GJ/L  may severely underestimate the torsional stiffness of thin-walled nanostructural members, due to neglecting element warping deformations. In order to investigate the exact element torsional stiffness considering warping deformations, this paper presents a matrix stiffness method for the torsion and warping analysis of beam-columns. The equilibrium analysis of an axial-loaded torsion member is conducted, and the torsion-warping problem is solved based on a general solution of the established governing differential equation for the angle of twist. A dimensionless factor is defined to consider the effect of axial force and St. Venant torsion. The exact element stiffness matrix governing the relationship between the element-end torsion/warping deformations (angle and rate of twist) and the corresponding stress resultants (torque and bimoment) is derived based on a matrix formulation. Based on the matrix stiffness method, the exact element torsional stiffness considering the interaction of torsion and warping is derived for three typical element-end warping conditions. Then, the exact element second-order stiffness matrix of three-dimensional beam-columns is further assembled. Some classical torsion-warping problems are analyzed to demonstrate the established matrix stiffness method.The typically-used element torsional stiffness  In recent decades, research on nanomaterials  has made great progress and is widely used in scientific research and industrial production. Some formations of oxides, metals, ceramics, and other substances have been discovered. These nanomaterials obey the fundamental laws of the classical physics governing the macroworld . Nanomatiu, id, and iq are element-end axial displacement, translational displacement, and rotation angle, respectively; iF, iV, and iM are element-end axial load, shear force, and bending moment, respectively, as shown in Ke] is the element stiffness matrix for flexural-axial problems.Although a number of computer programs such as Abaqus, OpenSees, Ansys, and MSC.Marc are readily available for the structural analysis of thin-walled nanostructural members, approaches to obtain the exact solutions ,7,8,9,10E denotes the elastic modulus; A denotes the cross-sectional area; I denotes the cross-sectional moment of inertia; L denotes the element length; and Tc, Qc, Sc, and Cc are coefficients in the element stiffness matrix as functions of factor λ, formulated as Equation (4); λ  denotes a dimensionless factor for the axial compression force Pc, defined as Equation (5).For first-order e.g., ,22) and  and 22])Using the element stiffness matrix, many different types of analyses can be conducted, including the traditional analysis for the element deformations and internal forces e.g., ,22), as , as 22])G denotes the shear modulus; J denotes the St. Venant torsion constant; yyI and zzI denote the cross-sectional moments of inertia about y-axis and z-axis, respectively; λy and λz denote axial force factors associated with the y-axis bending and z-axis bending, respectively; The matrix stiffness method can also be used for three-dimensional analysis of beam-columns. For second-order analysis, Ekhande  presenteGJ/L as torsional stiffness in engineering practices, which may severely underestimate the torsional stiffness of thin-walled nanostructural members without considering warping deformations. In order to present the exact element stiffness matrix for the second-order analysis of three-dimensional beam-columns considering torsion and warping, this paper investigates the interaction of the axial force, torque, and bimoment of torsion members. The exact element torsional stiffnesses are derived and shown to be significantly larger than the typically-used GJ/L. The exact element stiffness matrix associated with the element-end torsion and warping deformations can be obtained, and the exact three-dimensional element stiffness matrix can be further assembled. Some application examples of the exact element stiffness matrix for torsion and warping are also presented.However, these researchers neglected the interaction between torsion and warping and the axial force effect in torsional analysis. They directly used In addition, the temperature is non-negligible, which can also influence the bifurcation buckling of thin-walled structures, and the thermoelastic analysis of structures has already been presented in many papers. However, this paper is focused on solving the mechanical buckling of thin-walled members. Therefore, we list some papers ,25 on thTo develop the matrix stiffness method, an equilibrium analysis of an axial-loaded torsion member  is conducted, as shown in Besides the usual assumptions of the linear theory of elasticity, the following assumptions ,26,27,28The global cross-sectional deformation assumption, i.e., the cross-section is assumed to be perfectly rigid in its own plane while free to warp out of its plane;The classical Kirchhoff–Love’s thin plate bending assumption, i.e., straight lines normal to the mid-surface of the thin-walled plates remain straight, inextensible, and normal to the mid-surface after deformation;Each thin-walled plate is assumed to have null mid-surface membrane shear strains (Vlasov’s hypothesis) and null transverse extensions.φ and the cross-sectional bimoment ωB is formulated as Equation (7). The St. Venant torque Ts is formulated as Equation (8)  at the two cross-sections with distance x of the element short segment are in the opposite direction, and they combine to a bimoment increment formulated asσn on the top section is inclined to produce a shear stress σndAsdφρ in the horizontal plane. Therefore, the Wagner effect can be considered by taking moment of the shear stress about the shear center S, derived as Equation (12).σn = P/A denotes the assumed uniformly-distributed axial stress from the axial force; Sρ denotes the distance of a point in the cross-section to the shear center S; and ISp, denotes the polar moment of inertia about the shear center S.In addition, due to the torsion of the short segment .λc/t denotes a dimensionless factor for the effect of axial force and St. Venant torsion, defined as Equation (16); the subscripts “c” and “t” are associated with conditions of PISp,/A ≥ GJ and PISp,/A < GJ, which can be defined as generalized “axial compression” and “axial tensile” situations, respectively.Then, combining Equations (10) and (13) and considering Equation (7), the governing differential equation of equilibrium for this torsion-warping problem can be established as Equation (14), which can be derived to a dimensionless form as Equation (15).Q1, Q2, Q3, and Q4  are deformation combination factors defining the possible deformation curve.For null distributed torque (τ = 0), the general solution for the differential equation of equilibrium is given byIn the following, the derivations will focus on the generalized “axial compression” situation. It is noted that the derivations and the results for the generalized “axial tensile” situation is very similar to the “axial compression” situation, and the main difference is the use of hyperbolic trigonometric functions instead of trigonometric functions [as shown in Equation (17)].Based on Equations (7) and (13), the element stress resultants (bimoment and torque) can be derived as Equations (18) and (19), respectively.Analysis associated with the element deformations and stress resultants of torsion members can then be conducted using Equations (17)–(19).An element stiffness matrix, showing the relationship between the element-end deformations (angle and rate of twist) and the corresponding stress resultants (torque and bimoment), is formulated based on the previous section. Then, a matrix stiffness method for torsion and warping is established for the analysis of torsion members, including the approximate element torsion-warping stiffness matrix for simpler applications and the torsional stiffness analysis for three typical element-end warping conditions.Equations (17)–(19) can be used in the analysis associated with the element deformations and stress resultants of torsion members. In a matrix stiffness method, attentions are focused on the element-end torsion/warping deformations and the corresponding stress resultants, formulated in matrix forms for simplified and systematic analysis.The element-end deformations (angle and rate of twist) are formulated in a matrix form as Equation (20) based on Equation (17). The element-end stress resultants (torque and bimoment) are formulated in a matrix form as Equation (21) based on Equations (18) and (19).The relationship between the element-end deformations (angle and rate of twist) and the corresponding stress resultants (torque and bimoment) is then formulated as Equation (22) by combining Equations (20) and (21).Ke] is the element stiffness matrix for torsion and warping, which can be simplified asTc, Qc, Sc, and Cc are the same as that in the element stiffness matrix for a flexural-axial problem, which are formulated in Equation (4).λc (Equation (16)) in the case of PISp,/A ≥ GJ.Equation (23) gives the element stiffness matrix based on the axial (compression) force factor PISp,/A < GJ, the element stiffness matrix can also be derived from the general solution in Equation (17) and is then formulated in the same form as Equation (23), with a change of the subscript “c” to “t”. However, the element stiffness coefficients Tt, Qt, St, and Ct are expressed in a form using hyperbolic trigonometric functions based on the axial (tension) force factor λt. In the case of Tc, Qc, Sc, and Cc correspond to the element-end stress resultant (torque or bimoment) for a unit element-end deformation (angle or rate of twist). These coefficients are transcendental functions of the factor λc for St. Venant torsion and axial force. By noting that GJ−PISp,/A on these element stiffness functions are plotted in Tc, Qc, and Sc increase (while Cc decreases) with the increase in GJ−PISp,/A .The element stiffness functions Tc, Qc, Sc, and Cc are given as Equation (25) based on the Taylor series of the functions at λc = 0. The approximations are plotted as the dashed lines in GJ−PISp,/A is in a range between −π2ωωEI/L2 and π2ωωEI/L2.The linear approximations of the transcendental element stiffness functions Ke] can be approximated as Equation (26) for simpler applications, which shows the linear influences of the warping constant ωωI, the St. Venant torsion constant J, and the axial force P on [Ke]. This approximated element stiffness matrix can also be derived by using an energy approach and assuming a cubic deformation shape function , we can derive the torsional stiffnesses of members with typical element-end warping conditions.By using the element torsion-warping stiffness matrix [Ke] related to the rate of twist and bimoment can be deleted to obtain a torsion stiffness matrix.For members with restrained warping at the ends (ktor for this restrained-restrained warping condition can be expressed asωB1 = ωB2 =0), the torsion stiffness matrix that relates the element-end twisting angles and torques can be derived as follows.For members with no restraint of warping at the ends (ktor for this free-free warping condition is derived asBy using the element stiffness functions in Equation (23), the torsional stiffness ωB2 = 0), the torsion stiffness matrix can be derived as follows.For members with restrained warping at one end . The stiffness function TF and its approximation are also plotted in By using the element stiffness functions in Equation (23), the torsional stiffness GJ/L for the torsional stiffness is only valid for members with the free-warping condition and negligible axial force effect. GJ/L. Therefore, in structural analysis, the torsional stiffness value should be carefully selected based on the element-end warping conditions.In summary, the torsional stiffnesses of members with three typical element-end warping conditions can be formulated asIn this section, some classical torsion-warping problems will be analyzed as application examples to demonstrate the established matrix stiffness method.Ks]. The eigenproblem for an elastic buckling analysis tries to solve a nontrivial solution for the global load deformation equation [Ks]Δ = 0. Therefore, an elastic buckling analysis can be conducted by setting the determinant of the assembled structural stiffness matrix [Ks] to zero (Equation (35)). The factor λc,cr at the buckling state can be solved, corresponding to a buckling axial force Pcr formulated as Equation (36).The elastic bifurcation buckling analysis can be directly conducted using the assembled structural stiffness matrix ) equals zero, and the buckling condition in this case can be solved asFor a torsion restrained column  equals zero, and the buckling condition in this case can be solved asFor a column with torsion and warping restraints only at one node . Therefore, the buckling state corresponds to the condition that Tc(λc) = 0. Based on Tc decreases to 0 as λc increases to π. Therefore, the buckling condition in this case can be solved asFor a column with torsion and warping restraints at one node  (case 4), the only unrestrained element-end deformation is the warping deformation at the other node. The relationship between this unrestrained warping deformation and the corresponding element-end bimoment can be defined by Sc(λc). Therefore, the buckling state is corresponding to the condition that Sc(λc) = 0. Based on Sc decreases to 0 as λc increases to 1.43π (π/0.7). Therefore, the buckling condition in this case can be solved asFor a column with torsion and warping restraints