephrotic humans = the GSC of albumin in nephrotic humans = GSC of albumin in healthy controls (rats).−1 (day 3), due to albumin overload, yielded an increment of ∼360 mg protein day−1 in excretion rate which corresponds to a protein fractional clearance of ∼0.02. In a system where glomerular permeability is essentially unaltered and with endocytotic mechanism by PTCs being nearly saturated but with normal fragmentation still occurring, the results imply that the GSC is at least an average of 0.02 for all glomerular-filtered protein, the majority of which is likely to be albumin. Again, this is in agreement with the two photon studies described above. It is notable that overload proteinuria would never occur if the only tubular uptake was the low capacity degradation pathway associated with a very low GSC.There is an interesting set of studies that induces reversible nephrotic-like albuminuria; that is overload proteinuria as studied in rats . In over−1 yet the results in in vivo is going to be considerably higher due to the presence of non-albumin plasma proteins that are also filtered and retrieved. If we assume that albumin is approximately 50% of the total protein present then we should be looking at PALB near 2 mg ml−1  where the receptor is far closer to saturation. The transition from normal filtration in vivo to overload/saturability is very tight and much tighter than exemplified by the line of best fit of the upper inflexion of the uptake curve shown in The immediate albumin filtrate concentration in the rat will be ∼1 mg mlThe functional interplay of the retrieval and degradation pathways has still to be worked out. We have noted, however, that the tertiary structure of albumin, beyond simple size and charge, is a critical determinant for albumin processing by the retrieval pathway by the kidney . It sugg−1 albumin resulted in transcytosis of intact albumin   the increase in fractional clearances of proteins that approach those of inert probes of the same hydrodynamic radii, 2) no inhibition of the degradation pathway (when considered in terms of the Park and Maack receptors), 3) hypoalbuminemia resulting in a ∼60% loss in plasma protein and 4) associated with only small changes, if any, in the glomerular permeability of inert transport probes .The results in In speculating the initial events leading to nephrotic proteinuria, we could easily envisage a factor(s) being produced by stressed podocytesmilieu. They may be too small to accurately measure directly in terms of GSC or fractional clearance. Yet these small changes would be amplified in terms of protein excretion in a synonymous way as in overload proteinuria. This mechanism sets the basis for understanding how a ‘podocyte related event’ can have a disproportionate influence on albumin/protein excretion determined by the PTCs. There is also an important biomarker that establishes the common phenomenology of overload proteinuria and nephrotic proteinuria; the lack of inhibition of the degradation pathway. This is exactly what occurs in overload proteinuria where there is a saturation of the retrieval receptor but the degradation pathway is still operative (Alternatively we could consider an overload proteinuria effect. The overload effect we described above does not explain nephrotic proteinuria. We could turn the thinking around, however, and say that if there is a very small change in the exclusion properties of the GFB and that the output of albumin filtration across the GFB is increased by 10–20% then this would mimic overload proteinuria . The chaperative .m value for the retrieval pathway in m value may result in some albumin loss into the urineIt is interesting to speculate how a small increase in GFB permeability can ultimately generate hypoalbuminemiaWe view the initial etiology of the nephrotic state and particularly in nephrotic podocytopathies is a small aberration in the permeability of the GFB. This will result in an overload effect on the PTC with subsequent loss of filtered protein including albumin into the urine. Hypoproteinemia will be set in play. This evolution of nephrosis is probably never measured because, as noted earlier, partial nephrosis is rarely if ever observed. Rather what is seen is the end point of these overload effects where the fractional clearances of proteins particularly albumin and IgG, approach those of their Ficoll counterparts of the same radius and hypoproteinemia has reduced the plasma protein concentration by two-thirds ensuring that overload effects on PTC uptake will not occur . The reaIn terms of understanding proteinuria/albuminuria there has been two schools of thought. The dominant thinking for many years has been that proteinuria/albuminuria is an essentially glomerular problem associated with deficiencies in charge and size selectivity of the filtration barrier; increasing leakage of the GFB will give rise to increasing proteinuria. This idea started when Wearn and Richards  discoverin vivo of the GSC for albumin (The second school of thought, featured in this review, suggests that many of the conclusions of previous studies are incorrect and the studies themselves misleading. It started when apparent charge selectivity to albumin transglomerular transport was demonstrated not to exist. Dextran sulfate was originally used as a model of negatively charged albumin and its clearance in the kidney could be compared to its uncharged counterpart dextran. It turned out that the results of these clearance studies were confounded by the fact that dextran sulfate is avidly taken up by endothelial cells and desulfated. When the sulfate substitution on the dextran sulfate