An alternative technique for the determination of ultimate moment capacities of round prestressed concrete poles is presented in this paper. The traditional concrete compression area of poles is determined using the conventional iterative design process, the compatibility and equilibrium approach. Moreover, the area of annulus defined by the neutral axis is multiplied by an area reduction factor. It is possible to use numerical methods to calculate the centroid of the reduced annulus. However, the concrete compression zone and strain variation in the longitudinal reinforcement makes the calculation awkward. This paper addresses the problem from a different perspective. No reference to area reduction factor is required in the course of locating the neutral axis. The results are compared in a very satisfactory manner with that obtained using PCI approach. The purpose is to provide formulations that can be easily computerized and implemented on relevant software for the design of round prestressed concrete poles.

This research applied the method of direct acoustic admittance measurement to investigate effects of the flame character and the burner acoustic admittance on the flame-driving behavior. The acoustic admittance data were obtained by directly measuring the amplitudes of pressure and velocity oscillations, and the phase relations between the pressure and velocity oscillations. Results of this research reveal that the flame-driving characteristic is mainly dominated by the nonreactive burner admittance not the flame. Whether the flame is to drive or damp the acoustic field is detcrmined by the acoustic admittance of the burner system. The flame equivalence ratio plays a role mainly in determining the magnitude of the flame driving/damping.

In this study, the problem of creeping flow relative to an isolated porous spherical shell has been examined. The Brinkman-extended Darcy equations and the Stokes' equations are utilized to model the flow in the porous region (shell region) and free fluid region (inside the core and outside the shell), respectively. The stress jump boundary conditions at the porous media/free fluid interfaces are included and the exact solution has been found. The drag experienced by the porous shell has been discussed for various jump parameters and shell thickness.

This paper presents a time-domain control methodology, which is named as the robust Kalman-filter-based frequency-shaping optimal feedback (KFBFSOF) control method, to treat the active vibration control (or active vibration suppression) problem of flexible mechanical systems under simultaneously high frequencies unmodelled dynamics, residual modes, linear time-varying parameter perturbations in both the controlled and residual parts, noises (input noise and measurement noise),and noise uncertainties. Two robust stability conditions are proposed for the flexible mechanical system, which is controlled by a KFBFSOF controller and subject to mode truncation, noise uncertainties, and linear structured time-varying parameter perturbations simultaneously. The advantage of the presented KFBFSOF control methodology is that it can make the controlled closed-loop system to obtain both good robustness at high frequencies and good performance at low frequencies. Besides, the proposed robust stability criteria guarantee that the designed KFBFSOF controller can make the controlled flexible mechanical system to avoid the possibilities of both spillover-induced instability and time-varying-parameter-perturbation-induced instability. Two examples are given to illustrate the application of the presented control methodology to the active vibration control problems of a simply supported flexible beam and of a flexible rotor system.

This paper applies the method of self-similar potentials to analyze the dynamic behaviors of the problems of mode-I, mode-II, and mode-III cracks propagating along the x-axis with constant speed, while the constant speeds of both crack tips are not the same, called nonsymmetric crack expansion. It is assumed that an unbound homogeneous isotropic elastic material is at rest for time t < 0. However, for time t ≥ 0, a central crack starts to extend from zero length along the x-axis. On the crack surfaces of x ≥ 0, there exists uniform distributed load such that the rightmost crack tip propagates with speed ms, while the leftmost crack tip with speed s, where m > 1 and is constant. First, the complete solutions of the mode-I, mode-II, and mode-III problems are obtained. After the complete solutions are found, attention is focused on the crack surface displacements and dynamic stress intensity factors. The results of this study show that the DSIF is equal to the static SIF when the crack-tip speed is zero, and DSIF is zero as the crack-tip speed approaches the Rayleigh-wave speed. Moreover, for the special case m = 1 which indicates that the velocities of both crack tips are the same, the DSIFs of this three modes in this study is the same as those of Ref. [15].

A theory is proposed to derive the resultant elastic potential energy change in the systems accompanying phase transformation for a finite concentration of inclusions. Based on the thermodynamic equilibrium, this elastic potential energy due to the interaction of ellipsoidal inclusions is employed to evaluate the critically applied stress that induces phase transformation. It is shown that the material system with the disc inclusions is easier to get the stress-induced transformation than that with other shapes. The inclusions with 3-D random orientation are the most effective one to have toughness increments. The height of the transformation zone strongly depends on the volume concentration and the shape of the inclusions. As compared with the experimental data, the theory is in an acceptable range of accuracy.

1. Figures 13 to 18 should be deleted from the paper.

2. The first sentence on page 118, “The corresponding and traction field at r = 1, 10 and 20 for the case of m = 0 are depicted in Figs. 12(a), (b) and (c),” should be modified as: “The corresponding radiation patterns of displacement and traction field at ksr = 1, 10 and 20 for the case of m = 0 are depicted in Figs. 12(a), (b) and (c), respectively.”