at one node , the buckling state is corresponding to the condition that the denominator als zero . It can It is noted that this section gives the same results as that from classical analyses of these torsion-warping-axial buckling problems ,5,7,8, bThis example analyzes a classical torsion-warping problem ,5, the tKs] is the structural stiffness matrix associated with the unconstrained deformations φ2, L  ) are both governed by the element stiffness matrix [Ke] [Equation (23)]. By combining the stability stiffness matrices of element I .The governing equation is formulated as Equation (42), where , as shown in Equation (46). The integration of cross-sectional total torque (T = Tmid) from the element end to the midspan can be formulated as Equation (47).For this problem particularly, the element analysis can be conducted for the element I or II. The integration of different torque components (St. Venant torque and warping restraint torque) is analyzed ,5. The iGJ/ωωEI), the ratio of the accumulation of warping restraint torque decreases, while the ratio of the accumulation of St. Venant torque increases. Therefore, for thin-walled cross-sections with relatively large St. Venant torsion rigidity GJ , the St. Venant torque could be dominant in the total torque; in contrast, for cross-sections with relatively small St. Venant torsion rigidity GJ , the warping restraint torque could be dominant in the total torque.Then, the ratios of Equation (45) to Equation (47) and Equation (46) to Equation (47) are formulated as Equations (48) and (49), respectively. Equation (48) represents the ratio of the accumulation of St. Venant torque along the element length to the accumulation of total torque, and Equation (49) represents the ratio of the accumulation of warping restraint torque to the accumulation of total torque. The two ratios are plotted in relationships with the factor for St. Venant torsion. As shown in It is noted that this section gives the same results as that from the analysis in Trahair et al.  and ChenBased on the torsion-warping stiffness matrix, this section establishes the element stiffness matrix of axial-loaded three-dimensional beam-columns with a symmetric cross-section.x, y, z) denote the local coordinate systems for the three-dimensional element.The element-end displacement vector of a three-dimensional beam-column with a symmetric cross-section is considered in Equation (50) to include the displacements, the torsional and rotational angles, as well as the rates of twist, as defined in The associated element-end stress resultant vector is considered in Equation (51) to include the forces, bending moments and torques, as well as the bimoments.x-y and x-z planes and (2) the torsion-warping stiffness matrix (Equation (23))established in this paper, the element stiffness matrix of three-dimensional beam-columns can be obtained:The element stiffness matrix of three-dimensional beam-columns relates the element-end displacement vector in Equation (50) to the element-end stress resultant vector in Equation (51), and it is, therefore, a 14-degree-of-freedom stiffness matrix. McGuire et al.  noted thIn view of Equation (52), the stiffnesses associated with the y-axis bending, the z-axis bending, and the torsion and warping are uncoupled. Therefore, row/column 2, 6, 8, 12 of Equation (52) represents the z-axis bending of the element, row/column 3, 5, 9, 11 of Equation (52) represents the y-axis bending, and row/column 4, 10, 13, 14 of Equation (52) represents the torsion and warping.For beam-columns with the three typical element-end warping conditions discussed in ktor in the 4 and 10 rows/columns should be determined using GJ/L).In view of Equation (53), the torsional stiffness This 3D element stiffness matrix can be readily used to solve the exact solutions of the bifurcation buckling problem of 3D frames as well as the out-of-plane buckling of funicular arches.This paper presented a matrix stiffness method for the analysis of torsion and warping that is particularly important in beam-columns with torsional deformations. The main works and conclusions are summarized as follows:Equilibrium analysis of an axial-loaded torsion member was conducted based on the equilibrium conditions of torque and bimoment of an element short segment, and a governing differential equation of equilibrium for the angle of twist along the member was established. The solution of the governing differential equation can be used to analyze the element deformations (angle and rate of twist) and stress resultants (torque and bimoment).The exact element stiffness matrix of the torsion member was formulated, showing the relationship between the element-end torsion/warping deformations and the corresponding stress resultants. A dimensionless factor for the effect of St. Venant torsion and axial force was defined. The element stiffness matrix for torsion and warping was readily used for the bifurcation buckling and second-order analysis of axial-loaded torsion members. The exact element second-order stiffness matrix of three-dimensional beam-columns was further assembled.GJ/L for the torsional stiffness is only valid for members with the free-warping condition and negligible axial force effect. For members with restrained warping at the ends, the torsional stiffness can be significantly larger. For a notable axial force effect, the torsional stiffness may be reduced. Therefore, in the analysis of thin-walled nanostructural structures, the torsional stiffness value should be carefully selected based on the element-end warping conditions and the axial force level.Based on the element torsion-warping stiffness matrix, the exact element torsional stiffnesses considering the interaction of torsion and warping were derived for three typical element-end warping conditions. The commonly-used expression


The VIABAHN stent graft is often used for traumatic and iatrogenic vascular injuries. In this case, vascular injury at both edges of the VIABAHN stent graft was noted 4 months after endovascular repair for idiopathic superficial femoral artery (SFA) rupture. The longitudinal compression of the SFA with a decrease in hematoma size was assumed to exceed the flexibility of the stent graft. Thus, the use of stent grafts for vascular injuries with giant pseudoaneurysms may result in late vascular injuries at both edges of the stent graft. Therefore, cautious assessment of its indications and strict surveillance are required. Its use has been correlated to the risk of late graft occlusion. However, the occurrence of late vascular injury due to the use of a stent graft placed in an appropriate position has been deemed rare.The VIABAHN®  stent graft has been determined to be effective for the treatment of peripheral artery disease (PAD) in the superficial femoral artery (SFA). Moreover, it is often used for traumatic and iatrogenic arterial injuries due to its minimally invasive property and long-term outcomes.In this study, we report a case of a patient who presented with late vascular injuries at both edges of the stent graft (VIABAHN) after endovascular treatment for idiopathic SFA rupture with a giant pseudoaneurysm.A 61-year-old Asian woman was admitted to the emergency room at our institution due to exacerbating pain in the right thigh and swelling 2 days back. Two months prior to admission, the patient reportedly underwent patch-closure surgery for a ruptured sinus of Valsalva and re-surgery for re-rupture of the sinus of Valsalva after 1 month . The patient was taking warfarin after cardiac surgery.The patient’s level of consciousness was found to be normal, and she was not in a state of shock. Significant swelling and pain in the right thigh were observed. However, no signs of infection were noted, and there was no ischemia in the right lower extremity. Blood examination findings upon admission were as follows: white blood cell count, 7350/µL; C-reactive protein level, 5.23 mg/dL; and procalcitonin level, 0.21 ng/mL. As per her contrast-enhanced computed tomography (CT) scans, a giant pseudoaneurysm (with a diameter of 10 cm) was found in her right thigh with extravasation from the main trunk of the distal SFA. A shift and lumen narrowing of the SFA due to pseudoaneurysm displacement were observed. The three tibial vessels were clearly visible.The patient was then diagnosed with SFA rupture and immediately underwent surgery using the endovascular stent graft approach. A 7-Fr sheath was placed into the right SFA with a right common femoral artery antegrade puncture. A 0.018-in stiff wire was cannulated into the below-the-knee artery, and the VIABAHN  stent graft was placed, without predilatation, from the distal SFA  to the above-the-knee popliteal artery  superior to the patella  from the middle SFA to the below-the-knee popliteal artery, embolization of the behind-the-knee popliteal artery , embolization of the proximal edge of the stent graft , and ligation of the SFA at the peripheral area of the proximal graft anastomosis , and imaging results. Furthermore, the presence of true aneurysms was also dismissed, as per her CT findings. Therefore, the diagnosis was idiopathic SFA rupture, which is determined to be a rare condition. To the best of our knowledge, only seven other cases presenting with this condition have been reported worldwide. The cause of idiopathic SFA rupture remains to be unknown. In young patients, some investigators have reported connective tissue disorder or congenital arterial abnormalities, such as Ehlers–Danlos syndrome, as the possible causes of idiopathic SFA rupture. Meanwhile in elderly patients, atherosclerosis is considered to be its primary contributing factor. But in this case, systemic atherosclerosis might have caused both rupture of the sinus of Valsalva and SFA.Pseudoaneurysms are often caused by trauma and infections, or they may be iatrogenic. In this case, an endovascular approach was determined as the most suitable treatment because of the presentation, history of cardiac surgery, intake of warfarin, and presence of a giant pseudoaneurysm. We opted to not select open surgery considering the risks of general anesthesia complications, massive bleeding, and wound infections. Because the rupture hole and pseudoaneurysm were large, coil embolization and thrombin injection were deemed to be inappropriate. Percutaneous insertion of a stent graft was instead selected because there were healthy landing zones, both proximal and peripheral of the rupture site, greater than or equal to 2 cm, which did not come in contact with the joints.An appropriate treatment regimen is chosen based on the rupture site, cause of the rupture, presence of an infection, peripheral ischemia, age, and performance status. There are also reports examining late vascular injuries at stent graft edges after using the VIABAHN stent graft across the joints (non-stenting zone) or using an oversized VIABAHN. However, no study has assessed the occurrence of late arterial injuries in patients with PAD using a proper size VIABAHN stent graft for SFA with proper landing zones , avoiding the hip or knee joints (non-stenting zone), which is attributed to excellent flexibility and the graft material.Although there have been a number of reports on late vascular injury after bare stent insertion into SFAs avoiding knee joints, most cases are due to the fracture of the stent body, and not the edges, owing to physiological loading; thus, the deformity and torsion of the distal SFA could be due to muscle movement with knee joint motion.In this case, the cause of arterial injuries at the edges of the stent graft might have been the insertion of the stent graft into the SFA in the state of being extended and bended by the giant pseudoaneurysm due to idiopathic SFA rupture. CT scan at 19 days after endovascular treatment revealed a tendency for repositioning and longitudinal compression of SFA due to a decrease in hematoma size, and CT scan after 4 months revealed further repositioning and compression of the SFA and significant meandering of the stent graft ( Recently, endovascular repair using stent grafts across the knee joint was performed for popliteal artery aneurysms. However, the long-term outcomes of placement at the same site are yet to be clearly elucidated, and this treatment option was not selected in this case due to the risk of recurrence of vascular injury at the stent graft edges. Exclusion and bypass procedures from the SFA to the below-the-knee popliteal artery were selected. However, autovein was not used considering the risk of coronary artery bypass and good runoff below-the-knee artery.The surgical bypass procedure and stent graft reinsertion have been considered for new pseudoaneurysms.In this patient’s case, the rupture of an infectious aneurysm was not completely ruled out but was thought to be unlikely as there were no findings of an infection throughout the treatment process and also the hematoma quickly disappeared after the exclusion procedure alone, without drainage or long-term use of antibiotics.Thus, cautious follow-up, including assessments of graft occlusion, infection, anastomotic pseudoaneurysms, and blood vessel rupture at other sites, is necessary.As per the findings of this study, it was determined that endovascular treatment using stent grafts for arterial ruptures with giant pseudoaneurysms may result in new arterial injuries at the stent graft edges due to the repositioning and longitudinal compression of the artery associated with the decrease in hematoma size after treatment. Therefore, cautious assessment of indications and strict surveillance are a must.