was lowered to inhibit cell uptake then ‘charge selectivity’ disappeared. Examination of other negatively charged transport probes that were inert and not taken up by kidney cells clearly demonstrated that charge selectivity did not exist. In terms of size selectivity, macro-changes have never been observed and that apparent ‘large pores’ were the result of PTC uptake of probes like dextran and Ficoll with radii >40 Å. If macro-changes in size selectivity do not exist and if charge selectivity does not exist then the GSC for albumin should be similar to uncharged inert probes of similar radius and this is exactly what was discovered in the first direct measurements  albumin . The rel albumin . In fact albumin . This no albumin  receptorThe scavenger/retrieval role of the PTC now takes on broader dimension. Not only does it retrieve water and NaCl but it also plays a critical role in the retrieval of plasma proteins that are leaked through the glomerular filter.Finally, when it is reported that albuminuria/proteinuria exists it is probably of one of two forms; it will be either nephrotic with accompanying hypoproteinemia or it will be a reflection of changes in the degradation pathway with more intact material being excreted.


During the spring thawing, the decrease of soil-ice interface strength by temperature may lead to slope instability. For this reason, some researchers have explored the relationship between temperature and soil-ice interface strength. However, previous studies have not systematically explored the change law of strength at the soil-ice interface from negative temperature to 0 °C. Therefore, direct shear tests were conducted at different shear temperatures and different moisture contents. The effects of temperature and moisture content on strength, cohesion, and internal friction angle are analyzed, while the shear failure mechanism of specimens at different temperatures is discussed according to the location of the shear failure surface. The results show that: Shear properties of soil ice specimens are related to the unfrozen moisture content. The strength of the sample decreases with increasing temperature, and the change in strength is most significant from − 2 to − 0 °C. The strength reduction in this range is from 21.8 to 74.8%, and the higher the moisture content the more obvious this phenomenon is. The shear index tends to decrease with the increase of unfrozen water content, and the greater the increase of unfrozen water, the faster the decrease of both, especially in stage 2. When the temperature is higher than − 5℃, the failure surface is located above the soil-ice interface, and the strength of the specimen is similar to that of the frozen soil. When the temperature is − 10℃, the shear damage surface appears at the soil-ice interface, and the strength of the specimen is determined by the strength of the soil-ice interface. The soil-ice interface is a special interface formed by the cementation of soil and ice layers, which is widely distributed in the natural permafrost layer. Due to the special connection and structure between the ice and soil layers, the mechanical behaviors at the interface are largely different from those of general frozen soil. In the past 40–50 years, with the increase in temperature in the Qinghai-Tibet region, slope instability in cold areas has occurred frequently. The decay of the strength of the freeze–thaw and soil-ice interfaces may also be an important factor in slope damage. Therefore, it is necessary to study the shear properties of the soil freeze–thaw interface to understand the influence of the strength change of the soil-ice interface on slope stability.1. Extensive studies have shown that slope failure in permafrost areas is associated with the thawing of the active layer2, and the decay of strength at the freeze–thaw interface is the key to causing slope instability3. Moisture content and temperature are the two most important factors affecting the strength of frozen soil6. Meanwhile, a large number of studies have confirmed that the change of strength is mainly realized by changing the state of water through temperature9.Researchers have recently conducted numerous studies on slope instability due to active layer melting in permafrost regions11 found that the strength of the freeze–thaw interface was sensitive to the changes in temperature and moisture content through indoor tests. Ge Qi et al.12 studied the relationship between freeze–thaw interface shear strength and freeze–thaw cycles under low-temperature conditions. Wang et al.13 found that the variation of the strength of the freeze–thaw interface at different moisture contents was mainly related to cohesion. In addition, some researchers have studied the effect of soil type and confining pressure on strength14.Therefore, a series of indoor tests have been conducted by researchers to study the variation pattern of freeze–thaw interface strength under different conditions. Peng et al.15 conducted a series of tests with temperature-controlled direct shear apparatus to evaluate the effect of the thermal cycle on the strength of soil–the pile interface. Zhao et al.16 conducted a series of tests to analyze the effect of the rising frozen temperature, moisture content, and normal stress on the mechanical properties of the frozen soil-pile interface during the thawing process of permafrost. Janipour et al.17 investigate the interface shear strength characteristics between soil and concrete under static loads. Noroozi et al.18 evaluated the mechanical behavior of the interface between the filter soil material and the asphalt concrete core in the laboratory.In