In delayed or neglected cases of flexor tendon injury, reconstruction of flexor digitorum profundus (FDP) is usually performed using free tendon graft due to the retraction of tendon ends and shortening of the tendon. Flexor digitorum superficialis (FDS), palmaris longus or plantaris tendons can be used as a free tendon graft [1–3].This is a case report of female patient 17 years old with neglected cut of her left Ring finger's FDP and FDS tendons zone II with suspected concomitant digital nerve injury, the injury was neglected for 10 years in the patient's non-dominant hand.Upon exploration unusual finding of spontaneous healing of the proximal stumps of FDS and FDP tendons raised the idea of doing the repair one stage without free graft by using pedicled intra-synovial graft from the sublimis tendon to reconstruct the FDP tendon.The patient after 4 months follow-up and after completion of the physiotherapy program regained the ability to actively flex her finger to near full flexion with improved function and cosmesis.Delayed flexor tendon reconstruction in neglected cases is still offering good results even after long periods of delay provided that the finger's joints are still supple and mobile. •In neglected cases of flexor tendon injury, reconstruction of FDP tendon is usually performed using free tendon graft.•The presence of two repair sites with free graft in between carries the risk of rupture or adhesion formation around the tendon graft.•In this case report we succeed in doing the repair one stage without free graft.We used pedicled intra-synovial graft from the sublimis tendon to reconstruct the FDP tendon. Paneva-Holevich has modified a technique for staged flexor tendon reconstruction by using pedicled intra-synovial graft from sublimes tendon which offers less incidence of adhesions if compared to the free tendon graft ,6,7.In our case report we did one stage surgery using pedicled intra-synovial graft from sublimes tendon after spontaneous healing of the proximal stumps of FDP and FDS tendons.The author has nothing to disclose. No funding was received for this article.Written informed consent was obtained from the patient for publication of this case report and accompanying images. A copy of the written consent is available for review by the Editor-in-Chief of this journal on request.


All-inside anterior cruciate ligament reconstruction (ACLR) is a novel technique that has gained attention due to its minimally invasive and graft-saving properties. However, studies comparing MRI-based graft maturity between all-inside and standard ACLR are lacking.This study focused on the functional, knee laxity, and MRI-based graft maturity characteristics of all-inside and standard single-bundle ACLR.Randomized controlled trial (RCT).n = 27) or standard reconstruction group (n = 27). Using the same rehabilitation strategy. The Tegner, International Knee Documentation Committee, and Lysholm scores were recorded at postoperative months 3, 6, and 12 to assess functional recovery. MRI was conducted to measure the signal/noise quotient (SNQ) of the intra-articular graft to assess the maturity. A higher SNQ indicates lower graft maturity. Knee laxity was assessed using GNRB arthrometer at the postoperative month 12.Fifty-four patients were randomly assigned to an all-inside reconstruction group (p < 0.05). There was no statistical difference in graft SNQ between the two groups at postoperative months 3 and 12 (p > 0.05). Both groups exhibited the highest SNQ in the middle region of the graft, followed by the proximal region, and the distal region. Functional scores improved significantly for both groups and had no statistical difference (p > 0.05). The knee laxity was higher in the all-inside group (p < 0.05) at postoperative month 12. There was no correlation between the functional scores and graft maturity in both groups (p > 0.05).The graft SNQ of the all-inside group was significantly higher than that of the standard group at postoperative month 6 (All-inside and standard single-bundle ACLR show good functional outcomes; however, knee laxity was relatively higher in the all-inside ACLR group than in the standard ACLR group. Moreover, both techniques exhibited poor maturity in the middle graft region and the best in the distal region. Graft maturity with all-inside ACLR is inferior to that with standard ACLR in the early postoperative stages. There is no correlation between knee function and graft maturity.ChiCTR1800018543.Clinical trial registration numbers: Date of registration: 09/23/2018. Anterior cruciate ligament (ACL) tear is one of the most common sports-related injuries . The oddMany studies have concluded that there is no difference between all-inside and standard ACLR in terms of functional recovery and that all-inside ACLR is better than the latter in terms of pain reduction and slowing tunnel widening –10. LuboThis was a prospective randomized controlled trial and was approved by the hospital’s ethics review committee and registered at the national clinical registration center.Patients who underwent single-bundle ACLR in our department between September 2018 and July 2019 and meet the criteria were included in the study. All patients were randomly assigned to the all-inside ACLR group or the standard ACLR group using the random number table method. The sample size was determined via power analysis. The inclusion criteria for participants were history of trauma and MRI showing ACL discontinuity or hyperintense signal, age 18–45 years, body mass index (BMI) of 18–28, and informed consent and signed relevant documents. The patients were excluded if they had any of the following: injured > 12 months ago; combined collateral ligament, posterior cruciate ligament, cartilage injury of more than 3 degrees and a meniscus grade III injury; history of ACLR surgery; generalized ligament laxity; severe underlying diseases or uncooperative follow-up; and bilateral ACL injury.The Tegner, International Knee Documentation Committee (IKDC), and Lysholm scores were recorded at postoperative months 3, 6, and 12 to assess functional recovery. At postoperative month 12, a GNRB arthrometer  was used to assess knee laxity. Arthrometer measurement was performed as follows, under knee flexion of 30°and at 134-N load, the tibia anterior movement of the ACLR and normal contralateral knee was measured separately. The difference is obtained by subtracting the two measurements. The measurement was repeated three times and the value with the greatest difference was taken. The greater the anterior laxity, the more relaxed the knee joint. Data collectors and patients were not aware of the surgical approach. Only the surgeon was aware of the grouping. The study procedure is indicated in Fig. The signal/noise quotient (SNQ) of the proximal, middle and distal regions of interest (ROI) of the intra-articular graft was measured using 3.0-T MRI  at postoperative months 3, 6, and 12 to observe the graft maturity law. All patients were required to rest for 60 min before measurement. Using fat-suppressed proton density (FS-PD) sequence. Thickness was 3 mm. Repetition time (TR) was 3140 ms. Echo time (TE) was 36 ms. Flip angle was 150°. Scan time was 118 s. Matrix was 320 × 256. Field of view (Fov) was 160 mm. From a sagittal image (Pd-tse-fs-sag) to select the interface that clearly showed graft and quadriceps tendon ROI. The graft ROI was located at the proximal , distal , and the middle (between them) regions. The ROI for the quadriceps tendon was located 2 cm above the patellar attachment point [Graft SNQ =  / Background mean signal value.The MRI scan was perform at a single magnet for all patients, and the whole process was completed by a fixed, blinded radiologist. The location of the proximal, distal, and middle ROI have been presented in Fig. All operations were performed by Prof. W Li who is an experienced orthopedic trauma surgeon specializing in sports medicine. The patient was placed in the supine position after lumbar anesthesia, with the lower leg hanging over the edge of the bed. A subpatellar anteromedial and anterolateral approach was used to perform knee arthroscopy, and a planing knife was used to clean the subpatellar fat pad and synovial membrane. The injury was explored, the ACL tibial and femoral stumps were cleaned, and the posterior edge of the lateral wall of the intercondylar socket of the femur was exposed. Using an anteromedial approach, the entry point was prepped using a microfracture cone and marked slightly posteriorly above the midpoint of the lateral wall of the resident’s crest.The all-inside reconstruction technique was performed via an anteromedial approach using an inverted drill  to drill a 2-cm hole in the femoral bone tract in a retrograde fashion Fig. . StandarThe autologous tendons were collected for graft preparation before drilling the tunnel. The size of the graft was used to determine the size of the tunnel. Both techniques were performed by making an oblique incision on the skin at approximately 3 cm medial to the tibial tuberosity to expose the semitendinosus and gracilis tendon, and the tendon was retrieved using a tendon retriever. All-inside reconstruction involved only a single harvest of the semitendinosus tendon, and the extracted tendon was double-folded to create a four-stranded single-bundle structure . IBM SPSS 21.0 statistical software was applied to outcomes data analysis. The measurement data were expressed as mean ± standard deviation (M ± SD). For comparison between two independent groups, the data met the normal distribution criteria using independent samples t-test and Pearson correlation analysis. The data did not meet the normal distribution using Mann–Whitney U-test and Spearman correlation analysis. A n = 27) or the standard group (n = 27). During the follow-up, one patient in the standard group and two in the all-inside group were lost follow-up as they lived abroad. At the end of the follow-up period, the all-inside and standard groups included 25 and 26 patients, respectively. There were no statistical differences in the demographic data of the two groups (p > 0.05). The graft length was significantly shorter in the all-inside group than in the standard group (p < 0.001). There was no statistically significant difference in the graft diameters of the two groups  preoperatively and at postoperative months 3, 6, and 12. Knee stability was restored in both groups at postoperative month 12, but arthrometery showed that the difference in the tibial anterior movement of the ACLR and normal contralateral knees was significantly greater in the all-inside group than in the standard group . The graft SNQ increased in both groups in postoperative month 6 than at postoperative month 3. Nonetheless, the increase in the all-inside group was more obvious and reached a peak; the difference between the two groups was significant (p < 0.05; Fig. There was no statistical difference in the graft SNQ between the two groups at postoperative months 3 and 12 p > 0.0. The grap > 0.05; Table Correlation analysis was conducted for the functional scores and SNQ of the proximal, distal, and middle ROIs of the all-inside and standard groups at postoperative months 3, 6, and 12. There was no correlation between the functional scores and graft maturity for both groups (Intra-articular biohealing of grafts after ACLR includes a cell necrosis phase within 3 months postoperatively, cell regeneration revascularization and graft remodeling phase after three months –18. This. [The results also revealed that the standard ACLR graft SNQ gradually increased during 1-year follow-up. Nonetheless, Van Dyck et al.  conclude. , 24, 25.According to our study, all-inside and standard ACLR exhibited the worst maturity in the middle graft region, followed by the graft near the femoral tunnel; the tibial side showed the best maturity. Putnis et al.  determinNotably, knee laxity after all-inside ACLR is not better than that after standard ACLR, which may be related to returning to exercise prematurely after all-inside ACLR. Darren et al.  showed tThis study had several limitations. First, the small sample size resulted in low statistical power; therefore, a study with larger sample size is required to draw more reliable conclusions. Second, data on clinical outcomes and SNQ follow-up time were limited, making the exact change tendency at 2–5 years unclear. Third, the SNQ based on MRI mostly reflected the water content in the graft, which is an indirect evidence of graft maturity. Finally, there was no standard method of measuring the SNQ of ACLR graft; therefore, we chose the quadriceps tendon for tissue normalization; however, this does not rule out deviations in measurement technology.All-inside and standard single-bundle ACLR show good functional outcomes; however, knee laxity was relatively higher after all-inside ACLR. In addition, both techniques exhibit the worst maturity in the middle graft region and the best in the distal region. All-inside ACLR graft maturity is inferior to standard ACLR at the early postoperative stages as it requires a more conservative rehabilitation strategy. There is no correlation between knee function and graft maturity.


Pancreatic transplantation is a definitive treatment for selected patients with insulin-dependent diabetes. It is a technically challenging surgery, and vascular complications are the most common cause of pancreatic graft failure. Although rare, pancreas transplants present higher rates of pseudoaneurysms at the vascular anastomosis than other visceral transplants. We present a case of a simultaneous pancreas–kidney transplant complicated with graft failure and common iliac artery pseudoaneurysm that was successfully treated through endovascular techniques.A 34-year-old White woman presented with abdominal pain and a history of type 1 diabetes mellitus, end-stage renal disease, and two previous pancreas transplantation failures. The first was a simultaneous pancreas–kidney transplantation performed 7 months prior that was complicated by pancreas graft thrombosis within 1 month and required graft resection. Five months later, she underwent a second pancreas transplantation with another pancreatic graft thrombosis requiring graft resection. Abdominal angiotomography revealed a pseudoaneurysm in the right common iliac artery at the point of the previous graft anastomosis. The patient was successfully treated endovascularly with a covered stent in the common iliac artery.Stent graft implantation for the treatment of common iliac artery pseudoaneurysm as a complication of simultaneous pancreas–kidney transplantation is a safe and feasible procedure. Simultaneous pancreas–kidney transplantation is a definitive treatment for selected patients with insulin-dependent diabetes and end-stage kidney disease that aims to reduce the long-term complications of both conditions. Pancreatic transplantation is a technically challenging surgery, and vascular complications, such as graft thrombosis, are the most common cause of pancreatic graft failure . AlthougWe present a case of simultaneous pancreas–kidney transplant complicated with graft failure and giant common iliac artery pseudoaneurysm that was successfully treated through endovascular techniques. Written informed consent was obtained from the patient for publication of this case report and any accompanying images. A copy of the written consent form is available for review by the Editor-in-Chief of this journal.A 34-year-old White woman with type 1 diabetes mellitus diagnosed since youth and end-stage renal disease requiring dialysis therapy over the last 4 years was referred to our hospital with abdominal pain in the right iliac fossa. She had a history of two previous pancreas transplantation failures. The first was a simultaneous pancreas–kidney transplantation performed 7 months prior that was complicated by pancreas graft thrombosis within 1 month and required graft resection. Five months after the first transplant, she underwent a second pancreas transplantation in the right iliac fossa using the same right common iliac artery anastomosis. On the second day postprocedure, her case was complicated by another pancreatic graft thrombosis requiring graft resection. Her renal function was still preserved, but insulin was reintroduced to control glucose levels.At our emergency department, she complained of abdominal pain in the right iliac fossa without signs of peritoneal irritation, bleeding, or hemodynamic instability. Abdominal angiotomography was performed and revealed a pseudoaneurysm in the right common iliac artery at the point of the previous graft anastomosis measuring 5.15 × 4.21 × 4.10 cm with a 1.07 × 0.30 cm neck (Fig. Endovascular treatment was proposed, and the patient was transferred to our interventional radiology unit. Under local anesthesia and mild sedation, we achieved percutaneous access to the right common femoral artery with a 7-French sheath. On initial arteriography, a right common iliac artery pseudoaneurysm was identified Fig. . To coveThe patient was discharged after 2 days with normal renal function and a lack of symptoms. She was followed up for 5 months by interventional radiology and nephrology teams without complaints.Vascular complications are one of the most common and feared complications of visceral transplants, requiring challenging clinical and surgical management –4. PseudThe clinical presentation of pseudoaneurysms is highly variable, ranging from asymptomatic cases detected on follow-up imaging to the appearance of symptoms such as abdominal pain, endocrine insufficiency, hematuria, palpable and pulsatile masses, hemodynamic instability, and sepsis. They may present as late bleeding, first with mild digestive hemorrhages, called sentinel bleeding, and evolve to an arteriovenous or arterioenteric fistula and eventually to a fatal hemorrhage . Graft cIn suspicion of pseudoaneurysm or early complications, ultrasound is usually the diagnostic method of choice, as it is readily available and cost effective . With DoThe decision between endovascular treatment and open surgical repair must be individualized, depending on the location of the lesion, the patient’s clinical status, the presence of abscesses, and the experience of the multidisciplinary team. In some situations, graft resection with deviation of blood flow is necessary to maintain perfusion of the lower limb. In the present case, endovascular treatment was chosen because the patient was hemodynamically stable and had multiple previous open surgeries in the right iliac fossa, which could increase the risk of vascular and intestinal injuries. In addition, whenever feasible, an endovascular approach is preferred as a less invasive alternative with a clinical success rate of 71.4% in patients with hemorrhage after pancreatic transplant and should be considered for first-line therapy in this population .The technique used in the article involved a balloon-expandable stent, although there was no difference between the types of coated expandable or self-expanding stents in the final treatment result. The difference is in the precision in stent release when the balloon is expandable. When there are arteriovenous fistulas, selective microcatheterization and embolization with particles or coils become the options of choice . The usePseudoaneurysms, although rare, can be sources of bleeding after pancreas transplantation. The diagnosis requires a high degree of suspicion, and initial imaging is necessary for confirmation and planning of the procedure. Treatment can be endovascular or surgical, depending on the clinical presentation.