addition, Yazdani et al.19 used uniaxial tests to research the relationship between strength and water content of ice-rich frozen silty sand. Gao et al.14 conducted shear tests at positive temperature and shear tests at the soil-ice interface for different types of soils, respectively, and found that the strength of the soil-ice interface is higher than that of the thawed soil, and the type of soil affects the strength of the freeze–thaw interface. Shi et al.20 conducted shear tests at the soil-ice interface at negative temperatures and found that temperature affects shear characteristics.With the continuous research on the mechanism of strength change of frozen soil freeze–thaw interface and soil-structure interface, researchers have also started to pay attention to the effect of temperature on the strength of the soil-ice interface. Du et al.Most of the past studies are about the freeze–thaw interface and soil-structure interface. Researchers have a preliminary understanding of the mechanical mechanism of freeze–thaw interface and soil-structure surface based on a large number of experiments. The research on the soil-ice interface has also started, and the preliminary analysis of the influence of internal and external factors on the shear properties of the soil-ice interface has been carried out. Although some studies have explored the effects of temperature, initial moisture content, etc. on the strength of the soil-ice interface, these studies have not systematically investigated the development of the strength of the soil-ice interface at different temperatures. In addition, the experimental studies on the soil-ice interface are still relatively few, and the influence of the change of soil-ice interface strength on the stability of permafrost slopes and the shear characteristics of the soil-ice interface still needs further research.Therefore, this paper investigates the change law of soil-ice interface strength during the thawing process by improving the test method and conducting a controlled temperature direct shear test on the soil-ice specimens. The direct shear test of the soil-ice interface was carried out at different initial moisture contents , different temperatures , and different normal stresses . The variation laws of shear strength and shear index with temperature and moisture content were revealed, and the mechanism of shear strength decay of specimens during the thawing process was discussed. In addition, the changes in the location of the shear failure surface during the thawing process were also analyzed. And these discussions and analyses are of great significance for understanding the mechanical decay mechanism of the soil-ice interface, and also can provide some theoretical basis for the study of slope stability in a cold region.The soil, used in the direct shear test, is silty clay, which was all taken from a slope at the Qinghai-Tibet plateau. The grain size distribution and physical properties of the soil are shown in Fig. First, the dried soil samples were mixed with water according to different requirements, then wrapped with cling film and left for 24 h to ensure that the moisture content of the specimens remained consistent. Then, soil-ice specimens with a diameter of 61.8 mm and a height of 20 mm were prepared with a cutting ring. The specimens were formed as follows: first of all, a 10 mm thick pad was placed at the bottom of the cutting ring and covered with a layer of cling film. Then, the cutting ring was filled with distilled water and frozen at − 10 °C for 10 h. After the specimen was frozen, the ice that expands on the surface of the cutting ring was scraped off with a blade and the pad under the cutting ring was removed. Subsequently, the prepared filler is filled into the cutting ring in layers. After the soil sample was filled, the specimen was placed in an environmental box. The temperature was then set to − 10 °C and frozen for 5 h to allow the specimens to be formed, and the formed specimens are shown in Fig. The molded soil-ice specimen was placed into the cooled shear box. The shear box was then wrapped with an insulating film, leaving only the top surface as a channel for heat transfer. The specimen wrapped with the insulating film was placed in the environmental box, and the temperature was controlled to reach the designed shear temperature so that the specimen starts to heat from top to bottom. Once the temperature at the soil-ice interface of the specimen reached the design value, the shear test can be started. To ensure that the test temperature at the soil-ice interface reaches the shear design value, a comparison specimen was present during the test. Except for the one-meter thermometer buried at the soil-ice interface, all the operations of the comparison specimen were the same as those of the test specimen. When the temperature of the comparison specimen reached the shear design value, the shear test can be started.The initial moisture contents of the specimens were determined to be 9.0%, 16.0%, and 23.0% according to the corresponding moisture contents of the less-ice, more-ice, and ice-rich permafrost soils, respectively. To simulate the influence of the site environment on the shear strength of the soil-ice interface, the forming temperature was set to − 10 °C, and the shear temperatures were set to − 10, − 5, − 2, − 1, and 0 °C. Considering the sensitivity of the soil-ice interface to temperature, the rapid-shearing test was applied in this study. The normal stresses were determined as 100, 200, 300, and 400 kPa, and the shear speed was set to 0.8 mm/min. The test scheme is shown in Table The test apparatus consists of