The density and density profile, compressive strength, modulus of elasticity, acoustic properties and thermal conductivity were determined in the produced boards. These studies show that when food components are used as binding agents in the manufacture of lightweight wood fiberboards, the properties obtained can be comparable with those of commercial boards manufactured using synthetic agents.In this study, the possibility of using adhesives of natural origin for the manufacture of wood fiber-based lightweight panels was investigated. The boards, of a density ranging from 150 to 250 kg/m Properties of wood-based panels to a significant extent depend on the bonding agent used in their manufacture. In the production of building materials, synthetic resins, primarily formaldehyde resins, i.e., phenol–formaldehyde (PF), phenol–resorcinol–formaldehyde (PRF), melamine–urea–formaldehyde (MUF), urea–formaldehyde (UF) and isocyanine resins, have been applied for years. The advantages of formaldehyde resins are connected with their relatively low price and the good physicomechanical properties of materials manufactured using these resins ,7,8,9,103. A reduction in density results in an almost proportional decrease in the amount of used resin [3, swelling in thickness may be crucial in this case. However, this parameter is less significant in the case of lightweight panels dedicated to thermal or acoustic insulation. Thus, the aim of this study was to evaluate the applicability of compounds or products containing starch in the manufacture of lightweight panels from wood fibers.Construction boards are not only medium- or high-density boards, i.e., structural panels, but also lightweight panels, i.e., those of low and very low density dedicated to thermal and acoustic insulation ,17,18. Ted resin . Howevered resin ,21,22,23ed resin ,25. Natued resin ,39,40,41ed resin ,43,44. S3; pH, 8.09.Under laboratory conditions, panels were manufactured from pine fibers typically used in the manufacture of dry-formed fiberboards. Panels were bonded using a commercial urea–formaldehyde resin (UF), rye flour and potato starch solutions and two starch types: E 1404 (oxidized starch) and gelatinized starch. Basic properties of UF resin  used in the tests were as follows: viscosity, 650 mPa·s; solid content, 69%; gel time at 100 °C, 69 s; density, 1.282 g/cmThe rye flour solution was prepared from 100 g water and 25 g rye flour of approx. 3.5% moisture content. The solution was mixed until a homogeneous blend was obtained; it was left for 20 min and after being mixed again it was used as an adhesive. In turn, the potato starch solution was prepared as follows: 4 g of starch was added to 100 g of water. After being thoroughly mixed to produce a suspension, the solution was heated while mixing continuously to ensure its thickening. Fibers of 7.2% moisture content were coated with bonding agent solutions at 12% adhesive dry mass for UF resin, 12% and 6% adhesive dry mass for the rye flour solution and 2.4% and 3.6% adhesive dry mass for the potato starch solution.3. Pressing time was dependent on the type of the applied bonding agent. More detailed data are given in In turn, starch was applied onto fibers with approx. 12% and 28% moisture content. Three days before the bonding process, fibers of a 7.2% moisture content were sprayed with water in an amount ensuring an increase in the fiber moisture content to 15% and 30%. After thorough mixing, 0.5 kg batches were placed in a double-sided plastic bag and sealed tightly. At least twice daily, the mass in the bags was mixed thoroughly. The fiber moisture content prior to bonding amounted to approx. 12% (± 0.2%) and 28% (± 0.34%). Starch was applied in the amount ensuring a bonding rate at 12% or 20%. Starch was applied using a pneumatic lacquer spray gun. The mat was formed from such prepared material. After formation, the mat was pressed at 0.8 MPa at the temperature of heat press platens of 180 °C . Panels were manufactured at a thickness of 25 mm and design density ranging from 150 to 250 kg/mDensity (ρ) based on PN-EN 323:1999 .v)—the value of compression stress at 10% true strain according to PN-EN 826 [Compressive strength (fN-EN 826 .Modulus of elasticity (E) according to PN-EN 310 .2 mV), U is the voltage (mV), d is the partition thickness (m), Tg is the temperature of specimen surface in the heating chamber and Tc is the temperature of specimen surface in the cooling chamber.Thermal conductivity was determined using a measuring system presented in a study by Mirski . In thisAcoustic insulation capacity—sound absorption coefficient according to PN-EN ISO 10534-2:2003  within tDensity profile—determined using a laboratory DAX profile measurement gauge by GreCon .One of the panels was assigned only for the determination of thermal conductivity, with the panel being dried to constant mass at a temperature of 103 °C, while the other two were used in the other tests and conditioned at 20 ± 2 °C and relative humidity of 60 ± 5%. After the conditioning period (14 days), panels with approx. 7.4–8.2% moisture content the following parameters were assessed in terms of the following:Density, compressive strength and the modulus of elasticity were determined on samples of 100 mm × 100 mm. Tests were performed on 10 specimens. Recorded results were analyzed statistically. Statistical calculations were conducted using the Statistica 13.0 software .2. Next, the specimens were subjected to the joinery test (nonstandardized chisel test) in order to evaluate bond quality. The best results (the largest number of fragments of pulled-out wood) were recorded for starch solutions, while potato starch and rye flour solutions proved to be slightly weaker bonding agents.Prior to the tests, pastes were prepared from the selected agents and used to bond two beech wood specimens. Specimens were bonded at 120 °C under the pressure of 1.4 N/mmProperties of panels manufactured from fibers with the initial moisture content are presented in 3, respectively. Thus, in their case, an increase in density by slightly below 40% causes an increase in strength by over 500%. In contrast, all of the panels manufactured using potato starch and rye flour exhibit much lower compressive strength. The greatest compressive strength was recorded for panels produced with a 12% share of rye flour. This is approx. 20% lower than the value for panels manufactured from wood fibers resinated with UF adhesive. A decrease in the bonding rate with the rye flour solution from 12% to 6% leads to a reduction in compressive strength by almost a half. The compressive strength of panels manufactured from wood fibers bonded with the potato starch solution is generally much lower than that of panels manufactured with a share of rye flour.The view of the mold before and after pressing is shown in Strength properties to a considerable degree depend on panel manufacturing conditions. Considerably better results are obtained when applying longer pressing times and lower heat press platen temperatures rather than vice versa. Unfortunately, at a very long pressing time of panels at 180 °C, a very strong darkening of the panel surface is observed, resulting from fiber degradation. The use of potato starch solutions is problematic, since their dry mass is very low; thus, in order to obtain a minimal required bonding rate (the ratio of the dry mass of the active component to the dry mass of fibers), a very large amount of the solution needs to be applied. Although panel density was reduced to accelerate water vapor evacuation, this is a relatively long process. However, the obtained compressive strength levels for these variants do not have to be considered as very negative, as in terms of the bonding rate the obtained results were comparable or even more advantageous than those recorded for the variant with the application of UF resin. This is obviously assuming the linear dependence of strength on resination rate, which may not be considered as certain. The obtained values of thermal conductivity range from 0.037 to 0.039 W/(m·K), and it is rather difficult to assess the effect of applied modifications. Since these values are consistent with the reports of manufacturers of insulation boards produced from pulp, it needs to be assumed that this parameter is more strongly dependent on fiber quality and panel density, rather than the applied fiber bonding agent.Since satisfactory results were obtained for physicomechanical properties when applying relatively long pressing times, it was decided to reduce the amount of water introduced to the system. For this purpose, panels were again manufactured from pine fibers under laboratory conditions; however, this time it was decided not to introduce the adhesive in the liquid form. Gelatinized starch and oxidized starch (E1404) were used as bonding agents. When the adhesive is applied in the powder form, it is relatively easy to modify the amount used, and the level of mat moisture content before pressing is much more advantageous. Other tests, not described in this paper, were also performed, in which starch was applied on dry fibers and moisture content was increased to reach 12–15%. However, it was observed that when employing such a method of starch application, some of the grains are left in the mixer. Data in 3 was recorded in the case of panels manufactured from pulp with a greater moisture content. In turn, for pulp with an approx. 12% moisture content in the subsurface zone, the density of subsurface layers slightly exceeds 220 kg/m3 and is only slightly higher than in certain areas located deeper. This dependence may explain such advantageous properties of E0 panels at compression perpendicular to grain despite their low initial moisture content. Thanks to the more uniformly distributed density at the cross-section, they effectively transfer external loads.A change in the method of fiber bonding agent application has no effect on the level of thermal conductivity. In this case, the value of λ was within the range of 0.037–0.038 W/m·K, and no influence of the applied board manufacture parameters was observed. In turn, the level of moisture content in the pressed mat has a significant effect on the density profile of manufactured panels. As it results from values presented in Panel E0 was selected for tests assessing the sound absorption coefficient. Although such panels did not exhibit the highest strength parameters, their manufacturing process ultimately seems the easiest to perform, since the fibers used to form the mat contain the lowest amounts of water. In the application of food components as bonding agents in the manufacturing process of lightweight insulation panels from wood fibers, the obtained properties may be comparable to those of commercial panels produced using synthetic agents.The compressive strength of manufactured panels to a considerable extent depends on the ratio of the dry mass of the active agent to the dry mass of fibers.Modification of the bonding rate is dependent on the feasibility of preparation of a solution with a specific concentration from a given flour or starch.Since low concentration solutions are required, pressing time needs to be extended due to the considerable increase in moisture content of the formed pulp.It seems that although this problem has not been fully solved, the use of starch in the powder form on fibers with a high moisture content has considerable potential applicability.These tests are preliminary studies aiming at the identification of easy-to-use, readily available compounds that exhibit bonding properties in contact with water. Although a wide range of such products has not yet been launched on the market, numerous similar studies indicate considerable potential applicability, particularly for these agents, which are food components. This is because they are easily processed and are fully biodegradable, which in combination with wood makes it possible to manufacture products completely neutral in terms of their environmental impact. It seems that the trend towards the application of such agents to produce insulation materials is promising, since such panels are not required to exhibit high mechanical strength; as a result, meeting the relevant requirements is easier. On the other hand, lightweight panels of very low density are more difficult to produce because they fail to mask errors in the bonding and forming technology in terms of uniformity of adhesive application and variable density. Nevertheless, such production processes are deserving of further investigation, which may be aided by the following results from this study:


However, determining how to control semi-active devices is still a challenge. The dynamic mean ankle moment arm (DMAMA) provides a suitable biomechanical metric, as its simplicity matches that of a semi-active device. However, it is unknown how stiffness and locomotion modes affect DMAMA, which is necessary to create closed-loop controllers for semi-active devices. In this work, we develop a method to use only a prosthesis-embedded load sensor to measure DMAMA and classify locomotion modes, with the goal of achieving mode-dependent, closed-loop control of DMAMA using a variable-stiffness prosthesis. We study how stiffness and ground incline affect the DMAMA, and we establish the feasibility of classifying locomotion modes based exclusively on the load sensor. (2) Methods: Human subjects walked on level ground, ramps, and stairs while wearing a variable-stiffness prosthesis in low-, medium-, and high-stiffness settings. We computed DMAMA from sagittal load sensor data and prosthesis geometric measurements. We used linear mixed-effects models to determine subject-independent and subject-dependent sensitivity of DMAMA to incline and stiffness. We also used a machine learning model to classify locomotion modes using only the load sensor. (3) Results: We found a positive linear sensitivity of DMAMA to stiffness on ramps and level ground. Additionally, we found a positive linear sensitivity of DMAMA to ground slope in the low- and medium-stiffness conditions and a negative interaction effect between slope and stiffness. Considerable variability suggests that applications of DMAMA as a control input should look at the running average over several strides. To examine the efficacy of real-time DMAMA-based control systems, we used a machine learning model to classify locomotion modes using only the load sensor. The classifier achieved over 95% accuracy. (4) Conclusions: Based on these findings, DMAMA has potential for use as a closed-loop control input to adapt semi-active prostheses to different locomotion modes. Current passive prosthetic feet are offered in a variety of stiffness categories that are determined based on the user’s body weight. This prescription method results in prominent stiffness variability within single categories, both across manufacturers and even within single manufacturers ,2,3,4. STo optimize the walking performance of persons with amputation, prosthetic foot properties need to adapt. Active control through a powered prosthetic foot may provide this capability, but the added weight, height, power requirements, and cost of such devices present barriers to adoption . An alteHowever, determining how to control stiffness adaptation is a challenge. Changes could be controlled using open-loop mapping—e.g., predetermined stiffness for every speed and locomotion mode—or using closed-loop adaptation to achieve some biomechanical outcome. For the closed-loop case, an appropriate target measure is needed that can be measured in real time using onboard sensors. Various biomechanical measures could fill this void, such as rollover shape ,20, anklThe dynamic mean ankle moment arm (DMAMA) measure  was deveBiomimetic closed-loop control of DMAMA using a semi-active prosthesis requires both the ability to vary DMAMA  across locomotion modes  and the ability to estimate the locomotion mode in real time to permit such adaptation . PreviouThis study explores the potential for DMAMA to be used as a real-time control input for semi-active prosthetic devices. We aim to understand how DMAMA can be influenced by changes in prosthesis forefoot stiffness and by locomotion mode, specifically level ground, ramps, and stairs. We further investigate whether it is feasible to use a universal control law to exploit these relationships for control, or whether subject-dependent control laws are necessary, by evaluating subject-dependent and subject-independent trends across stiffness and locomotion modes. Additionally, we explore what it would take to employ such a measure in a real time, “outside of the lab” environment. We demonstrate a method to monitor both DMAMA and locomotion mode outside of the lab using only a prosthesis-embedded load cell sensor.In this work, we first investigate whether stiffness or locomotion mode has a significant effect on DMAMA. We then evaluate how well overall trends can explain intra-subject DMAMA fluctuations. Our first hypothesis is that there will be a positive linear sensitivity of DMAMA to forefoot stiffness across all locomotion modes . Our second hypothesis is that there will be a positive linear sensitivity of DMAMA to the ground incline across all stiffnesses . Finally, we use linear discriminant analysis (LDA) to test the classification accuracy of locomotion modes using only prosthesis geometric measurements and the inputs from a six-axis pylon-embedded load cell. Throughout this work, we discuss how DMAMA is affected by changes in stiffness and locomotion mode and evaluate the potential of this measure to be used in real-time control of semi-active prosthetic devices.Four adult participants with transtibial amputation (a subset of a larger study) participated in this focused test of the prosthesis-embedded technology. These participants  wore a novel prosthetic foot and a suite of prosthesis-embedded and wearable sensors that recorded the signals necessary for field-based analysis of prosthetic limb mechanics. Each participant used a prosthesis for daily ambulation and walked without the help of an ambulatory aid. Prior to beginning the study, each participant signed a written consent form approved by the University of Wisconsin–Madison Health Sciences institutional review board.For this study, prosthetic forefoot stiffness variation was supplied by the variable-stiffness foot (VSF), a custom-built prosthesis previously published by Glanzer and Adamczyk . The parWe instructed participants to walk through a circuit in a campus building that required five locomotion modes: level ground, ramp ascent (+5° incline) and descent (−5° incline), and stair ascent and descent . All participants walked this circuit once using each of three stiffness conditions. The participants walked the course at their preferred walking speed, as the goal of the study was to see how participants would ambulate with various prosthesis stiffnesses in a real-world environment. The circuit resulted in 50 ± 14 level ground, 11 ± 3 up ramp, 11 ± 3 down ramp, 11 ± 1 up stairs, and 11 ± 1 down stairs strides per trial that were used for analysis. Transition steps between locomotion modes were excluded from this analysis. This experimental setup can be viewed in The calculation of DMAMA requires the sagittal ground reaction force (GRF) and ankle moment. In a gait laboratory, these measures are typically found from force plate and motion capture data . HoweverWe fit the participants with a full-body set of inertial measurement units (IMUs)  that was used to discern changes in locomotion mode during the trials. We calibrated a kinematic skeleton model reconstructed from IMU data  using each participant’s anthropometric measurements along with standing and walking calibration trials.We used the load cell measurements as an approximation for the ground reaction force and moments in the sagittal plane. Prior to beginning the study, we measured the orientation and location of the load cell relative to the user’s prosthetic pylon and body segments using motion capture  in the laboratory. Functional movement calibration trials were used to calibrate the model in Visual3D, allowing the determination of the biological joint centers. We determined functional joint centers for the knee and hip from the motion capture trials using the Gillette algorithm . The ankle joint center for the prosthetic side was chosen as a fixed physical location on the prosthesis comparable to the biological limb, as there is no fixed prosthetic ankle rotational axis. It should be noted that this method of defining the prosthetic ankle joint axis reduces across subject variability but may result in an offset error relative to a biological DMAMA measure. The joint centers’ positions relative to the center of the load cell were determined using the average across the function movement trial. This step was taken to orient the load cell forces in the pylon/shank reference frame prior to all calculations. We recorded the locations of the knee and ankle joints relative to the load cell for use in joint moment calculations.We measured the force and moment about the load cell origin directly from the load cell embedded in the prosthetic pylon. To correct the load cell’s orientation and placement relative to the participant’s shank, we found the rotation matrix , and the amount of variance explained by the model (R2).We evaluated the effects of both stiffness and ground incline on DMAMA. We first evaluated subject-independent relationships using a linear mixed-effects model where either the stiffness or the ground incline was the fixed effect and the different participants were random effects. We separated the data by locomotion mode and ran a linear regression on the mean DMAMA values across participants and stiffness settings, allowing subject-dependent intercepts to account for the random effect. We determined the subject-independent sensitivity of DMAMA to each variable (slope of the best-fit line), the statistical significance of sensitivity , and the amount of variance explained by the model (R2). We also evaluated the correlation strength of the subject-independent best-fit line on each participant’s stride-by-stride data to determine how well the overall (subject-independent) best-fit line explained the variance in individual participants’ data. To represent this, we calculated the R2 value usingNext, we evaluated subject-dependent relationships between stiffness and DMAMA and between ground incline and DMAMA. We ran a linear regression on each individual participant’s data and determined the subject-dependent sensitivity of DMAMA to each variable (slope of the best-fit line), the statistical significance of sensitivity  and MATLAB 2019a .p < 0.05) in the DR, LG, and UR conditions (sensitivity range: 2.299–3.749 percent foot length per stiffness increment) but not in the DS (p = 0.124) or US (p = 0.522) condition , level ground (LG), up ramp (UR), down stairs (DS), and up stairs (US) . There wondition .p < 0.05) in all four participants in the DR condition (sensitivity range: 3.132–4.422 percent foot length per stiffness increment), all four participants in the LG condition (1.963–3.824 percent foot length per stiffness increment), three of four participants in the UR condition (2.070–3.032 percent foot length per stiffness increment), three of four participants in the DS condition (2.906–4.263 percent foot length per stiffness increment), and one of four participants in the US condition (sensitivity: 5.405 percent foot length per stiffness increment) (p < 0.05) of DMAMA to stiffness for one participant in the DS condition (−2.479 percent foot length per stiffness increment).We then evaluated the subject-dependent sensitivity of DMAMA to stiffness in each of the locomotion modes . There wcrement) . There wp < 0.05) of DMAMA to the ground incline in the low- (sensitivity: 0.540 percent foot length per degree incline) and medium- (0.450 percent foot length per degree incline) stiffness conditions but not in the high-stiffness condition  in all four participants in the low-stiffness condition (sensitivity range: 0.214–0.965 percent foot length per degree incline), all four participants in the medium-stiffness condition (0.181–0.699 percent foot length per degree incline), and two of four participants in the high-stiffness condition (0.169–0.535 percent foot length per degree incline)  .Considering only the ground incline conditions, the sensitivity of DMAMA to stiffness was the greatest in the DR condition (sensitivity: 3.749 percent foot length per stiffness increment), moderate in the LG condition (2.993 percent foot length per stiffness increment), and the least in the UR condition (2.299 percent foot length per stiffness increment). Similarly, the sensitivity of DMAMA to the ground incline was the greatest in the low-stiffness condition (0.540 percent foot length per degree incline), moderate in the medium-stiffness condition (0.450 percent foot length per degree incline), and the least in the high-stiffness condition (0.250 percent foot length per degree incline). We formalized this interaction by fitting a multiple linear regression with incline, stiffness, and incline-times-stiffness interaction terms to the DMAMA data for the three ground incline conditions and three stiffnesses . This muWe then evaluated the accuracy of the subject-dependent LDA classifier. After the initial input of 54 features extracted from a 300 ms window, forward feature selection led to the elimination of 32 features, leading to a model that used 22 features. We then optimized the window size, leading to the selection of a 300 ms window. This final subject- and stiffness-dependent LDA model accurately classified 90.9% of down ramp steps, 98.2% of level ground steps, 87.6% of up ramp steps, 96.8% of down stairs steps, and 96.6% of up stairs steps . This leBecause DMAMA is a summary measure computed from traditional joint mechanics data, these joint mechanics are also available within the analysis. We provide these data as an illustrative supplement in Biomimetic closed-loop control of semi-active prostheses demands a biomechanical control target that can be measured and manipulated in environments “outside of the lab.” DMAMA is a suitable metric to drive semi-active control due to simplicity and can be measured in real-world scenarios. In this work, we established