a direct shear apparatus, an environmental chamber (accuracy: ± 0.5 °C), an automatic data acquisition system, and a cylinder loading system, as shown in Fig. Figure The phenomenon observed from the shear test illustrated ice content plays a crucial role in the whole process of strength change with the increase in temperature.From Fig. Figure From the test results, the effects of both shear temperature and moisture content on shear strength are related to the state and distribution of water. The strength of the specimen at 0 °C decreases with increasing moisture content, which is similar to that of the thawed soil. Changes in the unfrozen moisture content and ice content of the soil under negative temperature are important factors affecting various other changes. In turn, the change of strength is a comprehensive reflection of various physical changes. At present, the unfrozen moisture content is a physical quantity that can be quantified more accurately. Therefore, this paper reveals the variation rule of soil-ice interface strength based on the unfrozen moisture content prediction model.21. In this paper, the curves of unfrozen moisture content with temperature for powdered clay at the initial moisture content of 23.0%, 16.0%, and 9.0% were plotted using the Anderson prediction model22, as shown in Fig. As we all know, unfrozen moisture content is closely related to temperature and the initial moisture content. Based on a lot of experiments, a series of prediction models of unfrozen moisture content with temperature have been proposed by some researchers23. It can be noted that the change in unfrozen moisture content is most pronounced in stage I and stage II. And the shear strength of the specimens in this paper also changed very significantly in these two stages, which indicates that the unfrozen moisture content is closely related to the strength change of the specimens. In addition, it can also be found in Fig. From Fig. 25. At a moisture content of 23.0%, as the temperature drops from 0 to − 2 °C, a large amount of unfrozen water freezes into ice causing the pores between the soil particles to be filled. Subsequently, as the temperature continues to drop, the pore ice content increases further. The volume expansion caused by the unfrozen water freezing into ice at this time will squeeze the soil particles and force them to move. And the mutual slip between soil particles disrupts the cementation between ice and soil particles, which weakens the contribution of ice cementation to cohesion, and thus cohesion decreases with temperature. It should be noted that due to the low content of unfrozen water when the temperature is lower − 2 °C, the cohesion is mainly influenced by the ice cementation effect, and the role of matrix suction can be ignored. It can be found that the variation of cohesion with temperature is closely related to the state of water in the specimen.From Fig. From Fig. When the moisture content is 9% or 16%, the generated pore ice does not fill the pores, while the pressure on the soil particles grows as the pore ice increases. The higher the pressure between soil particles, the higher the coefficient of friction, so the internal friction angle varies with the ice content. At a moisture content of 23.0%, the ice almost fills the pores of the soil. The increase of ice content at this time may make the pressure in the soil exceed the limit value, which causes displacement of the soil particles as well as ice particles. And the slippage between the particles will significantly increase the mechanical bite force between the grains. Therefore, when the moisture content is 23.0%, the lower the temperature, the greater the internal friction angle.From Fig. From Fig. According to the variation of shear strength and shear index under different conditions, it can be discovered that the effects of cohesion and internal friction angle on shear strength are not the same under different factors. When the moisture content is low (≤ 16.0%), the cohesion of the specimen starts to decrease with the increase in temperature, while the internal friction angle remains unchanged, and the change in the strength of the specimen is mainly due to the decrease of the cohesive force. When the moisture content is 23.0%, with the increase of temperature, the cohesion of the specimen increases and then decreases, while the friction angle decreases slowly and then decreases rapidly. At lower temperatures (< − 2 °C), the cohesion and the friction angle of the specimens increased with the moisture content, and the strength of the specimens was influenced by both together. At higher temperatures (− 2 °C ~ 0℃), the cohesion of the specimens increased slowly and then rapidly with the increase of moisture content, while the internal friction angle increased rapidly and then decreased. Although the change of cohesion was also significant, the strength of the specimens was more influenced by the internal friction angle.The results of the shear test illustrate the shear failure surface appears at the soil-ice interface and in the upper soil layer, as shown in Fig. The specimen consists of two parts: the ice layer and the soil layer. The strength of the specimen is determined by the strength of the ice layer, the strength of the frozen soil, and the strength of the soil-ice interfaces. In general, the strength of ice far exceeds the strength of soil, so the shear strength of the specimen is determined by the strength of the upper frozen soil layer and the soil-ice interface.It should be noted that the test results may be influenced by the size of the specimen and the location of the soil-ice interface