a method to calculate DMAMA from only a pylon-embedded load cell sensor and prosthesis geometric measurements, eliminating the need for force plates and motion capture systems. Prior work has established that wearable technology  could evaluate trends in DMAMA with similar sensitivity to laboratory-based measurements. The present work extends this to approximate the ideal case of a direct force and moment measurement fully embedded in the device that uses the data. Both these cases relied on approximations of the location of the ankle joint relative to the sensor, introducing potential offset errors in DMAMA. In contrast, future implementation of a load cell directly embedded in an ankle-foot prosthesis would result in specified, rather than measured, geometric relationships between the load cell and the ankle and would therefore eliminate the approximation steps required in this work and reduce variability and uncertainty in the measurements.The results suggest a significant positive sensitivity of DMAMA to forefoot stiffness in down ramp, level ground, and up ramp conditions but no clear effect on stairs due to high variability. Additionally, the linear fit showed a strong correlation between stiffness and DMAMA in the down ramp, level ground, and up ramp conditions. Within the ramp conditions, the significant sensitivity of DMAMA to stiffness varied across the ground incline, with DMAMA being less sensitive to stiffness as the incline increased. This leads us to partially accept our first hypothesis (increase in DMAMA with stiffness): DMAMA does increase with forefoot stiffness in walking-like conditions but perhaps not on stairs. The high variability of DMAMA on stairs may reflect qualitatively different stair ambulation strategies across individuals, which have made it difficult to study this behavior in other real-world studies as well .We next looked to understand how well subject-independent sensitivity trends could explain individual participants’ DMAMA values. In down ramp, level ground, and up ramp conditions, the subject-independent sensitivity provided a moderate-to-strong fit for individual participant data, while it provided a weak fit in the down stairs and up stairs conditions. No locomotion mode showed that a generalized sensitivity provided a strong fit for all participants. Even in the ramp and level ground conditions, where all individual participants had a significant positive sensitivity, participants had differing levels of sensitivity. Thus, we conclude that increasing forefoot stiffness can increase DMAMA in all participants, but the quantitative sensitivity is subject dependent.The results also suggest a significant positive sensitivity of DMAMA to the ground incline in the low-and medium-stiffness settings but not in the high-stiffness setting. Additionally, the linear fit showed a strong correlation between ground incline and DMAMA in the low-stiffness condition and moderate correlations in the medium- and high-stiffness conditions. As expected, because of similar interaction trends seen in the stiffness analysis, the sensitivity to the incline decreases as stiffness increases. This leads us to partially accept our second hypothesis: DMAMA does increase with the ground incline but with reducing sensitivity at higher stiffness settings.We again looked to understand how well subject-independent sensitivity of DMAMA to the ground incline could explain individual participants’ DMAMA values. All participants exhibited a positive sensitivity of DMAMA to the ground incline. The goodness of fit of the subject-independent sensitivity to individual participants was highly variable across participants and ground inclines. Across all stiffness conditions, the subject-independent sensitivity provided some explanations of trends seen in participants 1 and 4, though the goodness of fit was highly variable. However, the subject-independent sensitivity provided a poor explanation of trends seen in participants 2 and 3. Therefore, as above we conclude that increasing the ground incline can increase DMAMA in all participants, but the quantitative sensitivity is subject dependent.Some of the effects above could be explained by the interaction of our two independent variables, stiffness and ground incline. Our multiple regression revealed a negative best-fit coefficient for the interaction between stiffness and ground incline, though the coefficient did not reach the threshold for statistical significance. Nevertheless, an interaction with a negative sign could explain the trends observed above: that sensitivity of DMAMA to either independent variable is reduced at higher values of the other. This interaction can be thought of as a kind of saturation, as opposed to compounding (interaction with a positive sign). This effect makes sense mechanistically: both a stiffer forefoot and an uphill slope move the center of pressure toward the toe and lead to earlier heel lift-off and toe-only support, such that a high value of either leaves little room for additional forward movement due to the other. It is possible that avoiding such situations is itself a valuable goal in prosthetic ankle control; if so, a controller that reduces stiffness on uphill slopes and increases stiffness on downhill slopes would serve to keep DMAMA lower in support of this goal.The positive sensitivity of DMAMA to stiffness could be explained if participants maintain similar leg kinematics regardless of forefoot stiffness. If leg kinematics remain constant, stiffening of the forefoot will cause a more rapid forward shift of the COP during foot roll-over, therefore increasing the DMAMA value. The observed variation in DMAMA with stiffness suggests that persons with amputation may not change their kinematics to maintain the same DMAMA value when stiffness changes. This demonstrates that DMAMA can be varied in persons with amputation by manipulating stiffness. Direct evaluation of how well kinematics are retained is a component of the broader study from which these pilot data were collected and will be explored in future work.The moderate stride-by-stride fluctuations in DMAMA  and the behavior in up stairs vs. down stairs conditions were highly subject dependent. Thus, some level of subject-dependent tunability in a DMAMA-based controller would likely be beneficial—for example, adjusting the range and sensitivity of the target DMAMA setting to variations in slope.To determine whether real-time detection of the locomotion mode was possible with limited instrumentation, we evaluated the ability of an LDA classification algorithm to detect the locomotion mode based only on measures from a pylon-embedded load cell and shank segment geometric measurements. Our results showed that the LDA model was able to accurately classify locomotion mode for 95.72% of steps. The LDA model feature selection process revealed which sensors and signal features were the most useful in determining the locomotion mode. This selection process eliminated all features that were derived from the sagittal force, but kept features derived from the remaining eight sensor inputs. The selection process kept all feature types , indicating that all feature types provided some value in determining the locomotion mode. The model’s selection of two features that were derived from sensor input that required some calculation before passing into the model (The LDA model disproportionately misclassified up ramp (UR) trials in comparison to other classes, where up ramp misclassifications were primarily classified as level ground (LG) steps. We expect that this error is primarily attributed to the class imbalance in the data set, where we collected 50 ± 14 level ground and 11 ± 3 up ramp steps per participant. This class imbalance also explains why a comparable percentage of level ground steps are not misclassified as up ramp steps. Previous work has shown that the feature space between level ground and up ramp steps is similar, which may provide an explanation for the increased difficulty discerning these steps using our model ,34. OtheThis analysis was intended to investigate the feasibility of locomotion mode classification from a single sensor  using a simple machine learning model. However, we believe that more advanced machine learning algorithms may be able to improve subject-dependent model accuracy or provide subject-independent classifiers . Our resThe results of this study support both key hypotheses: that there would be (1) positive linear sensitivity of DMAMA to forefoot stiffness across all locomotion modes and (2) positive linear sensitivity of DMAMA to the ground incline across all stiffnesses. These results suggest that both stiffness and locomotion mode influence DMAMA, and therefore both variables need to be considered when targeting DMAMA changes. Stiffness can be considered in real time using the known stiffness value from the foot’s existing control system. Additionally, the locomotion mode can be determined using only a pylon-embedded load cell and prosthesis geometric measurements as inputs into a machine learning model. These results suggest that DMAMA could feasibly be used as a control parameter for a variable-stiffness foot and potentially for other semi-active prosthetic devices.The core practical application of these findings is to use variations in stiffness to enable some level of adjustment of a prosthesis to the user’s behavior and the terrain. The sensitivity of DMAMA to stiffness was modest in this experiment, suggesting that this effect might be limited in its application. However, current prostheses allow no adaptability at all; if emerging technology can offer any benefit, it will be an improvement on the current experience of prosthesis users. Furthermore, the VSF was originally conceived to also adjust parameters such as energy storage and return from the prosthesis and stability in qualitatively different behaviors like standing vs. walking ; if contOne additional use case for a controlled VSF could be a DMAMA-matching controller that seeks to match this metric on the prosthetic side to values measured in real time from the intact limb. Such measurements could be made using force-sensitive insoles  or even This study used a subset of data from a larger study to explore this concept of DMAMA as a possible real-time control variable. The study was limited by the small number of participants who used all the sensors required to compute DMAMA and track and label the different terrains. The small sample limited our ability to make conclusive claims about the biomechanical results of DMAMA. A study with a larger sample, perhaps also including tests that implement the controllers suggested by this analysis, could further explore subject-independent vs. subject-dependent control laws and classifiers and could establish an “existence proof” of the closed-loop control motivated through this study.Each participant wore the same VSF for this study, preventing our ability to match the length of the VSF prosthesis to their prescribed prosthesis or intact foot. A practical implementation of a VSF in clinical use would need to accommodate multiple sizes of prostheses for users of different heights, but this was not possible with only one prototype.Another potential limitation was the treatment of footwear in this study; participants wore their chosen athletic shoe on their biological foot and no shoe at all on the prosthesis. The use of preferred (rather than standardized) shoes could have affected the gait biomechanics across individuals in this study. Using a prosthesis with no shoe is uncommon in daily use but is common with certain high-performance prostheses  and running prostheses . In future uses of advanced prostheses such as the VSF, it could become more common as users seek to maximize the function of their prostheses without letting a shoe affect its properties .Lastly, this study allowed participants to walk at their preferred speeds across all conditions so that the tests would be ecologically valid relative to the eventual use of these results in uncontrolled real-world scenarios. Testing in these conditions provided us with information about how stiffness and ground incline changes would alter DMAMA in unconstrained environments; however, these results may be limited when evaluating how DMAMA is affected in speed-constrained studies or in conditions of varying speed. Additional studies of natural gait across locomotion modes or with combinations of speed and mode could build on prior findings of how DMAMA varies with speed on level ground .In conclusion, we found that DMAMA has significant positive sensitivity to both stiffness and ground incline and negative sensitivity to their interaction. Sensor data from a pylon-embedded load cell and limited subject-dependent prosthesis geometric measurements provided sufficient information to compute DMAMA without external instrumentation and also to classify locomotion modes through a machine learning model. DMAMA may be a suitable practical control parameter for a semi-active variable-stiffness foot, enabling real-time alteration of prosthetic forefoot properties for improved biomechanics in comparison to passive prostheses.