during shear. However, the authors still found some regularity based on the multiple sets of experiments. According to the shear test results in Figs. 26 who believe that the variation trend is related to unfrozen water content. According to the shear failure morphology, we found that the shear failure surface is not always in the soli-ice interface. As the temperature increases, the weak surface of the specimen moves from the soil-ice interface to the upper soil layer. This finding is consistent with that of Gao et al.14 who also found that the weak surface of the specimen is located in the freeze–thaw interface or upper soil layer. In addition, it can be inferred from the shear damage morphology that the strength of the soil-ice interface is greater than that of the soil layer when the temperature is above − 5 °C, while the strength of the soil-ice interface is less than that of the upper soil layer when the shear temperature is − 10 °C. This finding is consistent with that of Chen et al.27 who believe that the variation trend is related to unfrozen water content.The objective of this study is to evaluate the interface shear behavior and strength of the soil-ice interface under different temperatures and moisture content. The shear failure surface of samples is also evaluated based on the shear morphology of specimens. The pattern of change in the strength of the specimens at moisture content and temperature is similar to that of the general permafrost soil. The development of strength is closely related to the variation of unfrozen water content. The cohesion decreases first and then increases with increasing temperature when the moisture content is 23%. This finding is consistent with that of Du et al.During the spring thaw season, the frozen soil is influenced by the external air temperature, and the heat energy is transferred from the surface to the interior, and the temperature near the outside is significantly higher than the temperature inside the permafrost. Therefore, the strength of the upper soil layer has been significantly decreased before the heat energy is transferred to the ice layer. At this time, the weak surface of the permafrost does not occur at the soil-ice interface but is located in the upper soil layer. From the available results and findings, it can be seen that the strength of the soil-ice interface is higher than the general permafrost at temperatures above − 5 °C. Therefore, the weak surface of the slope will not appear at the soil-ice interface during the spring thaw but will be located in the upper soil layer.The strength of the sample decreases with increasing temperature, and the change in strength is most significant from − 2 to − 0 °C. The strength reduction in this range is from 21.8 to 74.8%, and the higher the moisture content the more obvious this phenomenon is. The strength of the sample increases with increasing moisture content at negative temperatures, and the increase is more significant at lower temperatures. With the increase of moisture content from 9.0 to 23.0%, the strength increases by 55.1%, 48.4% and 46.8% at − 10 °C, − 5 °C and − 2 °C, respectively. The shear strength of the specimens decreased with the increase of moisture content at 0 °C due to the presence of large amount of free water.The changes of cohesion and internal friction angle are closely related to the unfrozen water content, and the changes of strength are influenced by both of them together. The shear index tends to decrease with the increase of unfrozen water content, and the greater the increase of unfrozen water, the faster the decrease of both, especially in stage 2. It should be noted that when the moisture content is 23%, the cohesion and the internal friction show the alternating growth and decline trends, and the angle of internal friction has a higher influence on the strength than the cohesion.When the temperature is higher than − 5 °C, the specimen failure surface is located in the upper soil layer, slightly above the soil-ice interface. At this time, the strength of the soil-ice specimen is determined by the strength of the upper soil layer, and the strength of the soil-ice interface is greater than the strength of the soil layer. When the shear temperature is − 10 °C, the soil-ice interface is the shear failure surface. At this time, the strength of the soil-ice interface is smaller than the strength of the upper soil layer.According to the experimental results presented here, the following conclusions can be drawn:


To investigate the operation principles of the automatic tube current modulation (ATCM) of a CT scanner, using a dedicated phantom and the CT dosimetry index (CTDI) phantom.vol and table position, image noise, and water equivalent diameter (WED) of each phantom CT image, respectively. ImQuest was also used to derive the detectability index (d’) of five different materials  embedded in the Mercury phantom.The Mercury 4.0 phantom and three different configurations of the CTDI dosimetry phantom were employed. A frequently used clinical scanning protocol was employed as a basis for the acquisitions performed with all phantoms, using both scanning directions. Additional acquisitions with different pitch and examination protocols were performed with Mercury phantom, to further explore their effect on ATCM and the resulting image quality. Different software named DICOM Info Extractor, ImageJ, and imQuest, were used to derive CTDIvol value is the WED values of the phantom structures laying ahead towards the scanning direction, for a length equal to the effective width of the X‐ray beam. Furthermore, it was also exhibited that ATCM does not fully compensate for larger thicknesses, since images of larger WED phantom sections present more noise (larger SD) in all four phantoms and in Mercury 4.0 phantom smaller detectability (d’).It was exhibited with all four phantoms that the scanning direction greatly affects the modulation curves. The fitting of the dose modulations curves suggested that for each table position what determines the CTDIMercury 4.0 is a dedicated phantom for a complete and in‐depth evaluation of the ATCM operation and the resulting image quality. However, in its absence, different CTDI configurations can be used as an alternative to investigate and comprehend some basic operation principles of the CT scanners’ ATCM systems. Traditionally, the primary goal of all AEC systems was to adapt the dose to obtain a uniform image quality irrespective of the scanned patient size.vol values) for each patient, depending on the attenuation characteristics of each anatomical region scanned.In old CT scanners, a rough dose adaptation for different patient sizes was performed manually, by selecting appropriately the tube potential (kVp), tube current (mA), and exposure time per rotation (t), but none of these parameters could be varied during scanning. In modern MSCT scanners, a fixed value of rotation time and kVp is still used, though the kVp value can also be automatically selected prior to the CT scan when the Automatic Tube Voltage Selection (ATVS) feature is available. However, mA can be continuously adapted during scanning, allowing for the modulation of tube current, which effectively leads to the modulation of dose  are currently installed, which employ different ATCM systems.Given the fact that the Mercury 4.0 is a relatively large, heavy, and expensive phantom that cannot be afforded by many medical physicists around the world, the proposed alternatives of CTDI dosimetry phantom were also utilized for ATCM performance evaluation of the same CT scanner in order to check whether, in absence of the Mercury phantom, the CTDI phantom can partly substitute it for the evaluation of ATCM systems.23 and CT# = −90 HU) and is made up of five cylinders and four cone‐rings  whose dimensions are described in Table The Mercury 4.0 phantom and three different configurations of the CTDI dosimetry phantom were used to evaluate the ATCM system of a Siemens Somatom Definition Flash CT scanner. The main material of the Mercury 4.0 phantom is virgin ultra‐high molecular weight polyethylene (density = 0.93 g/cmThe first CTDI phantom configuration used in this study (henceforth referred to as CTDI‐P1) is shown in Figure All phantoms were initially scanned utilizing the most often used clinical protocol for helical scanning (Thorax CT‐Chest Navigation) and two different scanning directions, that is, from head‐to‐feet (HF) and from feet‐to‐head (FH). The standard collimation of 64 × 0.6 mm was used, with a selected pitch of 0.8 and a planned scan length that was extending beyond the phantom boundaries by about 4 cm . It must be noted that the SPRs on which the scans were designed, were extending at least 10 cm beyond both ends of the phantom. The standard convolution kernel for chest was used (Br38) and a reconstructed slice thickness (ST) of 2 mm was selected. However, in the context of the investigation of the effects of different acquisition and reconstruction parameters, additional scans and reconstructions were made with Mercury 4.0 using: a) Pitch factors of 0.6, 1, and 1.2 (HF direction only), b) two additional clinical examination protocols for abdomen, one helical and one axial (both scanning directions). Details about the scan acquisitions are presented in Table imquest@duke.edu) that has been produced for evaluation of image quality with the Mercury phantom (and the ACR 464 phantom) according to the methodology described in the AAPM TG233 report, was used in this study for producing: a) the water equivalent diameter (WED) of each phantom with respect to the Z‐axis value (table position), b) various image quality indices, including the detectability index (d’) which is the only image quality parameter that will be presented in this paper (applicable only to the Mercury phantom). A free software  and in‐house build macros were also used to read the CT number and standard deviation (SD) of the reconstructed CT images. A free software , to present the dose modulation rather than mA modulation only, to also account for the effect of rotation time and pitch. For each one of these ATCM curves (henceforth referred to as dose modulation curves) two additional curves were constructed: WED (secondary Y‐axis) versus table position, and SD of each image (secondary Y‐axis) versus table position. All these curves were superimposed on scan projection radiograph images, where the positions of the first and the last image positions of each scan series are also depicted, so as to be able to observe the dose modulation (primary Y‐axis), with respect to the different phantom sections. In this way, the result of dose modulation was visually studied in terms of simultaneous variations of CTDIvol, WED and SD with respect to the position of the CT image within the phantom. The main results of the above analysis are also presented in additional graphs in terms of CTDIvol versus WED and in Table For all phantoms, the ATCM modulation curves were created in terms of CTDI3vol modulation curves for the acquisitions ID No 1 (HF) and ID No 2 (FH) using the Mercury phantom are depicted in Figure The CTDIvol peak is observed in the Air gap 1 region, about 20 mm prior to the abrupt increase in WED observed in Section 0. On the other hand, the dose