In this research, annealing (ANN) using plasma-activated water (PAW) was first employed to modify potato and pea starches. Compared with the conventional ANN using distilled water (DW), the granular morphology of two starches was not significantly affected by PAW-ANN. The results of X-ray diffraction (XRD) and Fourier transform infrared (FTIR) spectroscopy showed that PAW-ANN could reduce the long and short-range ordered structure of potato starch while improving the long and short-range ordered structure of pea starch. Differential scanning calorimetry (DSC) analysis indicated that PAW-ANN lowered the gelatinization enthalpy of potato starch and increased the gelatinization enthalpy of pea starch. The analysis of viscosity and dynamic rheological characteristics illustrated that PAW-ANN reduced the peak viscosity and improved the gel strength of starch pastes. PAW-ANN represents a novel modification method for modifying the structure, reducing the viscosity, improving the gel strength of starch, and is very promising for applying in starch-based hydrogels and food additives. Starch is a significant carbohydrate source for most organisms, which has been widely applied in the food industry, such as papermaking, adhesives and biodegradable plastics, sweeteners, pasta products, and some fermented products . Native Annealing (ANN) is usually to treat starch for a period of time at a high moisture content  and a low temperature below 60°C). As a green modification method, ANN has received extensive attention because it only involves water and heat . It can 0°C. As aPlasma-activated water (PAW) from plasma treatment of distilled water (DW) has the advantages of uniform function, green and environmental protection. PAW will generate an acidic environment and generate active substances, such as hydrogen peroxide, nitric acid, and peroxynitrite, which will cause changes in the pH, redox potential, and conductivity. Recently, PAW has been widely used in food and agriculture yields, such as promoting plant growth, preservation of seafood, and broad-spectrum sterilization . HoweverIn this research, PAW was first employed during ANN to modify potato and pea starches. After being modified with PAW-ANN, the changes of structure and properties of potato and pea starches were investigated. It is worth mentioning that this green modification method is only involved in plasma, water, and heat. This research will not only propose a new way of starch ANN treatment, but also further expand the application of PAW in starch modification.Native potato starch (13.2% moisture and 21.6% amylose) was obtained from Qinghai Weston Potato Industry Group Co., Ltd. . Native pea starch (11.3% moisture and 30.8% amylose) was obtained from Henan Enmiao Food Co., Ltd. . A hydrogen peroxide quantitative kit was obtained from Shanghai Sangon Biotech Co., Ltd. . Nitrite and nitrate determination kits were obtained from Beyotime Biotechnology Institute . Other reagents used in this experiment were of analytical grade.2O2, NO2−, and NO3− were determined by corresponding assay kits. Notably, PAW was suggested to be further used within 12 h.The PAW was obtained by treating 100 ml of distilled water through plasma processing equipment  for 120 s. The power was 750 W and the working gas was compressed air 0.18 MPa) ( MPa (8).After drying at 45°C, the moisture content of potato and pea starches  is about 5%. Then, the moisture content of them was adjusted to 60% using DW or PAW, respectively. The starch samples were put into the reactor with sufficient stirring and then heated in an air oven at 50°C for 12 h. Finally, modified starches were obtained through washing, drying, and sieving and marked as DW-Potato, DW-Pea, PAW-Potato, and PAW-Pea, respectively.Morphology of native and modified starch granules was observed using a scanning electron microscope . A small amount of dry-based samples were adhered to the double-sided conductive adhesive and coated with gold in the ion sputtering device  for 120 s . The imaA suspension of starch samples  was obtained by using a solvent of glycerol and water  and observed under polarized light microscopy (PLM) . The images were observed and taken at ×200 for all starch samples.Before analysis of X-ray diffraction (XRD), starch samples needed balancing water over a saturated NaCl solution at room temperature for 1 week . Starch −1, resolution 4 cm−1, scanning time 64 s. The spectra were analyzed by OMNIC8.2 . The absorbance ratio at 1,047/1,022 cm−1 (R1047/1022) can be obtained by deconvolution with a peak width of 38 cm−1 and enhancement factor of 19 and employed to characterize the short-range ordered structure of starch molecules (Fourier transform infrared (FTIR) spectra of starch samples were determined by using an FTIR spectrometer . Before the test, potassium bromide should be dried at 105°C for 6 h. A certain amount of samples and potassium bromide (mass ratio: 1:100) were taken in an agate mortar, mixed, and grinded for about 2 min. The ground mixed samples were placed in a tableing mold and compressed with a tableting machine, where the pressure was maintained within the range of 10 MPa for about 1 min and then taken out for testing. Test conditions: scanning wave number 4,000–400 cmThermal properties of native and modified starches were obtained by using a differential scanning calorimeter . The starch  and distilled water were added with a microliter syringe to a total weight of 12 mg into an aluminum pan. The samples were equilibrated at room temperature for 12 h. Subsequently, the sample pans were gradually scanned from 30 to 120°C at a heating rate of 10°C/min .Pasting characteristics of the starch samples were obtained by using a rapid viscosity analyzer . Starch samples were suspended in distilled water to make a total weight of 28.0 g  and then analyzed using RVA Standard Procedure 1 profile. The samples were equilibrated at 50°C for 1 min, heated to 95°C within 222 s, held at 95°C for 150 s, cooled back to 50°C for 228 s, and then equilibrated at 50°C for 2 min. During this process, the paddle speed was kept at 960 rpm for the first 10 s, and then kept at 160 rpm. The pasting curves and parameters were obtained by RVA system software.The dynamic rheological properties of starch pastes from the RVA experiment were analyzed by using a rheometer . Sample pastes were transferred to the rheometer plate  and equilibrated for 5 min at 25°C. The strain was set at 1% and the frequency was taken at 0.1–20 Hz. G′, G″, and tan δ (G″/G′) of samples were recorded.post-hoc Duncan's multiple range tests (p < 0.05) was conducted to determine the significant differences between mean values using the SPSS 26.0 Statistical Software Program .All experimental data were recorded as the means ± SDs. ANOVA followed by +, H2O2, NO, NO2−, NO3−, ONOO−, and so on, which led to the changes in the pH value, conductivity value, and ORP value of the aqueous solution  can indicate the relative content of the short-range ordered structure of starch. R1047/1022 of native and modified starches followed the order: Potato > DW-Potato > PAW-Potato; Pea < DW-Pea < PAW-Pea  curves of native and modified starches are shown in The pasting characteristics curves of native and modified starches are presented in The dynamic rheological properties of native and modified starch pastes are displayed in In summary, PAW-ANN as a completely novel starch modification method, had an obvious effect on the structure, thermal, pasting, and rheological properties of potato and pea starches. The impact of PAW-ANN on the granular morphology of two starches was not significantly different from the conventional DW-ANN. Compared with DW-ANN, PAW-ANN increased long and short-range ordered structure, and gelatinization enthalpy of pea starch while decreasing these indicators of potato starch. In addition, PAW-ANN resulted in the lowest peak viscosity and the highest gel strength of the paste for two starches. Although, the exact mechanism was not clear, an acidic component in PAW might be an important factor for starch modification according to previous studies and above experimental results. This research showed that PAW-ANN provided innovative insights and novel technologies for the production of new modified starches applied to starch-based hydrogels and food additives.The original contributions presented in the study are included in the article/supplementary material, further inquiries can be directed to the corresponding author/s.YY contributed to the conception, design, and funding of the study. BP, XJ, and YH organized the database. BP wrote the first draft of the manuscript. BN and MS contributed to writing-review and editing. All authors contributed to the article and approved the submitted version.We are grateful to the National Natural Science Foundation of China (32101945), the Program for Science and Technology Innovation Talents in Universities of Henan Province (20HASTIT037), and the Henan Youth Talent Support Project (2020HYTP046).The authors declare that the research was conducted in the absence of any commercial or financial relationships that could be construed as a potential conflict of interest.All claims expressed in this article are solely those of the authors and do not necessarily represent those of their affiliated organizations, or those of the publisher, the editors and the reviewers. Any product that may be evaluated in this article, or claim that may be made by its manufacturer, is not guaranteed or endorsed by the publisher.