modulation curve for the FH scanning direction (FH scan) seems to present a considerable delay regarding the WED increase observed in Section 0, since the first CTDIvol peak is found in Section 1. The difference between these two CTDIvol peaks is about 60 mm. When moving towards the other sections of the Mercury phantom in the HF direction, it can be seen that the spatial difference between the two dose modulation curves varies and decrease to about 40 mm, while at some points along the phantom the two curves intersect. The dose modulation curve (HF scan) starts to move abruptly downwards about 30 mm before the end of the thickest section of the phantom (Section 9), whereas the dose modulation curve (FH scan) starts descending abruptly about 20 mm after the end of the phantom, without reaching the minimum CTDIvol of the dose modulation curve (HF scan) observed in the Air gap 2 region. Despite the differences observed between the dose modulation curves for the two different scanning directions, the CTDIvol and DLP values for the two respective acquisitions were practically the same, as can be seen in Table When looking at both modulation curves starting from positive table position values (Head) and moving to negative table position values (Feet), it appears that the dose modulation curve for the HF scanning direction (HF scan) precedes the WED increments present in the WED curve along the HF direction. For example, the first CTDIvol decreases with WED decrease at the cone rings  and remains relatively constant at the cylindrical regions . On the contrary, the dose modulation curve (HF scan) exhibits a mixed behavior, for both cylindrical and conical parts of the phantom. For example, the CTDIvol is relatively constant in Section 2 (cone ring) instead of increasing, it increases in Sections 5 and 7 (cylinders) instead of remaining constant, and it increases in Section 8 (cone ring) as it should. This behavior can be better appreciated in Figure vol data points of Figure vol values of the dose modulation curve (FH scan) follow an exponential‐like variation in response to the respective WED variations, in contrast to the dose modulation curve (HF scan) which looks more like a four‐step function.However, when considering that the dose modulation curve (FH scan) results when the Mercury phantom is scanned in the opposite direction than the dose modulation curve (HF scan), it becomes rather clear that the changes in the dose modulation curve for the FH scanning direction variations also precede the respective WED variations. A closer look at the modulation curves reveals that the dose modulation curve (FH scan) follows closer to the WED variations, as the CTDIvol values observed in the dose modulation curve (FH scan), resulted in smaller SD values. The variation of SD values with WED depicted in Figure In Figure vol and DLP values for the two respective acquisitions were practically the same , peak before the phantom middle and decrease thereafter, reaching the minimum slightly before the end of the phantom. When looking at the other dose modulation curve (FH scan) in the FH direction, the same observations apply.The dose modulation curves derived using the CTDI‐P1 phantom and opposite scanning directions are depicted in Figure me Table . The twovol and DLP values for the two respective acquisitions with CTDI‐P2 were the same  rather than the LAT direction (X‐axis).The dose modulation curves derived using the CTDI‐P2 phantom are depicted in Figure vol values of Figures vol values are larger when scanning the first half of the phantom, compared to the respective values arising when scanning the second half, which has been recently reported in terms of mA modulation by Papadakis and Damilakis.In Figure vol and DLP values for both scans were practically the same , in steps of 16, 20 and 24 mm.The response of the ATCM system dose at the interfaces may be better comprehended by looking at the dose modulation curves for the axial acquisitions ID No 8 and 9), shown in Figure  and 9, svol values of the dose modulation curves of the helical acquisitions considering the respective WED curves. Several relationships were tested and the one that gave the best fit was the following:1,w2,w3,w4 are fitting parameters (w1+w2+w3+w4 = 1), z is the table position of each reconstructed slice and d is the range of images ahead that is taken into account in the CTDIvol  value selection.To understand how this modulation works, it was attempted to find a relationship that fits the CTDIvol data from the Mercury phantom using the WED diameters of the five cylinders and comparing them with data from dose modulation curves in actual patients. The difference is that combining Equation  ahead of the table position that correspond to each CT image towards the scanning direction (HF or FH) are considered for the mA modulation of the scans derived with pitch 0.8, and it is not only the maximum and minimum WED value within this table positions’ range that is considered, since the best fitting was obtained with none of the w1,w2,w3 and w4 being zero (a single set of fitting parameters was used for both modulation curves). The results of this fitting method for the acquisitions performed with the Mercury 4.0 phantom and the abdomen protocol (ID No 6 and 7) are shown in Figure Equation  is the sEquation  with EquEquation  we attemThe dose modulation curves for acquisitions ID No 3 (pitch = 0.6), ID No 1 (pitch = 0.8), ID No 4 (pitch = 1.0) and ID No 5 (pitch = 1.2) with the Mercury 4.0 phantom are shown in Figure vol values were larger, leading to either smaller noise values or/and larger Task Transfer Function (TTF) values, since d’ is a function of TTF and noise power spectrum (NPS), and increases for larger TTF values and smaller values of noise.Finally, results of the image quality evaluations of the acquisitions of Table 4The scanning direction greatly affects the modulation curves, since from the fitting procedure it is suggested that at each table position the WED values of the phantom structures that lay ahead (towards the scanning direction) within the X‐ray beam width range, are accounted to select the right tube current and consequently the right dose. The effect of the scanning direction on dose modulation was verified using all four phantoms. Therefore, even when scanning a perfectly symmetrical phantom, like the CTDI‐P1 and CTDI‐P2 the dose modulation curves for opposite scanning directions may differ in other CT scanners as well.According to the results presented the main findings of this study are the following:The current adjustment of ATCM modulation strength does not fully compensate for the larger phantom thickness, since the phantom regions with larger WED present larger noise. This was also verified using all four phantoms.The dose modulation is affected by the pitch selection, and with smaller pitch values the ability of the ATCM system to follow the WED changes along the z‐axis is improved, since the effective X‐ray beam width is reduced.vol values at the uniform part of the smallest diameter cylinder where the noise is determined. It must be noted that as can be seen in Figures vol values of the helical abdominal and axial abdominal scans were respectively more than twice and three times the CTDIvol value of the thorax protocol.Apart from WED, the dose modulation is greatly affected by the noise level required for each examination protocol. As can be seen in Table Despite what suggested by the results of the fitting procedure, it was notable that a unique set of fitting parameters that could reproduce the dose modulation curves for all the phantoms could not be found. The exact way that the mA values are calculated is not known but most probably it is based on a proprietary algorithm, that may use the WED values  of the SPR images, to determine which is the most attenuating phantom segment that is included within the X‐ray beam width, so as to adjust the mA for the maximum WED and also apply angular modulation at different rotation angles.The variation of SD values with WED depicted in Figures It must be noted that for all phantoms, the air gaps before and after the phantom were intentionally used, to investigate how the ATCM responds in sudden changes, that is from air to a medium and from a medium to air. Apart from this, air gaps are also useful for detecting the min–mA allowed.The main limitation of this study is that the results were obtained from a single modern CT scanner and therefore cannot be generalized to other CT scanners of the same or other manufacturers, though it is reasonable to assume that a similar behavior should be expected. A second limitation is that there are many additional parameters that are expected to affect dose modulation and the resulting image quality  which could not be presented in this paper because of space limitations. However, this study highlighted that the scanning direction is an important parameter which should be considered when the ATCM operational parameters of other CT scanners are explored, and also have set a framework for the measurements that should be performed when comparing different CT scanners and examination protocols.It must be noted that the Mercury phantom is a better fit to the modern methods required to evaluate image quality in CT scanners using iterative reconstruction algorithms, since they are nonlinear and the conventional image quality metrics like the SD are of limited utility.5The Mercury 4.0 phantom (in conjunction with the ImQuest software), is a dedicated phantom for a complete and in‐depth evaluation of the ATCM operation and the resulting image quality of modern CT scanners, using the new methodologies proposed in the AAPM TG233 report. However, in its absence, the CTDI configurations can be used instead, to investigate and decode the basic operation principles of the ATCM systems of CT scanners from different manufacturers, especially when the imQuest software is available to facilitate the WED calculations. If the imQuest software is not available, there are other free software (ImageJ and DICOM Info Extractor) which can be used instead to obtain the SD value curves as a function of dose modulation curves.For the CT scanner tested, it was seen with all four phantoms that the operation of ATCM system is intensely affected by the scanning direction and that image quality is reduced with increasing phantom thickness. From acquisitions performed in the Mercury 4.0 phantom only, it was seen that the dose modulation and the resulting image quality are also affected by the pitch selection and the examination protocol selection.Ioannis A. Tsalafoutas and Shady AlKhazzam conceived the idea, designed and performed the data acquisition and analysis, and wrote the manuscript. Huda AlNaemi and Mohammed Hassan Kharita contributed to interpretation of data, revising and approval of the manuscript's final version.The authors have no conflict of interest to declare.Table A1Click here for additional data file.Figure A1Click here for additional data file.Figure A2Click here for additional data file.Figure A3Click here for additional data file.Figure A4Click here for additional data file.Figure A5Click here for additional data file.Figure A6Click here for additional data file.Figure A7Click here for additional data file.